Diesel Engines
Diesel Engines
Diesel Engines
A.J. Wharton, CEng, FI MarE
1 Engine Types
Cycles and Timing
Gas Exchange Processes
Engine Parts
5 Operating Systems
Safety and Operation
This third edition of an established book has become necessary to enable the marine
engineer to keep abreast of new developments in design and manufacture of marine
diesel engines. As the most efficient means of producing power from low grade fossil
fuels, the diesel engine is predominant as the propulsion machinery for ships. It should
be remembered that the majority of the world's produce is still transported by sea.
The book continues to meet the requirements of Marine Engineer Officers preparing
for examinations. The wider coverage is supplemented by additional diagrams to aid
understanding and retention of important details and to facilitate easy reference.
Engine descriptions are limited by the size of the book but are sufficient to show present
trends in current engines. Some earlier large engines have been retained, as these
illustrate some of the major changes in past development.
Modern designers have the benefits of computer aided design and finite element
analysis together with improved materials and manufacturing processes, all of which
have contributed to the highly efficient and reliable engines available today. New
developments in turbocharge systems improve fuel combustion efficiency over a wide
range of engine speeds and allow the effective use of previously wasted energy to
improve the total energy output. It is not surprising that engines from different
manufacturers appear to have many features in common.
It is the adding of new technology to well proven practice that ensures reliability and
gives ship-owners the confidence to order new engines even before they have been
proved in service.
The author wishes to express his appreciation for the assistance given by the following
manufacturing companies, which were most generous in providing information and
allowing the use of their material.
ABB, Asea Brown Boveri Ltd; MAN-B & W, MAN GHH (Great Britain) Ltd;
SEMT Pielstick, NEI Allen Ltd Crossley Engines; Stork-Wiirtsilii Diesel Ltd; New
Sulzer Diesel UK Ltd; Wiirtsilii Diesel Oy. He is also grateful for the help of the
Information Centre, Institute of Marine Engineers.
= kilogram
= newton
= metre
= newton/sq metre
= degrees celsius
I inch=25.4mm=0.025m
I foot=0.3048m
I square foot=0.093m1
I cubic foot=0.028mJ
I pound mass (lb)=0.453kg
I UK ton (mass) = 1016kg
I short ton (mass) = 907 kg
I tonne mass = IOOOkg
I pound force (lbf)=4.45N
I ton force (tonf)=9.96kN
I kgf=9.81 N
I Ibf/in1= 6895N/m1= 6.895kN/m1
I kgf/cm1= I kp/cm1= 102kN/m1
I atmos= 14.7Ibf/in1= 101.35kN/m1
I bar= 14.5Ibf/in1= IOOkN/m1
Note: For approximate conversion of pressure
lOOkN/m1= I bar = I kg/cm1= I atmos
I tonf/in1= I5440kN/m1 = 15.44MN/m1
I HP=0.746kW
Engine Types
Marine engines are required to operate continuously, reliably and safely in unmanned
engine rooms and with extended periods between planned overhauls. They may be
expected to operate for considerable periods at low power without ill effects, and to be
tolerant oflow quality fuels while maintaining very high thermal efficiency. Continuing
development and improvement in design are necessary to meet these demands and new
generations of engine models are produced to take advantage of advances in research
and experience. Considerable improvements in design and performance of turbochargers and charge air systems in recent years have contributed substantially to
increases in engine power and economy. The supply of charge air at reduced power is
sufficient to maintain efficient combustion, arid at full power surplus exhaust gas
energy can be diverted to operate 'power take off systems which will supplement the
main or electrical power output.
Changes in world trade may lead to the emergence of different ship types with special
demands upon their power systems. Manufacturers produce a range of cylinder sizes
and numbers in each model so that a wide selection of power or other parameters is
available. Further sizes are developed when a particular demand becomes evident.
Practically all new merchant ships are powered by diesel engines, and some existing
large steamships have been re-engined with diesel power to improve their economy and
extend their useful life.
Basically, marine diesel engines can be divided into two main types: large, slowrunning direct drive engines with limited numbers of cylinders, or medium to high
speed engines driving through reduction gears. Cylinder sizes do not necessarily
distinguish between these; slow-speed engines are available with cylinder bores down
to 260 mm, while medium speed engines are produced with bores up to 620 mm.
Although there are distinct differences between both types, much of the subject
matter in this book is relevant to all engines. Where necessary their characteristics are
dealt with separately.
This category refers to engines operating at between 55 to 150rpm. Large, slow-speed,
direct drive main engines operate exclusively on the two-stroke cycle and are of
crosshead construction which allows complete isolation between the cylinders and the
crankcase. They are best able to tolerate low-quality fuels and burn these successfully
to obtain the highest thermal efficiencies. Slow piston speed and fewer working parts
make them very economical in lubricating oil, and give low rates of wear and
remarkable reliability. Although there are now few manufacturers producing these
engines, they dominate the market particularly in ocean-going ships.
A variety of size and numbers of cylinders are available, to suit all power
requirements. In addition to the standard models they are produced in long-stroke
versions with stroke/bore ratio up to 3.8: I and at speeds down to 55 rpm, allowing the
use oflarge, slow propellers for high efficiency in vessels such as bulk carriers and large
tankers. Slightly shorter stroke engines operating in the speed range of 100 rpm are
designed for high speed container ships. Short-stroke models cater for ships with
limited draft, propeller size or engineroom height. Many of the basic engine
components are common to all versions.
Current models of two-stroke engines are illustrated here, together with some older
models which are still in operation and have distinctive features of which the student
must be aware.
Fig. 1.1 shows a Sulzer RTA 84C engine which is a
typical modern slow-speed, two-stroke, crosshead type, long stroke engine. It has a
boreof840mm, a stroke of2400mm and an operating speed of 100rpm. It is available
with between four and twelve cylinders and is particularly produced for the large, fast
container ship market. Its design and construction is similar to other engines in the
RTA Series 2 which offer a number of cylinder sizes down to 380mm.
A deep single-wall, box type bedplate fabricated from welded steel plat.es and
castings, and substantial welded 'A' frames surmounted by cast iron cylinder jackets
boIted together to form a cylinder block, together give a rigid overall construction. The
structure is pre-stressed by long vertical tie bolts.
Cylinder liners are of alloy cast iron. A stiff collar at the upper end resists the heavy
gas load; this lands on the cylinder jacket. It is bore cooled with water flow rates
regulated to maintain correct temperatures throughout the liner. The lower end is
uncooled within the scavenge space. Multi-level cylinder lubrication is used to reduce
liner wear rates.
The solid forged steel cylinder cover is bore cooled to reduce thermal stress and to
conduct heat from the fuel injector pockets. A centrally positioned large exhaust valve
cage has its valve seat intensively cooled, taking water from the cover. The valve is
manufactured of Nimonic 80 A alloy and is rotated by vanes fitted to its spindle. It is
opened by hydraulic pressure from a cam driven actuator and is closed under the action
of an air spring.
The piston has an alloy steel crown and has five compression rings fitted in
chromium plated grooves. There is a short cast iron skirt. The piston is oil cooled using
both the 'shaker' method and small jets to propel oil into bore holes close to the
underside of the crown and behind the ring grooves. Oil for cooling is supplied and
returned through a bore in the piston rod from swinging links at the crosshead. The
single piece cross head has the piston rod bolted to its upper surface and a continuous
full-length lower half, top end bearing, which is of white metal and lubricated with high
pressure oil. Guide shoes are attached to each end of the crosshead.
A semi-built up crankshaft has slim webs to allow large bearing areas; main bearing
caps are secured by jackbolts from the engine frames. The main camshaft is gear driven
and is fitted with servomotors to re-time fuel pump cams and the air start distributor
when operating the engine astern.
Fuel pumps are of the cam-driven valve-timed type, with variable ignition timing to
adjust timing and maintain efficient combustion at low speed. Each pump supplies
three uncooled fuel injectors placed symmetrically in each cover. Hot fuel oil circulates
the valves when not injecting.
The engine has through-scavenge
and constant pressure turbocharging
with a high
efficiency, uncooled turbocharger
supplemented at very low speeds by two constant
speed, electric driven auxiliary air blowers.
Fig. 1.2 shows a Sulzer RT A Series I engine. These have many similarities to the
Series 2 engines which were developed from it. This engine has water cooled pistons.
Water is supplied and returned through telescopic glands. These are entirely separated
from the crankcase to avoid any possibility of contamination.
A separate piston
cooling water circulation system is used.
The crosshead has split top end bearings. The piston rod is attached to the centre
block of the crosshead between the two top end journals. Each bearing has a flexible
support to limit load concentration.
Thin-wall aluminium-tin
bearing shells are fitted.
MAN-B & W MC Engines
Fig. 1.3 is of a MAN-B
& W K90 MC-·C engine.
This is a large crosshead type two-stroke engine with a bore of 900 mm, a 2300 mm
stroke and an operating speed of 104 rpm. It is constructed with between six and twelve
Developed as one of the extensive range of the manufacturer's
MC engines, it is of
the power and speed best suited to large, fast container ships. The increase in running
speed is obtained by a slight decrease in engine stroke. High thermal efficiency is
maintained by an increase in mean effective pressure.
Construction can be considered generally as typical for the whole range. The engine
bedplate is of rigid box form, fabricated from steel plates with main bearing supports
of cast steel. Welded 'A' frames are assembled into a frame box which contains the
crankcase, the crosshead guides and also supports the wheels for the chain drive of the
camshaft. A cast iron cylinder frame accommodates
the scavenge space between the
cylinder jacket and the diaphragm, both of which are water cooled. Long pre-stressed
tie bolts are fitted between the top of the frame and the underside of the bedplate
The cylinder liner is of alloy cast iron, its upper flange lands on top of the frame and
has bore cooling. It is secured by a forged steel cylinder cover which is also bore cooled
and is shaped internally to accommodate
most of the combustion space. Cylinder
lubricating oil is injected at one level in the liner. Pistons have a chrome-molybdenum
alloy steel crown with hard chrome-surfaced
ring grooves in which four compression
rings are fitted. In this particular model a protective layer oflnconel is welded to part of
the crown surface to prevent high temperature corrosion. The piston is oil cooled, oil
/.3 MAN-B
& W K90 MC-C
being supplied by a telescopic gland to the crosshead and then through the piston rod.
It is returned from the crosshead to a slotted pipe in the crankcase. A short cast iron
skirt is added. The crown is bolted to the piston rod at an inner support ring.
Surface hardening reduces wear on the piston rod at the diaphragm gland. The rod is
bolted at the top of a cylindrical crosshead which is of large diameter and incorporates
a full length bottom half bearing shell. Floating guide shoes are attached at each end.
The crankshaft may be either semi-built up or of welded construction, with large
journals and pins. All crankcase bearings are of white metal. Main bearings have thick
shells, crankpin (bottom end) and crosshead (top end) bearings have thin-wall shells.
White metal is also used for the guide surfaces. The exhaust valves are operated
hydraulically under oil pressure from earn-timed actuating pistons. They have
compressed air springs which allows them to be rotated by vanes. The valve spindles
are usually manufactured by the hot isostatic pressure (HIP) method, a compound
Nimonic and austenitic steel part construction. Valve housing is cooled at its seat and
spindle guide bush but its upper duct is uncooled to avoid low temperature corrosion.
Fuel pumps are cam driven and timed by the plunger helix. An adjustable barrel
allows variable ignition timing to maintain combustion efficiency at low speeds and can
be adjusted to match the ignition quality of fuels. Pump timing is changed for astern
operation by a link connected at the cam roller guide which is activated by compressed
air. Each pump supplies three identical fuel injectors for the corresponding unit.
Injectors are uncooled but they circulate hot fuel oil directly while their needle valves
are in the closed position.
The engine operates with a constant pressure exhaust system, with uncooled
turbochargers. Two auxiliary blowers are fitted to operate at low charge air pressure or
at low engine revolutions. A number of waste heat recovery and power take-off systems
can be operated under running conditions.
& W S 26 MCE Engine Shown in Fig. 1.4, this is the smallest of the
MC range of engines. It has a cylinder bore of260mm, a stroke of980mm, operates at
a speed of 250 rpm and it is produced as in-line engines with four to eight cylinders.
Being the smallest crosshead type two-stroke engine available, it can be considered to
compete directly with medium speed four-stroke engines of equivalent power for the
propulsion machinery in ships of limited size. The moderate speed and relatively long
stroke enable it to burn low quality fuels successfully and allow it to be coupled directly
to a propeller, eliminating the required space, weight and cost of reduction gearing.
Separation of the crankcase from the cylinders limits the consumption and deterioration of lubricating oil compared with trunk piston engines. A greater height of
engineroom will be required.
The engine bed plate is of cast iron with transverse members supporting steel backed
white metal main bearings. The crankcase is formed from one single casting which
includes frames and guides together with the scavenge trunk and the diaphragm. Long
pre-stressed tie bolts passing through frames and transverse members secure the
bedplate and crankcase framebox together.
The cylinder liner is of alloy cast iron and is surrounded by two part water jackets.
The lower jacket lands on the substantial top flange of the frame box and supports the
liner collar. The upper jacket section, fitted above the collar, cools the combustion
space and passes water to the cylinder cover. The one piece, forged steel cover is bore
cooled and is secured on the top of the liner by long studs from the frame box. The
construction of most of the working parts is similar to those for larger MC engines. The
piston crown is of alloy steel with four compression rings, is oil cooled and has a short
cast iron skirt. The piston rod passes through a diaphragm gland and is fitted to a single
bearing crosshead. The crankshaft is solid forged and includes a thrust collar. All
crankcase bearings are of white metal.
The hydraulically operated exhaust valves have air springs. The camshaft is chain
driven and is mounted within the framebox. Fuel pumps are cam driven with helix
control; each supplies the two uncooled injectors per cylinder. Reversing is carried out
by adjustment of the cam follower linkage.
The engine operates with a constant pressure turbocharge system with one auxiliary
blower automatically started during low-speed running. A compressed air jet system
can accelerate the turbocharger in an emergency.
B & W KEF Engine
This engine shown in Fig. 1.5and produced up to twenty years
ago includes many features which have been improved and developed in present
models. Exhaust valves were operated by pushrods and were closed by helical steel
springs. Through scavenge was used and a pulse exhaust system required grouping of
1.4 MAN-
B & W S 26 MCE engine
cylinder exhausts into groups ofthree or four. Turbocharge efficiency was moderate by
present standards.
Bedplates were fabricated although some were cast iron. Semi-built up crankshafts
and split crosshead bearings were used. Reversing was carried out with a large
servomotor which rotated the whole camshaft to re-time fuel pump and exhaust valve
cams. Heavy fuel was used and it was necessary for fuel injectors to have a separate
cooling system.
These engines have given good service over many years.
Shown in Fig. 1.6, this engine was produced in 1968. It
showed many developments which are present in modern engines but illustrates a
different philosophy in combustion chamber arrangements.
No exhaust valve was fitted and exhaust timing was controlled by a long piston skirt
which uncovered exhaust ports in the cylinder liner. This required a form of cross/loop
scavenge. The close proximity of both scavenge and exhaust ports made liner design
more difficult and increased stress and wear in this area.
Cylinder covers were in two concentric, water-cooled parts. The centre section was
of cast iron and contained the valve pockets; the outer section formed a 'strongback' of
cast steel. In later models forged steel, one piece covers with bore cooling were
introduced. A single, central fuel injector was fitted; this was water cooled for burning
heavy fuel. Cam-driven, valve-timed fuel pumps were used and apart from the air start
distributor, these alone were re-timed for reversing. A servomotor rotated the main
camshaft for this purpose.
The bed plate was fabricated from steel plates with steel castings at main bearing
supports. Lower stroke/bore ratio allowed bedplates to be more shallow with strong
longitudinal box sections at each side. Constant pressure exhaust system was used,
with water cooled turbochargers. Charge air pressure was augmented by a system of
under-piston charging.
Many of these engines were produced for service aboard ships and for shore plant.
DOXFORD J Type Engine
Fig. 1.7 shows a single acting, opposed piston engine
of this type which was the last opposed piston model produced by the company.
Various cylinder sizes up to 760mm were marketed; it had a combined stroke of
2180 mm and a speed of 119 rpm. Engines were built with from three to nine cylinders.
Two pistons are fitted per cylinder. The lower or main piston is oil-cooled and has a
long skirt fitted; it uncovers the scavenge ports at the bottom of its stroke. It is
connected via a normal crosshead and connecting rod to the centre crank for the unit.
The upper or exhaust piston is inverted in the top ofthe liner. It also has a skirt fitted, is
water-cooled and uncovers the exhaust ports at the top of its stroke. This piston is
connected by a beam to two side-rods with crossheads and is connected with each ofthe
two outer cranks of the unit. Two pistons moving in opposite directions give good
primary balance. Because all the gas load is transmitted directly to the crankshaft, no
tie bolts are required. The bedplate is also of lighter construction and is manufactured
of fabricated steel plates.
The crankshaft is of built-up construction. It has three cranks per unit. This causes
bending and high stress. The cylinder liner is in three parts bolted together, the centre
combustion chamber being of cast steel. This contains all the valve pockets.
The fuel system works on a common rail; the fuel pumps, which are operated by a
chain driven crankshaft, maintain the system under pressure. Two injectors are fitted
per cylinder, and are controlled by a cam-operated mechanical timing valve.
The engine operates with through scavenge and pulse turbocharging. It was unable
to match the high pressures and fuel economy of more orthodox two-stroke engines.
Most of these are designed to operate on the four-stroke cycle and are of trunk piston
construction. They are much lighter and smaller than equivalent slow-speed engines,
but require gearing or some other means to reduce the drive speed for ships' propellers.
Medium speed generally means between 400 and 1000 rpm.
Smaller engines are robust and can be highly rated. Their power/weight and
power/size ratios make them particularly attractive where engineroom size is limited.
They also require less strength built into the ship's supporting structure. Long-stroke
versions are produced to give greater fuel economy on heavy fuels.
Medium speed engines are widely used in ferry type vessels and those of limited size.
Rapid progress has brought many different models into production and a wide choice
exists in size, speed and power. They are well adapted for re-engining existing ships,
fitting well into existing engine space and giving ample power with increased economy.
Flexible couplings and torsional vibration dampers are fitted at the couplings between
engines and gearbox drives.
Plant systems may incorporate a number of similar engines of this type, giving great
flexibility for part-load operating or maintenance. With controllable pitch propellers
there is no need to fit the more expensive reversible engines.
Modern engines can operate successfully on all but the heaviest of fuels with
surprisingly low wear rates. Smaller and lighter parts, with simpler construction, give
easier maintenance. Shell bearings are in general use.
Engines may operate on either the pulse or constant pressure turbocharge system.
As four-stroke engines they are self-aspirating at slow speeds. Pulse charging will
incorporate pulse converters to reduce the size of exhaust manifolds and improve
turbine efficiency. This system may have some advantage where rapid acceleration is
required, while constant pressure gives a greater efficiency overall. Turbochargers may
require water cooling with the high exhaust temperatures which may be produced.
Medium and high speed engines may be constructed in either in-line or Vee
configuration, and many manufacturers supply all but the largest-bore models in either
form. As the name Implies, Vee engines are constructed with two banks of cylinders
arranged at an angle using a common crankcase and bedplate with considerable saving
in size and weight. Twice as many cylinders can be accommodated in a given length,
although the width of the engine is increased. Due to the increased power from a Vee
engine, however, the crankshaft must be of adequate strength. A variety of methods
have been utilised to connect two pistons to each throw of the shaft: most common is
the side-by-side arrangement with two bottom end bearings.
The Vee formation gives economy in the common space used for the exhaust system
and turbochargers. Two camshafts are used, but in multi-cylinder engines starting air
valves may only be necessary on one bank of cylinders.
& W L58/64 Engine
Fig. 1.8 shows an engine of this type. With a
cylinder bore of 580 mm and 640 mm stroke, operating at 400 rpm, this is one of the
larger medium speed four-stroke engines and the largest ofa series of three of the same
design by the same manufacturers. All are in-line construction with between six and
nine cylinders. They operate well on heavy fuels and are designed for ease in
maintenance with extended periods between overhauls.
The engine has a stiff cast iron, monoblock frame as its main strength member. The
solid forged crankshaft is underslung with its main bearings supported by long tie bolts
extending to the top of the frame. They are also braced on each side by horizontal side
bolts. Balance weights are fitted to each web of the crankshaft.
Each cylinder liner is supported by a jacket ring which lands on the engine frame and
directs intensive circulation of cooling water to the top part of the liner. Cylinder
covers are in nodular iron and are water-cooled;
a strong deck plate gives vertical
support to a thin flame plate for good conductivity of heat. It contains air and exhaust
ducts together with all necessary valve connections. Cover tie bolts pass through the
jacket ring and are secured under the diaphragm plate in the main frame. All tie bolts
are hydraulically tightened to pre-stress the main parts.
A composite piston has a forged steel crown with an aluminium alloy skirt in which
the undrilled gudgeon pin is secured. Pistons are oil cooled by the shaker method, oil
passing from the crown through bore holes to cool behind the ring grooves. Three
compression rings are fitted in induction hardened grooves. The top ring is plasma
coated, the others being chrome plated. A single oil control ring is fitted at the top of
the skirt.
The connecting rod is jointed near the top end, allowing the piston to be removed
without disturbing the bottom end bearing. Each cover has two exhaust valves, fitted in
water cooled cages. They have Stellite faces and vanes on their spindles cause them to
be rotated by the gas stream. Two air inlet valves are fitted in uncooled cages; they are
rotated by Rotocaps.
A single, central fuel injector is water-cooled and can be pre-heated before starting
on heavy fuel. The cam-driven, helix-controlled fuel pump has its control edge profiled
to cause uniform firing pressure at all loads. A rocking lever at the pump follower
allows manual adjustment for variable ignition timing.
Constant pressure turbocharging
is used with uncooled, high efficiency turbochargers supplying adequate charge air over the full range of load. Ample exhaust
energy is also available for waste heat recovery.
ZA 40 S Engine
This engine, shown in Fig. 1.9, has a cylinder bore of
400 mm, a stroke of 560 mm, and normal speed of 5 I0 rpm. The suffix S denotes it as a
longer stroke development of the earlier ZA 40 engine with which some parts are
identical, although the connecting rod is shorter to maintain the same engine height.
The longer stroke gives improved fuel economy with slightly reduced speed and allows
the use of lower grade fuels.
It is produced in many cylinder numbers, ranging from six to nine for in-line, and
twelve to eighteen for Vee form engines.
The main engine frame is a single piece, grey iron casting of rigid construction. It has
main bearings, secured by tie bolts in both vertical and horizontal
directions. The main bearings are thin-wall, steel-backed, aluminium-tin,
type. They support the solid forged crankshaft which is fully balanced and has large
bearing surfaces.
The cylinder liner has a deep, bore cooled top flange for strength, landed on the
engine frame. Its lower end is uncooled. Controlled lubrication is injected near midlength of the liner. A cylinder cover of special cast iron is water-cooled,
its lower
landing having bore cooling holes drilled. Air and exhaust passages pass through the
cover. There are two exhaust and two air inlet valves per unit. No valve cages are used,
the valve seats being pressed directly into the cover. The seats are bore cooled. Exhaust
valves are of Nimonic alloy with Rotocaps fitted.
The composite piston has a steel crown with a grey cast iron skirt. Three
compression rings are fitted in chromium-plated grooves in the piston crown, all rings
being barrel faced and chrome-plated. One oil control ring is fitted near the bottom of
the skirt.
The piston is oil-cooled, using the shaker action, with bore holes which give efficient
cooling while maintaining stiffness ofthe crown and ring grooves. Oil is taken from the
top end bearing.
A special feature of this engine is its rotating pistons. Each one is slowly rotated by
the action of small hardened pawls which engage a toothed ratchet ring as they are
moved through the swing of the connecting rod in which they are spring loaded. The
rotation ensures an even temperature around the piston crown, allowing reduced
piston and skirt clearance and improved lubrication of piston rings and liner.
To allow for piston rotation, the top end bearing is spherical. A steel sphere on the
connecting rod works in a bronze bearing. This has greater surface area contact and
therefore a lower bearing pressure than a gudgeon pin.
The connecting rod is bolted to a palm at the bottom end bearing, allowing it to be
disconnected during piston overhaul without removal of the bearing. Steel-backed,
aluminium-tin thin-wall bearings are fitted in the bottom ends.
Fuel pumps are of the helix type. The plunger is shaped to give variable ignition
timing to maintain maximum combustion pressure at all loads. Pumps can circulate
hot fuel oil.
The turbocharging system uses a single-pipe exhaust with pulse converters giving a
combination of pulse and constant pressure. A bypass is fitted which is open at part
load, allowing some charge air to mix with the gas stream and ensuring good matching
ofthe turbocharger, without surge. In systems not fitted with variable ignition timing, a
waste gate is fitted to the air receiver to blow off excess air when operating above 85%
offull power. Two-stage fresh water circulated charge air coolers allow pre-heating of
charge air at low power. High efficiency of turbochargers will allow use of some
exhaust energy for power take-off or waste heat recovery.
Fig. 1.10 shows a Vee version of this. It is
one of the wide range' of medium speed engines produced by this company for marine
and land applications and available as in-line or Vee configurations.
The cylinder bore is 400 mm, with a stroke of 460 mm, and speed of 520 rpm. A
longer stroke 2--6 B version has stroke increased to 500 mm, giving greater fuel
economy and higher power at the same rpm. Cylinder numbers are from six to nine for
in-line and from ten to eighteen for Vee versions.
The main engine frame is fabricated from steel plates and castings. This contains the
crankcase which is completed by a steel-bottom oil pan. It also supports the cylinders
and jackets. The cylinder liners are of alloy cast iron, bore-cooled in their upper
section. Cylinder covers are of cast iron with water cooling. They are secured to the
frame bossing by tie bolts which also locate the jackets. Exhaust and air passages pass
within the cover with two air inlet valves and two exhaust valves with Rotocaps fitted in
water cooled valve cages.
The composite piston has a crown of alloy steel with three compression rings in
grooves; one further compression and two spring loaded oil control rings are fitted near
the top of the aluminium alloy skirt. The piston is oil cooled by the shaker method, with
oil taken via the floating gudgeon pin. The connecting rod has a thin shell bottom end
bearing with serrated butts angled at 40°. In Vee engines, pairs of connecting rods from
the same Vee have bottom ends fitted side by side on the same crankpin.
The crankshaft is solid forged in chrome-molybdenum alloy steel by the continuous
grain process. Thin shell main bearings of copper-lead overlaid with a tin alloy are
underslung from division plates in the frame and are secured by vertical and horizontal
A cam-operated, helix-controlled fuel pump supplies a central fuel injector. The
injection pressure exceeds 1000 bar, ensuring rapid combustion of heavy fuels.
Injectors are water cooled from a separate system. The gear driven camshafts are fitted
with both ahead and astern timed cams. Reversing is carried out by servomotors
which move each camshaft axially. Only one bank of cylinders is fitted with air start
The turbocharge system uses modular pulse converters to achieve a compromise
between a pulse and a constant pressure system. By-pass and wastegate valves prevent
surge and limit maximum pressure.
Fig. 1.11 shows a Wiirtsilii R 46 engine. It is a
medium speed, four-stroke, trunk piston engine which has been specifically designed to
operate efficiently on low grade oil fuels. This is accomplished by the use of twin
injection of fuel, together with changes in the operating systems.
The engine has a cylinder bore of 460 mm, a stroke of 580 mm, and designed speeds
of 450,500 or 514 rpm. It is available with between four and nine cylinders in-line or
with between twelve and sixteen cylinders in Vee form.
The main structural strength member is the engine block. Manufactured in nodular
iron from a single casting, this contains the transverse frames, cylinder jackets,
camshaft housing and most of the crankcase. A light welded steel oil sump is attached
below. Main bearing caps, which support the underslung crankshaft, are clamped by
hydraulically tensioned bolts, two from below and two horizontally. Cylinder liners
are of alloy grey cast iron and each has a deep, thick collar at its upper end which lands
on top of the engine block. The collar gives added strength and is bore cooled, passing
water from the jacket to the cylinder head. The cylinder head of cast iron is a rigid,
symmetrical box-like form with an intermediate deck giving intensive water cooling of
valve seats and injectors. Air inlet and exhaust gas ducts pass through the head. There
are two inlet and two exhaust valves per unit, fitted directly to the head. All valves have
Rotocaps fitted, are Stellite-faced and land on hardened steel seats. Two fuel valves are
fitted. The main injector is positioned in a central pocket with its fuel supply passing
through a bore in the head. A pilot injector is placed in a pocket at the front of the head,
angled down at 45°. The cylinder head is secured to the engine block by four long studs.
The composite piston has a steel crown with two compression and one oil control
rings, the top ring grooves having hardened surfaces. There is a skirt of nodular cast
iron in which the gudgeon pin is fitted. Oil cooling for the piston is taken from the top
end bearing, a honeycombed surface giving efficient cooling with high rigidity in the
crown. Some oil from the piston passes to a groove at the top of the skirt for cylinder
Connecting rods are in three pieces allowing the rod to be separated from the bottom
end bearing so that piston or bearing may 'be overhauled independently. The
crankshaft is solid forged and has balance weights fitted on each web. All bearings are
designed for large areas with low pressures to allow use of softer bearing metals. Steel
backed copper-lead bearings have a nickel barrier layer with an overlay of
tin-antimony which prevents corrosion of the thick copper-lead pad.
As stated earlier, the fuel system uses twin (or pilot) injection (see page 90). This gives
easy starting and smooth, efficient combustion ofJow quality fuel over the full range of
engine power.
Each earn-driven fuel pump has helix-control and is designed to supply the pilot
charge at lower pressure and earlier than the corresponding main charge. The quantity
of pilot charge is limited by its discharge chamber capacity.
The 'SwirlEx' turbocharge system has double skin exhaust connections to the
turbocharger, and swirls the gas mechanically to give a combination of pulse and
constant pressure charging.
1.11 Wartsilii Vasa R 46 engine
To improve operation on low quality fuels further a high compression ratio is used.
A central water cooling system with two-phase charge air coolers allows charge air
temperature to be raised when operating at low power. Independent circulating pumps
and a preheater ensure the en_gineis raised to working temperatures before slow
turning, prior to starting.
Cycles and Timing
The term 'cycle' refers to one complete sequence of operations required to produce
power in an engine. This cycle of operations is continuously repeated while the engine
is running. For a diesel engine it consists of four operations within the cylinder:
Compression of a charge of air
Injection of fuel which then ignites
Expansion of the hot gases formed during combustion
Expulsion of the used gas to exhaust.
The cylinder is then recharged with air and the cycle is repeated.
Diesel engines can be designed to complete this cycle once during each revolution
and this is termed the two-stroke cycle, or alternatively to take two engine revolutions
to complete-the four-stroke cycle. An engine can only operate on the cycle for which it
was designed.
Engine stroke is measured as the full distance through which the piston moves between
each end of its travel. It can be seen that it must move through two complete strokes
(one up and one down) during each revolution of the engine.
Engine timing refers to the relative time or position of the crank, at which each
operation during the cycle is commenced and is completed. It is measured as the angle
through which the crank has been rotated from a datum position such as top or bottom
Two-stroke cycle Practically all large, slow-speed, direct drive marine diesel engines
operate on the two-stroke cycle (Fig. 2.1). As its name implies a two-stroke cycle takes
place in two consecutive strokes of the engine piston, or one revolution of the
crankshaft. Thus each operation in the cycle is repeated during every revolution of the
engine. The two strokes of the cycle may be termed: Compression stroke and Power or
expansion stroke. Operations take place in a fixed order and must occur when the
piston reaches a corresponding position in its stroke. These positions are shown as
volumes on an indicator diagram which relates them with pressure within the cylinder.
It is convenient to express them in terms of angles of crank position measured from top
dead centre (TDC) or bottom dead centre (DOC) and they may be shown as a circle on
a timing diagram. (Numbers have been added for reference.)
Actual timing may differ between engines due to construction and design differences
such as: ratio of connecting rod length/crank length, stroke/bore ratio, engine speed,
engine rating etc.
Completion of scavenge. Air is entering the cylinder, expelling exhaust gas
and recharging it for the next combustion. Scavenge and exhaust are open.
2./ Two-stroke cycle and timing diagrams
Post~scavenge. Scavenge ports have closed and some air within the cylinder
may leak to exhaust. In some engines 2 and 3 are made to coincide to
eliminate leakage of air.
Compression. Exhaust has now closed and the air trapped within the
cylinder is compressed by the upstroke of the piston to raise its temperature
sufficiently to ignite the fuel.
Fuel injection takes place and combustion occurs causing a rapid rise in
pressure. The period for which this continues depends upon the fuel pump
setting and power to be produced.
Expansion. Combustion completed, the hot gases expand forcing the piston
downwards and converting the heat energy from combustion into work on
the piston.
Exhaust blowdown. Exhaust has opened allowing gas to pass to exhaust
manifold, and pressure drops rapidly in cylinder.
Scavenge. Scavenge ports have opened and air enters to expel the remaining
exhaust gas.
Scavenging then continues for the next cycle.
Position Irepresents bottom of stroke (DDC). Position 5 represents top of
stroke (TDC).
Four-stroke cycle The majority of medium and high speed diesel engines for main or
auxiliary drive operate on the four-stroke cycle, which takes place during four
consecutive strokes, or two complete revolutions, of the engine. The four strokes may
be termed: Compression stroke. Power or expansion stroke. Exhaust stroke. and
Aspirating or air induction stroke (Fig. 2.2). Numbering the operations in sequence on
the timing diagram:
Completion of aspiration.
Compression. Air inlet valve has closed, air in cylinder is now compressed to
raise its temperature for combustion of fuel.
2.2 Four-struke cycle and timing diagrams
Fuel injection. Combustion takes place with corresponding rise in pressure.
Period controlled by fuel pump setting.
Expansion. Combustion completed, gas pressure does work on piston
during downward stroke.
Exhaust. Exhaust valve opened, piston expels exhaust gas on upward stroke.
Overlap. Air inlet valve opened while exhaust remains open. The length of
this is increased in supercharged or high speed engines.
Aspiration. Exhaust valve closed, piston draws air into cylinder during
downward stroke.
Aspiration continues for next cycle.
4 and 9 are TDC positions. I and 7 are DDC positions.
An indicator diagram is a pressure/volume graph taken from the pressures in the
cylinder of a working engine. It is plotted as a continuous line showing all pressure
changes during one complete engine cycle. Indicator diagrams can be of great
assistance to students in understanding the operation of an engine. Examination
candidates are required to be familiar with the information which may be obtained
from the shape and dimensions of the diagram.
In many modern engines sophisticated measuring devices may be fitted to record
cylinder pressures as an oscilloscope trace or printout. Usually these are not complete
(closed) indicator diagrams and they are described later in this chapter. For some slowspeed engines it is possible to obtain indicator diagrams from each cylinder by using an
engine indicator.
Briefly, this consists ofa spring loaded piston on which
the spring is compressed in proportion to the cylinder pressure acting on it (Fig. 2.3).
The piston moves a pen attached to an arm through a parallel link mechanism. A card
with a sensitised surface is wrapped around the indicator drum. The indicator cock for
the chosen cylinder is first blown through to clear it of carbon and the indicator is then
connected to it. The cord on the indicator drum is attached to some form of engine
stroke synchronising mechanism from the crosshead or a cam. The cock is now opened
and the indicator pen is held against the card, tracing a diagram for one cycle of the
engine. Pressure is recorded to a vertical scale according to the stiffness of the indicator.
compression spring. Corresponding cylinder swept volume is recorded on a horizontal
scale due to rotation of the drum by its cord. By turning the indicator cock to a vent
position, a horizontal line representing atmospheric pressure is added to the diagram.
This can act as a pressure datum line.
Four types of indicator diagram can be obtained. These are illustrated for a twostroke engine. Power card is taken with the indicator drum rotation in phase with
piston movement (Fig. 2.4). The area within this diagram represents the work done
during the cycle to scale. This may be used to calculate the power produced or the mean
indicated pressure (MIP) for the cylinder. Irregularities in the shape ofthe diagram will
show operational faults. Maximum or peak pressure may be measured to scale between
the atmospheric line and the highest point on the diagram. Power cards for two- and
four-stroke cycles are shown in Figs. 2.1 and 2.2, drawn vertically to relate to the
corresponding timing diagrams. It can be seen that in the four-stroke cycle, pressure
changes are so low during the last two strokes that these appear as straight lines.
Compression diagram (Fig. 2.5) is taken in a similar manner to the power card but
with the fuel shut off from the cylinder. The height of this curve shows maximum
compression pressure. If compression and expansion lines coincide, it shows that the
indicator is correctly synchronised with the engine. Reduction in height ofthis diagram
shows low compression which may be due to a worn cylinder liner, faulty piston rings,
2.5 Compression diagram
insufficient scavenge air or leaky exhaust valve, any of which will cause poor
Draw card or out of phase diagram (Fig. 2.6) is taken in a similar manner to the power
card, with fuel pump engaged but with the indicator drum 90° out of phase with piston
stroke. This illustrates more clearly the pressure changes during fuel combustion. Fuel
timing or injector faults may be detected from its shape.
Light or Weak spring diagram (Fig. 2.7) is again similar to the power card and in
phase with the engine, but taken with a light compression spring fitted to the indicator
showing pressure changes during exhaust and scavenge to an enlarged scale. It can be
used to detect faults in these operations.
Some irregularities in the shape of an indicator diagram illustrate incorrect engine
operation. If the same fault appears on diagrams from all cylinders, the cause must be a
general one such as fuel condition or supply, turbocharger system, exhaust, engine
cooling or even incorrect camshaft timing. Indicator diagrams affected by some of the
more common faults are illustrated.
FUEL IGNITION TIMING Provided the fuel and engine temperatures are
correct and the injector is operating normally, the time between injection of fuel and
ignition (the ignition lag) is almost constant. Consequently ignition timing faults
signify a corresponding timing fault in the injection. They will affect the shape of the
power diagram, but are far more easily detected on an out of phase diagram.
This will cause an abnormally high peak pressure in the
cylinder at about the top of the piston stroke (Fig. 2.8). A heavy shock load will be
transmitted to the running gear and bearings with a corresponding knocking sound.
Although thermal efficiency is high and exhaust temperature reduced, the shock load
and consequent vibrations may cause damage.
Causes of early ignition may be incorrect fuel pump timing, broken or wrongly set
injector springs, incorrect fuel condition, or overheating of parts within the cylinder.
2.8 Diagram showing early ignition
It can be seen from Fig. 2.9 that this causes a low peak pressure
which occurs well after top centre of the piston. Power is lost since the fuel is not burned
correctly to transmit power at the most effective part of the stroke. Combustion may
continue during the expansion stroke (see Afterburning) and may be incomplete,
giving loss in energy produced, high exhaust temperature and smoke.
Late ignition may be due to excessive injector spring setting, poor atomisation, high
viscosity or poor quality fuel, fuel pump leaking or incorrectly timed,low compression,
insufficient supply of combustion air, or undercooling of parts within the cylinder.
This refers to slow or late combustion of fuel during the
expansion stroke of the piston and is shown by a rise in the expansion line during the
latter part of the stroke (Fig. 2.10). Exhaust temperature and pressure will be high, with
burning fuel and carbon passing to exhaust. This may burn exhaust valves and foul the
exhaust system, with risk of turbocharger surging or uptake fires. High temperatures
within the cylinder cause a deterioration in lubrication and possible damage to liner
surface and piston rings. There may also be burning of the piston crown.
can be detected by a loss in power in the
affected cylinder together with smoke at exhaust and high exhaust temperature. There
may also be a knock or pressure wave in the injection system. An indicator diagram
taken from this cylinder will show fluctuations of pressure during the expansion
process due to secondary burning of fuel leaking from the valve (Fig. 2.11). A higher
expansion line at the exhaust opening indicates afterburning.
Loss in power will be due to incorrect combustion since the fuel pressure wave from
the fuel pump must refill the space left by fuel leaked from the injector; this may also
cause the knock. Hot gas from the cylinder may blow back into the injector tip,
forming carbon and choking the atomiser. Carbon may also form on the outside of the
atomiser due to burning of the dribble of fuel.
The injector should be changed and tested. To reduce the possibility of leaking
injectors, good maintenance of fuel pumps and injectors should be carried out with
routine changing of the injectors. Fuel must be purified and filtered and must be
supplied at the correct temperature. Fuel injectors must also be maintained at the
correct temperature.
PAR TL Y CHOK ED FUEL VALVE may be due to contamination ofthe fuel in
which debris may choke the small atomiser holes in the injector. Alternatively it may be
caused by a leaky injector allowing hot gas to blow back into the injector causing
carbon to form and choke the injector. Overheating of injector nozzle may also cause
buildup of carbon. There will be Ii loss in engine power. There will probably be
hammering in the fuel pipes between fuel pump and injector and this may lead to
rupture of fuel pipe. A choked valve can be confirmed by indicator diagram power and
draw cards (Fig. 2.12), and reduced exhaust temperature.
The remedy is to change the fuel injector, clean the fuel system and ensure correct
centrifuging and filtering of fuel, and maintain correct fuel valve temperature.
2.12 Diagram showing a choked injector
2.13 Diagram showing low compression
is shown by the low pressures on a compression diagram
(Fig. 2.13). It is probable that first evidence was from faulty combustion. Low
pressures may be due to a lack of air being supplied (turbocharger faults, choked
scavenge ports or inlet valve, etc); or that air is escaping past the piston during
compression, which can be due to worn or broken piston rings, worn cylinder liner, etc,
but blowpast is then usually detected during the expansion stroke.
Blowpast is a term used when combustion gases at high temperature leak past the
piston between the piston rings and cylinder liner. This will cause rapid local
overheating, removing the lubricating oil film and causing further damage to the
surfaces and possible piston seizure. Burning gas will enter scavenge spaces in twostroke engines causing fouling, overheating and even scavenge fires. In four-stroke,
trunk piston engines these gases will enter the crankcase with severe risk of a crankcase
Blowpast may be caused by excessive cylinder liner wear; lack of cylinder
lubrication; worn, broken, stuck or poorly maintained piston rings; worn piston ring
groove landings allowing rings to cant and jam; carbon jamming rings in grooves. It
will be aggravated if the engine is overloaded.
The remedy is to gauge cylinder liner and renew if necessary; overhaul piston; clean
ring grooves and gauge them; machine or fit new groove inserts as necessary, and renew
piston rings with correct clearances. Maintain cylinder lubrication and avoid overload.
on a slow running engine can be checked by
means of a light spring indicator diagram. This will not give an accurate timing check,
but by comparison with a normal diagram or one taken during original engine trials, it
may be seen if valve opening is early or late (Fig. 2.14).
Early opening of exhaust valve will cause a loss in power since pressure is released too
soon. It will give a high exhaust temperature
with risk of overheating
Late opening will reduce the scavenge efficiency by reducing blowdown. It will also affect the supply of energy to the turbocharger.
2.14 Light spring diagram showing
early or late exhaust valve opening
2./5 Light spring diagram
showing choked exhaust
CHOKED EXHAUST The pressure during valve opening does not fall as rapidly
or as far as it should (Fig. 2.15) and prevents correct blowdown and recharge of the
cylinder with air. It causes loss in power, high exhaust temperature, smoke and possible
surge in the turbocharger,
and is the result of contamination
due to incorrect
combustion or excessive cylinder lubricating oil.
For economy and efficient running of an engine equal power should be produced from
each cylinder. This is then said to be 'power balanced' and it may be necessary to make
some minor adjustments to achieve this.
Power produced is related to the quantity of fuel injected and balancing is carried out
by small adjustments to individual fuel pump controls. Adjustments are limited to
ensure units are not overloaded, exhaust temperatures
are not excessive, and that
pump controls still cut off when brought to stop. Fuel pump rack positions and exhaust
and cooling return temperatues from each cylinder should be noted. Equal exhaust
temperatures on their own do not necessarily show an accurate balance, particularly in
engines. Figures do, however, tend to give some indication of balance
when they follow a pattern similar to that recorded at test bed trials.
Power balancing can be checked by measurement
from indicator diagrams by
calculating the area or mean height from each cylinder ..
For medium and high speed engines it may be impractical to use a normal engine
indicator due to the high velocities of parts causing vibrations in springs or drive
mechanisms. If other means are not available, power in the engine may be related to
peak or maximum pressure in the cylinder. Some manufacturers recommend that the
'ignition jump' (the difference between peak pressure and maximum compression
pressure for each cylinder) may be used to check power balance. A number of
assumptions have to be made regarding the operating efficiency of the engine. In order
that these can be justified, it is particularly important that regular and correct
maintenance is carried out on the engine and that any deviation from normal running
is noted, investigated and corrected at the earliest opportunity.
Fuel injection equipment is particularly important and fuel pump settings,
clearances and timings must be checked and maintained during periods when the
engine is out of service. Fuel injectors must be changed regularly, cleaned and tested to
ensure trouble-free operation. The injector is the most likely part of the system to be
subject to faults in service. A fault in one injector will cause loss in power in the affected
cylinder, but may also mean that other cylinders are subjected to overloading as the
engine governor attempts to maintain normal total power or speed.
If an engine operates in an unbalanced condition, some bearings and running gear
may become overloaded; this may cause overheating and bearing failure. Overload in
cylinders may cause piston blowpast, with the corresponding dangers of overheated or
seized pistons. Unbalance will also set up vibrations which, if maintained for
prolonged periods, will cause fatigue from the fluctuating stresses induced. This may in
turn lead to fatigue cracking of metal in bearings, fracture of bearing studs or bolts,
cracks in crankshaft and bedplate, and slackening or failure of holding-down bolts. A
watch keeper may ascertain that running conditions are normal by observation of the
relevant temperatures and pressures, particularly exhaust and cooling return temperatures, lubricating oil and turbocharge pressures. The exhaust should be clear of smoke
and there should be no unusual noise or vibration.
This instrument will measure and record the maximum pressure within the cylinder. It
does not identify the position within the cycle at which this occurs. The one shown in
Fig. 2.16 is similar in principle to the engine indicator but uses a beam type spring
which is less susceptible to vibration. The indicator is attached to the indicator cock:
when the cock is opened, pressure raises the indicator piston which records maximum
beam deflection on a dial pressure scale. If fuel is shut off from that cylinder, the
maximum compression pressure can be recorded and the ignition jump calculated.
Modern measuring instruments and recording equipment make it possible to monitor
all necessary pressures, speeds, temperatures and even wear rates within an engine.
By connecting a pressure sensing transducer to the indicator cock it is possible to
record pressure in the cylinder continuously. Pressures can be plotted on an
oscilloscope trace to a horizontal scale of time or corresponding crank angle (Fig.
2.17). A printout of this diagram appears comparable to an accurate, constant base
scale out-of-phase diagram. Other important points such as top centre position, point
offuel injection, etc can be superimposed. Peak and compression pressures can be seen
and equated for power or balance.
Combustion monitoring may be used to maintain efficient conditions when fuel of
different quality is used. Model diagrams should be made for the engine during shop
testing and initial trials in the ship; these can then be compared to subsequent working
Gas Exchange Processes
In an internal combustion engine this is the removal of exhaust gas from the cylinder
after combustion and its replenishment with air for subsequent combustion. Efficient
scavenging is necessary for good combustion and it is required for the very first
working cycle of the engine. The passage of scavenge air will also assist cooling of the
cylinder, piston and valves.
SCAVENGE OF FOUR-STROKE ENGINES is carried out mainly by the
pumping action of the piston during two of its strokes. The piston expels exhaust gas
on its upward stroke and draws in air on the subsequent down stroke. The scavenge
diagram (Fig. 3.1) shows that there is a period of overlap during which exhaust and
inlet valves are both open together. Provided the turbocharger supplies air at a pressure
greater than that in the exhaust manifold, a flow of cooling air will pass through the
cylinder during this period.
Scavenge is positive and very efficient, even in very high speed four-stroke engines.
SCAVENGE OF TWO-STROKE ENGINES Two-strokes rely on a charge of
scavenge air under pressure sweeping through the cylinder and expelling the exhaust
in front of it. This process must take place while both scavenge and exhaust
connections are open and the piston is near the bottom of the cylinder. Even in slow
running engines this allows only a very short period of time for scavenge to be
completed. Some mixing of air and gas will occur, but this must be kept to a minimum
and scavenging can be improved by supplying a volume of air in excess of the cylinder
volume. the excess passing to the exhaust system. Scavenge air in a two-stroke engine
must be supplied to the cylinder at a higher pressure than that in the exhaust manifold;
in modern engines this is provided from exhaust gas driven turbochargers, the air being
cooled to increase its density and cooling capacity.
Earlier two-stroke engines used a variety of methods to supply scavenge air. These
included the use of reciprocating pumps, rotary blowers and compression from the
underside of the piston on its downward stroke. All these added to maintenance work.
absorbed part of the power produced by the engine and were only able to supply
limited quantities of excess air.
In large, slow-running engines the air enters the cylinders through a circumferential
row of scavenge ports cut in the liner and uncovered by the piston as it moves towards
the bottom of its stroke. They are closed at a symmetrical timing on the following
upstroke. The exhaust valve must open before ports are uncovered to allow a period of
Blowdown is the release of gas pressure at the end of expansion as the exhaust valve
opens. Gas passes to the manifold carrying its remaining energy to the turbocharger.
Cylinder pressure drops rapidly to a value below that in the scavenge trunk. allowing a
fresh charge of air to enter.
There are three basic methods of scavenging within the cylinder (Fig. 3.2).
(a) Loop scavenge in which air passes over the piston crown and rises to form a loop
within the cylinder, expelling gas through exhaust ports cut in the same side of the liner
above the scavenge ports.
(b) Cross scavenge where air is directed upwards. passing under the cylinder cover
and down the opposite side. expelling gas through exhaust ports on that side.
(c) Uniflow or through scavenge in which air passes straight up through the length of
the cylinder. forcing the exhaust through ports or valves at the top of the cylinder.
In all these the swirl or direction of the air is assisted by the port edges being angled in
one or two planes and by the shape of the piston crown. In cases (a) and (b) a piston
skirt or exhaust timing valve will be necessary to prevent scavenge air leaking to
exhaust while the piston is at the top of its stroke. Case (c) tends to give the highest
scavenge efficiency with the least mixing of air and gas. It may also be used with greater
stroke/bore ratio and in opposed piston engines. It avoids the difficulty of high
temperature gradient between adjacent scavenge and exhaust ports in (a). or the
temperature difference on opposite sides of piston and rings in (b).
Uniftow or through scavenge with a large centrally positioned exhau'lt valve is used
in most modern, slow-speed, large engines. It allows design of efficient long-stroke
engines and is less susceptible to damage from the products of combustion from low
quality heavy fuels. Absence of exhaust ports allows the liner to be of simpler
construction, long piston skirts are not necessary and cylinder lubrication is
satisfactory and economical. The swirl imparted by the port edges causes the air to
rotate as it rises through the cylinder.
The scavenge air trunk forms part of the engine structure and must be of sufficient
size and shape to supply even quantities of air at constant pressure to all cylinders. It
must be maintained in a clean condition and be regularly inspected. Pressure relief
valves are fitted for safety.
Here the charge of air in the cylinder before compression commences is above
atmospheric pressure. Pressure charging (sometimes called supercharging) forces a
greater mass of air into the cylinder for combustion and consequently a greater
quantity of fuel can be burnt efficiently. Power output is increased giving a higher
thermal efficiency with no increase in size ofthe engine. Pressure charged engines must
be designed to operate safely under the high loads and temperatures which may be
In most modern engines high air pressures are supplied by means of exhaust gas
driven turbochargers and consequently the general term turbocharged engine is used
for both two- and four-stroke engines.
Most modern engines are turbocharged to increase the power output and efficiency
with only a relatively moderate increase in size, weight and initial cost. The efficiencyof
the system is increased by fitting a charge air cooler after the compressor (Fig. 3.3). This
will cool the air at constant pressure, increasing its density before supplying it for
compression in the engine cylinders. The mass of air per cycle can now be increased and
the quantity of fuel injected can be raised to give a corresponding increase in engine
output and thermal efficiency.
Exhaust gas from the cylinders operates the gas turbine, giving up some of its heat
energy to do so. The turbine drives a directly coupled air compressor, which draws air
from the atmosphere, compresses it and then cools it in the charge air cooler before
supplying it to the engine through scavenge ports.
The turbocharger is 'free running' and at any engine speed the exhaust gas energy
must cause the turbine to run at a stable speed at which the compressor will supply the
correct mass of air to the engine. It is then said to be 'matched' to that engine, and it
must remain matched over the full main operating power range. When an engine has
two or more turbochargers operating in parallel they must all be matched. Matched
systems may become unstable and surge if fouling or damage occurs.
A turbocharge system may be designed to operate on either the pulse or the constant
pressure exhaust system.
The main energy to power the turbine in this system is derived
from the pulse or impulse energy in the pressure wave formed during blowdown from
each cylinder. These waves travel through the manifold to the turbine nozzles where
they are converted to kinetic energy at high velocity to rotate the turbine blades. This
system gives a rapid buildup of turbine speed when an engine is started or
To maintain the pulses, exhaust connections are of limited diameter, no sharp bends
are used and the turbocharger is fitted close to the engine. The peak pressure in pulse
waves may be higher than the pressure in other cylinders whose exhaust valves are also
open. To prevent a back-flow into these, it is necessary in multi-cylinder engines to subdivide the exhaust between a number of manifolds, each connected to a separate nozzle
box at the turbine. Up to three cylinders can use each manifold without interference
depending on their firing order. A divided exhaust system reduces turbine efficiency
and may require more than one turbocharger per engine. The use of pulse converters in
modern engines aIlows individual cylinders to exhaust to a common manifold, giving
higher turbine efficiency and a more compact system which aIlows rapid acceleration
when manoeuvring.
Here all cylinders exhaust into a common
receiver large enough to convert pulses to a steady pressure. This system has one
connection to the turbine which, due to the common nozzle ring and steady pressure
conditions, can be designed to operate at a high thermal efficiency. High efficiency
turbochargers can operate effectively down to about 25% of fuIl engine power.
This system is slower in its buildup of pressure when starting, and insufficient air is
available for two-stroke engines during manoeuvring or operating at low speed and
low power. To overcome this difficulty modern engines have electric driven centrifugal
compressors fitted to augment the air pressure in the scavenge trunk when necessary.
These are driven by constant speed, low power motors controIled to start automatically if scavenge pressure drops. They draw additional air through the main
turbocharger air intake and charge air cooler discharging it through non-return valves
to the engine scavenge trunk.
Fig. 3.4 shows a section of a turbocharger as fitted to large bore engines. It consists of a
single stage, axial flow exhaust gas driven turbine mounted on a common shaft with a
centrifugal air compressor. The turbine has a nozzle ring foIlowed by a rotating disc
with a single row of turbine blades. These blades are attached to the disc by fir-tree
shaped roots and they have free room to expand when heated. Binding wires are fitted
to the blades to reduce vibrations. Blades and nozzles are manufactured of heat
resisting steel or nickel alloy. The turbine casing is in two parts, both of cast iron with
adequate water cooling spaces. There are inspection and cleaning covers to these
spaces which are circulated with fresh water from the engine cooling system. A drain
valve is fitted in the exhaust gas space.
The air compressor consists of a radial flow impeller disc together with an inducer,
both of aluminium alIoy. The impeller discharges air through a diffuser to a volute
casing. Compressor casing, also in two parts, is of cast aluminium and is uncooled. Air
is drawn from the engineroom atmosphere through inlet filters which may be removed
for cleaning. Air inlets are streamlined and fitted with insulation internalIy to reduce
Turbine nozzle ring, air diffuser, impeller and inducer will be replaceable to allow
matching between turbine and compressor, and between turbocharger and engine.
Two labyrinth type seals are fitted to the shaft, one between thrust bearing and air
compressor and the other between turbine and bearing. They are sealed with air under
pressure from the compressor discharge through internal passages and restriction
plugs. Air from the glands then passes to the atmosphere or assists cooling of the
turbine disc. The seals prevent possible oil leakage into the turbine or compressor, or
exhaust gas into the corresponding bearing oil. Some air will pass down the back of the
impeller through a labyrinth arrangement, along the turbine shaft assisting cooling,
and leaving with the turbine exhaust gases.
Two shaft bearings are fitted, one at each end allowing accessibility and cooling. End
thrust is taken at the compressor bearing, allowing the turbine bearing free thermal
expansion of the shaft. Bearings may be of either plain sleeve type with copper-lead
bushes on hardened steel shaft sleeves, or alternatively, ball or roller type. These reduce
friction drag but are susceptible to vibration and fatigue, both when running and also
by vibrations from outside sources transmitted while the engine is stopped. They must
be fitted in resilient mountings which use springs and oil damping in both axial and
radial directions. These bearings also have a limited fatigue life and they must be
renewed at stated intervals (say 8000 hours).
Lubrication of the bearings may be by various means. Ball and roller bearings may
be lubricated by selfcontained gear type pumps operated from the shaft and drawing
oil directly from the independent bearing sump. Oil level must be maintained in these
sumps and the oil should be renewed at stated intervals.
Alternatively, the bearings may be lubricated by external systems, either by
connections from the engine lubricating system, through a fine filter, or by an
independent system of pumps, cooler, filters, oil sump and alarms. Independent
systems allow a choice of turbine oils to be used. All lubrication systems must maintain
adequate lubrication when tilted to an angle of 15° in any direction or a temporary tilt
of 221°.
Updraught in the exhaust system may cause the turbocharger to rotate while the
engine is out of service. Care must be taken to ensure no damage is caused to bearings
from lack of lubrication or by vibration transmitted from other systems while the
engine is stationary. Sleeve type bearings are less susceptible to this damage. With large
turbochargers fitted with sleeve bearings lubricated from the main engine system, a
gravity tank must be fitted to ensure adequate lubrication in the event of pump failure
or stopping of engine driven pumps.
Some models of turbochargers are designed with both rotor shaft bearings situated
on the centre section, allowing the turbine and compressor to overhang at each end.
This may reduce the overall size and weight and economise in the shape of air inlet and
exhaust connections. It will require larger bearings and the oil to them must be cooled.
Central bearings are less accessible for maintenance.
do not have their turbine casing water
cooled. In general construction they are similar to those already described, but some
changes are necessary to accommodate the higher gas temperatures. The casing is
manufactured from cast steel and insulated to conserve heat and meet safety
requirements for hot exposed surfaces. Fig. 3.5 shows a cross-section of an Asea Drown
Doveri, ADD, VTR 4 uncooled turbocharger as fitted to many large two-stroke
The turbine bearing will require cooling: in this case it is protected by insulation and
also has a water cooled casing which maintains acceptable oil temperatures.
Higher working temperatures will eliminate the possibility of condensation and acid
3.5 Uncoo/ed turbocharger (Asea Brown Bover; ABB VTR4)
corrosion within the gas casing. The final outlet gas temperature may be up to 10°C
lligher than that from cooled turbines and this is attractive for large, slow, two-stroke
engines which produce limited exhaust temperatures. This allows more efficient uSe of
• waste heat boiler and therefore adds to the total heat recovery and thermal efficiency
of the plant.
Under operating conditions turbocharge systems may become fouled, causing reduced efficiency, loss in power and
maybe surging. Compressor surfaces can be contaminated by oil and dust drawn from
Ihe engineroom
passing through the intake filters and adhering to
working surfaces. This gives loss in compressor efficiency and chokes the charge air
can be removed while the engine is running by water washing. A
smalI spray of clean, fresh water is injected into the air stream at the compressor inlet:
water droplets are drawn through and scour the surfaces clean. Water drains must be
open and the quantity of water limited since some ofthis, together with contaminants,
may pass into the engine cylinder. Additives to improve the cleaning effects of the water
must not be used in trunk piston engines where it is possible for them to pass through
the cylinder to the crankcase oil. Hydrocarbon
based additives must not be used since
there may be an explosion risk, and they may also add further combustibles to the
cylinder which have bypassed the governor control.
Fouling of the exhaust system and turbine is the buildup of the products from
combustion of fuel, used cylinder oil and its additives, ash and any other noncombustibles present in the fuel. These will create hard deposits which interfere with gas flow,
may cause local overheating, reduce turbine efficiency and even destroy balance of the
Water washing may also be used to clean the gas side of the system. A small water jet
sprayed into the gas flow, upstream of the protection grids for the turbine, will dampen
and soften the deposits, reducing adhesion and enabling them to be blown clear by the
gas stream. The introduction of water to very hot surfaces can cause thermal shock, so
before water washing the turbine engine power must be reduced so that temperatures in
the system are lowered. Water drains must be open and great care exercised since
removal of deposits may affect the balance of the turbine rotor.
Examination and more comprehensive cleaning of parts by physical and chemical
means can be carried out during overhaul.
In modern two-stroke
engines a charge air cooler is necessary.
Compression will raise the air temperature and a charge air cooler is fitted to reduce the
temperature of the air between the turbocharger and the engine inlet manifold, causing
increased air density at lower induction temperature. The engine is maintained at safe
working temperatures and the lower compression temperature reduces stress on piston
rings, piston and liner.
Increased density will raise scavenge efficiency and allow a greater mass of air to be
compressed; more fuel may now be burned giving an increase in power.
Fig. 3.6 shows a section of a charge air cooler. The air makes a single pass through
the cooler and, for efficient cooling, its velocity should be low and cooling area large.
This is obtained by making the air inlet connection divergent; the outlet is convergent
to restore air velocity after cooling.
3.6 Charge air cooler
Condensation of moisture in the compressed air will occur during cooling and a
drain is fitted to the outlet side air casing to allow this condensate to be removed. A
moisture eliminator may also be fitted to remove entrained water droplets from the air
stream. The drain should be kept open and its discharge noted; this will also indicate if
a cooling water leak has occurred.
The cooler consists of a tube stack of aluminium brass tubes rolled and solderbonded into two brass tube plates. Cast iron water boxes are attached to the tube plates
and allow salt water circulation within the tubes to make two passes. One tube plate is
secured to the casing while the other is free to move axially as thermal expansion
occurs. The air seal is maintained by means of a fitted rubber joint ring. An air vent is
fitted to the top water box to remove air which may have been released from the salt
water system. Corrosion plugs may be fitted within the water space.
Thin copper fins are soldered to the outside of the cooler tubes. The air will pass
between these plates which greatly increase the area of heat transfer. There are two side
plates of mild steel or aluminium alloy.
Temperatures and pressures are recorded at each inlet and discharge. Discharge air
temperature should not exceed 55°C since engine temperatures-notably
the exhaust
temperatures-will increase, with loss in efficiency due to reduction in air density.
Air at very low temperatures wiIl also cause thermal shock when in contact with hot
liners and pistons.
Some medium speed engines' turbocharge systems may be adapted to improve low
speed operation. A by-pass valve may automatically return some charge air to the
compressor inlet to improve acceleration when running up to speed. Engines designed
to operate with high power at reduced speed may have an excess of charge air at full
speed and require a blow-off valve to open at 85% full speed.
In recent years the efficiency of turbochargers and their systems has been increased
considerably. In addition to increased engine power, greater stability is possible at low
speeds. Part of the exhaust gas energy may be used to drive power units which increase
the overall thermal efficiency of the plant.
Two stage turbocharging is currently only used in high speed four-stroke engines. Two
turbochargers are fitted in series, each acting as one stage of the charge air compressor.
They are matched to avoid surge at all speeds. They are of normal construction and
there is no mechanical linkage.
The first stage compressor is driven by the second stage exhaust gas turbine. The air
then passes through a first stage charge cooler before continuing to the second stage
compressor, which is driven by the first stage gas turbine. The air proceeds through a
second stage charge cooler and then to the engine inlet.
The advantages of two-stage compression with an intercooler are outlined on page
108. Two stage turbocharging gives higher charge air pressure and mass. Both charge
air coolers can be fresh water cooled and if a central cooling system with temperature
controls is utilised, further total energy savings may be made.
Undercooling of air occurs when it is cooled to a temperature below the dew point at
that pressure. To prevent undercooling and excessive condensation, air temperature
should not be taken below ZQ-Z5°C.
Some measure of cooler efficiencycan be ascertained from the difference between air
discharge temperature and cooling water inlet temperature under normal running
conditions: a rise in this indicates fouling of the cooler. An increase in the air pressure
drop indicates fouling of the air passages, while an increase in water pressure drop
indicates fouling of the water side.
When necessary, charge air coolers can be cleaned while out of service. Water side
may be cleaned by removing water boxes and brushing through tubes. If this does not
remove the scale, acid cleaning may be carried out, after which it must be flushed
through. After cleaning, a pressure test should be carried out on the water spaces.
Air passages may be blown clear with a compressed air or steam jet and specially
shaped brushes may be used. If these do not remove the dirt, the stack may be
immersed in a hot degreasing fluid and then blown clear.
Some medium speed and high speed engine systems circulate the charge air coolers
with fresh water. Additionally, to improve total heat recovery and prevent undercooling at reduced engine load or speed, the charge air cooler is split into two parts or
Tappet clearances are necessary to allow for thermal expansion of the valve spindle
length at working temperature, and to ensure that positive closing of the valve
continues as it wears or seats during use. Clearances should normally be set while the
engine is cold and the cam follower is off the cam peak. Wear in parts of the valve
operating gear will tend to increase clearances.
Excessive tappet clearance will cause the valve to open late and close early in the
cycle and will reduce its maximum lift. It will also cause noise, and eventually damage,
from the impact of working surfaces.
Insufficient clearance will cause the valve to open early and close late with increased
maximum lift. In extreme cases, it may prevent the valve from closing completely as it
expands or beds in. This in turn will cause hot gases to blow past valve faces, causing
valve burning, low compression etc.
type of valve is now fitted to most large two-stroke engines. The pushrod, rocker lever
and tappet are all replaced by an hydraulic system. The exhaust cam follower raises the
piston of a single acting, high pressure oil actuating pump. As the piston is raised it
delivers oil at high pressure through a reinforced pipe to a hydraulic piston attached to
the top of the valve spindle, forcing the valve to open. When the cam peak has passed, a
helical return spring forces the follower down onto the earn profile, lowering the
actuating piston and releasing the hydraulic oil pressure. This allows the valve spring to
force the valve closed, returning oil from its hydraulic cylinder. Each hydraulic valve
system is continuously charged with oil from the main engine lubrication so that any
leakage is immediately made up.
Some early valves ofthis type had helical valve springs, but present practice is to use
an air spring or pneumatic piston in which air under pressure is further compressed as
the valve opens. This pressure energy is stored in the air spring and used to close the
valve. Air springs are charged continuously from the main air start supply through a
reducing pressure valve. This ensures the appropriate exhaust valves are closed before
the engine can be started.
This system has many advantages over push rods. It reduces overall weight, allows
the engine designer to place the camshaft in the most effective position, reduces the
power absorbed and inertia losses, eliminates side thrust on the valve spindle and
spring surge or vibration. It also allows the valveto be rotated by the gas stream as
described above.
AIR INLET VALVES are necessary in four-stroke engines and similar in
construction and operation to a spring loaded, pushrod type of exhaust valve. Inlet
valves should have an area equivalent to that of the exhaust valves. The passage of cool
inlet air will maintain them at moderate temperatures so the design of valves, seats and
pockets will not require such intense water cooling as exhaust valves. Seats and
landings do not require such high-temperature and corrosion resistant materials.
Rotocaps or guide vanes may be fitted to maintain even temperatures and remove
deposits. Air inlet valves may be fitted directly in cylinder covers without valve cages.
Engine Parts
Sectional views and brief descriptions of complete engines were given in Chapter I and
details of the major parts, their operation and maintenance can now be given in greater
detail. Size and shape of particular parts will differ according to engine type and
manufacturer, but most of this information will be general to all.
These are designed to give an extensive working life with long periods between
overhauls. They must maintain low wear rates and friction losses from the sliding
motion of the piston rings under fluctuating pressure and temperature. Materials for
liners must provide adequate strength and fatigue life, readily transfer heat, resist
abrasion and corrosion, be able to retain a film of lubricating oil on working surfaces,
and have a rate of thermal expansion compatible with adjacent parts.
To meet these requirements at limited initial cost,liners are cast in pearlitic grey cast
irons to which alloying elements such as vanadium and titanium are added to enhance
strength, wear and corrosion resistance. Actual specifications depend upon the engine
type, its rating and fuel. Chromium plating of the internal surface can be carried out to
reduce wear and corrosion further. The chrome must either be deposited as a porous
structure or else be etched to retain lubrication. Chrome plated liners have extended life
but the initial cost is greater.
Chromium plated piston rings must not be used with plated liners.
The upper end of the liner forms a flange of sufficient strength to support it. This
flange is secured between the cylinder cover and the jacket or engine entablature,
leaving the lower end of the liner free for thermal expansion. The liner thickness must
give adequate strength to resist the internal gas load but is limited by the necessity to
transfer heat rapidly to reduce the thermal stress. Thickness may be reduced towards
the lower end where pressures and temperatures are less, but external flanges and
bossing may be added to accommodate seal rings, lubricator connections, etc. Some
earlier model liners had steel hoops or rings shrunk on or a close fit to the outer
surface: this increased strength with only moderate increases in size and weight.
In large two-stroke engine liners (Fig. 4.1) scavenge air ports will be cut or cast in
near the lower end. Engines using loop or cross scavenge systems will also need exhaust
ports. All port edges are shaped to direct the flow. Cylinder liners in four-stroke
engines require no ports and are of more simple shape.
To remove excess heat, limit thermal stress and maintain a surface temperature at
which an oil lubrication film can be retained,liners require an external cooling system,
described in Chapter 5.
This is often referred to as jacket cooling and is carried out by circulating fresh water
between the outer surface of the liner and the cylinder jacket or sleeve in which it is
supported. The water space extends over the upper part length of the liner. Silicone
rubber rings fitted into grooves in the liner make a watertight seal at the lower end while
still allowing freedom for thermal expansion ofthe liner. The upper end is sealed by the
landing of the liner flange.
Water enters at the lower end of the jacket, flowing upwards and leaving at the top to
pass on to cool the cylinder cover. Passages may be fitted within the cooling spaces to
ensure positive circulation and assist venting of any entrained air.
BORE COOLING is the term used when a number of individually small holes are
bored within the thickness of a part so that water can be passed close to the heated
surface (Fig. 4.3). Straight holes drilled at an angle through a curved thickness give
very effective cooling with little loss in strength of the part. As cooling takes place near
the hot surface, thermal stress is greatly reduced.
Bore cooling is used in most modern liners to maintain acceptable temperatures in
the liner at its top flange where the metal thickness is greatest. The small holes angled
upwards maintain the velocity of the cooling water.
CYLINDER JACKETS For large engines these consist of individual iron
castings, supported on the engine framing or entablature and bolted together to form a
rigid block into which the cylinder liners are fitted. The jackets are held in compression
by the tie bolts (see below in this chapter).
In smaller engines the jackets may be formed in one large casting or fabricated in box
form and termed the cylinder block. It may act as the main strength member of the
engine construction. Cleaning access to the cooling spaces must be provided.
Cylinder liners require adequate lubrication to reduce piston ring friction and wear.
The oil film also acts as a gas seal between liner and rings and as a corrosion inhibitor.
Some oil is forced into piston ring grooves, reducing wear on the piston surfaces and
rings, and keeps them free to move within their grooves and follow the contour of the
4.3 Bore cooling of cylinder liners: (a) section at X; (b) section along one bore
hole; (c) section of a development through centrelines o.f bore holes
In large crosshead type engines the cylinder space is completely isolated from the
crankcase and a separate lubricating system is fitted to supply a measured quantity of
oil to each liner. Special cylinder oils are used for engines burning heavy fuels: they
have a high alkalinity (usually 70 TON) to neutralise acids formed from the
combustion of sulphur present in the fuel. Oils must maintain their viscosity at high
temperatures, resist oxidation and carbonisation,
spread readily yet cling to working
surfaces. They prevent scuffing and wear while keeping piston rings free from carbon
Alkaline agents will be expended during use and any. oil drainings from the cylinders
must not be reused.
Oil is injected through a number of holes drilled in the liner, usually six or eight of
these displaced circumferentially around the liner at a chosen vertical position within
the piston stroke. Oil is supplied by pressure pulses from positive displacement
mechanical lubricators driven from the engine camshaft and regulated to deliver oil at
the required rate. By linking the pump drive to the engine throttle setting levers, the oil
quantity can be related to the power produced. Lubricator quills are connected to the
oil holes and each contains a non-return valve to prevent blowback of hot gases.
Ideally lubricators should be timed to inject oil between the piston rings as they pass.
In practice, due to the elasticity of the .system this accuracy is difficult to achieve and if
4.4 Cylinder lubricating quill with accumulator
timing is out of phase the oil may be swept away. In consequence timing is generally not
attempted. Accumulators may be fitted to the lubricator quills so that a quantity of oil
will enter each time the cylinder pressure drops below that in the accumulator, thus
ensuring oil does not enter the cylinder when exposed to hot gases.
The vertical position of the lubrication points will depend upon the engine design.
They should be clear of the combustion space with its high pressures and temperatures
but should also be clear of ports in the liner since unused oil may be lost through these.
Distribution of oil around the liner circumference may be aided by oil gutters adjacent
to the lubricator points and angled downwards to assist flow by gravity, while reducing
the piston ring chipping effect. Lubricating oil is spread over the length of the liner by
the piston rings during their stroke.
Some large two-stroke engines use multi-level cylinder lubrication.
Some oil is
injected near the top of the liner to neutralise acids formed during combustion; the
majority of the cylinder oil is injected at 10% of the stroke below this to ensure
complete oil wetting of the liner surface at mid-stroke, where the piston rubbing speed
reaches its maximum and abrasive wear may occur.
Mechanical lubricators should be operated before starting the engine to ensure
priming of the connections and injection of oil for the first engine' movement.
The supply should be temporarily increased during running-in periods for new
cylinder liners or new piston rings.
In opposed piston or exhaust piston engines, two sets of lubricator points are fitted
to each liner, one set situated within the stroke of each piston. Due to the higher
operating temperature, it is usual to supply a relatively greater quantity of oil to the
upper or exhaust piston.
In trunk piston engines some cylinder lubricators may be fitted as in crosshead
engines. Due to smaller cylinder bores, there are usually fewer points around the liner.
Since most of these engines are four-strokes, there will be no ports in the liner through
which oil may be lost and temperatures are reduced during the induction stroke,
making lubrication less demanding. The majority of the cylinder lubrication will be
effected by oil splashing from the crankcase into the open lower end of the liner. This
would cause uneven excessive lubrication so oil scraper or control rings are fitted to the
piston skirt to spread this circumferentially and return excess oil to the crankcase. In
some engines, oil may be bled from the piston cooling return, to pass through small
holes in the piston skirt and increase cylinder lubrication. Used oil from the cylinders
will also drain into the crankcase and consequently a common oil is used for both
cylinders and crankcase. The oil must be compatible for cylinder lubrication as well as
for cooling and crankcase lubrication of bearings. For trunk piston engines running on
heavy fuel the oil must be alkaline (30/40 TBN). Alkalinity will be reduced with
contamination from combustion products and must be monitored. A stable value
should exist when loss in alkalinity is balanced by that in the fresh oil added as 'make
up', to maintain the crankcase oil level.
Operation of an engine with insufficient cylinder lubrication will cause high wear
rates to liner and piston rings. Corrosion of the liner may increase when burning heavy
fuel. Loss of oil seal to the piston rings will cause blowpast of hot gases causing local
overheating, rapid breakdown of surfaces and possibly piston seizure. In trunk piston
engines there is also danger of a crankcase explosion. Excessive lubrication will cause
carbon deposits, piston rings sticking in grooves allowing possible breakage or
blowpast. There will be fouling of the exhaust system including turbocharger and
contamination of scavenge spaces.
Wear in a cylinder liner is mainly due to friction, abrasion and corrosion, although
under severe conditions adhesion may also occur. Each ofthese may have a number of
Frictional wear takes place between the sliding surface of cylinder liner and piston
rings. It will depend upon the materials involved, surface conditions, efficiency of
cylinder lubrication, piston speed, engine loading with corresponding pressures and
temperatures, maintenance of piston rings, combustion efficiency and contamination
of air or fuel.
Corrosion occurs mainly in engines burning heavy fuels, particularly with high
sulphur content. It is caused by acids formed during combustion and these must be
neutralised by the use of alkaline cylinder oil. SuIphuric acid corrosion may be caused
in the lower part of the liner if the jacket cooling water temperature is too low. This may
allow vapour present after combustion to condense. The moisture formed will absorb
any sulphur present to form sulphuric acid. This can be prevented by maintaining
jacket temperatures above the corresponding dewpoint.
Water vapour will be present from the combustion of hydrogen together with any
water present in the fuel. It may be increased if water passes from the charge air cooler.
Abrasion may take place from the products of mechanical wear, corrosion and
combustion-all of which form hard particles. Ash may be present in some heavy fuels,
as well as fines (aluminium compounds added as a catalyst during the refining process
and not fully removed from the residual fuel) which may act as abrasives.
Adhesion or Scuffing is a form of local welding between particles of the piston ring
and the liner rubbing surface, resulting in very rapid wear. If may occur if the
lubricating oil film between ring and liner is removed due to excessive temperature,
insufficient supply or incorrect distribution of oil, piston blowpast, etc. Engines
operating on some low sulphur grades of fuel may be prone to scuffing damage.
Cylinder liners should be gauged internally at fixed intervals during
overhaul (6~8000
hours) to measure accurately the increase in bore. Continuous
records of gaugings should be kept for each cylinder.
The liner must be cleaned and inspected. General appearance of the surface may
show whether lubrication has been adequate. A careful examination must be made for
surface cracks, particularly around the combustion space.
The liner is now gauged with a micrometer and extension bar which has been
calibrated against a master gauge. The liner should preferably be cold, but ifthis is not
possible the gauge must be at the same temperature as the liner to cancel expansion
Gaugings are taken at a number ofvertical positions (4 to 6) over the area swept by
the piston rings. Readings are taken in fore and aft and in athwartships directions. To
ensure readings are taken at corresponding points, a template may be used. Gauging
figures are noted as total wear from original and mean rate of wear since the last
recording was made.
The pattern of wear over the length of the liner will differ according to engine type
but in single acting, two-stroke engines it tends to be greatest at the top of the stroke
adjacent to the combustion space, where pressure and temperature are greatest. This
reduces towards the lower end of the stroke, but will increase at the exhaust and
scavenge ports when relative pressure on port bars is increased and blowpast may
remove lubricating oil film (Fig. 4.5).
4.5 Cylinder liner wear pattern
The rate of wear varies over the life of the liner. It is high during the initial running-in
period after which it should reduce to an almost constant rate for most ofthe useful life
of the liner. Finally the rate will progressively increase as wear becomes excessive (Fig.
Normal wear rates differ but an approximate figure of 0.1 mm per 1000 hours is
acceptable for large engines. Wear is increased if the engine is overloaded. Maximum
wear before renewal is usually limited to 0.6-0.8% of original bore diameter, or less
when the manufacturers advise. The time when this figure will be reached can be
anticipated from wear records for advance ordering of replacements.
4.7 Gauging a cylinder liner
4.8 Record of cylinder liner wear
After gauging, any ridges on the liner should be ground off. These may be evident at
the top of the ring travel and at port bars. Ridges may be due to broken piston rings or
where the piston has been raised to readjust compression. If new piston rings are fitted
the liner should be deglazed. To allow running-in, cylinder lubrication should be
increased temporarily and maximum load limited. Ports must be cleaned, sharp edges
removed, lubricators tested and a close inspection made before putting back into
If the cylinder is operated with excessive wear, the rate of wear will rapidly increase.
Gas blowpast may remove the lubricating oil film, piston rings may distort and break,
piston slap may cause scuffing. Compression is reduced causing incorrect combustion
with fouling of the exhaust system. Unburned oil may be blown into scavenge spaces
giving rise to scavenge fires.
circumference. After the cylinder cover is hardened down, a water test must be carried
•• on the jacket, lubricator points and seal rings.
The new liner should be regauged after final landing to check any distortion. These
pugings should be recorded to assist later esimates of wear, etc. Lubricator
aJDnections are refitted and lubricators tested.
New piston rings should be fitted with a new liner and all running-in procedures of
Kduced fuel and increased lubrication should be carried out.
For engines in which cylinder liners land on a head ring on the entablature, the lifting
Ib'ongback may be secured directly to the head ring and this, together with the liner,
may then be jacked from the entablature without the necessity for a crossbar.
The cover (or cylinder head) forms the top part of the combustion space and it must be
ofsufficient strength to withstand the gas load at maximum pressure. The cover lands
on the top of the cylinder liner flange and is secured to the cylinder block by a number
ofcover studs and nuts. These are tightened hydraulically to maintain a gas-tight seal
under the arduous fluctuating pressure and temperature conditions.
The cover must also locate and support all the valves required to operate the engine
together with their securing studs. In a large two-stroke engine there will be a large
central exhaust valve and cage, two or three fuel injectors, a starting air valve, relief
valve and indicator connection. Medium speed four-stroke engines generally have two
exhaust valves and two air inlet valves per cylinder, one centrally placed fuel injector,
an air start, pressure relief and indicator valves.
It can be seen that the cover is of intricate design: it should transfer heat readily,
resist bending, be symmetrical in shape and have a coefficient of thermal expansion
similar to the liner. Large two-stroke covers are machined from solid steel forgings
drilled to give bore cooling. They are circulated with fresh water from the jacket
cooling system which maintains moderate temperatures and allows intensive cooling
ofthe exhaust valve seating while making separate fuel valve cooling unnecessary. (See
Figs 4.10 and 4.11).
4.10 Cylinder cover for two-stroke engine (M AN-B & W)
4.11 Cylinder coverfor four-stroke engine (MAN-B & W)
Earlier large engines had covers in two parts, the part holding the valves being of cast
iron held in position and supported by a cast steel strongback to give added strength
with effective water cooling.
With their greater number of valves, four-stroke engine covers are usually formed
from a casting which may be of pearlitic or nodular graphite cast iron, or in some cases
cast steel. Deep sections are used to resist bending under the peak pressures in highly
rated engines. The flame plate is of limited thickness to ensure efficient cooling while
the top deck plate gives adequate strength. An intermediate deck may be included in
some designs. The deep box-like construction allows air inlet and exhaust passages
(and in some cases fuel pipes) to be accommodated within the main casting.
In crosshead type engines a diaphragm is fitted to isolate the lower surface of the
cylinder and scavenge box from the crankcase. In large two-stroke engines the
diaphragm may form part of the cylinder jacket.
Isolation of the cylinder space from the crankcase prevents contamination of the
crankcase lubricating oil by residues from the combustion of heavy fuels, cylinder and
ring wear, corrosion or gas blown past the piston rings. It allows the use ofspecial high
alkaline cylinder oils (see cylinder lubrication), while the main crankcase oil can be
selected for better bearing lubricating and cooling properties.
The piston rod must pass through the diaphragm and a piston rod gland is fitted.
This consists of inward sealing metallic or teflon packing and oil scraper rings. Each
ring has three or four segments which are a good fit on to the piston rod surface and are
beld by a coiled garter spring. There must be sufficient circumferential butt clearance
between segments to allow contact with the rod.
The gland consists of two sections; rings in the upper section act as seal rings for
scavenge pressure and scrape off any residue or dirt from the piston rod on its
downward stroke. This contaminated oil residue should be conveyed to the sludge
tank. The lower section rings act as oil control rings to scrape excess crankcase oil from
the piston rod during the upward stroke. This oil is returned via drains to the crankcase
system. A void or vent space is left between the two sections and the drain from this
should be inspected regularly to ascertain the efficiency of the gland.
Maintenance of the gland consists of maintaining correct clearances of ring
segments in circumferential (butt), axial and radial directions; checking tension of
garter springs; keeping drains and vents clear. Particular care must be taken during
removal of piston and rod to prevent damage to the gland.
Incorrect, or lack of, maintenance may lead to contamination of the scavenge space
with oil, loss of scavenge air, contamination of crankcase oil and damage, with possible
overheating of the piston rod, leading to danger of a hot spot within the crankcase and
risk of explosion. Diaphragms in some engines contain water cooled spaces.
4.12 Piston rod gland and diaphragm
Piston cooling and cylinder jacket cooling must be maintained for a period after the
engine is stopped in order to allow a gradual reduction in temperature and consequent
_rmal stress.
Fig. 4.13 is a section ofa MAN-B & W K90 MCC engine, oil cooled piston. The
chrome-molybdenum steel crown is cast in one piece and has adequate internal cooling
spaces while an inner ring supports the crown and transmits the gas load directly to the
piston rod flange to which it is rigidly bolted. A short cast iron skirt is secured to the
..werside of the piston, sealing off the cooling space. The Inconellayer welded to the
top of the crown is not used in all MC range engines.
Cooling oil enters from the central tube within the piston rod and flows under
the centre of the crown before passing through bore cooling holes in a radial direction
to the toroidal space at the piston edge, swirling down behind the piston ring grooves
before returning to the crosshead by the concentric space outside the tube in the
piston rod.
Four piston rings are fitted in grooves with chrome-plated horizontal surfaces.
A water cooled piston, as fitted in some Sulzer RTA Series I engines. is shown in Fig.
4.14. The piston crown is cast in alloy steel and has internal bore holes which extend
close to the working surfaces including the outer edge and behind the piston ring
grooves. Water pressure together with the shaker action ensures circulation within
these bores. Water enters and leaves through telescopic pipes attached to the piston;
the sealing glands and standpipes for these pass through the diaphragm to a casinz
which completely isolates them from the crankcase to avoid the possibility of
contamination in the event of a leakage.
Five piston rings are fitted in chrome-plated grooves. A short cast iron skirt of
slightly larger diameter containing a bronze wearing ring is secured between the piston
and rod. The assembly is secured to the piston rod flange by stud-bolts waisted for part
length and fitted in extension sleeves to give greater resilience.
Most recent models of Sulzer RT A Series 2 engines are oil cooled, oil passing up the
bore in the piston rod and projected by jets into the bore holes. Oil leaves down the
central tube to the crosshead.
Surface temperatures of working parts within the combustion space of a diesel
engine will vary with design. In two-stroke engines they tend to be symmetrical because
of the centrally positioned exhaust valve. Maximum temperature and thermal cycling
will affect the wear and tear on parts and their maintenance.
Fig. 4. I5 shows temperatures at critical positions within the combustion chamber of
a MAN-8 & W 6826 MC engine operating at full power on heavy fuel.
4.15 Temperatures in combustion space (MAN-B
& W)
Before removing the piston it will first be necessary to remove the cylinder cover, valves
and connections. Carbon deposited around the top of the liner should be cleaned off
before attempting to lift the piston.
Engine turning gear should be used to turn the piston to top dead centre position.
Screw threads in lifting holes in the piston crown must be cleaned and the piston lifting
plate or brackets securely bolted into place; the crane is then attached. After removal of
its locking device, the piston rod nut or nuts securing it to the crosshead are slackened
hydraulically and removed.
The weight of the piston and rod is taken on the crane and then turning gear is used
to turn the engine and lower the crosshead clear of the piston rod. Care must be taken
to protect all threads from damage. The piston and rod may now be lifted by the crane.
In some engines the piston rod gland is freed from the diaphragm and lifted with the
rod, in others special devices prevent damage as the rod is lifted. Where piston cooling
connections are fitted, care must be taken to avoid damage or misalignment.
When piston and rod are clear of the liner they may be landed for removal of rings,
inspection and cleaning. The piston should be inspected externally for wear, corrosion
and cracks. The crown should be gauged to reveal any burning or corrosion. Careful
examination should be made for fatigue cracks. Piston ring clearances are gauged and
then the rings removed by means of the retractor or tensioner. Ring grooves 8R
cleaned, examined and gauged for wear or taper, particular care being taken witlt
flatness of landing surfaces.
Internal cooling spaces of the piston should be cleaned and surfaces inspected, whidl
will require removal of the piston rod and seals. Locking devices on piston studs arc
dismantled and nuts slackened hydraulically. Studs must be inspected and tested. AI
surfaces must be cleaned and new seals fitted. After reassembly, a pressure test of
cooling spaces should be carried out.
The piston rod should be inspected for wear and cracks, particularly at its
connection with the crosshead.
If new piston rings are to be fitted, they must be gauged for butt clearances in the
unworn part of the liner, and axial clearance and free movement in grooves must be
checked. Ring gaps should be spaced at 180 in alternate grooves.
After cleaning and gauging of liner, etc the piston and rod may be replaced. A guide:
ring is placed at the top of the liner to assist entry and all necessary surfaces are
lubricated. During lowering, equal care must be taken of threads, glands and cooling
systems. The maker's instructions should be carried out, particularly regarding
slackening and tightening of all bolts and studs. Locking devices must be fitted
The utmost safety must be observed at all times. Lifting gear must be kept in good
order and all parts made secure.
Skirts are fitted to pistons of both two- and four-stroke engines, but for different
reasons. In large crosshead type, two-stroke engines using uniflow scavenge, a short
skirt may be fitted to act as a guide to stabilise the position ofthe piston within the liner.
It is made of cast iron, usually very slightly larger in diameter than the piston to ensure
that if any piston slap takes place the steel piston will not contact the liner wall or
damage the surface. Soft bronze bearing rings may be caulked into the skirt, standing
proud and assisting the running-in process. These rings can be renewed if necessary.
For two-stroke engines operating with loop or cross scavenge, long skirts are
required to blank off the scavenge and exhaust ports in the liner when the piston is
above these.
In trunk piston engines the piston skirt has the gudgeon pin fitted within it and
therefore transmits the gas load from the piston to the gudgeon, or top end bearing.
Due to the absence of crosshead guides in these engines, the skirt acts as a guide for the
piston and transmits the side thrust from the connecting rod to the liner walls. Skirts
will require lubrication and may have oil control rings fitted.
Four.stroke skirts may be made from aluminium alloy or cast iron. The gudgeon pin
fitting causes the mass of metal within the skirt to be unsymmetrical, leading to unequal
expansion when hot. To ensure circularity of the skirt at working temperature it is
machined to be slightly oval when cold.
A piston rod acts as the main strut to transmit the heavy gas load downwards from the
piston to the crosshead and running gear of the engine. The length of the rod will
depend upon the length of engine stroke and its design. It is machined from a steel
forging and is round in section for strength and rigidity. There is a flange at its top end
where bolts or studs attach it to the piston. The lower end is attached to the crosshead,
either by a bolted flange or by passing through the crosshead with a nut underneath.
The piston rod passes through the diaphragm gland and must have a smooth
working surface to allow bedding-in of gland packing rings. This may be surface
treated to minimise friction and wear at the gland.
For oil-cooled pistons, the coolant is transmitted upwards through a centre bore
hole in the piston rod. It returns through a concentric tube fitted within this.
These are fitted to practically all modern four-stroke engines and may vary in size from
small high-speed engines up to cylinder bores of about 600 mm in the medium speed
Medium speed engines of modern design have composite pistons which comprise a
thin-sectioned alloy steel piston crown supported on an aluminium alloy skirt, or
trunk, giving a strong, light structure capable of resisting high temperatures and
corrosion but with low inertial losses. The steel crown may be cast or forged and is
shaped to give adequate cooling spaces. Inner surfaces are cooled by the shaker method
in which oil from the gudgeon pin bearing passes to the piston and relies on the
reciprocating motion to splash it against hot surfaces and to force it through passages
where necessary. The oil drains down inside the skirt to the crankcase.
Piston compression rings are fitted in grooves in the cylindrical part of the steel
crown. Landings of grooves may be plated or surface hardened to reduce wear. The top
surface of the crown may be recessed to allow clearance for inlet and exhaust valves
which will be open during the overlap period of the engine cycle. A number of
prestressed bolts are screwed into the crown to secure it to the skirt.
The skirt of silicon-aluminium alloy will contain the gudgeon pin bearing and its
outer diameter may be oval to allow for thermal expansion (see above). It must be of
sufficient length to dissipate the side thrust to the liner. Oil control rings may be fitted at
the top of the skirt to remove excess lubricating oil splashed from the crankcase.
Small engines may have cast iron pistons and skirts manufactured from either grey
iron or spheroidal graphite iron, giving strength and wear resistance with low thermal
Aluminium alloy pistons have been used and can be cast integral with the skirt for
economy. They have low inertia and give good strength and wear under moderate
conditions. The material work-hardens to give low ring groove wear, although an
insert may be required at the top ring where higher temperatures and corrosion occur.
4.17 Trunk piston (M AN-B
& W)
The high coefficient of thermal expansion requires greater piston clearance. These
pistons are not used when burning heavy residual fuels as their strength, corrosion and
abrasion resistance is reduced at elevated temperatures.
Fig. 4.17 shows a composite trunk piston and skirt as fitted to the MAN-B & W
SS{64engine. It consists of a forged alloy steel crown ofthin section fitted by long studs
to a forged aluminium alloy skirt. Piston ring grooves are induction hardened to
sufficient depth for regrinding if necessary. Three compression rings are fitted in the
crown, the top ring being plasma coated and the lower ones chromium plated. One oil
control ring is fitted in the top of the skirt.
The piston is oil cooled by the shaker method, taking oil through a spring loaded
catcher which rides on the top of the connecting rod. Oil is thrown to the underside of
the centre of the crown and then passes through drilled holes into the space behind the
ring grooves. It finally returns to the inside of the skirt to drain into the crankcase.
Materials for diesel engine piston rings must have good strength, elasticity and wear
resistance with low friction, and must maintain these properties at high working
temperatures. They must resist corrosion, readily transfer heat, and have thermal
expansion compatible with the piston in order to maintain ring groove clearances.
For large two-stroke engines most rings are cast and machined from pearlitic grey
cast iron. This may include some chromium, molybdenum, vanadium, titanium, nickel
and copper. In some engines spheroidal graphite iron is used, which has greater
hardness and tensile strength.
The cross-section of piston rings is rectangular with small radii on all edges. This
allows an oil wedge to build up on the outer surface and prevents sticking at the back of
the ring groove. The section may vary adjacent to the ring butts.
Rings are machined either circular or to a cam shape from which they expand to
form a circle at working temperatures and avoid port chipping. A radial cut is made
and the shape of ring joint ends may be butt (vertical), which gives a robust joint for top
rings; scarfed (diagonal), giving better gas seal but less robust; or a form of lap or
bayonet joint which gives a good gas seal but is more vulnerable to breakage and is only
used in some lower rings (Fig. 4.18).
Piston compression rings are fitted in corresponding ring grooves machined in the
piston. They will bear heavily on the lower surface or land of the groove during the
power stroke and these lands must remain true or the rings will distort. To reduce wear
and fretting, they are either chromium plated, heat-treated or have hard wear rings
inserted. Sufficient thickness of treatment should allow regrinding of surfaces to fit
oversized rings if necessary as wear occurs.
Piston rings are retracted outwards to pass over the piston and released into the
grooves which have been oiled. The rings should then float freely within the whole
depth of the grooves. A special retractor should be used to give even bending without
twisting the ring section. Excessive bending should be avoided.
In use, piston rings convey heat to the cylinder liner and act as a gas seal between this
and the piston. They must be free to follow the liner surface irrespective of transverse
movement and will build up an oil wedge on the liner, reducing wear and spreading the
cylinder lubrication. The outward pressure is initially due to elasticity in the
compressed ring but is increased by gas pressure which acts on the back of the ring.
This pressure is greatest in the top rings; the temperature is also higher here and these
rings will have higher wear rates than others lower in the piston.
Ring clearances are necessary to allow movement and thermal expansion. Axial
clearance in the groove must be gauged to allow gas pressure to pass to and from the
back of the rings. The clearance will increase with wear and the landings must not be
allowed to taper. Clearances vary with engine size and rating but for a large engine may
be 0.4 mm for the top ring to 0.2 mm for lower rings.
Piston ring and groove wear will inevitably be uneven and rings must not be moved
from one groove to another during overhaul. If wear is excessive the used rings must be
discarded and new ones fitted.
Circumferential clearance at the ring joint is necessary to allow for thermal
expansion but should not allow excessive blowpast of gas. It is measured before fitting
by inserting the ring into the least worn (lower) part of the liner bore. This clearance
wi\1also vary with the shape of joint used, engine size and rating, but for large engines
approximate figures may be from 0.5% of cylinder diameter for moderate ratings to
t .0% for higher ratings. Circumferential wear of rings can be measured by reduction in
the width ofthe ring section and by increase in butt clearance at the corresponding liner
Excessive wear will reduce engine efficiency and may cause hot gas to blow past the
rings, removing lubrication, causing ring distortion and breakage.
Four-stroke engines with smaller cylinder bore and higher mean piston speeds
require rings with good wear resistance, low friction and inertia, and the ability to bedin rapidly to the liner surface. Cast iron rings may be used in most cases, but for highrated medium speed engines, to avoid scuffing and corrosion, ring coatings, inserts or
special profiles are used. Sets or stacks of matched rings are selected for particular
engines or fuel quality.
The top ring, which is subjected to the greatest load and temperatures, may be
chrome plated or plasma coated. Plasma coating consists of a layer of very hard
powdered metals such as molybdenum and chromium, fused to the rubbing surface of
the rings. Plasma coating is carried out with finely powdered metal in a stream of gas so
hot that all metal molecules become electrically charged causing fusion with the ring
surface layer to form a very hard rubbing surface. Chromium or plasma coated rings
must not be used in chromium plated liners. To assist buildup of oil wedge lubrication
and accelerate bedding-in and conformity, these rings may have their surfaces
barrelled or grooved. Their cross-section may be bevelled to balance working
Lower compression rings may have bronze inserts or copper plating to ensure rapid
Also referred to as scraper rings. They may be fitted
below the compression rings on the piston, or alternatively to the skirt. Their cross-
4.19 Trunk piston rings: (a) Barrel face used with plasma, chrome or copper
surface plating; (b) Taper face used with plasma. chrome or copper; (c) Serrated:
usedfor chrome or copper; (d) Plain: can be copper plated; (e) Grooved: copper
plated; (f) Bronze insert; (g) Single edge oil scraping; (h) Double edge with oil
drain holes; (j) Conformable: low inertia supported by spring loading
section is shaped with a downwards wedge or lip which spreads the oil on the upstroke
of the piston but scrapes off excess oil on the downstroke. They are light in section to
conform readily to the liner contour. Bevelled relief of the back surface of the ring may
be applied to reduce the ring pressure. Low-inertia oil rings are of reduced section,
supported by an outward loading spring. For improved life the scraping edges may be
chromium plated.
Rings may have drain holes which allow rapid removal of oil to the back of the ring
groove from where it can pass through holes to the inside of the skirt.
As the name implies, this rod connects the top and bottom end bearings, facilitating
conversion of the reciprocating forces on the piston into rotating power in the
Rods are machined from a steel forging shaped at each end to accommodate the
appropriate bearings. An oil hole is bored through the centre of the rod to allow
passage of lubricating oil between the bearings - downward in crosshead engines but
upward in trunk piston engines. The length of the rod is as short as the design allows, to
limit the overall height of the engine.
In slow, cross head type engines the rod acts mainly as a strut under axial loading and
is made of round section. Modern practice is to forge one-half of the top and bottom
end bearings integral with the rod. Earlier models had flat palms to which the bearings
were attached.
In medium and high speed trunk piston engines the rod is also subjected to
transverse bending forces from inertia of parts and an I shaped section girder is more
efficient. Medium speed engine connecting rods may be shaped with a flat flange bolted
4.21 Two-stroke connecting rod showing crosshead, top and bottom end bearings and
guideshoes (MAN-B & W)
connection, either adjacent to the bottom end bearing or at some point along the length
of the rod. This allows separation so that the piston may be withdrawn without
disturbing the bottom end or, conversely, the bearing can be dismantled without
affecting the top end or piston. This also reduces the minimum headroom required
above the engine for overhaul.
These are fitted to crosshead engines and are vertical sliding bearings which locate and
maintain alignment of the crosshead over the whole length of engine stroke. They are
subjected to fluctuating load from the transverse components of the connecting rod
reaction. Guide bars or surfaces are secured to the frames adjacent to the unit and have
either cast iron or steel bearing surfaces. Guide slippers (or shoes) are attached to the
ends of the crosshead and may be free to articulate; they are white metal lined with oil
grooves lubricated from the crosshead.
Guide clearances must be checked periodically and should not exceed 0.7 mm for a
large engine. Excess clearance will cause noise, wear on bearings and glands, uneven
loads and fatigue.
A crosshead consists of a forged steel block secured to the foot of the piston rod. It
includes the journal for the top end bearing which acts as a hinge by which the pistonthrust is deflected via the connecting rod to rotate the crank. As stated above, the
transverse component of this force is transmitted to the guides, and the guide shoes also
form part of the crosshead assembly. Further attachments are required for lubrication
and cooling connections (Fig. 4.20).
This bearing transmits the full gas load from the piston to the connecting rod and
crankshaft. The load fluctuates throughout the engine cycle. The two-stroke cycle
causes an almost continuous downward force, tending to squeeze the oil film from the
underside bearing surfaces. Movement of the bearing is limited by the small angle
through which the connecting rod articulates and reverses direction during each engine
revolution. Consequently there is insufficient sliding velocity between surfaces to
establish an oil wedge or spread an oil film. To distribute the oil over the lower bearing
surface, oil gutters are cut in an axial direction in the surface of the bottom shell. These
are spaced closely to allow an oil film to remain between them as the bearing swings. Oil
holes connect the gutters to the oil supply; they do not extend to the end of the bearing
but small drain holes may promote circulation of oil for cooling. Gutters are not
necessary in the top shell.
Most top end bearings are of the thin-walled shell type (insert bearings) with a steel
backing supporting low friction materials such as white metal or aluminium-tin alloy.
The latter will require a superfine surface finish on the journal and an increase in oil
supply pressure.
To ensure an ample supply of cool, clean oil to the crosshead bearings, the main
lubricating system supplies the crosshead directly by telescopic glands or swinging link
connections. Excess oil then passes to bottom end bearings, guides and to piston
Modern designs have a single bottom half bearing shell over the full length of the
crosshead which is machined in cylindrical form. A transverse slot at mid-length in the
top surface has the piston rod flange bolted to it. The top half of the bearing shell and
keep are similarly slotted to allow clearance of this flange as the bearing articulates.
Large diameter journals increase the effective bearing area, reducing pressure in the
lower half and eliminating bending.
Older crosshead designs used split top end bearings. The foot of the piston rod was
stepped and threaded to pass through the centre block of the crosshead, being secured
by a nut at the underside. Two identical bearings were situated, one on each side of this
block. The connecting rod was either forked or given a large palm to support these
bearings. This system acted as a simply-supported beam, the heavy piston load at its
centre causing bending in the supported journals. Heavy, fluctuating pressure was
concentra ted at the inner edge of each lower half bearing leading to fatigue and damage
in the bearing metal. Uneven wear between adjacent bearings may cause unbalanced
forces, misalignment and damage to running gear.
To overcome these disadvantages, flexible bearing supports were fitted allowing
them to align with the flexing of the journals or pins, reducing the concentration of
stress at their inner edges.
An alternative design had the piston rod bolted to the top of the crosshead, which
was profiled on its lower centre section. An additional bearing pad was placed between
the palm of the connecting rod and this profile.
In trunk piston engines the function of a crosshead is carried out by a gudgeon pin in
the piston skirt. The pin is of steel with hardened, polished bearing surface. It may be
fixed in the skirt or free to float and rotate, with circlips to limit axial movement.
The connecting rod has a single sleeve bearing bush pressed into an 'eye' at its top
end. The piston load is reversed during the induction stroke of the four-stroke cycle
and this allows oil to spread over the lower surface of the bearing. The oil supply in
these engines is made by way of oil holes in the crankshaft to the bottom end and thence
through the connecting rod. A non-return valve may be fitted in the foot of this rod to
prevent back-flow of oil under gravitational force at the bottom of each stroke.
These transmit similar loads to those in the crosshead, but there is the benefit of
continuous sliding movement between crank pin and bearing surfaces which
establishes an oil wedge and fluid-film lubrication. Large diameter crank pins will
improve the rubbing speed and the rigidity of the bearing which has its length limited
by the distance between each pair of webs on the crankshaft. If necessary, the webs can
be undercut (recessed) to give added length to the bearing, provided the metal
remaining still meets classification requirements.
In large two-stroke engines the bearings will be thin walled, with white metal lining.
Medium speed engines use a variety of thin-wall bearing materials such as
lead-bronze, copper-lead, white metal, aluminium-tin etc, all of which have good
fatigue resistance. Thin-wall or insert bearings have steel shells and must be forced
(nipped) into their keeps to prevent fretting on the back surfaces and to assist heat
transfer. Harder bearing materials may require a thin,' deposited overlay of very soft
alloy to aid rapid bedding in. They may need larger bearing clearances to allow for
thermal expansion, have hardened, polished journal surfaces and an increased supply
of finely filtered oil.
Some designs have the bottom end bearing as a separate unit, bolted to the
connecting rod. This allows the piston and rod to be withdrawn for overhaul without
disturbing the bearing. It also reduces the height required for removal of the piston.
Alternatively angled butts may be used. These are situated at an angle to the normal
transverse position and allow the connecting rod width to pass more easily through the
cylinder after dismantling. In this case the butts have serrated mating surfaces to
absorb side thrust and shear when the engine is running.
Vee type engines have the connecting rods from two cylinders connected to each
throw of the crankshaft. The most common method is to have each bottom end take
half the width of the crank (termed side by side). Other arrangements may be
articulated (master and slave), or fork and blade.
A crankshaft consists of a number of cranks or throws, which are rotated by piston
forces transmitted through the connecting rods and bottom end bearings. In-line
engines have one crank for each cylinder. Every crank is made up of two crank webs
joined by a common crankpin to which the bottom end bearing is fitted. Each web is
also connected to a journal to form part of a continuous shaft. A main bearing will
support the shaft at each journal. One end of the shaft will be flanged to transmit the
engine power and the other end is termed 'free'. Gear or chain drive wheels are added to
drive camshafts and other mechanisms if they need to be synchronised with the
crankshaft movement. Crankshafts of direct drive engines may incorporate a thrust
bearing (see below).
The crankshaft receives and transmits the full working power of the engine and is
subjected to fluctuating bending, torsion and shear stresses. The material used must
have high strength, a long fatigue life and form good bearing surfaces. Carbon or lowalloy steels are used, depending on size.
Medium speed and smaller engine shafts are usually manufactured as a solid forging
and may be made by the continuous grain process. In this a single billet of steel is forged
with its principal crystal grain in the longitudinal direction. Each crank throw is forged
from this so that the grain is maintained continuously along journals, webs and pins for
the whole length of the crankshaft. This process gives improved resistance to fatigue.
With short stroke engines, solid forged shafts are very rigid. They have overlap
between crank pins and journals which gives added strength. Oil holes are drilled
between journals and pins to transmit lubricating oil from the main bearings to the
adjacent bottom ends.
Slow running, two-stroke engine crankshafts are too large for a single forging and
most are of semi-built up construction (Fig. 4.22). Each throw of crankpin with two
webs is either forged or cast and then machined. Holes are bored in each web to allow
them to be shrunk on to adjacent journals which have also been forged and machined
separately. The webs are heated to the correct temperature and then cooled slowly,
shrinking them on the journals to form a continuous length of crankshaft with correct
angles between units to correspond with the firing order of the engine.
Shrinkage fits are well able to transmit the total engine torque without slippage. No
dowels or keys are ever fitted as these act as stress raisers and initiate fatigue cracks. No
4.23 Welded crankshaft for large engine (M A N-B
& W)
oil holes are drilled in built up shafts. Overlap is not possible since the webs require
adequate material for strength after the shrink fit holes are bored.
Large crankshafts can be of solid welded construction. A number of parts are forged
or cast separately and then welded together to form a complete shaft. Each part
consists of a crankpin and two webs, with a half length of journal on each web. Parts
are heat treated and rough machined; they are placed end to end at correct crank angles
and narrow-gap submerged-arc welding joins them to form complete journals. Post
heat treatment is followed by finish machining of journals and their central bore holes.
Welds are stress-relieved
and subjected to non-destructive
tests (Fig. 4.23).
Welded crankshafts allow a considerable saving in weight. Webs are of reduced
width and thickness allowing longer bearing length. Overlap permits large diameter
journals and pins, reducing bearing pressures. Shafts have a higher natural frequency
of torsional vibration which improves design.
Semi-built up shafts are adequate for current engine powers but solid welded shafts
will give advantages to designers of future generations of engines.
Specification of materials, shape, size, manufacture methods, heat treatment and
inspection of all crankshafts must conform to Classification Society rules. Fillet radii
are machined at each change in section at the ends of pins and journals.
In large multi-cylinder engines the crankshaft may be made up of two part lengths
joined at flanges with fitted bolts.
External balance weights may be attached to crank webs and holes bored through
large diameter pins to improve rotary balance.
MAIN BEARINGS Main bearings are required to support the crankshaft at each
journal. They support the weight of all working parts and are also subjected to heavy
loads, fluctuating in magnitude and direction. They must be secure, accurately bored
and aligned to prevent bending in the shaft. Each bearing is loaded by its adjacent units
and wear rates may not be equal. This must be checked and recorded regularly.
Weardown can be measured by bridge gauge and by crank deflections.
Large crosshead type engines use thick-shelled white metal bearings, each supported
in a transverse girder of the bedplate, giving it rigidity to support the high vertical and
transverse loads. The top half of each bearing is secured, either by studs and nuts to the
bedplate, or by jack bolts which transmit downward force from the engine frames.
Medium speed engines with high maximum cylinder pressures use steel backed thinwall bearings of harder materials such as aluminium-tin.
Main bearings are
underslung from the main frame or block with bolts or studs securing the lower half.
Horizontal side bolts are fitted to give added transverse rigidity.
In direct drive main engines it is convenient to position the
main thrust bearing within the engine. This makes use of the engine bedplate and
seatings to distribute the propeller thrust to the ship's structure. It also allows the
bearing to form part of the engine lubrication system.
A single collar tilting-pad type of bearing is usually fitted at the drive end of the shaft.
Smallcr engincs for gearbox or auxiliary drives do not require a thrust bearing.
There is little end-thrust generated in the engine and sufficient area on the radii on the
shaft journals and corresponding
main bearing bush ends will locate the shaft in its
correct axial position.
High-speed engines or those with few cylinders may be fitted with a
Iywheel, a solid wheel shaped to have high mass and inertia to store energy while the
engine is running. This energy is proportional
to the engine speed squared. If engine
speed falIs kinetic energy in the flywheel must also falI, and some is returned to the
engine shaft. In this way the flywheel absorbs energy during peak torque and
acceleration of the shaft and returns it during low torque and deceleration, smoothing
out fluctuations in engine speed.
The energy stored by a flywheel of a slow-running engine is limited and its effect is
negligible compared with the inertia stored in all the other large rotating parts of the
The term 'flywheel' may still be used, but in this type of engine it refers to the wheel
fitted to the end of the crankshaft, external to the crankcase and used mainly to engage
the turning motor during overhaul. It is also very convenient for marking crank angles
and top centre positions.
will be required in order to turn or
rotate the engMie when it is out of operation. It can also be engaged to lock the engine in
a given position.
The gear usualIy consists of a reversible, electric motor drive which can be engaged
to rotate the engine flywheel through a set of gears. The motor is ofIimited power and it
is necessary to open all cylinder indicator cocks to relieve compression while turning
the engine. The electric load on the motor sh0uld be monitored while turning as an
excessive load may give warning of an obstruction in the engine or propeller system.
The turning gear must be engaged before working within the engine. Blocking
devices must be applied--or electric fuses removed-to
prevent inadvertent movement
of the engine before personnel enter to work on it.
The gear must be disengaged when the engine is prepared for sea and safety cut-out
devices may be incorporated
in the manoeuvring gear to ensure this has been carried
Misalignment of an engine crankshaft occurs due to wear of main bearings or from
distortion of the engine bedplate, or possibly from damage to the ship's structure. It
can be detected by measuring deflections of crankshaft webs for each unit of the engine.
If misalignment exists the crank webs will open and close slightly as the engine is
rotated; this is measured by means of a clock or dial gauge fitted between adjacent webs
at a point in line with the outside of the journals furthest from the crank pin. A spring
extension rod will hold this in position.
The first measurement
is taken with the engine just beyond bottom dead centre
position with the gauge close to the side of the connecting rod. It is usual to set the
gauge to zero. The engine is now rotated by the turning gear and stopped at each
quarter turn where gauge readings are taken as plus or minus values. The final (fifth)
reading is taken near bottom centre, with the connecting rod on the opposite side of the
gauge to the first reading. The first and last readings are averaged to use as an
for bottom centre position. This procedure is repeated for each unit in
turn. Feeler gauges should be used to ascertain that the crankshaft has not lifted or
sprung in adjacent main bearings.
4.25 Record of deflections
It may be advisable to rotate the turning gear slightly in reverse direction after
stopping for readings; this will ensure free positioning of the cranks, particularly for
units adjacent to the turning gear.
All readings are recorded and these should be compared with previous values,
preferably with the ship in a similar load condition and at similar temperatures.
Total deflection vertically and horizontally is calculated for each crank. The vertical
total will be proportional to misalignment between the bearings due to weardown. The
horizontal total indicates side wear in the bearings. By plotting all vertical deflections
for the whole engine it is possible to obtain information as to which main bearings are
'high' and which are 'low'. This may be' assisted by bridge gauge readings from the
bearings, but these do not take possible distortion of the bedplate into account. It must
be checked that the journal is in contact with the bottom surface of the bearing.
Limiting values for maximum deflection are set by engine builders. These depend
upon the stiffness of crankshaft, engine stroke bore ratio, etc. They indicate the limits
to which misalignment may be permitted before renewal of bearings and realignment
are necessary.
Deflection measurements are readily taken and they relate directly to misalignment.
They should be taken at regular intervals and when damage to running gear in the
crankcase, propeller, shafting or ship's structure is suspected.
Excessive misalignment will cause bending of the crankshaft and webs with
fluctuating and alternating stresses, causing fatigue and the possibility of shaft failure.
It will set up vibration and cause damage to main bearings.
The bed plate of a large engine acts as the main strength part of its section, providing
rigid support for the main bearings and crankshaft. It is also a platform on which other
structural components such as frames or columns and guides may be accurately
mounted to support the engine cylinders and ensure the alignment of all working parts.
In high powered engines the bedplate must withstand heavy, fluctuating stresses from
working parts. It must transmit engine loads, including the propeller thrust, to the
ship's structure, distributing these over the necessary area, and may complement the
ship's strength and propeller shaft alignment. The bedplate also collects the lubricating
oil from the crankcase, returning it to the drain tank for recycling.
Fig. 4.26 shows part of a bedplate for a large cross head type of main engine. It is
fabricated from steel plates and castings welded together to form a deep longitudinal
box structure with stiffening members and webs to give additional rigidity. Lightening
and access holes are made with compensating sleeves, to maintain lightness with
strength. Transverse members or girders are fitted between each engine unit and at each
end of the bedplate. The central section of each is formed by a steel casting shaped to
support the main bearing and with holes for a pair of tie bolts. The casting has
substantial butt welds securing it to the main box structure.
The depth of the bedplate raises its top flange clear of the sweep of the bottom end
bearing in long-stroke engines, keeping overall width to a minimum. To limit overall
height, the centre oil pan or sump is lowered and will require a recess in the ship's
All welding in the bedplate must be to a very high standard, carefully controlled and
inspected. It must be stress-relieved, shot blasted and tested for flaws. All plate edges
must be correctly prepared, with double butt welds and complete penetration where
possible. Plates of different thickness should not be butt welded together. Bedplate
flanges are finally machined for landing on support chocks or for assembly of other
Fatigue cracks may commence at points of high stress or sudden change in section
and regular internal inspection must be carried out, particularly after heavy weather or
damage has occurred. All nuts, including holding down and tie bolts, must be checked
for tightness and to ensure fretting has not occurred between mating surfaces.
Earlier engines with lower stroke/bore ratio had bedplates formed by a longitudinal
box shaped girder at each side joined by transverse girders supporting main bearings
and tie bolts in steel castings.
Fig. 4.27
In crosshead type engines, frames or columns are used to support the cylinder block
from the bedplate. These are termed 'A' frames because oftheir shape; they are fitted at
each transverse girder and at both ends of the engine. 'A' frames are fabricated from
steel plates welded to form a hollow structure on each side. There is clearance to allow
access to main bearings. Transverse stiffening webs are fitted and flanges added for
necessary connections. Brackets secure the crosshead guides and other internal fittings.
Bottom flanges allow the frame to be accurately aligned and secured by fitted bolts and
studs to the bedplate and transverse girder. The top flange is fitted to the cylinder
block, scavenge trunk etc. Longitudinal bars and crankcase casing plates are attached
to the outer edges. Tie bolts pass vertically within the frame; the pre-stressing of these
maintains the frame in compression at all times.
In some designs, a number of frames with the crankcase casing are fabricated
together to form a box before attachment to the engine.
In medium speed trunk piston engines the main strength member may consist of a
single frame or block which incorporates the crankcase, bed plate, frames and even the
cylinder block. These may be formed as a single casting of either grey or nodular cast
iron machined on all necessary faces. Alternatively they may be fabricated from steel
plates and castings. Main bearings are under-slung from the frames and horizontal tie
bolts may add to the rigidity at this position. Typical frames for in-line and Vee engines
are shown in Fig. 4.29.
In most single-acting engines, except opposed piston engines, the main gas loads from
the cylinder covers are transmitted by long tie bolts. Two bolts are fitted to each
transverse girder and they pass up through the casting, through tubes constructed in
the engine frames and entablature to the top of the cylinder block where locking nuts
are hydraulically tightened to pre-stress the structure, maintaining the cylinder block
and frames in compression. Transverse locating bolts prevent vibration in the tie bolts,
which may be assembled in screwed sectional lengths to limit the access and removal
height. Tie bolt centres should be as close to the crankshaft axis as possible to reduce
bending stress on the girders and to prevent unbalanced loads being transmitted to the
welds (Fig. 4.30).
A main engine is secured to a foundation built into the ship's structure, specially
designed and stiffened to support rigidly the weight of the engine and the forces
generated during operation. Stiffening will extend outside of the engine to allow the
distribution of stresses through a wider area of structure. The top surface of the tank
top must be level and shaped to match the engine bedplate base.
Direct drive engines are jacked up accurately to align the main bearing centreline
with the propeller shafting. Holding-down bolt holes are then drilled and tapped in the
tank tops.
Cast iron or steel chocks are carefully machined and fitted
between the tank top and bedplate. In addition to chocks at each bolt, their total area
must be sufficient to support the engine. All chocks must be tight when the bolts are
hardened down (Fig. 4.31).
Alternatively an approved non-shrink epoxy resin chocking material may be cast
into the space between the engine and tank top (Fig. 4.32). This is not as strong as cast
iron, but by filling a larger area and by the intimate matching of surfaces left by casting
it will give excellent load-bearing and avoid the possibility of fretting which can occur
with metal chocks.
HOLDING-DOWN DOLTS or studs are screwed into the tank top and a nut
fitted and locked underneath. Seals must be used to ensure the integrity of the
watertight tank tops. Studs will normally have clearance in the bedplate to allow for
thermal expansion of the engine. In some cases some fitted bolts may be used to
transmit thrust at the drive end of the engine.
To improve the resilience of holding-down studs, their unthreaded section may be of
reduced diameter and the overall length increased. The top nuts are hydraulically
tightened in sequence to reach the correct tension in the studs and compression in the
chocks. Nuts must be hydraulically tested and chocks hammer tested at regular
intervals, and additionally after heavy weather or damage.
Side and end thrust (transverse and longitudinal) is transmitted through brackets
welded to the tank tops at the sides and ends of the bedplate. Vertical chocks or
packing pieces are fitted and locked between each bracket and the engine. Side brackets
are situated at the ends of each transverse member. End thrust is taken in a similar
Although the main forces are transmitted at the bedplate, further transverse struts to
secure large engines to the ship's structure are fitted at upper platform levels.
4.32 Engine seated on epoxy resin chocks (Sulzer)
A camshaft is necessary to operate the valves and fuel pumps which control the engine
cycle. Each valve or pump is actuated by a cam follower which rises or falls as the cam
rotates beneath it. The profile or shape of each cam is designed to give the correct
timing, speed and height of lift to its corresponding
High fluctuating forces are set up and cams are made of steel with hardened surface
to the profile. On large engines they are an interference fit on the camshaft and it is
possible to adjust or renew them hydraulically. Camshafts may consist of a number of
part lengths joined at flanges with fitted bolts. The whole shaft must be rigidly
supported with substantial bearings and adequate lubrication. Some engines have a
separate camshaft lubrication system to prevent any possibility of fuel leakage from
pumps passing into the crankcase oil.
Camshaft rotation must be accurately synchronised with the crankshaft and this
timing must be checked periodically and after any adjustments or repairs have been
carried out.
Vee type engines will require a separate camshaft for each bank of cylinders.
Reversible engines may have servomotors fitted to their camshaft to readjust the
timing of the cams when the engine is to run astern.
Four-stroke engines require inlet valve, exhaust valve and fuel pump cams for each
cylinder, and since this cycle is completed in two engine revolutions the camshaft
rotates at half engine speed.
Two-stroke engine camshafts rotate at full engine speed. They have fewer cams since
they do not require air inlet valves, and loop or cross scavenge engines do not need
exhaust valve cams.
A camshaft may be assembled from a number of similar part lengths; each wi\1have a
set of cams for one unit. Part lengths are connected at flanges. Each part length may be
dismantled separately to simplify overhaul, repair or adjustment.
CAMSHAFT DRIVES So that they remain in phase with engine timing,
camshafts are driven from the crankshaft. Camshaft bearings are supported from the
engine structure in a position convenient to the operating mechanisms that the shaft is
to control. Choice of this position together with engine size and type wi\1influence the
camshaft drive system to be used.
Further engine systems may, in turn, take their drive from the camshaft mechanism.
These may include starting air distributor, engine speed governor and cylinder,
lubricator pumps. Large long-stroke engines may have a moment balancer fitted. With
this weights are rotated at full engine speed and in some cases at twice engine speed, in
synchronisation with the crankshaft. This improves primary and secondary balance.
SmaIler engines may have power takeoff wheels to drive cooling and lubricating
pumps, air compressors etc.
Two forms of camshaft drive are in current use. A train of gear wheels is arranged in
positions and sizes to give appropriate speeds and directions required: these are used in
most medium speed engines and those in which the distance between shafts is limited.
Alternatively, a series of sprocket wheels connected by a roIler chain may be used. This
gives a reduction in weight, particularly where the distance between shafts is large.
Both methods have proved efficient and reliable. Systems for large two-stroke engines
are shown in Fig. 4.33.
CHAIN A roller chain forms a flexible drive, robust
but light in weight with very small friction loss; it is smaIl in width which adds little
extra length to the engine. It can accommodate a number of additional driven wheels
rotating at different speeds or even in opposite directions, and can be used to operate
engine driven pumps, etc.
A roller chain consists of alternate pin links and roller links. Each pin link consists of
two pins riveted between two side plates. A roller link consists of two rollers free to
rotate on two bushings which are a press fit into two side plates. The pins fit within the
bushings of two adjacent roller links (Fig. 4.34). Alternate links are all riveted together
to form an endless chain of correct length.
Chains have a very high factor of safety to prevent stretching. Any loss in
tension in the chain will be due to wear while in use. Wear takes place between pins and
bushings, between bushings and rollers, and between rollers and sprocket wheel teeth.
Wear in wheel teeth flanks can be checked by a profile gauge while wear in the chain
is measured by its extension.
While in place, the extension and loss in tension is indicated by the transverse play in
the chain span between wheels on the slack side of the drive. The chain is adjusted to
allow a limited transverse movement approximately equal to one link pitch.
To measure the extension, the total length over a given number of links on the
longest span, when drawn tight, is measured and the increase is calculated as a
percentage of the original length of an equal number of links. A chain should be
renewed when its extension approaches 2%. For a camshaft drive its life should also be
limited to fifteen years, irrespective of wear.
In addition to the driver (crankshaft) wheel and driven (camshaft) wheel, a chain
system will contain jockey and idler wheels. These may be used to alter the direction of
run ofthe chain and reduce unsupported lengths of chain. At least one jockey wheel has
the position of its centre adjustable: it can usually move about a pivot to adjust the
chain tension. Adjustment may be made by a screw or hydraulic jack arrangement or a
spring-loaded pulley. The adjustable wheel is usually located on the slack side of the
chain when going in ahead direction.
Adjustment is usually limited to the length of two chain links and when this has been
fully taken up the chain must be shortened by removing two links. To carry this out,
two adjacent pin links are removed with their common roller link. The ends of the
chain are then brought together and joined by riveting in a new pin link. On no account
should an old pin link be riveted a second time: it is always necessary to use a new link.
After tensioning adjustment the timing of the camshaft must be checked. If this is
incorrect, the camshaft flange is slackened and adjusted to synchronise the cam timings
with the crankshaft position. This may be checked either by putting one engine unit on
top centre position or from corresponding marks on the flywheel. After repositioning
the camshaft, the flange is retightened.
When in use the chain is lubricated by means of oil sprayers which direct a
continuous stream oflubricating oil into the chain bearings. This oil will lubricate and
cool the bearings and will cushion fluctuating forces.
Maintenance consists of periodic examination to check wear on the chain, teeth and
wheel bearings, free any seized rollers, check alignment of wheels by examining for
bright marks on side plates, and examine for possible fatigue cracks. When renewing a
chain, it must be ascertained that wheels are in correct alignment and that relative
positions of wheels and timing are correct.
Excessive tension on a chain will cause high loads and possible damage, while
excessive slack will cause vibration, add to cyclic stresses and possibly to fatigue
Where heavy loads are transmitted, duplex chains may be used. These consist oftwo
identical matched chains. Any adjustment carried out to one must be carried out
equally to the other.
Operating Systems
Marine engines must operate successfully on heavy residual fuel oils which may vary in
quality and analysis depending upon the source of their crude and the refinery
processes. Fuels from different sources may be incompatible and not mix successfully:
it is therefore advisable to store ships' bunkers from different ports in separate tanks to
avoid the risk of stratification and formation of heavy sludges.
Accurate and complete analyses are difficult to obtain but specifications should
confirm that certain parameters are not exceeded. Samples of the oil delivered should
be kept for reference.
Density of the oil must be given (usually measured at 15°C), since the consignment
will be measured by weight. It will also be necessary to know this to adjust centrifuges
to give efficient purification.
Viscosity is required, to calculate temperatures at which the fuel is treated and
injected into the engine.
Flash point for the fuel must not be below 60°C, although the actual value need not
be stated. This minimum is for safety reasons.
All residual fuels will contain contaminants, and to protect the engine some of these
must be removed or reduced. Solid contaminants may consist of rust, sand, dust and
refinery catalysts (catalytic fines), all of which are abrasives and will cause wear in fuel
pumps, injectors, cylinder liners, piston rings and exhaust valve seats. Liquid
contaminants will be salt or fresh water, which may also contain other soluble
substances; all of these are corrosive.
Fuel to be used is first transferred from storage tanks to a settling tank in which it is
heated to allow some water and sludge to settle out by gravity and be drained off. The
fuel is then passed through the purification system and discharged to a daily tank.
There are usually two daily tanks, used alternately, one in use while the other is being
recharged. Settling and service tanks are lagged to conserve heat.
A recommended standard of treatment for residual fuel to be used in a large engine
requires two centrifuges of adequate capacity, operating in series. The first acts as a
purifier to remove water, solubles, sludge etc while the second acts as a clarifier to
remove solids. The purifier must be fitted with the correct disc or dam to match the oil
density. The oil is heated before purification (max. temp. 98°C) and the rate of
throughput is limited to assist efficient separation. Both centrifuges must be cleaned
frequently. Such systems can operate effectively on oils of densities up to 0.99.
From the service tanks the treated oil is pumped through a pressurised fuel system to
the engine. With the oil temperatures necessary for high viscosity fuel, and the
possibility that a trace of water may still be present, it is necessary to maintain the
engine pump suctions and circulating connections under pressure to inhibit boiling,
gasification and cavitation.
The oil first passes to the primary or supply pump which raises its pressure to about 4
bar; this pressure is maintained in the circulating returns. The circulating or booster
pump draws oil from the primary discharge, raising its pressure to 10/12 bar and
delivering it through the heater, viscosity regulator and fine filter to the main engine
fuel pumps. The heater reduces the oil's viscosity (see below) for efficient combustion.
The temperature required will depend upon the oil quality, but to avoid fouling it
should not 'exceed 150°C. The fine filter is of stainless steel mesh to filter out particles
larger than 50 microns (0.05 mm), or less for smaller engines. Filters must be changed
over and cleaned regularly.
In large two-stroke engines the fuel injectors will circulate fuel during the periods
they are not actually injecting it to the cylinder. This ensures the system remains fully
primed and at uniform temperature. The circulated oil is returned to a buffer or venting
tank from which it passes either back to the low pressure part of the system before the
circulating pump suction, or into the service tank. The fuel oil system is shown in Fig.
All pumps are in duplicate and have safe pressure relief valves fitted. All pipes
carrying hot oil must be well lagged and fitted with trace heating if necessary. Fuel
pipes must be clipped and supported to reduce stress from vibration.
A tank containing diesel oil may be connected to the main heavy fuel system. This
will allow flushing with light oil before stopping the engine for long periods or for
Various safety devices must be included in the system with alarms to detect loss in oil
pressure, low tank level etc. Quick closing valves which can be operated from outside
the machinery space must be fitted to all tanks and to the main inlet to the engine. Fuel
pumps must have remote switches by which they can be stopped in an emergency.
There must be arrangements for venting and draining the system, cleaning strainers
etc, but utmost care must be taken. Drain trays and save-ails, where fitted, must be kept
clean; all joints must be kept tight with safeguards to prevent the possibility of hot oil
spraying on to heated surfaces.
Oil contained in tanks with open drains should not normally exceed a temperature of
50°C or, if this is lower, may not exceed a temperature 20°C below the flashpoint of the
When burning heavy fuel oil in a diesel engine it is
necessary to reduce the high viscosity of the fuel to a value at which correct atomisation
can take place in the fuel injectors. This will allow intimate mixing of fuel oil with
heated air for ignition and efficient combustion. Viscosity of a fuel may be reduced by
raising its temperature and it is passed through a heater to do this. Automatic control
of the heater may be regulated either to maintain a constant temperature, or to measure
and control the viscosity.
It is possible that oils of varying properties are contained within a ship's bunker
tanks or even one tank, when its contents are from a number of different sources.
Consequently it is preferable that the actual viscosity of the fuel is controlled within
close limits. In most cases this is regulated to have kinematic viscosity between 10 and
15 centistokes at 50°C (that is 50-70 seconds Redwood No. I, at 100°F).
The Viscotherm shown in Fig. 5.2 is an instrument which measures fuel oil viscosity
at the heater discharge and regulates the heater temperature to control this. It consists
of a small gear type pump rotated at a slow but constant speed (40 rpm) and is fitted
within the fuel supply close to the heater discharge. The pump draws fuel from the
system at a controlled rate and discharges it through a capillary tube. The form offlow
within a capillary tube at corresponding speeds is such that the pressure difference
1:ietweeneach end of the tube is directly proportional to the viscosity of the oil flowing
through it.
Pressures at the corresponding points are measured with Bourdon tubes and
compared to read as viscosity. In addition these pressures are fed to a differential
pressure transmitter which can automatically operate the heater control to maintain
fuel viscosity within close limits.
5.2 Viscosity regulator
All parts of the instrument are of stainless steel. In the event of a failure of this
control an oil bypass valve is included in the system and the temperature may then be
controlled by hand or by a thermostat.
The general indications of good combustion are similar in any operating diesel engine:
a clear exhaust, power produced and exhaust temperatures normal for the throttle
setting. There should be no uneven running, knocking from cylinders or the fuel
Viscosity, or resistance to flow, in a fuel oil is important when considering
combustion. It must be low enough to ensure correct atomisation at the fuel injector.
Since viscosity reduces as temperature is increased, it will be necessary to heat heavy
fuel oil to reduce its viscosity to about 15 cSt at 50°C before atomisation for
Atomisation is the splitting up of the fuel into very small droplets by the fuel injector
forcing fuel at high pressure through small atomiser holes. The droplet size will depend
upon the size of holes and the pressure difference between fuel pump discharge and that
of the compressed air in the combustion chamber, and consequently the size of droplets
may vary over the whole injection period. Atomised droplets have a high surface to
mass ratio giving good heat transfer from the hot compressed air in the cylinder
causing rapid evaporation and mixing.
Penetration refers to the distance the oil droplets travel into the combustion space
before mixing with the air and igniting. This will depend upon droplet size
(atomisation), velocity leaving the injector and the conditions within the combustion
chamber. It is desirable that fuel should penetrate into the whole of the combustion
space for good mixing, but droplets should not impinge on the internal surfaces before
burning. The number of atomiser holes and their position will decide the spray pattern.
Turbulence is the movement of compressed air and fuel within the combustion space
before combustion occurs. This may have several causes. Swirl is imparted to the air
during its entry at scavenge ports. It may'be further agitated by the fuel spray pattern
and the shape and movement ofthe piston crown. Turbulence will improve the mixing
of fuel and air for effective and rapid combustion. It is particularly desirable for rapid
combustion of heavy fuels in medium or higher speed engines,
Fig. 5.3 illustrates these processes.
is the term used to describe the combustion in
diesel engines and they are often referred to as compression ignition engines. The
combustion process may be considered as three consecutive phases. In phase one the
atomised oil droplets emitted from the fuel valve nozzle into the combustion space at
the start of injection will evaporate and mix with hot, compressed air and some
chemical changes will also occur. The mixture will reach an ignitable condition and
spontaneous combustion will commence. The time elapsed during this phase is termed
the ignition delay or ignition lag.
In phase two the ignition and start of combustion will set up a flame front which will
accelerate through the chamber, enveloping and burning all the other droplets present,
causing a very rapid generation of heat with a corresponding rise in pressure and
During the ignition delay, the injector continued to inject fuel and, if this has built up
a sufficient quantity, the rapid combustion and pressure rise will be quite violent,
causing detonation, the shock loading creating a noise termed diesel knock.
Following the rapid pressure rise in phase two, hot, turbulent conditions existing in
the combustion chamber will ignite and burn the remainder of the measured fuel
charge as it is injected. This is phase three; it is termed the controlled part of the
combustion process as pressure is regulated by the rate at which fuel continues to be
delivered. The cylinder pressure may start to reduce as the piston moves down after
passing over top centre.
Ignition quality of a fuel is the term used to denote its ignition delay, combustion
characteristics and tendency to cause knock. It depends mainly upon the form of the
hydrocarbon compounds in the fuel and no precise unit has been derived to measure
The most usual measure of ignition quality of distillate fuels is the Cetane number.
This is found by comparison of their 'knock producing' properties to those of
mixtures containing an equivalent percentage of Cetane when burnt in a test engine. A
high Cetane number indicates a short ignition delay. Cetane numbers are not quoted in
normal fuel specifications or analyses. Slow speed, two-stroke engines can operate
efficiently on fuels down to a Cetane number of about 24 but medium speed, fourstroke engines normally require a figure above 34; high speed engines need higher
figures. Another measure which is similar to Cetane number but found from different
parameters of the fuel, is termed its diesel index.
For residual or blended fuels the ignition quality can be expressed as a CCAI value
(calculated carbon aromacity index) or CII (calculated ignition index). The lower the
CCAI value, the better the ignition quality. A limiting figure of 870 is suggested. These
units take the viscosity into consideration.
The ignition quality of a fuel is particularly important for ease of starting an engine,
or when operating at reduced power for long periods. It can be improved by increasing
the compression ratio of the engine or by pre-heating the scavenge air. There will be
design or operational limits to both of these.
VARIABLE IGNITION TIMING (VIT) Ifan engine operates for long periods
at reduced power or speed, the residual heat in the main components of the combustion
chamber will decrease, causing a lower air temperature after compression. This will
lead to an increase in the ignition delay of injected fuel, which will cause knocking or
'rough running' in the engine, with consequent damage from shock loads and poor
combustion. This problem can be reduced by the use of variable ignition timing to
advance the start of injection, allowing for the longer delay but maintaining ignition
timing and the same peak pressure. In large, two-stroke engines. variable ignition
timing is automatically superimposed on the normal fuel pump setting from the engine
governor. Additional linkage from the governor will advance each pump setting over
the range of lower speeds. Fuel pumps are designed to carry out necessary adjustment
while the engine is running. This control can also be manually regulated to advance
normal governor settings ifit is known that fuel with low ignition quality is to be used.
In medium speed engines, variable ignition timing is built in by shaping the profile of
the pump plunger to give an earlier start of injection at lower fuel settings. Manual
advancing the timing for low ignition quality fuel, is fitted on some
This is a system by which fuels with low ignition quality can
be burnt smoothly and efficiently, even in medium or high speed engines. A small 'pilot'
charge of finely atomised fuel is briefly injected very early during compression. This
small charge is heated and vaporised, passing through its delay period and igniting just
before the main charge of fuel is injected. Combustion of the pilot charge will cause
high temperature and turbulence in the combustion space. This together with its flame
front will cause the main charge offuel to ignite and burn smoothly as it is injected. In
this manner the violent pressure rise usual with these fuels is avoided. Combustion is
efficient and the peak pressure is limited.
Pilot injection can he implemented by two different methods. In the first case two
separate injectors are fitted to each unit; both must be correctly timed. The pilot
injector has fine nozzle holes to atomise the fuel at a moderate injection pressure and is
timed to commence before the main valve. The main injector is situated in a position to
give an effective fuel spray and it injects the main, measured charge at normal fuel
injection pressure. This system with two valves is sometimes referred to as twin
The alternative method of pilot injection has one injector fitted to each unit. This
injector is designed to inject the pilot charge with fine atomisation and must also inject
the main fuel charge. To do this it requires variable nozzle holes and a two stage needle
valve lift. Fuel pump drive for this system may require specially shaped cam profiles.
Fig. 5.5 shows a section through a hydraulically operated fuel injector as fitted to a
large two-stroke diesel engine. The general design is similar for most engines and
consists of a spring loaded non-return needle valve operated hydraulically by a fuel
pressure wave from the fuel pump to discharge fuel at high pressure through an
5.4 Fuel injector needle valve and nozzle
atomiser nozzle. A typical fuel injector will consist of a valve body or nozzle holder to
which the nozzle or atomiser is secured by a retaining nut. The valve body contains the
spring and its compression nut, with an intermediate spindle if required. Surfaces
between the body and atomiser are ground and lapped to form an oil pressure-tight
seal. A dowel ensures alignment of the oil passages.
The needle valve is lapped into the bore of the atomiser and these must be kept as a
matched unit. As shown in Fig. 5.4, there are two chambers in the nozzle, the upper one
being charged with fuel oil from the fuel pump and sealed by the needle valve when
closed. The lower chamber, or sac, is sealed by the mitre seat of the needle valve and has
a number of small atomiser holes of correct size and pattern to atomise and distribute
the fuel spray into the combustion chamber.
Injector spring compression is adjusted under test and a compression ring fitted. It is
set to allow the needle valve to lift or open ata predetermined fuel pressure. The
intermediate spindle conveys the spring compression to the needle valve and may be
arranged to limit its lift. The valve will open when the pressure from the fuel pump
acting on the shoulder of the needle valve overcomes the spring compression. As the
needle valve lifts, oil flows to the lower chamber in the atomiser. The additional area of
the needle mitre now subjected to pressure causes the needle to lift rapidly, allowing
fuel at high pressure to pass through the atomiser holes into the combustion chamber.
When the fuel pump cuts off pressure, the valve will close under spring compression.
Since the full area of the needle is now exposed to pressure, closing will occur at a
pressure lower than that at which it opened. The action of the needle valve must be
rapid and positive with no oil leakage.
A fine edge strainer may be fitted at the fuel inlet and a priming or venting plug is
fitted to the fuel passage. Valves should be primed if the engine has been out of service
or during preparation for commencement of a voyage.
Fuel injectors must be overhauled at regular intervals to ensure correct operation
and combustion. The injector compression spring must be screwed back before
slackening the retaining nut. Parts are cleaned, inspected and renewed if necessary.
Lapped surfaces must be free of damage and correctly aligned, springs inspected for
distortion, atomiser holes must be clear and unworn. Defects in injectors, while in use.
may be choking due to dirt in the fuel or carbon building up at the atomiser. A leakinS
needle valve will cause secondary burning and reduce combustion efficiency.
After assembly the injector is tested with a test pump. Operating pressure and fuel
spray pattern are checked and there must be no leakages.
Fuel injectors are inserted into pockets in the cylinder cover and must form a gastight joint at their lower landing. They are secured by studs and nuts.
Fuel pipes between pumps and injectors are subjected to extremely high internal oil
pressure with cyclic pressure fluctuations which a.lsocause vibrations. The pipes and
their connections are therefore subjected to considerable stress and fatigue. Failure
may cause a very dangerous spray of high temperature, atomised oil which is both a
health hazard and a fire risk. Where such pipes are exposed, they should be enclosed in
a double skin tube. The outer skin will collect any oil leakage and return it to a safe
Medium and high speed engines may be designed with the whole fuel pump system
enclosed within protective casings. Some engines have the fuel pipe to the injector
placed in a passage cast within the cylinder cover to the valve pocket.
When burning heavy fuel the injectors will require to be cooled. In many engines this
is accomplished by circulating water or oil through additional passages in the injector
assembly. Due to the risk of contamination of this coolant with fuel, an independent
system is required. If atomiser tip temperatures are too high, carbon may form and
impede the spray; at high or low temperatures, corrosion may occur.
5.6 Uncoo/edfue/ injector (MAN-B
& W)
The design of modern two-stroke engines ensures the injectors are adequately cooled
by the transfer of heat to the surrounding bore-cooled cylinder cover. Each injector is
also fitted with a spring-loaded circulating valve which permits hot fuel at the
circulating pump pressure to pass through passages in the valve body before returning
to a buffer tank in the oil system. This maintains the valve at the correct temperature at
all times, allowing the engine to be manoeuvred on heavy fuel. The high pressure wave
of the engine fuel pump immediately depresses the spring, ensuring that the circulating
passages are sealed off before the high pressure lifts the needle valve and injects fuel
into the cylinder. A fuel valve of this type is shown in Fig. 5.6.
The ideal position for a fuel injector is in the centre of the cylinder cover, allowing a
symmetrical, conical spray pattern in the combustion chamber. This is achieved in
most four-stroke engines. Tn large engines with a centrally placed exhaust valve,
injectors (usually three) are placed symmetrically around the cover and charged from a
common distributor connection to inject equal quantities of fuel simultaneously.
Engine fuel systems are designed for normal working conditions. If an engine is run
for long periods at low power, combustion may be inefficient, leading to fouling and
possible wear. When long periods at low power are anticipated a set of low-power
injectors may be fitted which have reduced orifice area giving high atomisation and
peak pressure. It may also be necessary to adjust fuel pump timing. Engines with
variable ignition timing fitted can adjust to low powers automatically. They may also
be fitted with fuel injectors which maintain high efficiency over a wide range of
operating conditions, including slow speeds.
Fig. 5.7 shows a Jerk type fuel pump. This type of pump with small variations is used in
many diesel engines. One is fitted for each cylinder. The pump consists of a camoperated, single acting plunger of fixed stroke. Helical springs are fitted to return the
plunger on its down stroke and to maintain contact of follower on the cam.
No timing valves are required and fuel delivery commences at a fixed point on the up
stroke when the top edge of the plunger blanks off the suction port. A helix or scroll is
machined on the plunger and delivery of fuel ceases on the up stroke when the curved
surface of the helix uncovers the suction port. This allows fuel pressure above the
plunger to fall to the suction pressure through a vertical slot or hole.
The quantity of fuel delivered is regulated by the vertical length of the helix where it
is in line with the suction port. This setting may be altered by rotating the plunger.
A rack is fitted to the pump to engage with a pinion machined on the outside of a
sleeve. The sleeve fits over the plunger and has slots engaging with keys. In this way the
plunger may be rotated by movements of the rack (Fig. 5.8).
The fuel cam is designed to raise the plunger at the rate required to build up fuel
pressure and maintain this for the corresponding period to operate the fuel injector.
Since the pump only discharges on its up stroke, only one flank of the cam operates the
timing. The trailing flank of the earn returns the plunger to the bottom of its stroke to
allow the chamber to refill.
Timing is controlled by the relative angular position of the cam peak to the
crankshaft. It can be adjusted by moving the cam with respect to the shaft. Further
adjustment is made by raising or lowering the pump plunger with respect to its
follower. Raising the plunger will make cut off of the fuel suction port and
corresponding fuel injection early, while lowering the plunger will make these later.
Alternatively, the fuel pump casing itself may be lowered or raised on its mounting to
5.8 Fuel pump timing control
5.9 MAN-B
& W fuel pump with variable ignition timing
give the corresponding effect. The plunger must maintain sufficient clearance at the top
of its stroke.
Oil supply to the pump suction is by means of a continuously operating supply or
surcharge pump which causes flooding ofthe fuel pump chamber as soon as the suction
port is uncovered by the plunger.
In some pumps, a non-return spring-loaded discharge valve is fitted. This is arranged
to reduce pressure on its discharge side as it closes, ensuring positive seating of the fuel
injector needle and reducing cavitation within the pump.
A priming or vent plug is fitted to the discharge.
Possible faults in a fuel pump include wear of parts, such as plunger and barrel.
5.10 Sulzer valve-controlled fuel pump
These surfaces are a close fit, the small clearance allowing some leakage to lubricate the
plunger. Larger clearances are necessary for the higher temperatures when burning
heavy fuel. Care must be taken that no leakage from the pumps can enter the camshaft
lubrication system as this may lead to fuel contamination of the crankcase. Other faults
may be erosion of suction port or plunger surface. Cam"timing and relative position of
plunger should be checked periodically.
Fig. 5.9 shows a fuel pump for an MAN-B & W MC type engine. It operates on the
same principle of control by a helix on the plunger but has further adjustments which
make it possible to regulate the injection timing during operation of the engine. The
pump barrel is able to be moved within the pump casing and hot oil is circulated
between them. It also permits the barrel to be raised or lowered by means of a second
rack which rotates a coarse-threaded sleeve attached to the lower end of the barrel.
Movement of the barrel will alter the timing of the start of injection and this can be
advanced to improve combustion of low quality fuel or at low power. The variable
ignition timingHnkage allows individual fuel pumps to be adjusted if required; it can
also be linked to the engine governor control to adjust all pumps collectively. A shock
absorber is fitted to the pump chamber. The follower is raised to cut off fuel.
During the part of the cycle when the pump is not injecting high pressure fuel, hot oil
at the circulating pump discharge pressure passes through to circulate the injectors.
Fig. 5.10 illustrates a valve-timed fuel pump as used in large Sulzer RTA engines.
The pump plunger is raised and lowered by a follower on the cam. Spring-loaded
suction and spill valves control oil to and from the pump chamber and each of these can
be opened by lifting it from its seat by a push rod moved by a lever operated from the
plunger drive. Pivot points of the levers are positioned so that one push rod moves up as
the plunger rises while the other push rod moves down.
From the bottom of its stroke the plunger is moved upwards, the spill valve is closed,
but the suction valve is held off its seat by its pushrod. No high pressure oil delivery
takes place until the suction valve is able to seat as its push rod moves down. From this
point oil pressure is raised and injection continues as the plunger continues its up
stroke. The spill valve pushrod will force it off its seat, releasing the oil pressure and
ending injection. A non-return valve prevents loss of oil from the discharge. The
plunger chamber is recharged during the downward stroke.
The pivots of the levers are eccentric and by rotating these the timing can be altered.
Under normal operation the suction valve timing and start of injection is fixed, while
spill timing is used to control the power.
Variable ignition timing has links which superimpose control on the suction valve
pivot, giving adjustment of the start of injection according to fuel quality.
Circulating oil passes through the pump during the period for which it is not
injecting. The pump follower is lifted clear of the cam to shut off fuel.
Fig. 5.11 shows a lubricating oil system for a large main engine. Pressure pumps,
strainers and fine filters are in duplicate, one set being used while the other acts as
standby. Fine filters should be capable of being cleaned without interruption of the oil
flow. Mesh size will depend upon the bearing materials and clearances: in most large
engines it is 50 microns. Capacity of the system must be adequate for the type of
installation. If the engine has oil-cooled pistons the capacity and throughput will be
increased accordingly.
Lubricating oil pressure pumps draw oil from the engine drain tank through suction
strainers, the tank suction being clear of the lowest point to avoid picking up any water
or sludge which may have settled. The pumps discharge at pressure through the oil
cooler, ensuring that sea water at its lower pressure cannot leak into the oil system in
the event of a fault in the cooler. The oil then passes through the fine filters to the
engine. It will be distributed to all bearings, piston cooling, sprayers, exhaust valve
actuators, control systems etc.
Various sections of the lubricating system may require different pressures and to
accommodate this engine driven booster pumps may raise the supply pressure, while
pressure reducing valves and restricted orifices may reduce pressure or flow to other
Used oil drains to the bottom of the crankcase and passes through strainers by
gravity to the drain tank. Drain returns are kept remote from the pump suction and
must be submerged to reduce aeration and to make a safe seal. With oil-cooled pistons
5.11 Lubricating oil system
each piston oil return has its temperature monitored and it then passes through a sight
glass before returning to the crankcase.
The oil drain tank is usuaIly built into the ship's double bottom but it must be
surrounded by a coffer dam to prevent any contamination from leakages. It is fitted
with an air vent, level measuring gauge and sounding pipe. Central positioning of the
level gauge wiII reduce fluctuation in readings due to pitching and roIling at sea. The
tank must be of sufficient size to accommodate the fuIl charge of oil. Its interior
surfaces may be coated to prevent rusting due to condensation on its non-flooded
surfaces. The system should also have low pressure, high temperature and low tank
level alarms fitted.
A bypass centrifuge system is fitted to purify oil from the drain tank to remove water,
sludge and insolubles. This should be operated continuously at sea with a slow
throughput, the oil being preheated to 7o-90°C to assist separation.
When the engine is not in use batch purification of the whole charge may be carried
out. It is most important that water content in the oil is eliminated or kept to a
Extensive demands are made on oils for lubricating systems and their specifications
may differ according to the type of engine and the fuel used. In addition to lubricating
and cooling the bearings, the same oil may be required to lubricate piston rings and
cylinder liners, and neutralise acids entering the crankcase in trunk piston engines. It
may be required to act as the cooling medium for hot surfaces within pistons, and must
-tolerate high pressures in gear or chain drives and cams, and even high-speed bearings
in turbochargers. The oil may be used as a hydraulic fluid in exhaust valve actuators,
pressure control systems and servomotors.
System oils must inevitably contain additives to improve their properties and
prolong their life. Additives are chemical compounds which mix with the oil's
molecular structure so that they do not separate out during storage, use or treatment in
purifiers and filters. Their properties will be depleted with use, however, and to
minimise this it is necessary to limit maximum temperatures, thermal cycling and
contamination in the system. Topping-up with fresh oil to replace that consumed will
help to maintain a stable crankcase oil condition.
Oil for engines operating on residual fuels must be alkaline. For crosshead engines
10 to 15 TBN (SAE 30) oils should be adequate, but for trunk piston engines up to 40
TBN (SAE 40) is used depending upon the sulphur content of the fuel.
Besides correct alkalinity and viscosity, oils require rust and oxidation inhibitors, a
high viscosity index, detergent/dispersant properties, water repellants, and must resist
foaming and emulsification.
To monitor its condition, oil samples must be taken from the working system at
regular intervals and sent to an oil laboratory for detailed analysis, which will give
information on oil condition and depletion of additives. It may also assist in
diagnosing engine faults and advice on any remedial measures may be given. General
trends in condition are important, but any sudden changes or losses must be
investigated immediately.
A number of simple oil condition tests may be carried out on board ship with
portable equipment. A flashpoint test will warn of possible fuel oil contamination.
Colour titration tests can indicate loss in alkalinity. A variety of viscosity tests are
available. Water present can be detected, but special equipment is required if actual
percentages are to be measured. Inspection of the sludge/water discharge at the purifier
will give a good indication of condition. In blotter tests the slow migration of a drop of
oil on blotting paper will indicate the oxidation, carbon present and dispersive
properties when compared with similar drops of fresh oil.
MICROBIAL DEGRADATION OF OIL is the possible infection of an oil
system by micro-organisms. These live by consuming hydrocarbons in the base oil
together with nitrogen, sulphur and phosphorus from additives. Ifwater is present they
will multiply rapidly in the warm, agitated conditions in the oil system. Contamination
may occur from cooling water leaks, condensation in humid climates or accidental
entry of dirt and water during overhaul.
Infection by micro-organisms will form organic acids and more water while
depleting the additives. This will cause corrosion and wear of metal surfaces
(particularly bearing metals), and will create sludge and slime that choke oil filters. The
oil may be prone to emulsification and saponification.
Tests for microbial infection make use of a special gel which is either dipped in the oil
or has oil poured over it and is then incubated to develop a growth or culture. The
appearance of this culture is then compared visually with a standard coloured chart, to
indicate the degree of contamination if any. If the oil does not wet the gel, it may be
necessary to mix the sample with a small quantity of sterile water.
Contamination does not commonly occur if cleanliness and care are exercised and
water is eliminated from the system. Recommendations are that a maximum of 0.20/.
of water content must not be exceeded when using detergent oils.
-tolerate high pressures in gear or chain drives and cams, and even high-speed bearings
in turbochargers. The oil may be used as a hydraulic fluid in exhaust valve actuators,
pressure control systems and servomotors.
System oils must inevitably contain additives to improve their properties and
prolong their life. Additives are chemical compounds which mix with the oil's
molecular structure so that they do not separate out during storage, use or treatment in
purifiers and filters. Their properties will be depleted with use, however, and to
minimise this it is necessary to limit maximum temperatures, thermal cycling and
contamination in the system. Topping-up with fresh oil to replace that consumed will
help to maintain a stable crankcase oil condition.
Oil for engines operating on residual fuels must be alkaline. For crosshead engines
10 to 15TON (SAE 30) oils should be adequate, but for trunk piston engines up to 40
TON (SAE 40) is used depending upon the sulphur content of the fuel.
Besides correct alkalinity and viscosity, oils require rust and oxidation inhibitors, a
high viscosity index, detergent/dispersant properties, water repellants, and must resist
foaming and emulsification.
To monitor its condition, oil samples must be taken from the working system at
regular intervals and sent to an oil laboratory for detailed analysis, which will give
information on oil condition and depletion of additives. It may also assist in
diagnosing engine faults and advice on any remedial measures may be given. General
trends in condition are important, but any sudden changes or losses must be
investigated immediately.
A number of simple oil condition tests may be carried out on board ship with
portable equipment. A flashpoint test will warn of possible fuel oil contamination.
Colour titration tests can indicate loss in alkalinity. A variety of viscosity tests are
available. Water present can be detected, but special equipment is required if actual
percentages are to be measured. Inspection of the sludge/water discharge at the purifier
will give a good indication of condition. In blotter tests the slow migration of a drop of
oil on blotting paper will indicate the oxidation, carbon present and dispersive
properties when compared with similar drops of fresh oil.
MICROBIAL DEGRADA nON OF OIL is the possible infection of an oil
system by micro-organisms. These live by consuming hydrocarbons in the base oil
together with nitrogen, sulphur and phosphorus from additives. Ifwater is present they
will multiply rapidly in the warm, agitated conditions in the oil system. Contamination
may occur from cooling water leaks, condensation in humid climates or accidental
entry of dirt and water during overhaul.
Infection by micro-organisms will form organic acids and more water while
depleting the additives. This will cause corrosion and wear of metal surfaces
(particularly bearing metals), and will create sludge and slime that choke oil filters. The
oil may be prone to emulsification and saponification.
Tests for microbial infection make use of a special gel which is either dipped in the oil
or has oil poured over it and is then incubated to develop a growth or culture. The
appearance of this culture is then compared visually with a standard coloured chart, to
indicate the degree of contamination if any. If the oil does not wet the gel, it may be
necessary to mix the sample with a small quantity of sterile water.
Contamination does not commonly occur if cleanliness and care are exercised and
water is eliminated from the system. Recommendations are that a maximum of 0.2%
of water content must not be exceeded when using detergent oils.
Slow and medium speed engines use fresh water to circulate and cool the cylinder
jackets and covers. The same system may also circulate exhaust valves and water
cooled turbochargers if fitted.
In large, slow engines the cooling water pumps are independently driven and
controlled. Pumps are arranged in duplicate, each being sufficient to operate the system
alone while the other acts as standby. They should be changed over regularly. Pressure
relief valves are fitted and alarms to give warning of pressure loss, high or low tank
levels or excess temperature.
Fresh water coolers are circulated and cooled with sea water and have bypass valves.
Pressure of the fresh water must exceed that of the sea water to prevent any possibility
of sea water leakage into the system.
In medium speed engines the fresh water cooling system may be split, or comprise
two sections. Each incorporates a pump, salt water circulated cooler, temperature
regulators and bypass valve. Pumps are usually main engine driven. A high
temperature section circulates the cylinder 'jackets, covers, exhaust valves, the
turbocharger and the first stage of the charge air cooler. This section can be regulated
to maintain a fairly constant high temperature, irrespective of the engine speed or load.
An additional electric driven pump and a pre-heater are included to facilitate warming
of the engine before starting from cold. A low temperature section circulates the
lubricating oil cooler and the second stage of the charge air cooler. At low speed or
power, the temperature regulator in this section directs heat from the lubricating oil
cooling to be used to raise the temperature of the charge air before it enters the engine.
Fresh water may cause the deposit of scale on heated surfaces and although
demineralised or distilled water will prevent this, it may still cause 'pitting and erosion
and become acidic with use. Corrosion inhibitors and alkaline agents are added to the
system to protect metal surfaces. They must be soluble in water, but cannot be
poisonous if a domestic fresh water generator is present in the system.
Chemical additives such as nitrite-borates may be used; they create a thin passive
oxide surface layer on most metals. They do not harm water seal rings but will attack
any zinc present. With this treatment, therefore, no zinc anodes are fitted; galvanised
piping and all solder joints must also be avoided. A reserve concentration of the agents
is held within the water and this replaces any areas of surface layer which may have
been removed by cavitation or vibration. Frequent tests must be carried out to check
the concentration and alkalinity.
Soluble emulsion oils may be used in high speed engines, provided they do not have
small cooling passages such as those used in water-cooled exhaust valves. These oils
form a greasy protective film which adheres to surfaces and prevents corrosion. Its
tenacity gives better protection against cavitation than chemical additives. Emulsion
oils must be maintained within close concentration limits and the system kept clean:
any contamination causes deposits to build up, with the risk of choking of passages
causing overheating. The oil deteriorates in time and the system should be cleaned and
recharged annually.
It is possible for micro-organisms to exist in water containing either of these
additives. Any biocide added to kill these micro-organisms must be compatible with
the domestic fresh water generator if one is fitted.
Fig. 5.12 shows line diagrams for fresh water cooling of cylinder jackets and pistons
in a large two-stroke engine. These are in fact separate systems and only the jacket
water cooling is required for engines which have oil-cooled pistons.
WATER SYSTEMS form a closed circuit. Water
passing from the engine returns through a cooler to the pump suction and then back to
the engine (the order between cooler and pump may be reversed in some systems). A
header or expansion tank is placed at a reasonable height to a\1ow venting and
pressurising ofthe system. This is connected to both the engine discharge and the pump
suction lines. A heater is included, with a bypass. and the system can warm the engine
through when necessary. Alternatively, auxiliary engines' fresh water systems may be
coupled, to use waste heat from these for warming the main engine during stand-by
periods. A fresh water generator may be added to use further waste heat.
Cooling water enters at the lower end of the cylinder jackets, passing up to
connections from the top of the jacket to the cylinder covers and then to the exhaust
valve cages and seats. A restricted amount of water is taken from this discharge and
passes through the turbocharger cooling spaces, returning to the main discharge
through a baffle which creates a pressure difference in the line. An air separator is fitted
in the return and this discharges any trapped air to the expansion tank. Other ventins
connections are taken from the top of the engine system, the turbocharger and the
is also shown in Fig. 5.12. It is
independent of the jacket system, allowing piston temperatures to be higher and
preventing possible contamination of the jackets in the event of piston gland leakase
passing lubricating oil to the system. The circulating pump draws water from a drain
tank, passing it through a cooler to the piston cooling inlet glands. Cooling returns
pass through sight glasses and thermometers before flowing by gravity to the drain
tank. Air vents are fitted at high points. adjacent to the glands to eliminate 'water
hammer'. Leakage from glands may be drained back to the tank after passing through
an oil separator and inspection chamber. A heating coil fitted in the drain tank allows
heating when warming through the engine in preparation for sea.
Fuel valve cooling systems may be fitted as a separate fresh water system. A heater
can be incorporated in these to warm fuel injectors before starting on heavy fuel.
Fig. 5.13 shows a starting air system for a large two-stroke engine. Clean, dry,
compressed air at starting pressure (30 bar) is stored and supplied to the system from
air receivers. The main starting valve (automatic valve) is operated by a pneumatic
actuator, and similar actuator controls are fitted for slow-turning and to the
distributor or timing valves. The main starting valve can be locked shut and this must
be done before the engine turning gear is engaged. When this valve is open, air passes
through a non-return valve and flame arrester to the main manifold supplying pressure
to the cylinder valves, one of which is fitted to each cylinder of the engine.
The cylinder air start valves are normally held closed by a compression spring
together with cylinder pressure acting over the valve lid. Air from the manifold enters
these valves where it forms a pressure balance between the underside of the valve lid
and a balance piston of equal area on the valve spindle. Consequently this does not
cause the valve to open.
Cylinder valves are opened when operating air, transmitted from the distributor,
applies pressure to the larger operating piston on the valve spindle. As the valve is
forced open, starting air from the manifold enters the cylinder, applying pressure on
the piston and causing the engine to rotate in the corresponding direction. To close the
cylinder valves the connection from the distributor is opened to atmosphere, allowing
the spring to close the valve and return the operating piston.
Air from the main valve also passes through two connections to the distributor. One
depresses each timing valve in the distributor from its free position to engage it with the
timing cam. This line can be blocked for test purposes or by safety cutouts. Until air
pressure is applied, the timing valves are held clear of the cam by springs.
The second connection supplies air to the operating ports, from which it will pass
through any timing valves which are open, to the corresponding cylinder valves,
causing them to open. Air from the manifold will then enter that cylinder to start the
engine. The timing valve for each cylinder is synchronised with the engine position and
successive valves operate in the firing order.
On large engines a slow-turning valve can bypass the main starting valve and will
supply just enough air to rotate the engine slowly. This is used to turn the engine
through one revolution while the indicator cocks are kept open, when preparing for
sea. At any subsequent time, if the engine has not been used for about half an hour the
slow-turning valve should be used to turn the engine slowly for one revolution before
opening the main valve for a normal start.
Starting air is shut off from the engine as soon as sufficient starting speed has been
reached; fuel is then applied and engine speed increased.
A main engine must be capable of starting from any position and it is normal to have
starting valves in each cylinder. Multi-cylinder Vee type engines may only require
starting valves to be fitted to one bank of cylinders.
Adequate drains must be fitted to the air start system to keep it free of oil or water. A
safety or relief valve is fitted to the main air manifold. Flame arresters or bursting caps
or discs must be fitted to each cylinder valve connection from the manifold in reversible
engines. Non-reversing engines require one only, fitted between the non-return valve
and the engine.
Checks must be made that cylinder starting air valves do not leak. These can be
carried out while the engine is running by feeling the temperature of the air pipe
adjacent to the valve (see Chapter 7 on explosions in air start systems). Tests while the
engine is stopped require starting air to be applied and the main valve opened, after first
blocking off air to the timing valves. Pressure will escape into the cylinder of any
leaking air valve and this can be detected by listening at the open indicator cock.
STARTING AIR RECEIVERS are required to store compressed air for starting
and manoeuvring main engines, for starting auxiliary engines, and to supply
pneumatic control systems, exhaust valve air springs, air whistle and other auxiliary air
pressure systems. Air is normally stored at pressures up to 30 bar for the main engine
starting. Reducing pressure valves are fitted to regulate lower pressures for other
Two receivers are fitted and their total capacity must be at least sufficient to obtain
the regulation number of engine starts without recharging: 12 starts required for
reversible engines, 6 starts for non-reversing engines. Receivers are welded from good
quality steel, cylindrical in shape with dished ends. They must meet all regulations and
tests for pressure vessels. Compensating rings are necessary where any openings have
been cut. Welds must be stress-relieved and subjected to non-destructive testing. An
acceptable welded attachment for fitting a valve manifold to a receiver is shown in Fig.
The usual fittings required to a main starting air receiver are: safety valve, pressure
gauge connection, drain valve, filling valve, discharge to engine air starting main, air to
ship's whistle, auxiliary systems, pneumatic control system, etc. Although each
connection may be made separately to the shell ofthe receiver, it may be preferred to fit
a valve manifold with only one common connection to the shell.
5./4 Air manifold connection
The safety valve must be set to relieve the receiver of excessive pressure rise and must
have sufficient area to prevent accumulation of pressure from any reason. The pressure
gauge connection must be in direct communication with the internal pressure
irrespective of other valves being open or closed. The drain will remove the contents
from the lowest point in the receiver and should be of sufficient size to prevent choking.
If it is possible to isolate the safety valve from the receiver a fusible plug must be fitted
to release pressure and discharge the receiver in the event of heating due to a fire in the
vicinity. The fusible plug should have a melting temperature of 150°C and a connection
must be made to discharge the contents in a safe manner.
All stop valves should be slow-opening valves to prevent sudden buildup of pressure
when connecting any pipe system.
The filling connection from the air compressors must be independent of all others. A
manhole must be fitted for access to allow internal inspection and cleaning. This should
be either fitted to the cylindrical shell with its minor axis arranged longitudinally or else
situated within the dished end. A flat internal joint face is made for the manhole door
gasket (Fig. 5.15).
Care must be taken when filling the receiver that the air supplied is free of oil or
moisture and is not of excessive temperature. Drains are fitted on the filling lines and
these should be open when not under pressure; they are closed after the start of filling,
opened periodically during filling, and opened again when filling is complete. Similarly
drains are fitted and used in discharge lines. The drain on the receiver should be 'blown'
periodically and particularly before using air from the receiver. Discharge from the
drain should be observed to give an indication of conditions within the receiver and
also the possible carryover from compressor discharge. Discharge valves should be
opened slowly to prevent shock waves in pipelines, and drains should be open while
this takes place. When air is being used from the receiver the pressure should not be
allowed to fall too low in case an emergency should arise. In most starting systems,
5.15 Starting air receiver
unless manoeuvring the main engine, one receiver is kept closed at full pressure while
the other is in use.
When opening up the receiver for maintenance, care must be taken to ensure that it is
isolated by locked valves from any pressurised part of the system and that internal
pressure is completely discharged before opening the manhole door. This door is
opened inwards and, if undue force is required, it should be confirmed that no pressure
remains in the receiver.
Before entering the receiver its interior must be well ventilated and at least two closed
stop valves should be situated between the receiver and any pipeline under pressure.
Internal cleaning must be thorough and care taken to avoid debris entering
connections. It is usual to overhaul all fittings while cleaning takes place. Careful
inspection must be made for possible internal corrosion.
Before applying protective compound all surfaces must be clean and dry and
connections plugged. Many compounds give off noxious fumes and great care must be
taken in confined spaces: protective clothing and goggles should be worn and a second
person outside the manhole must keep watching in case the person inside is overcome.
The compound must be allowed to dry thoroughly, plugs must be removed and all
connections cleared before boxing up the receiver and recharging.
are required to charge the starting air
receivers and supply compressed air for a variety of uses. Safety regulations demand at
least two compressors, each capable of supplying all demands. One must have an
independent drive for emergency use. With the high pressures demanded and to obtain
the best efficiency at full or part load, reciprocating compressors are invariably used.
These are generally two-stage compressors, electrically driven.
Pistons for each stage may be arranged in line, either as trunk pistons or with
crossheads. Alternatively, to reduce size and weight, a tandem arrangement with one
piston mounted above the other and driven from a common crank may be used. An inline two-stage trunk piston compressor is shown in Fig. 5.16. It can be seen that the
second stage (h.p.) piston is smaller than that for the first stage (J.p.). This is to
accommodate the corresponding reduction in the volume of air with a common piston
5.16 Air compressor
stroke. Compression and oil control rings are fitted to each piston; cylinders and covers
are water cooled.
Each stage has light, spring-loaded non-return suction and delivery valves of lowinertia stainless steel plates. Limited lift reduces hammering and will create high air
velocity to maintain cleanliness.
Water cooled heat exchangers are fitted to act as an intercooler, which cools the air
after first stage compression, and an aftercooler to cool it again after the second (final)
stage. Both cooler discharge headers have pressure relief valves and drains. The
maximum air temperature at compressor discharge must not exceed 93°C and a fusible
plug or alarm may be fitted.
The compressor intake, which has an air filter, should be placed in a position free of
any oil vapour. Compressors should be started with all drains open, the drains being
operated intermittently during running and opened again on shutting down.
The crankcase must meet normal regulations for pressure relief and will hold the oil
reservoir and self-contained oil pump.
Automatic controls for compressors and their drains may be fitted.
Before overhaul the maker's handbook should be read and
any special instructions must be adhered to. Ensure that all air receivers are fully
charged and that remaining air compressors are operational. This will provide for any
possible emergency in which compressed air is required.
The compressor for overhaul is now isolated with electric switches off and fuses or
circuit breakers out. Discharge valve and all cooling water connections are shut. The
precise order for dismantling will depend upon compressor design and only a general
description is given here. All parts should be cleaned upon removal and carefully
inspected for signs of wear, damage or maloperation.
Crankcase is opened and drained, oil strainer and internals cleaned. Lubricating
pump overhauled and clearances checked. Cooling water spaces are drained, opened,
cleaned and inspected. Corrosion fittings are renewed if necessary. Cooling water
pump, glands and bearing are overhauled, water strainer cleaned. Air suction filter is
cleaned, suction and discharge valves for each stage are dismantled; valves, cages,
springs and seats are checked and renewed as necessary. When dismantling or boxing
up valves, ensure that set bolts for locating valve cages etc are slackened back while
cover nuts are tightened. Inspect valve pockets for signs of excess oil carryover.
Remove cylinder covers, dismantle bottom end bearings and draw pistons with
connecting rods. Clean, inspect and gauge piston, compression and oil control rings
and renew rings if necessary. Clean, inspect and gauge cylinder liners. Remove
gudgeon pins after checking clearances. Examine bearing surfaces: bearings should be
renewed if excessively worn.
Overhaul bottom and end bearings, check clearances and journal surfaces. These
bearings are usually of the thin-wall type and should be renewed if unduly worn. Oil
holes and passages must be checked and bearing shells must have a good nip in the
keeps. Clearances are taken by inserting feeler gauges or by jacking the bearings and
measuring lift by a clock gauge.
Main bearings are dismantled, inspected and weardown checked. Shaft may be
turned by hand to check alignment. All lubrication connections must be clear and pipes
secured. If a drive chain is fitted for pumps, this should be inspected. Before boxing up
cylinders and crankcase, the piston end clearances must be checked at top dead centre
Intercooler and aftercooler must be opened up. Water spaces should be cleaned and
examined for corrosion and relief valve or bursting disc checked. Air side is opened up,
cleaned and inspected and tubes cleaned internally. Drain pockets are cleaned. Drains,
relief valves and discharge valve should be overhauled.
If a cylinder mechanical lubricator is fitted, this should also be cleaned and
overhauled together with the lubricator points in the cylinder liner.
Holding-down bolts and shaft coupling bolts should be checked for tightness. It is
assumed that the compressor motor will also have been overhauled.
After boxing up all parts, cooling water may be applied. Ensure adequate venting
and that all parts are completely filled. Check for water leaks. Recharge crankcase with
fresh oil up to the required level. The compressor should now be barred over or turned
by hand with drains open. If this is satisfactory, connections can be remade and the
compressor run for ten minutes with drains open. Stop and feel round all bearings and
coolers. If satisfactory restart compressor, check drains for excess oil, check all
temperatures and pressures together with electrical load. Drain thoroughly before
returning to service. Check all pressure gauges when discharge pressure has built up.
Check vents from water spaces. If automatic controls are fitted, these must be tested
and put into service. Check automatic pressure or temperature cutouts or alarms,
together with remote reading pressure gauges.
Dividing the total compression of air
between a number of stages gives several advantages. For normal air start pressures we
only require two stages, but for higher final pressures more stages would be used. Air
temperature rises during compression and multi-staging allows the air to be passed
through intercoolers between stages, returning it to its original temperature and
bringing its compression (P-V) curve closer to that for an isothermal operation.
5.17 Two-stage compression
Temperatures are limited to values at which the lubrication of cylinder walls can be
Multi-staging with a number of separate pistons will give mechanical improvements
of compressor load and balance, reductions in size and mass, robust construction of
smaller, high pressure parts and reduction in clearance volume losses. Air discharged
from the first stage of compression is passed through the intercooler where it is cooled
at constant pressure, causing a reduction in temperature and volume. Air may make
one pass (usual in first stage) where temperature change is moderate, or two passes in
later stages. The cooling water makes a single pass in both cases.
The gain in efficiency can be illustrated on a pressure-volume diagram (Fig. 5.17) on
which it can be seen that isothermal compression (at constant temperature) causes least
area of diagram and therefore least power absorbed. Returning the compression line to
this curve between stages will reduce power absorbed in the second stage by the shaded
area shown.
A main consideration is the reduction in air temperature which makes lubrication of
later stage liners possible and generally adds to the safety of the system. Intercoolers
are fitted with pockets and drain valves which allow removal of moisture condensed
during cooling together with any excess lubricating oil mist carried over in the air.
An aftercooler is fitted after the final stage of compression. This does not improve
the compressor efficiency but aids removal of moisture and oil and reduces air
temperature before passing it to the air receiver.
An intercooler contains banks of copper tubes through which the air passes. These
are expanded into a brass tubeplate. If straight tubes are used, a header is required to
permit thermal expansion. U tubes in a cooler cause the air to make two passes and
tubes are free to expand. A cast iron outer casing permits circulation with cooling
5.18 Compressor
A relief valve should be fitted to the air connection and a bursting disc fitted to the
water casing. The bursting disc will relieve excess pressure in the water casing in the
event of failure of an air tube.
Little maintenance is required by intercoolers when in use. Correct temperatures and
adequate water circulation must be maintained and air drains opened when starting or
stopping the compressor and at frequent intervals while operating. During overhaul,
coolers must be cleaned internally and checked for corrosion.
The effect of operating with fouled intercoolers will be to raise the air temperature:
this will cause lubrication difficulties in the compressor, efficiency will rapidly fall off,
and there may be the risk of an explosion caused by overheated parts in the air system.
In order to manoeuvre a ship, the propelIer thrust must be reversible, by means of
reversing the propelIer drive or by altering propelIer pitch. Controllable pitch
propellers, or diesel-electric drive, allow the use of unidirectional (non-reversible)
engines. In systems of limited power using medium or high speed engines through
gearboxes, arrangements of clutches and reverse gears may be used. In many ships,
however, it is necessary for the main engine to be reversible and able to operate
efficiently in both ahead and astern directions.
To run in the astern direction, all the operations in the engine cycle may need
retiming. The starting air must first rotate the engine in the reverse direction and this
will require retiming the distributor to supply compressed· air to the appropriate
cylinders in the correct order. The retiming may be carried out by altering the position
of the distributor cam with respect to its drive from the main camshaft.
The number of other readjustments to be made and the methods used depend upon
the engine cycle and type. Four-stroke engines will require a change in timing of the
fuel pumps, a different change for the air inlet valves, and yet another for the exhaust
valves. To obtain alI these changes on the same camshaft a separate set of astern cams is
fitted. Each astern cam is fitted to the camshaft adjacent to its corresponding ahead
cam (Fig. 6.1). The reversing procedure is then carried out by moving the whole
camshaft axialIy, which moves the ahead cams clear of their followers which now
engage the astern cams. Ramps (sloping sides) fitted between corresponding ahead and
astern cams cause the folIower's rolIer to slide smoothly from one to the other. The
axial movement is carried out by a hydraulic cylinder fitted to the camshaft; locking
devices and safety cutouts ensure that the camshaft has carried out its fulI axial
movement and is in the correct position before the engine can be restarted. To maintain
alignment of the camshaft drive, a spline coupling may be necessary.
Turbochargers are of course unaffected by reversal of the engine, but engine driven
pumps must be reversible.
Large two-stroke engines have scavenge ports which control scavenge timing. This
must therefore be symmetrical and wilI thus be unchanged when reversed. Engines
operating with constant pressure turbocharge systems have almost symmetrical
exhaust valve timing. Consequently no change in timing is necessary for exhaust cams.
Fuel pump timing must be readjusted since it will be the opposite flank of the cam
which will now raise the pump plunger to deliver fuel. There are alternative methods
employed to chllflge the fuel pump timing without altering the main camshaft. Two
such systems are illustrated and described. In both cases the main camshaft drive
timing is not altered and therefore any other drives taken from this remain
synchronised with the engine while running astern. This is important if a balancer
system is fitted.
6.2 Reversing servomotor, Sulzer RTA engines
Sulzer RTA engines have oil pressure operated hydraulic 'lost motion' servomotors
on the camshaft which rotate the fuel pump cams to their astern positions (Fig. 6.2).
Fuel pumps and their cams are grouped in pairs along the camshaft and a servomotor
is fitted for each pair of adjacent cams. Oil pressure locates and secures each
servomotor in its correct position while the engine is running. A similar servomotor is
used to re-time the starting air distributor at its drive from the camshaft. 'Lost motion'
is the term used to indicate that the timing has been retarded, or moved back, through a
given angle with respect to the 'new' direction of rotation.
MAN-B & W MC engines have their fuel pump cams fixed directly to the camshaft
but the follower rollers can be displaced to alter the pump timing as shown in Fig. 6.3.
The link which displaces each follower is actuated by a pneumatic cylinder and piston,
powered by compressed air from the starting system. The link is self-locking in either
position and a sensor is fitted to each pump to cut off fuel delivery if the link is not
locked in its correct position.
6.3 MAN-B
& W reversing system
Earlier models of two-stroke engines use elaborate systems of servomotors to rotate
their whole camshaft to reverse the timing.
A governor automatically controls engine speed by regulating the fuel supply. Modern
practice requires a governor to be sensitive to small changes in speed and then be
capable of returning the engine to a set speed. If it maintains the engine at a set speed,
irrespective of changes in load and power, it is said to be isochronous. When fitted to
main engines, there must be the facility to adjust the set speed on the governor, which is
then said to be a variable speed type.
Earlier large, slow running, direct drive main engines were inherently stable under
normal conditions. Hand controls were adjusted to give a set speed regulated by the
propeller load. Only an overspeed cut-out was considered necessary to prevent racing
and damage in an emergency such as a shaft fracture, loss of propeller or the propeller
leaving the water in heavy weather.
Modern practice requires an engine to be fitted with a governor giving close speed
control for fuel economy, in particular for engines driving electric alternators. The type
and size of governor must be designed to match closely the characteristics of the engine.
M any diesel engines are fitted with mechanical-hydraulic type governors. These use a
hydraulic servo-piston which creates sufficient power to operate the engine fuel pump
controls without any loss in sensitivity of the mechanical speed-sensing ballhead.
Hydraulic oil pressure is taken either from the engine's lubrication system or from a
separate gear pump operated by the governor drive. This is usually preferred as it
allows the use of selected oil maintained in a clean condition. It will require oil pressure
boosters, charged from the starting air system, to ensure sufficient power to manipulate
the engine fuel pump controls when starting from rest.
Fig. 6.4 shows the working parts of a mechanical~hydraulic,
isochronous type
governor, with variable speed control. The ball head consists of two identical,
eccentrically pivoted, flyweights mounted on opposite sides of a rotating sleeve. Speed
of rotation is proportional
to engine speed and may be by direct drive from the
camshaft or through a stepup gear to increase governor speed and sensitivity. Dampers
are fitled in the drive to damp and reduce any torsional vibrations. A crank on each
flyweight bears on a sliding compression plate, which is held in equilibrium between
6.4 Mechanical-hydraulic engine speed governor
forces from the flyweight and a compressive force from compression of the governor
speeder spring. A conical or trumpet shaped spring is used as this gives a more effective
relationship between spring compression force and the centrifugal force on the
flyweights. If the engine speed slows, the centrifugal force on the flyweights decreases
allowing the moving sleeve to be lowered by the spring. If the engine speed increases,
the centrifugal force increases and flyweights raise the sleeve.
The sleeve is connected to a hydraulic piston valve by a rod and there is also a
compensation piston (pilot valve receiving piston) attached to this rod which is
subjected to a force if oil pressures on each side of it are unbalanced. This force wiIl
then add to, or subtract from, the speeder spring compression force at the ballhead, the
effect being to cause a slight, temporary change in engine speed (temporary speed
As the engine speed slows, forces on the ballhead move the piston valve down,
uncovering the port and allowing oil under pressure to pass into the servo system. The
pressure acts on the buffer piston, forcing it to the right, compressing the right-hand
buffer spring, displacing oil on that side and raising the power piston, which in turn
increases the engine fuel pump settings. Due to compression forces on the buffer
springs, the oil pressure on the right of the buffer piston wiIl be slightly less than that on
its left. These pressures are communicated to opposite sides of the compensation piston
giving it an upward force which adds to the compression on the speeder spring. This
assists in returning the piston valve to its shut position even though engine speed has
not yet returned to its set value.
The pressure difference across the compensation piston is only temporary because
the needle valve allows oil to pass slowly and balance both sides. The buffer piston
springs also return it to its mid-position.
The ballhead has now returned to its original setting (at the set speed), while the
controls are at a higher setting commensurate with the demand for greater power.
Conversely, if the engine speed increases the ballhead raises the piston valve,
releasing oil from the servo system. The buffer piston moves to the left: the power
piston moves down, reducing fuel settings to slow the engine. There is now a downward
force on the compensation piston and this together with the speeder spring
compression forces the piston valve down to close the oil passage. These changes are
also temporary until pressures equalise through the needle valve. The flow rate past the
needle valve is adjusted to match the engine's response characteristics (the time it takes
to respond to a change in the fuel setting).
Additional attachments to the governor system allow control of a permanent speed
droop, which is necessary if two or more engines are to operate in tandem through a
common geared drive. In this case their speeds are locked together, but their respective
power settings must be governed ..
Alternator drive engines wiIl require synchronisers to control frequency.
Maintenance of governors must ensure that working parts are in a free and unworn
condition. Oil passages must be clean. The oil supply must be maintained and
replenished when necessary. A heat exchanger may be fitted to the oil circuit to regulate
a constant oil temperature.
Definitions of terms used in governors and control engineering can be found in the
book Instrumentation and Control in this series, or from other standard works. The
term 'speed droop' has been used and a simple explanation of this follows:
Speed Droop is the change in operating speed to cause the fuel controls to move from
full to zero, or vice versa.
Temporary Speed Droop means that the governor will, after a short delay, return to
its original setting giving constant engine speed.
Permanent Speed Droop means that the governor will take a different position and
normally a different speed as load changes.
It is possible that an engine governor may fail, or not act quickly enough in an
emergency situation to prevent damage to the engine. Thus it is usual to have an
additional simple overspeed cutout fitted. This should act independently of the main
speed governor, and in the event of overspeeding it will immediately cut power off the
engine, either by raising fuel pump plunger followers clear of their cams or by opening
pump suction valves, either of which will shut off fuel and power. Cutouts are fitted at
each pump and they can also be used, if required, to shut power off one unit of the
engine without affecting the remainder of the safety system.
Power to operate the emergency system may be electrical, pneumatic or hydraulic.
The system may be utilised for other emergencies such as failure of vital engine
operating systems. It can be coupled to remote emergency stop buttons. Any override
of the overspeed trip can only be carried out manually at the engine.
Simple electronic control systems can be
designed to carry out the functions of a mechanical-hydraulic governor including
facilities to control load-sharing, synchronising, power sensing etc. They may readily
be incorporated as part of an overall control system to include starting and stopping
routines, and can be programmed to avoid running at critical speeds, to prevent
overspeed and for other safety measures. Fig. 6.5 shows a simple block diagram of the
A speed sensor such as a magnetic pickup transmits a signal representing the actual
engine speed and this signal is compared with one for the desired (set speed) value. The
error (difference) in these is amplified and passed to a controller and actuator which
generates sufficient mechanical force and power to operate the engine pump controls.
Safety and Operation
This section includes serious dangers which may exist if a diesel engine is not correctly
operated and maintained. Fortunately most of these rarely occur, but operators must
make themselves aware of such dangers, understand how they can develop and know
how they are prevented. Regulations demand the fitting of certain safety devices and
the operator must be familiar with their function and maintain them in good working
The cause of a crankcase explosion is a 'hot spot' or overheated part within or adjacent
to the crankcase of an operating engine. Under normal running conditions the air in a
crankcase will contain oil droplets formed by lubricating oil splashing from the
bearings onto moving surfaces. This mixture will not readily burn or explode.
Crankcase lubricating oil should normally have a high closed flashpoint (above 200°C)
and this must be maintained in order to reduce risk of explosions. The most common
cause of lowering the flashpoint is contamination with fuel oil.
Local hot spots may arise due to overheating of bearings, piston rod gland, timing
chain, hot combustion gas or sparks from piston blowpast in engines where no
diaphragm is fitted; or from fires in spaces adjacent to the crankcase, such as scavenge
trunks etc.
Such sources can be eliminated by proper maintenance, correct lubrication and oil
condition, cleanliness and by avoiding overloading the engine. The general use of white
metal bearing materials which have moderate softening and melting temperatures also
helps to avoid a rapid rise in temperature.
If a hot spot exists, some oil will come into contact with it and will be vaporised,
circulate to cooler parts of the crankcase and there condense to form a white mist of
finely divided oil particles well mixed with air. This mist is combustible within certain
concentrations. If the mist should now circulate back to the hot spot in such
concentrations, it will be ignited and a primary or minor crankcase explosion will
occur. This explosion causes a flame front and pressure wave to accelerate through the
crankcase, vaporising further oil droplets in its path.
The pressure shockwave may build up sufficiently by the time it reaches the
crankcase casing to rupture crankcase doors or panels, unless otherwise relieved. In
this event, the low-pressure wave following will draw air back into the crankcase where
it will mix with vaporised and burning oil to cause a secondary or major explosion of
such intensity as to cause widespread damage. It may start fires in the vicinity and
injure personnel.
Should the conditions of a hot spot arise within the crankcase, a watchkeeper may
detect them by irregular running, engine noise, increase in temperatures, by smell and
by the appearance of the dense white oil mist.
Detection by instruments may be by temperature sensitive probes within the
crankcase near the bearing oil returns, or more commonly by the use of a crankcase
mist detector (see Fig. 7.1). This operates visual and audible alarms in the event of a
white mist being formed at well below the concentrations required for explosive
Crankcase explosion relief valves (see Fig. 7.2) must be fitted to all but the smallest
crankcases. They open automatically at moderate pressures allowing the pressure of
the primary or minor explosion to be dissipated and preventing possible rupture of the
casing. The valves instantly close when the pressure drops and thus prevent the ingress
of air, eliminating the possibility ofa major explosion. Valves are fitted with wire gauze
to prevent the emission of flames and may have external deflectors to aim hot gases in
directions where they will do least damage.
In the event of a hot spot or explosive condition being detected, the engine must
immediately be slowed and should be stopped as soon as possible to allow the
overheated parts to cool down. It may be advisable to operate the engine turning gear,
with indicator cocks open, to prevent seizure of overheated parts.
Personnel should avoid the vicinity of relief valves.
On no account must the crankcase be opened until the parts have cooled, as such
action may allow the ingress of air and precipitate a major explosion. When parts have
cooled, inspection and maintenance or repairs may be carried out.
Subdivision of crankcases will inhibit the buildup of high velocities and pressures of
flame propagation through the crankcase from a primary explosion.
Crankcase doors should be of robust construction to prevent rupture. Any internal
crankcase lighting must be flameproof. Vent pipes fitted to crankcases should not be
too large; they must be led to a safe place, remote from the engine, and fitted with
gauze. Oil drain connections from the engine must extend below the oil level in the
sump. There must be no common communication between the crankcases of two
In some cases inert gas flooding systems may be fitted to crankcases. If these are
used, after cooling the crankcase must be well ventilated before personnel may enter
for inspection.
Fig. 7.1 is a diagrammatic view of a Graviner oil mist detector which may be fitted to
monitor samples of the air and vapour mixture taken continuously from the crankcase
of a diesel engine. Such a device will detect the presence of oil mist at concentrations
well below the level at which explosions may occur giving a warning in time to allow
avoiding action to slow the engine and prevent either serious damage or an explosion.
The detector consists basically of two parallel tubes of equal size, each having a
photo-electric cell at one end which generates an electric current directly proportional
to the intensity of the light falling on its surface. Lenses are fitted to seal the ends of each
tube but allow light to pass. Two identical beams of light from a common lamp are
reflected by mirrors to pass along the tubes onto the cells which are then in electrical
One tube is sealed to contain clean air and is termed the reference tube. The other, the
measuring tube, has connections through which samples of the crankcase vapour are
drawn by an electric extractor fan. If a concentration of oil mist is present in the
sample, light will be obscured before reaching the cell of the measuring tube: electrical
balance between the two cells will be disturbed and an alarm will be operated.
Sampling points should be fitted to each cylinder crankcase and their connections
are brought to a rotating selector valve which is driven from the fan motor. This
repeatedly connects each sampling point to the measuring tube in sequence. Sampling
connections should not exceed 12.5 metres in length and must slope to ensure positive
drainage of oil; they must avoid any loops which could fill with oil.
In the event of oil mist being detected the rotator stops to indicate which sampling
point is concerned. The instrument must be reset before the alarm ceases and sampling
will recommence its sequence.
The extractor fan is very small and after testing the samples are exhausted to
The detector should be tested daily and the sensitivity checked. Lenses and mirrors
should be cleaned periodically.
In this model the total mist concentration is measured with respect to clean air. An
alternative model draws samples through both reference and measuring tubes. A
mixture from all cylinder crankcases is passed through the reference tube while
comparison is made with samples from each cylinder crankcase and also from the
atmosphere. In this manner a general sample of all cylinders is compared with the
normal atmosphere and each individual sample is also compared against the average.
Fig. 7.2 shows a crankcase explosion relief valve which may be fitted to a diesel engine
crankcase. It consists of a light spring-loaded non-return disc valve of simple
construction. The valve disc is of aluminium alloy which reduces its mass and the
inertia to be overcome when opening or closing valves rapidly. The large diameter
spring will give sensitivity and allow the valve to float. The absence of a valve spindle
eliminates the risk of the valve jamming.
The valve landing must make a gas and oil tight seal when closed and a non-stick oil
and heat resisting rubber ring is fitted to the disc face.
An external aluminium valve cover secures the valve spring and acts as a deflector to
direct any gas emitted over an arc of 120 aimed where it can do least damage. Inside
the crankcase is a dome-shaped flame trap made of several layers of woven, mild steel
wire gauze. This projects into the crankcase where it will become wetted with oil mist or
splash from adjacent bearings. When wet with oil the gauze dissipates heat at a greater
rate and becomes more effective as a flame trap. Free area of the gauze must at least be
equal to the area of the open valve.
7.2 Crankcase pressure relief valve
The valve assembly is secured to an aperture cut in the crankcase by a number of
cover studs and distance spacers; these act as guides for the valve disc. The valve spring
is designed to allow the valve to open under an internal pressure of approximately
5 kN/m2 and will close automatically when pressure has been relieved.
Regulations demand that for engines of over 300 mm bore, one crankcase relief
valve of approved design is fitted to each crankcase and chain case. The combined area
ofthe relief valves should be not less than 115cm2 per cubic metre of crankcase volume.
The free area of each valve is to be not less than 45 cm2• For smaller engines a reduction
in the size and number of valves is allowed. These regulations also apply to the
crankcases of large air compressors, etc.
Crankcase doors should be robust to prevent damage or rupture before relief valves
operate to relieve pressure.
Valves will require little maintenance but should be tested periodically by hand; the
spring should be inspected and the gauze cleaned.
A scavenge fire may be caused by the ignition of unburned oil and carbon which has
been blown from the engine cylinder into the scavenge spaces. This may include
unburned fuel or cylinder lubricating oil and may be due to incorrect combustion
caused by a defective injector, faulty fuel pump timing, incorrect fuel condition, lack of
scavenge air, partially choked exhaust, low compression, afterburning, by operating
the engine at overload conditions, or due to defective piston rings, badly worn cylinder
liner, or by wrongly timed or excessive cylinder lubrication.
The oil wiII build up in scavenge spaces where it will become carbonised by further
heating and wiII then reach a condition in which it can burn in the presence of air. It
may be ignited by hot gases and burning particles from bJowpast of piston rings.
Indications of a scavenge fire are loss in power and irregular running of the engine,
high exhaust temperatures of corresponding units, smoke in exhaust gas, high local
temperature in scavenge trunk, surging of turbocharger, and sparks and smoke
emitted from scavenge drains.
If a fire is detected engine speed must be reduced, fuel shut off from the affected units,
cylinder lubrication increased and scavenge drains shut. A minor fire may shortly burn
out and conditions wiIIgradually return to normal. The affected units should be run on
reduced power until inspection of the scavenge trunking and overhaul of the cylinder
and piston can be carried out at the earliest safe opportunity.
Should a fire persist, if there is a risk of the fire extending or if the scavenge trunk is
adjacent to the crankcase with risk of a hot spot developing, the engine must be
stopped, normal cooling maintained and it may be advisable to engage turning gear to
prevent seizure.
Fire extinguishing medium should be applied through fittings in the scavenge trunk:
these may inject carbon dioxide, dry powder or smothering steam.
Measures must be taken to avoid the spread ofthe fire and if necessary cooling may
be applied to the outer surfaces. On no account must the scavenge trunk be opened up
while it is stilI hot or danger of an explosion may arise.
Personnel should avoid standing close to relief valves.
After extinguishing the fire and cooling down, the scavenge trunking and scavenge
ports should be cleaned and the trunking together with cylinder liner and water seals,
piston, piston rings, piston skirt, piston rod and gland must be inspected. Tightness of
tie bolts should be checked before restarting the engine. Fire extinguishers should be
recharged at the first opportunity and faults diagnosed as having caused the fire must
be rectified.
To prevent scavenge fires good maintenance and correct adjustment must be carried
out. Scavenge trunking must be periodically inspected and cleaned and any buildup of
contamination noted and remedied. Scavenge drains should be blown regularly and
any passage of oil from them noted.
Safety devices fitted to scavenge systems may include an electrical temperature
sensing device fitted within the trunking which will automatically sound an alarm in the
event of an excessive rise in local temperature (above 200°C).
Pressure relief valves consisting of self-closing spring-loaded valves are fitted and
should be examined and tested periodically. Other safety devices could include the
scavenge drains and the fitting of a fixed fire extinguishing system. With such a system
maintenance will include inspection and cleaning of injection spreaders, weighing of
gas containers and rotation of dry powder containers to keep the powder freeftowing.
The alarm systems should be tested daily.
When any fire occurs in the engine spaces, the fire alarm should be sounded and
assistance or advice summoned. Cleanliness in these spaces will help to prevent fires
Possible causes of an explosion in high-pressure starting air pipelines are the
continuous leaking of a defective cylinder non-return valve while the engine is
operating, or such a valve sticking in the open position during manoeuvring.
Under normal operation, some lubricating oil mist may be discharged from the air
compressor to the air start system. This oil may be from excess compressor cylinder
lubrication, from faulty oil scraper rings, or may even be suspended oil vapour
contaminating the engine room atmosphere and drawn in at the compressor suction.
Oil discharge is kept to a minimum by draining the aftercooler, air receiver and starting
system. If small quantities of lubricating oil do get passed into the starting air system,
they will deposit as a thin moist film over internal pipe surfaces but are not readily
If a cylinder non-return valve should leak while the engine is in operation, some hot
gas, possibly with unburned fuel and cylinder lubricating oil, may be blown through
the valve to the adjacent air manifold. With further heating from the leaky valve, this,
together with the already deposited oil film, will carbonise and form incandescent
carbon. If starting air is applied to the system while still hot, the high pressure air
coming into contact with the burning carbon may cause an explosion. Such an
explosion will cause a flame to pass back through the air start pipe system, evaporating
the deposited oil film and igniting it in the presence of air. Very high velocities and
shockwaves are generated which may rupture pipes and fittings.
Alternatively, if excessive oil has entered the air start system, a mixture of air and oil
droplets may be discharged through the open cylinder non-return valve during
starting. This spray may ignite due to high temperatures in the cylinder, causing a flame
through the still open valve to the air manifold.
To prevent an explosion, air start valves must be correctly maintained and
lubricated to ensure correct timing and free movement with positive closing. Oil in the
system must be kept to a minimum, pipe lines must be drained and cleaned internally
when necessary and oil discharge from air compressors must be kept to a minimum by
good maintenance.
To minimise the effects of such explosions, the air start manifold to each cylinder
valve must be fitted with a flame trap and ample relief valves; bursting caps or discs
must be fitted to relieve excess pressure. An isolating non-return valve is fitted at the
outlet from the main control valve.
A leaking air start valve can be detected while the engine is operating by local
overheating of the pipe adjacent to the valve. If this should occur, the engine should be
stopped at the first opportunity and the valve replaced. As a temporary measure, a
blank flange may be fitted to the air manifold connection to isolate this valve, but since
it will no longer operate it is possible for the engine to stop in a position from which it
will not readily restart. Warning of this situation must be transmitted to the bridge.
When the air start system is not in use it must be shut down and all drains should be
opened. Drains are closed after air pressure has been put on the system. Air start valves
should be lubricated, where such fittings are present, before the start of a voyage or
during a long voyage without use. Cylinder non-return valves may be tested for
tightness while in port by shutting isolating valves on the control air connections and
applying air to the start manifold. Escape of air through open indicator cocks will
disclose a leaking valve.
In the event of overheating of the discharge from the air compressor to the filling
line, an explosion would be possible between the compressor and the air reservoir.
Overheating may be caused by failure of compressor intercooler and circulating water.
In this case the high temperature within the high-pressure stage will make operation of
the compressor and its cylinder lubrication difficult.
Excess discharge temperature is detected either by an alarm system or a fusible plug
which will melt at 121°C to give warning.
In some engines it is possible to gag an air start valve. It must be stressed that gagging
does not stop the valve leaking and should not be adopted.
Fig. 7.3 shows a half-section ofa cylinder safety valve fitted to the covers ofa large twostroke engine to release excess gas pressure from the cylinder. The valve is of stainless
steel with a mitre seat and is loaded by compressing a helical spring. The lower end of
the spindle is radiused to allow the valve to align with its seat; valve lift is limited by a
shoulder at the top of the spindle. The spring keep or cap nut is locked in position to
regulate the correct spring compression and the valve is set to lift at not more than 20%
above the designed engine pressure.
Maintenance consists of cleaning and inspection at the same intervals as cylinder
overhaul. The valve and seat should be examined and reground if necessary, the spring
checked for its free length with no warping. After assembly the valve should be set and
pressure tested.
7.3 Cylinder relief safety valve
Lifting of a safety valve relieves the dangerous pressure in the cylinder and also
warns of incorrect conditions. The reasons for this must be ascertained and corrective
action taken. The high temperature gas and flame expelled may damage the safety
valve seat.
Valves may lift during manoeuvring or slow running for a number of reasons.
Ignition may be violent if engine speed is too slow, if a fuel injector has leaked during
priming, or a fuel pump setting is too high or incorrectly timed. Should starting air be
used as a brake when stopping an engine for reversal or during a 'crash stop', high
compression pressures at about top dead centre may cause safety valves to lift
Valves lifting during running may indicate an overload or timing faults including the
camshaft drive.
Safety valve area is only designed to relieve excess gas pressure and may be
insufficient to prevent damage in the event of cooling water or oil leaking into the
cylinder. Before starting an engine for the first time it should always be turned slowly
with indicator cocks open to expel any leakage.
Black smoke exhaust indicates incomplete combustion of fuel and unless the reason is
obvious, it is necessary to ascertain whether one particular unit or a number of units are
causing the smoke.
A particular unit making smoke may be detected by shutting fuel off from each
cylinder in turn, while observing the exhaust condition. This method will be
inconclusive if more than one unit is causing smoke. On some engines, test cocks are
fitted to the exhaust branches from each unit: by opening these in turn the exhaust gas
from each unit can be sampled. The gas may be allowed to blow onto a piece of wet rag
or damp paper on which a deposit will show from units causing smoke.
Ifsmoke is from one unit only, the probable cause is the fuel system for that cylinder.
Injector faults may be a worn or partially choked atomiser, leaking valve, broken or
incorrectly set spring. Fuel pump faults include incorrect timing, worn plunger, and
also worn, slack or wrongly positioned cam. Low compression results from a worn
cylinder liner, worn or broken piston rings or leaking cylinder valves. Insufficient
combustion air in one cylinder may show choked scavenge ports in a two-stroke, or
faulty air inlet valve in a four-stroke engine. Diagnosis of the particular fault can be
assisted by checking the exhaust temperature, use of combustion monitoring or by
taking indicator diagrams.
If all units are producing smoke, reasons may be fuel condition or lack of
combustion air. Fuel temperature and viscosity should be checked. Lack of air may be
due to fouling of the turbocharger either on the gas side or the air side, including inlet
filters and passages in the charge air cooler. An exhaust gas back-pressure may be
caused at a waste heat boiler, etc.
Overloading an engine or too high a rate of acceleration causes smoke. Turbocharger lag occurs as engine speed is increased, but this should be overcome as exhaust
energy builds up and accelerates the turbocharger.
A scavenge or uptake fire willcause smoke, but is usually obvious by other effects. If
an uptake fire is the cause ofthe smoke, it may have little effect on engine operation. It
can be caused by unburnt carbon carried over from incomplete combustion.
External surfaces in the vicinity of the fire should be cooled and any combustible
material in the immediate area removed. If the engine is stopped, starting air should
not be applied while the fire continues.
Remedial action will depend upon the cause: faulty injectors can be replaced, fuel
condition or timing rectified, turbochargers may be water washed, etc.
In trunk piston engines, smoke may also indicate the failure of scraper rings, causing
excessive lubricating oil to be burnt in the cylinder.
Any fluctuation in the level of lubricating oil in the engine sump or drain tank must be
investigated; it may be due to a number of causes. It may occur to a limited extent due
to movement of the ship----either pitching or rolling-particularly if the oil level gauge
is not situated at the centre of the tank. The reliability of the level gauge should also be
checked since float or transmission equipment may be faulty. A check must be carried
out to ascertain that no valves have been incorrectly opened or closed, allowing oil to
enter or leave the system.
If the oil level has risen it is possible that there is a leak into the system, which may be
water from leaks in the cooling system, such as glands or seal rings, allowing
contamination of the crankcase from cylinder jackets, water cooled pistons etc. There
may be a leak at the oil cooler, although the oil pressure should be greater than that of
sea water to prevent this. A test for water will indicate such a leak.
Although fuel contamination could occur under certain circumstances, it is unlikely
to be at such a rate as to be noticed at sump level. A flashpoint test on the oil will
ascertain if fuel is present. Choking of the oil filters or failure of the pumps may allow a
level rise, but this will be evident from the pressures within the system and perhaps
pressure alarms.
If the oil level has fallen there may be an oil leak in the system and this must be
searched for at all pumps, filters, pipes, joints and glands, at the crankcase and the
drain tank. The water seal in the oil purifier together with the purifier pump should also
be checked. Choking of oil drain grids within the crankcase will cause the oil level to
build up there and not return to the drain tank. The possibility of the pressure relief
valves allowing oil to leak from the system should be examined.
In trunk piston engines, failure of oil scraper rings may cause the consumption of
lubricating oil to be high. This should be evident by high exhaust temperatures, smoke
and blowpast from jammed piston rings. It may cause danger of fires or an explosion.
Excessive noise from a crankcase is generally local to one particular area and may have
a number of causes. Sudden, loud noises probably indicate the breakage of parts and
may cause overheating. In this case the engine must be slowed or stopped, but
crankcase doors must remain closed until safe entry can be made to investigate.
Worn or slack camshaft drive gears or chains tend to vibrate in that part of the
crankcase. Overheated bearings or glands also generate noise. Knocking may be heard
from slack bearings, particularly cross head guides, as a result of wear, misalignment,
slackening or breakage of bolts, studs or other fixtures.
Piston cooling system noises include 'water hammer' or knocking from pipes or
Lubrication faults of insufficient oil pressure, failure or choking of connections or
passages, reduction in lubricating properties due to excessive temperature or
deterioration, and jets of oil splashing against fast-moving surfaces will add to the
noise level.
Overloading the engine or early ignition will transmit shock loads through the
running gear and bearings.
In trunk piston engines, additional crankcase noise is heard in the event of piston
blowpast or an overheated piston or liner.
Any change or sudden noise pattern must be investigated.
This can be detected during operation by high local exhaust temperature with possibly
noise and smoke. The compression pressure and peak firing pressure in the cylinder will
be reduced.
If allowed to continue, damage to the exhaust valve and seat will increase due to
burning by high velocity, very hot gases. Partly burned fuel may pass to exhaust grids,
turbochargers, silencers and uptakes causing fouling, loss in efficiency, surging of
turbocharger, uptake fires or explosions under certain circumstances.
The valve must be changed at the first opportunity and until then fuel should be shut
off that cylinder. To prevent such leakages exhaust valves must be changed and
overhauled at regular intervals and valve tappet clearances checked. Excessive powers
must be avoided and combustion efficiency maintained.
The power and corresponding torque transmitted to the crankshaft by each cylinder of
a reciprocating engine varies throughout its cycle. The torque is greatest during the
expansion stroke and least during compression, and this is repeated from each cylinder
in turn. The varying cyclic torques tend to set up torsional vibrations in the engine and
propeller shafting. The frequency of any peak torque can be found as: number of
engine cylinders x number of cycles per unit time.
These vibrations are normally damped and their energy dissipated by the elasticity
of mechanical parts, oil films at bearings, and the work absorbed at the propeller. An
engine system, including all the shafting, working parts, bearings, propeller etc, is
manufactured from metal parts which have natural elasticity. There will be a natural
frequency for the whole system at which it will tend to vibrate if caused to do so.
Should the frequency of peak torques from the running engine coincide with, or be a
multiple of, the natural frequency for the system, a condition of resonance is set up.
Resonance may cause unacceptable torsional vibrations and consequent stresses to be
induced. Engine speeds at which this can occur are termed critical speeds.
To prevent the buildup of these high stressesj engines must not be run continuously
at, or close to, critical speeds. Warning of any critical speed within the operating range
of the engine must be given. Remote speed control systems must be programmed to.
avoid operating the engine at critical speeds.
Surging is usually first detected by noises from the air compressor side. The noises may
differ but may include howling, whining, knocking and banging. The turbocharger
may race intermittently; air discharge pressure will drop and behave erratically. Unless
the surging becomes frequent or co.ntinuous it should no.t harm the turbocharger or
engine, but loss of air will cause po.o.rcombustio.n with all its relative problems.
A sudden reduction in engine speed with the co.rresponding drop in demand for air
may cause surging, but this is tempo.rary and stable co.nditions are so.on re-established.
Turbo.chargers are designed to. avoid surging under normal operating conditions
and a surge will indicate abnormal change in either air or exhaust gas flow to or from
the charger. Turbochargers may be prone to surging if an engine is o.perated with one
or more units not producing power.
Provided no sudden change in engine po.wer has taken place, surging in a normally
stable system while running can usually be traced to fouling. If this is o.n the air side it
may be within the compressor, including the impeller and diffuser, intake filters, charge
air cooler or choked scavenge po.rts. On the gas side it may be in the turbine at the
nozzle ring or moving blades. It may also be due to choked grids o.r in a waste heat
As a tempo.rary remedy, stability can be reinstated by reducing the engine speed and
po.wer. Measures must be taken as so.o.nas co.nvenient to.remo.vethe fouling. The first
measure which can be carried out while the engine is operating is to water-wash the
turbocharger (see Chapter 3).
Sho.uld the turbocharger vibrate and slow do.wn, it should be sto.pped and examined
fo.r bearing damage o.r Io.ssin balance.
A scavenge fire may lead to surging, but this beco.mesobvious from its o.ther effects.
Abrasion, 48
'A' frames, 76, 78
Afterburning, 25
Aftercooler, 107
Air compressor, 106
Air cooler, 107
Air distributor, 103
Air filter, 107
Air inlet valve, 42
Air spring, 42
Air starting system, 103
Alignment, 73, 75
Aluminium bearings, 68, 69
Articulated bearing, 70
Aspiration, 20
Atomisation, 88
Austenitic steel, 40
Automatic valve, \03
Balanced rotor, 37
Balance of power, 27
Balancer, 82
Barred speed, 126
Barring motor, 73
Bearing materials, 68, 69
Big end bearing, 69
Black smoke, 124
Blade vibration, 34
Blended fuel, 89
Blocking device, 73
Blowdown, 20, 31
Blowpast, 26
Booster pump, 86
Borate treatment, 101
Bore cooling, 44
Bottom end, 69
Bursting disc, 104, 110, 122
Camshaft, 81
Camshaft drive, 82
Catalytic fines, 49, 85
Cavitation, 85, 96, 101
CCAI value, 89
Central cooling, 39, 101
Centrifuge, 85, 99
Cetane number, 89
Charge air cooler, 37
Chocking of bedplates, 79
Choked exhaust, 27
Choked injector, 26
Chromium plating, 43, 65
CII value, 89
Clarifier, 85
Cleaning turbocharger, 37
Clock gauge, 73
Cocktail shaker, 61
Column, 75, 79
Combustion, 88
Combustion chamber, 59
Combustion monitoring, 28
Common rail system, 9
Compressed air, 42, 104
Compression diagram, 22
Compression ignition. 89
Connecting rod, 66
Constant pressure charge, 34
Continuous grain process, 70
Control, III
Cooling systems, 16, 101
Corrosion in cylinder, 48
Corrosion inhibitors, 101
Coupling, II
Crankcase, 78, 117
Crankcase explosion, 117
Crankcase lubrication, 99
Crankcase noise, 125
Crankcase relief valve, 119
Crankshaft, 70
Crankshaft deflection, 73
Crank web, 70
Critical speed, 116, 126
Cross head, 68
Crosshead bearing, 68
Cross scavenge, 31
Cycles, 19
Cylinder cover, 53
Cylinder jacket, 45
liner, 43
lubrication, 45
safety valve, 123
wear, 48
Damping, 114
Deflections, 73
Deglazing, 51
Degradation of oil, 100
Density of fuel, 85
Detonation, 89
Dewpoint, 39, 48
Diaphragm, 54
Diesel Indell, 89
Diesel knock, 89
Direct drive, I
Draw card, 23
Duplell chain, 84
Early ignition, 24
Electronic governor, 116
Emulsification, 100
Emulsion oil, 101
Engine indicator, 22
Engine seating, 78
Engine stroke, 19
EpollY resin chocks, 79
Ellcess air, 31
Ellhaust ports, 10,31
Ellhaust temperature, 40
Ellhaust valve, 40
Ellplosion in air start, 122
Ellplosion in crankcase, 117
Fire in scavenge, 120
Fir tree root, 34
Flame arrester, 104
Flash point, 85, 117
Fluid-film lubrication, 69
Flywheel, 73
Fork and blade, 70
Four-stroke cycle, 20
Frameboll, 78
Fresh water cooling, 101
Fretting, 56, 69, 78, 79
Friction in cylinder, 48
Fuel contamination, 85
Fuel ignition, 89
Fuel injector, 90
Fuel pump, 94
Fuel system, 86
Fuel treatment, 85
Fuel valve cooling, 93, 103
Fusible plug, 105, 107
Gauging a liner, 49
Gear drive, I
Governor, 113
Gudgeon pin, 69
Guides, 68
Helill pump control, 94
Holding down bolts, 80
Hydraulic ellhaust valve, 41
Hydraulic governor, 114
Ignition delay, 89
Ignition jump, 28
Ignition lag, 24, 89
Ignition quality, 89
Inconel, 40
Indicator, 22
Indicator diagram, 21
Inhibitors, 101
Injector, 90
Inlet valve, 42
Intercooler, 107, 109
Isochronous governor, 113
Isothermal compression, 108
Jacket cooling, 43, 101
Jerk pump, 94
Kinetic energy, 73
Kinetic viscosity, 87
Late ignition, 25
Leaking ellhaust valve, 125
Leaking injector, 25, 92
Leaking start valve, 104, 122
Light spring diagram, 23
Loop scavenge, 31
Lost motion, 113
Low compression, 26
Low inertia ring, 66
Low speed running, 90
Lubricating oils, 100
Lubrication system, 98
Main bearing, 72
Matched turbochargers, 33
Mean indicated pressure, 22
Mechanical-hydraulic governor, 114
Medium speed engines, II
Microbial degradation of oil, 100
Micro organisms, 100, 101
Mist detector, 118
Moment balancer, 82
Multi-stage compression, 108
Natural frequency, 126
Needle valve, 90
Nimonic alloy, 40
Nitrite-borate, 101
Oil control ring, 65
Oil cooled piston, 57
Oil mist detector, 118
Opposed piston engine, 9
Oscilloscope, 28
Out of phase diagram, 23
Overlap, 21, 30
Overspeed, 116
Peak pressure, 28
Penetration, 88
Pilot injection, 90
Piston cooling, 56
Piston rings, 63
Piston rod, 61
Piston rod gland, 55
Piston skirt, 61
Plasma coating, 65
Post scavenge, 19
Power balance, 27
Power card, 22
Pressure charging, 32
Pul~e charging, 33
Punfier, 85, 99
Push rod, 40
Relief valve, 101, 110, 119, 122
Reservoir, 104
Resin chocks, 79
Reson~nce, 126
Reversmg, III
Roller chain, 83
Rotating piston, 15
Rotocap, 40
SAE numbers 100
Safety valve, i23
Saponification 100
Scavenge, 30 '
Scavenge fire, 120
Scraper ring, 65
Scuffing, 49
Seal rings, 52
Secondary burning, 25, 92
Semi built-up shaft, 70
Settling tank, 85
Shaker, 61
Shell bearing, 68, 69, 72
Shrink fit, 70
Side by side bearings, 70
Slow turning, 104
Sludge, 85
Sodium, 40
Speed droop, 115
Spherical bearing, 15
Split cooling, 101
Split crosshead, 4, 69
Starting air explosion, 122
St8fting air receiver, 104
Starting air system, 103
Stellite, 40
Stratification, 85
Sulphur, 48, 100
Supercharge, 32
Surge, 121, 125, 126
Tappet 41
TON n~mber 46 48 100
Teflon 55 ' , ,
Temp~rary droop 116
Thermostat, 87 '
Thinwall bearing, 68, 69
Through scavenge, 31
Thrust bearing 72
Tie bolts 78 '
Top end: 68
Torsional vibration, 126
Trunk piston, 61
Turbocharger, 34
Turbocharge system 32
Turbulence 89
Turning ge~r, 73
Twin injection, 90
Two stage turbocharging 39
Two-stroke cycle 19
Uncooled fuel valve, 94
Uncooled turbocharger, 36
Underco.oling, 39
Under-pIston charge, 9
Uniflow scavenge, 31
Valve cooling, 93, 103
Valve-timed pump, 98
Vanadium, 40
Variable ignition timing, 90
Variable speed governor, 113
Vee engines, II
Vibration, 101, 126
Viscosity, 87, 88
Viscosity regulator, 87
cooled piston, 58
cooling system, 101
hammer, 103
treatment, 101
washing, 37, 126
spring diagram, 23
Weardown, 72
Wear in cylinder, 48
Wear on rings, 64
Welded crankshaft, 72
Whitemetal bearings, 68, 69, 117
White mist, 117
Planning capacity of generating
tool shops, for example,
would be expected to have considerable
This is because of setting-up and taking-down operations, and the infrequent use of certain machines.
It should never be assumed that driving motors
will operate at their full-load (nameplate)
Machinery designers tend to use standard frame size
motors in their equipments. They inevitably choose
the next largest motor to meet their drive requirements. It would not be unreasonable
to assume that
every drive in an installation will be 'over-motored'
in this way.
5.3 Planning capacity of generating
Generating capacity must be sufficient to meet peak
power demand, even if the peak only occurs for a
few hours a year. There must be some in-built
margin for contingencies
such as plant failures and
abnormally high loads. Due allowance must also be
made for routine maintenance.
Future load expansion should not be ignored. In
certain applications
there may well be a regular
annual rise in energy requirements.
An increase of
7 % per annum equates, when compounded,
to an
doubling of load demand every 10
years (EIO = Eo (1 + 0.07)10 = 1.96 Eo). This sort
of growth rate would not be unreasonable
for small
communities in developing areas.
The timing of power plant additions must be
carefully planned and expenditure on extra capacity
should be deferred until the need arises. Designs
must be flexible enough to allow for planned expansion with the minimum of disruption
to existing
plant. It is usual to provide, at the outset, a 10-20 %
margin of capacity over and above that required by
the annual peak demand. Where a number of small
power stations feed individual communities,
may pool their generating resources (using tielines)
to share power reserves between them. Tieline costs
may be retrieved by levying direct (power-demand)
charges, and the capital cost of any new plant added
to the pool may be shared [7].
Where supplies have to be maintained after the
loss of a generating unit, load shedding is inevitable
unless some reserve capacity is available. This problem is not unique to the isolated power station. It
applies wherever only part of a load demand can be
met from available generating
plant. One must,
therefore, always consider the security or reliability
of any proposed system.
Security of supply is defined by the ratio:
and is normally expressed on an annual basis. Public
Utility supplies usually have excellent reliability. In
the United Kingdom, for example, the National
Grid has a reliability of better than 98 %. This
means that consumers experience annual outages of
less than 2 hours. If a private generating plant is to
have reliability approaching this level, it must have a
greater number of generators than that apparently
required to meet the anticipated maximum demand.
It would be quite unrealistic to expect full security of
supply from a base-load system (or a peak-lopping
system) which has only one generating unit.
All plant needs regular overhauling and maintenance. When this is being undertaken the remainder
of the installation
is at risk. More so where the
annual load pattern is 'flat' and shows no marked
in demand.
A well-designed
generator plant is unlikely to experience total failure
- since the units should fail individually.
durations will depend upon the nature of failures,
the availability of spares, and the skill of repair staff.
Evidence, from such sources as the Annual Working
Cost and Operational Reports of the Institution of
Diesel and Gas Turbine Engineers, indicates that it
would be rash to expect any type of generating plant
to have an availability greater than 98 %.
5.3.1 Multi-generator
power plants
West [8] suggests that, in a multi-set installation of
similar units, 'reliability may be calculated according
to the laws of chance, bearing in mind that while one
set is being maintained, the likelihood of outage for
the remaining installation increases'. This likelihood
may be expressed as the number of hours of anticipated outage in a year. Table 5.1 permits determination (for any assumption
hours) of the number of hours in a year when the
number of sets available is at least one less than the
number required to meet the maximum demand.
The hours of risk will reduce if the load pattern
contains periods when load is at a sustained lower
The following example [8] shows how Table 5.1 may
be used to determine the annual hours of risk for a
6-unit installation (where 4 are required to meet the
maximum load demand).
It is assumed that each
unit requires 1000 hours of maintenance
per year,
therefore 5 sets only are available for 6000 hours
each year, and all 6 sets are available for 2760 hours
each year.
From Table 5.1:
for N = 5 and r = 4: the hours of risk, per year, are
for N = 6 and r = 4: the hours of risk, per year, are
Load considerations
Table 5.1 Non-availability
of plant due to breakdown
(Average number of hours each year when the number of
sets available is less than the number required to meet the
maximum demand)
* = less than 1 hour
Therefore, total hours of risk is
(6000/8760) x 33 + (2760/8760) x 2
= 23 hours per year
Using Table 5.1 and this method of calculation one
can assess the effect of various maintenance outages.
Table 5.2 shows the average hours of risk for an
installation in which 4 generators are required to
meet the maximum demand. It considers the effect
of maintenance periods varying from 100 to 1000
hours per annum and shows that, for all reasonable
assumptions of maintenance outage, the minimum
number of sets required for adequate reliability is 6.
Table 5.2 The effect of varying
annual maintenance
Diesel power stations playa very important role
in the development of electricity supplies in rural
areas. Indeed, they are often used to build up local
demand in villages and small townships. They may
start with just a single generator. Expansion to meet
agricultural and industrial load growth is possible
with very modest capital outlay. Load growth may
eventually justify the construction of an interconnecting transmission grid - between regional generating centres and expanding load centres. If larger
thermal or hydro power stations come to be installed
at a much later stage, the diesel stations may be used
for peak power or standby duty.
Figure 5.6 shows a typical layout for a small
rudimentary power station consisting of selfcontained, canopied diesel-generator
sets (2)
housed in a simple shed with a metal roof and fenced
sides (5). Two of the three sets are on base-load duty
and the third is on standby. The capacity of the fuel
oil storage tanks (1) is governed not only by the
specific fuel consumption of the sets but also by
economic factors relating to fuel deliveries and bulk.
purchase incentives. Item (3) is a low voltage distribution board, and the pole-mounted step-up transformers (4) are for two outgoing high-voltage lines
(typically 11kV).
Perry [9] gives an interesting description of the
large-scale use of diesel generating plant in isolated
areas. The total capacity of the diesel generator
plant operated by the State Energy Commission in
Western Australia had reached 170MW, by 1985. In
certain areas, the annual load growth since 1976 had
been between 10 % and 33 %.
Among the many points taken into consideration,
when selecting sites for power stations, were:
• the ability to distribute power from a station
through at least four outgoing feeders;
• the ability to initiate fuel-saving measures by
utilizing engine waste to generate process steam
or to be used in lithium-bromide chillers for air
conditioning; and
• a site's potential for development over a period of
not less than 25 years.
Stations are planned to form an integral part of the
business and commercial centres of the future development towns of the north-west territories, the
intention being to encourage total energy. schemes
from the outset. Design philosophy has been to
Figure 5.6 Example of the layout of a small diesel power
station (Courtesy: ABB - Asea Brown Boveri Ltd)
'modularize equipment as much as possible to
minimize on-site installation costs and to allow units,
where required, to be removed and reinstalled in
another location with a minimum loss of installation
equipment and material'.
An eventual total of nine generating sets is envisaged for each station. Unit sizes are initially based
on the need to supply off-peak load with no fewer
than two sets. The nine units eventually installed
will comprise:
• four of the original selected size; followed by
• five, whose individual capacity is approximately
twice that of the original size.
For example: 4 x 1 MW and 5 x 2MW.
A 'safe generating capacity' (SGC) caters for
system demand.
The SGC = (the installed capacity of the station) (the capacity of the largest machine) - (a further
margin of 15% of the remaining installed generating
The latter margin allows for site derating due to high
ambient temperatures and low atmospheric pressures. A typical 14 MW station would have an SGC
2. when more than two units are in service, the load
that each carries is restricted to 85 % of its
capacity at normal ambient conditions;
3. when three units of equal size are loaded to more
than 85 % of their nameplate rating, a fourth
machine, of twice the rating of the running
machines, is introduced. The load carried by the
incoming machine should not then exceed 40 %
of the total system load.
The spinning reserve in Table 5.3 is defined as the
capacity of the station to pick up the immediate loss
of the largest unit in service. It is calculated as:
The spinning reserve has a twofold purpose:
1. to cover such incidents as significant generating
loss; and
2. to ensure that any subsequent fall in system
frequency is sufficiently slow as to allow the
automatic load-shedding
controls time to
operate, and thus minimize disturbance to the
station and to the loads it serves.
(14) - (2) - (12 x 0.15)
5.3.2 Programming expansion of power plants
Table 5.3 describes the Energy Commission's
operating policy for the stations. The objectives are
to ensure that:
We have mentioned the prudence of planning for
future expansion when designing a power station.
Expansion may take the form of:
1. a minimum of two units are in service at all
times - to guarantee integrity of supply;
• a larger generator to replace the original, in a
single-set standby installation; or
Load considerations
Table 5.3 Operating
Machines in service
x rating (MW)
4 x
4 x
4 x
policy for diesel stations
System load
in Western Australia
Spinning reserve
% load on units
System load on
largest set (%)
• additional units, to cater for anticipated load
demand increases, in multiple generator stations.
(The timing of these additions is important, as
one would not wish to commit capital resources
too early in any programme.)
We have seen that most loads have repetitive chronological patterns. These patterns must be carefully
analysed, in order to predict future trends. The
collection of historic load data must, therefore,
always be the starting point.
Many techniques are used in load prediction.
(The lEE publication [10] Chapter 4 will give the
reader a good introduction to the subject.) Dale [7]
describes a dynamic-programming technique for the
economic selection and timing of generator-plant
additions. He uses a minimum-percentage-reserve
concept to determine the date at which generating
plant should be added to a power system. Although
he deals with large systems, the principles apply to
any scale of operation. The diagrams that follow
have been adapted to show load and capacity ordinates of reduced scale.
Figure 5.7 shows the peak-load trend forecast for
a utility system (curve 1). It is obtained by extrapolating the curve through the actual annual peaks for
the preceding few years. The system is assumed to
have had 8 generator units, initially (at year 0):
• two of 2.5 MW capacity;
• two of 2.0 MW capacity; and
• four of 1.25 MW capacity.
Figure 5.7 Predicted load growth, and related future
requirements [7]
The annual installed-capacity curve (shown dotted
in Figure 5.7) is derived by adding the required
percentage reserve (20 %), to the peak-load trend
Figure 5.8 shows two alternative ways of adding
generating capacity. In each case, the installed
capacity is never less than the required level established in Figure 5.7. Plan 1 proposes large units
installed at three-yearly intervals, whereas Plan 2
advocates a smaller-sized unit being added each
year. Relative capital and running costs will determine which is the more economically attractive
Load elements
of many thousand possible paths may be determined, by just one run through the program.
5.4 Load elements affecting plant size
and performance
The type and rating of any generating plant must
depend upon the nature and size of the load it is
required to serve. When considering the characteristics of a particular load, it is necessary to identify:
Figure 5.8
Alternative generator-expansion plans [71
If choice of unit size was confined to these two
options (very unlikely, in practice), the number of
ways in which they could be combined assumes very
large proportions - as shown in Figure 5.9. L represents a large-unit addition sufficient for two years;
and S a small-unit addition sufficient for one year.
Very probably, one of the many paths shown in
Figure 5.9 will be more economical than either Plan
1 or Plan 2. Dale goes on to explain that 'previous
methods of trying to find the optimum generationexpansion plan have been based, almost without
exception, on procedures or computer programs
that evaluate the total cost of following one particular path, or plan, over a long-range study. By
running through this program for each of, say, 10 or
20 different plans, the so-called optimum plan can
be found. Using the procedure he describes, the best
Figure 5.9 Some possible ways of increasing installed
capacity, with two available unit sizes [71
1. those elements which demand close-tolerance
power parameters (e.g. technical loads such as
computer and telecommunication systems); and
2. those elements likely to impinge upon these
demands by adversely affecting generator steadystate and/or transient performance. These might
• step-change loads (induction motor starting);
• non-linear loads (rectifiers and switched-mode
power supplies - SMPS);
• cyclically varying loads; and
• regenerative loads.
We shall now examine the characteristics and effects
of some of these load elements.
5.4.1 Motor loads
Perhaps the most common problem associated with
the correct sizing of synchronous generators is that
related to the starting of induction motors.
Industrial installations inevitably contain numbers
of induction motors. These are often of sufficiently
large size as to dictate the capacity of the generating
plant required to meet the 'impact' loads imposed
during their starting. Nowhere is the problem more
acute than in marine electrical systems, where loads
are essentially concentrated, and contain a high
proportion of induction motors; and where generators are operated near to full capacity. Also, the
operational and economic advantages offered by
3-phase systems are exploited to the full, by the
almost universal use of the more robust direct-online started squirrel-cage motor. Applications include: pump, fan, and separator drives; deck auxiliaries such as cargo winches and windlasses; and oil
transfer pumps on tankers.
The problem is that of limiting voltage variations
to a minimum while coping with the transient currents and torques imposed by the ship's system.
Careful consideration must be given to the resulting
reactive and active power demands, and to equating
these with the required generator reactance and
prime mover output, respectively. Marine classification societies also require that, with one generator
out of action, the remaining plant shall have sufficient reserve capacity to permit starting of the
largest motor in the ship. This must be achieved
Load considerations
without causing any motor to stall or any devices,
such as contactors, to fail due to excessive voltage
drop in the system.
Motor characteristics
The equivalent circuit of a single-squirrel-cage induction motor is shown, in its simplest form, in
Figure 5.10. Here, RI and XL are the combined
rotor and stator resistance and leakage reactance
values, respectively, expressed in stator terms. XM is
the magnetizing reactance across the supply voltage
V. The 'added' resistance, given by the term Rz
((1 - 5)/5), is zero when 5 = 1, and infinity when
5 = O. The internal mechanical power per stator
phase is equal to the power absorbed by this resistance [11].
An induction motor must run at a speed less than
synchronous speed in order to produce a driving
torque. 5 is a ratio known as the fractional slip or,
more usually, as the slip. It determines the amount
by which the rotor 'slips backwards' relative to the
rotating stator field to give the induced rotor e.m.f.,
and hence the current, required to produce driving
At rotor standstill, 5 = 1; and at synchronous
speed, 5 = O. At no-load when the motor is running
freely the slip is very small and usually of the order
of 0.5 % (0.005). When the rotor shaft is connected
to a mechanical load the torque required to drive the
load necessitates an increase in rotor current above
the no-load condition. This can only be achieved by
an increased rate of cutting of the rotating (stator)
field by the rotor conductors - to give a greater
induced e.m.f. and, hence, more rotor current.
There is, therefore, a speed corresponding to each
value of load torque at which the rotor current is just
sufficient to produce the required torque. Rotor
speed falls (the slip increases) as load torque
increases. One expects the slip to be between 0.03
and 0.05 at normal load. The magnetizing current
through the parallel branch XM may be of the order
25 % to 50 % of the full-load current.
Where a motor obtains its supply from a constant
voltage source, such as a very 'rigid' public utility,
the relationship between electrical input and torque
output is as shown in the performance curves of
Figure 5.11. These curves have the following features:
The current, and therefore the kV A demand
(since a constant voltage supply has been assumed),
has a maximum value at standstill. It falls almost
linearly, and gradually, until the motor reaches
about half speed. Thereafter, it drops more rapidly
as synchronous speed is approached.
The value of the 'added' resistance, R' z [1 - 5)/S]
in the equivalent circuit of Figure 5.10, is zero at
motor standstill. The secondary circuit is closed and
the machine is equivalent to a transformer whose
secondary winding is short-circuited. A very large
current is therefore demanded from the power
source during starting. Whilst this is, effectively, a
short-circuit current (hence the term locked-rotor
current), it is not as high as it would be in a
short-circuited transformer. This is because the
leakage reactance of the motor is so much greater
than that of an equivalent transformer. Nevertheless, starting currents of the order of 5 to 6 times
full-load value are not unusual for most 3-phase
induction motors. Actual ratios of starting-to-rated
current will vary with motor size and design, and
may be as low as 4, and as high as 10. The magnitude
of the starting current is independent of applied load
torque and is the same value whether the motor is
started with or without load. As rotor speed
load elements
increases, the secondary induced e.mJ. decreases.
This results in less current demand until, at synchronous speed, there is no e.m.f. and no current drawn.
The starting power factor is very low (of the order
of 0.3 to 0.4) and increases during the run-up
period, to reach its peak just before nominal operating speed. Thereafter, it drops very rapidly to zero
at synchronous speed. Full-load power factor is of
the order of 0.8 to 0.9 and, of course, is always
lagging. When the load on a motor is reduced its
power factor falls.
The active power requirement during run-up is
the product of kVA (which follows the current
curve) and the power factor. The rising power factor
during the run-up means that the peak power occurs
at approximately to 80 % speed and is of the order of
3 to 3.5 times the nameplate rating of the motor.
When the motor is supplied from a generator, this
instantaneous peak power has to be provided by the
generator's prime mover.
The motor torque characteristic is generally of the
same shape as the power curve but its peak value,
known as the pull-out torque or stalling torque, is
slightly displaced from it. At higher values of slip (5
approaching 1), the torque is seen to decrease. This
is because the rotor power factor decreases at a
greater rate than that at which the rotor current
The dynamics of starting
The relatively large currents during the starting
cycle place heavy mechanical and thermal stresses
on windings - conditions which are tolerable for
limited periods only. It is important, therefore, that
the motor is driven up to its normal operating speed
as quickly as possible, consistent with the requirements of its driven machinery. Run-up time is
significantly influenced by the inertia and the dynamic characteristics of the driven load. The torque/
speed characteristic of the driving motor must be
matched to the requirements of the driven load.
Among the features of particular interest to the
machinery designer are the starting, and maximum,
torque capability of the motor. During the run-up
period the difference in torque between that produced by the motor and that required to turn the
load is available to accelerate the combined rotating
masses of motor and load (see Figure 5.12). This
accelerating torque may be expressed as a function of
the speed:
where J is the combined moment of inertia of the
motor and the load coupled to it and Wo is the motor
speed, in mechanical radians per second. Also
Figure 5.12 A graphical procedure for determining run-up
speed as a function of time
where N is the speed, in revolutions per minute
Load considerations
Run-up times are largely governed by the combined inertia of motor and load, and by the operating speed. They need to be determined for each
specific application - using, perhaps, the graphical
integration procedure described above. Start times
on loaded machines rarely exceed 5 seconds and are
likely to be within the range of 0.5 to 2.5 s, for
motors up to 100 kW.
On certain machinery, it is possible to employ
control devices, such as unloading or bypass valves,
to reduce the starting torque required. For example,
it is usual, on marine tankers, to start cargo oil
pumps with their valves closed; the valves are only
opened when the pumps have run up to full speed.
It is possible to have transient starting torques
which may persist for several periods of the power
supply frequency. They are affected by:
• the instant at which the first phase closes, and the
delay, if any, in closure of the other two phases;
and by
• the residual flux from a previous duty cycle [13].
Pulsations may be superimposed on the steady-state
torque/speed characteristic during fast-run-up conditions. The torque reversals that may be produced
form 'loops' of the type shown in Figure 5.13. These
reduce the accelerating torque available from the
motor and prolong the machinery run-up time.
Their effects may not be too disturbing on slow-start
applications involving high-inertia machinery as
they may die away before the machine has gathered
Voltage reduction effects
Referring to the equivalent circuit of Figure 5.10, it
can be shown that:
1. at a given slip, and neglecting magnetizing current, motor current varies directly as voltage; and
2. input power and motor torque both vary as the
square of the voltage.
A fall in voltage to 50 % of nominal would thus
mean that the available torque is reduced to 25 % of
nominal. Figure 5.14 [14] shows a typical torque/
speed curve for a pump (curve a), superimposed on
the torque/speed curves of alternative designs of
induction motor, at various values of reduced voltage. If the pump were coupled to the first motor
(Figure 5.14(a», it would not accelerate to full
speed unless the supply voltage is restored to a level
above 90 % of nominal. Similarly, the lowest permissible voltage level for a successful start with the
second motor (Figure 5.14(b», would be approximately 70 % of nominal.
Input power to a motor varies as the square of the
voltage. If the supply for this second motor was
obtained from a generator of comparable rating,
and the voltage depression on starting was 30 %, the
active power required at the instant of start is 49 %
of that which would have been required had system
Load elements
voltage been 100 %. Using the efficiency of the
generator, the momentary active power required
from the prime mover may then be determined. As
a corollary to this, the less the allowable voltage
depression on starting, the more the momentary
power that is required from the prime mover.
An induction motor may behave in one of several
ways during a voltage depression, and behaviour is
directly related to the driven-load requirements
1. it may stall completely and fail to restart because
the voltage cannot recover as a result of the
current drawn by the motor, and by others
connected to the same point in the system;
2. it may stall momentarily before recovering to a
reduced speed level as the voltage is restored to a
value lower than normal;
3. it may stall or lose speed and then run up to
normal speed.
If the selected motor is designed to respond as in 3,
to anticipated voltage depressions, it is more than
likely to behave as in 1 or 2 under abnormal
For any given motor and load characteristics,
there is a minimum voltage at which the motor can
run up to speed from a particular slip. Figure 5.15
shows this voltage for a motor operating against two
load characteristics [15]. Motor reswitching characteristics are discussed later in this section.
We have already established (in Sub-section 4.3.2
of Chapter 4) that generating sets represent a highimpedance power supply source and that, when
relatively large wattless loads are switched on to
them, instantaneous voltage dips will occur. A
Figure 5.15 Minimum voltage required for re-acceleration
of a motor against (a) linear load and (b) constant load
torque characteristics
typical voltage response characteristic was shown in
Figure 4.38 and its features discussed in the accompanying text. Figure 5.16 [16 and 17] shows the
current and speed characteristics for a typical motor
switched to a constant voltage, zero impedance
supply source such as the 'infinite busbar' of a public
utility (solid line curves). The dotted curves show
the effect of switching the same motor to a higherimpedance generator, whose voltage recovery response is similar to that of Figure 4.38. The run-up
period is approximately extended by the response
time of the generator's excitation control system.
The faster this response is, the less significant are
the effects of the voltage dip on the motor's starting
Transient voltage disturbances can be a source of
particular trouble in two areas: lighting and contactor switchgear. The light output of lamps depend,
very much, upon the supply voltage. Filament lamps
are inherently more sensitive than fluorescent lamps
to fluctuations of voltage at frequencies causing
flicker. It is therefore common practice to install
fluorescent lamps in ship's accommodation spaces
because they will accept two to three times as much
voltage fluctuation as filament lamps, for the same
visible flicker.
High-pressure mercury vapour and high-pressure
sodium lamps are installed on deck and in areas
where a higher level of flicker may be tolerated.
Both types are more sensitive to rapid voltage
fluctuations than filament lamps. One solution is to
employ a combination of fluorescent and vapour
lamps. Ordinary sodium lamps are insensitive to
flicker but their use will be more often limited by
their colour characteristics [18].
BS 5424: Part 1 (IEC 158 - 1) requires contactors
to close satisfactorily at 85 %, and drop-out between
75 % and 10 % of rated voltage. Some thought must
therefore be given to the supplies for contactor coils
in systems where voltage dips greater than 25 % are
possible. Unless suitable measures are taken, there
is a risk that contacts will chatter, arc and, finally,
Figure 5.16 Effect of voltage depression
on motor run-up
Load considerations
weld-on. One possible solution is to isolate the coils
of contactors in essential circuits through small
constant voltage transformers. Contactor coils operated from a d.c source are another possibility.
More severe difficulties are likely to be experienced
when large voltage dips coincide with contactor
closing operations. The closing force may be
reduced to such an extent that it is unable to
overcome contact spring pressures.
Estimating the voltage dip
The magnitude of the voltage dip at a generator's
terminals, following load switching, is a direct function of the sub transient and transient reactances of
the machine.
The expression only gives an approximation because
most modern excitation control systems are sufficiently fast-acting to lop-off the bottom of the
transient voltage 'trough' (point B in Figure 4.38)
Generator manufacturers supply performance
data (usually in the form of curves supported by
application notes) enabling determination of voltage
dips for given impact loads. Examples of such data
are given in Figures 5.17 and 5.18.
Figure 5.17 shows typical 3-phase voltage dip
characteristics, based on results using ultraviolet
recorder measurements. The performance curves
are for a specific frame size. The terms 'minimum'
and 'maximum' voltage refer to the bottom and top
ends of the standard range of voltages available
from the particular winding applied to the machine.
Typically, these may be 346 volts and 480 volts,
respectively. The 'maximum rated kVA of the frame' is its standard (industrial) continuous maximum
rating. In all instances, the impact current must
never exceed the machine's declared overload capability. The overload ratings may vary between those
equivalent to 1.5 pu and 4.5 pu full-load impedance
at zero power factor for periods of up to 15 seconds.
BS 2949 (the specification for rotating electrical
machines for use in ships) requires generators to
accept 50 % excess current for 15s, at the approximate rated voltage. Marine Classification Societies' rules generally call for a 50 % overload, at 0.6
to 0.8 pf, for periods up to 2 minutes. This is one of
the reasons why industrial machines are derated for
marine use. Because of this derating the per-unit
overload capability of marine generators may appear to be better than their industrial equivalents.
The actual overload performance (in terms of current or kV A) is the same in both cases.
Figure 5.18 illustrates yet another approach to the
presentation of impact performance data. The curves relate to the full range of 4-pole low voltage
machines available from the manufacturer. t:.S (in
kVA) is the impact load, assuming connection to an
infinite bus bar system. Stype is the maximum rating
(kVA) listed in the manufacturer's technical data for
the generator type and frame size. Figure 5.18(a)
gives the voltage dip (t:.V) related to per-unit
applied load, at 0.8 power factor lagging, for socalled 'short' and 'long' generators. The manufacturer uses the ratio:
Load elements
Figure 5.18
Manufacturer's typical curves for determining
voltage dip on a range of 3-phase, 4-pole, low voltage
generators (Courtesy: A van Kaick Neu-Isenburg
Voltage dip is largely independent of the load
already carried by a generator, particularly if this is
a mixed passive load, e.g. lighting and heating. But
any motors running on the system at the time will
experience a speed change, which will cause them to
draw more current. This increased load current,
when added to the starting current of the incoming
motor(s), causes the voltage dip to exceed its expected value. An allowance may therefore have to
be made for any dynamic preloading when estimating the voltage dip - especially in borderline cases.
A dynamic (motor) preload of 50 % of the generator
rating will increase the voltage dip by some 10 %
A few words of caution on methods of measuring
voltage dip are apposite at this stage. Perhaps the
most versatile instrument for monitoring and displaying time-varying phenomena is the cathode-ray
oscilloscope. Available in either analogue or digital
reading forms, it offers the possibility of storing fast
transient data for retrieval at a later date. This data
may then be permanently recorded on either slowspeed pen or galvanometer-type X-Y recorders.
Utraviolet (uv) recorders may also be employed to
provide time-domain records, but their response to
step functions is largely governed by the damping of
the galvanometers they use. With underdamping,
the response is fast but contains an overshoot and a
decaying oscillation. With overdamping, the time to
reach a final deflection is excessive. A damping
factor of 0.64 is usually suitable for a.c. signals [20].
Pen recorders are, generally, not fast enough for
Load considerations
directly recording the magnitude of short-time transients. Similarly, most digital voltmeters are unsuitable for measuring transient conditions. Much depends upon the measurement systems they employ;
and the intervals at which signals are sampled.
Remember that one could be dealing with generator
voltage response times as low as 80 ms (and certainly
below 1 s), with most modern machines. The use of
conventional switchboard-type moving-iron, induction, and moving-coil rectifier instruments should
never be entertained. It is surprising how often one
is still asked to accept their evidence, on site performance tests. When faced with no other alternative,
apply a 'correction factor' of between three and
three and a half times the observed voltmeter reading.
Limiting voltage dip
Several lines of approach are possible:
1. Where the installation contains a number of
motors (and particularly where one or two relatively large motors constitute the major part of the
load), it may be feasible to rearrange starting
sequences so as to minimize impact currents on
the generators. The motors with the largest starting currents should be run-up first. See the case
study, Appendix 5.1.
2. Use a generator of lower transient reactance.
One way of achieving this is by using a larger
frame size machine. Alternatively, order a low
reactance machine, specifically designed for the
starting duties proposed. A machine of this type
may suffer from the disadvantage of having a
lower than normal continuous rating for its frame
size. The economics of either solution must be
compared with the additional costs of (3) below.
3. Use one of the reduced-voltage starting systems
described later in this chapter. A careful assessment of each application will be necessary. While
starting current may be reduced, the starting
torque is, unfortunately, also reduced in the
same proportion. This could be a serious limitation with heavy-start, high inertia loads.
4. Replace some, if not all, of the squirrel cage
motors in the system with those of the slip-ring
type, which can provide large starting torques,
without drawing excessive starting current. This
is not quite the panacea it may seem, at first
sight. First, there is the economic penalty of
added cost. Then there is the possibility that the
prime mover rating may have to be increased to
cater for the higher active power demand resulting from the higher starting power factors (see
the sub-section on prime mover sizing later in this
Motor types
Induction motors are subdivided into two groups:
cage motors, and wound-rotor (or slip-ring) motors.
1. Cage motors are highly robust, reliable machines
offering simplicity of operation in their single
speed direct-on-line started forms. Rotor construction varies from the diecast aluminium singlecage arrangement used on small motors, through
the various forms of deep-bar rotor used on
medium and large machines, to double-cage versions (and even 3-slot arrangements) employing
bars of copper, cuprous alloys, or nickel alloys.
2. Slip-ring motors were the forerunners of cage
machines. As their description implies, they have
rotor windings which are similar to the stator (or
primary) winding. Access to the rotor windings,
through shaft-mounted slip rings, means that
starting performance can be controlled by external resistance. The external resistance permits
repeated starting against high inertia loads, besides giving a limited amount of speed control.
Any losses and heat dissipation during starting
are confined to the external resistance.
Conventional methods of starting
Discussion is limited to 3-phase machines. Motor
protection devices are omitted from the diagrams
that follow, for convenience of presentation.
Starters may be classified in one of two categories:
1. those in which full line voltage is applied to the
motor windings - known as full-voltage starters
2. reduced voltage starters, in which the voltage
applied to the windings is some fraction of the
line voltage. This gives a corresponding reduction in starting current, but at the expense of
greatly reduced starting torque (which is proportional to the square of the voltage applied).
Full voltage starters
Direct-an-line (d.a.l.) starting. This is the simplest method of starting a cage induction motor. A
contactor is used to switch full line voltage directly
to the motor terminals (see Figure 5.19(a)). The
motor stator windings are normally connected in
delta. It is of course possible to start slip-ring motors
in this way. The method is seldom, if ever, used
since no advantage is gained over the cheaper cage
machine. Starting torques are of the order of 0.75 to
1.0 times full-load torque and inrush currents are
between 4 and 6 times full-load current. Typical
values of the starting kVA/rated kW ratio are given
in Table 5.4. Direct-on-line starting is customarily
applied to small and medium power drives, but it is
not unusual to have LV motors of up to 350 kW
rating started in this way.
Load elements
Solid-state starting. This is primarily employed
on loads with a rising torque/speed characteristic.
These starters give a 'soft' and stepless start by
controlling stator voltage through anti-parallel thyristors fitted in each supply line, and phasecontrolled by appropriate triggering. An alternative
method is to connect the thyristors across the
secondary of a gapped-core inductor whose primary
is in the supply line. This gives variable-impedance
contro!. Voltage reductions of between 25 % and
70 % rated value, with corresponding torques of
between 6 % and 50 % of the d.o.!. value, can be
achieved. Extended run-up times may be arranged
to give smooth starting with low inrush currents.
First costs are higher than the electromechanical
d.o.!. starter, but may be justified where very frequent starting is necessary.
Rotor resistance starting. This method is confined to wound rotor machines. Resistors are connected in series with the rotor windings (see Figure
5.19(b)). They are graded in sections, or steps,
whose values follow a geometric progression - similar to the pattern employed on d.c. shunt
motor starters. The motor is started with all steps in
circuit. As it runs up, the external resistance is
reduced, step by step, in a controlled sequence,
until, at normal speed, the rotor winding is shortcircuited at the slip rings. The series resistance limits
the starting current and improves the starting torque. Where a machine is required to start against
full-load torque, the resistance is arranged to give
about 1.5 times full-load torque with about 1.5 times
full-load current. Points to note are: the starting
power factor is high (typically 0.9); peak power
demand usually occurs at switch-on. This will influence the rating of a generator's prime mover.
Alternative forms of wound-rotor starter use liquid
resistors, or star-connected choke coils of low resistance and fairly high inductance, chosen to suit the
given rotor.
Reduced voltage starting
Primary resistance starting. This system is
mainly used for small drives and is much favoured in
American practice. A resistance, usually graded in
two steps and controlled by contactors, is inserted in
each line to the stator terminals of the machine. It is
designed to reduce the stator current by a fraction
(x) of its direct-on-line value, as the motor drives up
to normal speed x approaches unity. The starting
torque is x2 of that obtainable with direct-on-line
Star-delta starting. This is the most widely
applied method of starting cage motors on reduced
voltage (Figure 5.19(c». The system requires that
all six ends of the three phase windings are brought
out to the motor terminals. The starter circuit is
arranged so that the stator windings are first connected in star (where the phase impedance is three
times that in the delta condition) until either a
steady speed is attained or a preset time has expired.
The windings are then connected in delta for motor
running. Motor acceleration should proceed as far
as possible before the transition is made so as to
minimize the transient current peaks during changeover to the delta mode. Starters for small motors
may employ hand-operated changeover switches.
Where contactors are used, some form of interlocking is necessary in order to give the proper starting
Star-delta starting is less effective in reducing the
peak kW demand, particularly if the changeover
occurs at relatively low speed. The power requirement is then of the same order as the maximum that
would occur if the motor were switched direct to the
supply in delta (see Figure 5.11).
An interesting variant of the star-delta system is
the 'closed transition' Wauchope type starter. This
improves the motor's starting characteristic by inserting resistances in series with the windings, when
these are in 'transition' to the delta stage. The
resistances are short-circuited before the motor runs
up to full speed, in delta. They not only reduce the
high peak current demanded by the motor during
the transition from star to delta (see Figure 5.20),
but they also give smoother acceleration [21].
Auto-transformer starting. An auto-transformer,
which has fixed tappings along its windings, is
interposed between the supply and the motor terminals (see Figure 5.19(c)). Usually, no more than
three tappings are necessary - typically, at 50 %,
70 % and 80 %. The lowest voltage tap is used to
feed the motor at standstill. As it speeds up, the tap
position is changed to the next highest level until full
Load elements
Generator and prime mover sizing
The kV A elements of system load must be supplied
by the generator. It should be rated to cope with the
steady-state and transient demands made upon it
without exceeding its designed overload capacity or
the permissible voltage dip for the application. The
active power demand is met by the prime mover,
which should be matched to the continuous maximum rating of the generator, and should be capable
of meeting the peak kW demands of all impacting
loads without excessive and intolerable speed
Assessment of the generating set's capacity must
be based upon load analyses which take account of
the steady-state and transient conditions relating to
every stage of the installation's operations. The case
study in Appendix 5.1 of this chapter shows how a
typical assessment is made. Figure 5.21 illustrates
the stages in such an exercise.
Figure 5.20 Variation of line current during star-delta
line voltage is eventually applied to the motor
terminals. The auto-transformer is then switchedout. The transformer need only have a short-time
rating, but its leakage impedance should be small
enough to avoid undue limitation of the starting
current [13]. Hand-operated air-break switches may
be used. Fully automatic control is achieved by
means of contactors and timing relays. In the Korndorfer variant of this starting system a part of the
transformer winding is arranged to remain temporarily in series with the motor windings at the
changeover to full voltage. Thus, the supply to the
motor is not interrupted in the transition from
reduced to full voltage. The lowest tapped-voltage
provided must be consistent with the starting torque
requirement of the load. If this tap provides a
fraction x of nQrmal voltage, the starting torque and
current are each x2 times that obtainable with direct
switching. For example, 25 % of the direct-switching
values would be obtained for a 50 % tap. Autotransformer starters may be more expensive and
bulky than their star-delta alternatives but they do
provide higher starting torque.
Table 5.5 [17] gives a comparison of typical starting characteristics for various forms of starting.
1. Phasor OA represents the standing load (in
kV A, at 0.95 power factor) on the generator,
when a motor is started.
2. AB is the transient kV A demand. Its magnitude,
and the angle it subtends above the x-axis (determined by the power factor), will vary with the
type of motor and the starting method used.
Different values for AB may apply when sizing
the generator and when sizing the prime mover.
For example, the starting characteristics in Figure
5.11 show that the maximum kV A (5.5 pu)
occurs at starting, with a corresponding power
factor of about 0.3. The maximum power occurs
at about 80 % synchronous speed and corresponds, approximately, with 4 pu kV A at 0.65
power factor. The latter values should be used
for plotting AB when prime mover selection is
being considered. Figure 5.21 is therefore presented in two constructions: (a) for generator
assessment, and (b) for prime mover assessment.
Diagram constructions for generator and prime
mover assessments will be one and the same
when considering slip-ring motor starting
because the peak power demand is likely to be at
starting (see the sub-section detailing autotransformer starting p. 156).
Table 5.5 Comparison of typical starting characteristics (per unit values)
Starting method
A uta-transformer
Slip ring
Starting kY A
Peak kY A
Starting torque
Peak kW
Load considerations
where Pd is the shortfall in engine capacity, or the
power deficit and J is the polar inertia of the
system's rotating masses (combined engine and
6. The projection of OC onto the abscissa (in either
diagram) represents the continuous output required from the engine, once the motor is running at normal speed on designed load.
Motor reswitching transients
Figure 5.21 Assessment of generating plant sizes for
impact loadings
3. OB is the 'vectorial' addition of OA and AB, and
represents the resultant transient load imposed
on the generating unit during the motor run-up
period. Its magnitude (see Figure 5.21(a» must
be within the rated overload capability of the
generator .
4. Phasor AC is the running load of the motor (1 pu
kV A). The angle it subtends above the abscissa
(the x-axis) corresponds to the motor's designed,
full-load power factor.
5. Projection of OB (in Figure 5.21(b» onto the
abscissa gives the transient output required from
the prime mover. This 'momentary' demand
should not exceed the engine's declared overload
capacity - which is usually 10 % of its ISO continuous rating. If the demand marginally exceeds
the overload rating the inertia stored in the
system's rotating masses is generally sufficient to
cover the deficit with only a minimal reduction in
Situations may arise, such as on board ship, where
essential services are provided with a maintained
supply. This supply may be derived from two separate sources, feeding into a sectionalized duplicate busbar system incorporating 'main' and 'reserve' busbars, which may not necessarily be synchronized. Changeover conditions are classified as
an intentional supply interruption. There may,
however, also be quite unintentional interruptions
to a running motor's supply, e.g. physical shock or
excessive vibration, causing a starter's contactor
contacts to open momentarily [22]. Any motor so
disconnected from its supply source immediately
loses its driving torque and begins to decelerate. The
rate of deceleration is inversely proportional to the
combined polar inertia of the motor and its driven
machinery, and to the load torque. The motor
behaves as a generator producing a voltage at its
open-circuited terminals. If it is then reswitched to a
supply source, surge torques and currents will result
- after even the shortest interruption. The magnitude of the peak current is governed by:
• the motor load;
• the supply disconnection period; and
• the difference, in both phase and magnitude.
between the residual stator e.m.f. and the supply
voltage at the instant of supply reconnection. The
greater the phase, and magnitude differences
between the two voltages, the greater the reswitching transient will be.
Shoesmith [22] has reported on a series of motor
reswitching tests conducted at the Admiralty Engineering Laboratory. The tests led to the following
1. Provided contactor contact disturbance does not
exceed three cycles of the nominal system frequency, the reswitching inrush currents will not
greatly exceed the full-voltage starting current
2. After short-term supply interruptions, such as
switch contact opening, current peaks on reclo-
Load elements
sure are unlikely to exceed 2 to 3 times existing
load current.
3. Conditions will be more severe on supply
changeover schemes - particularly where rapid
changeover is attempted; peaks of the order of
13 times full-load current may be experienced
on reconnection, at rated voltage.
Other examples of fast reswitching are the various
electrical braking methods employed to stop motors
or to reduce their speed. Current surges of the order
of 16 to 20p.u. may be experienced. See Figure
5.22. Some plugging systems insert resistances in the
lines to the motor terminals when reconnecting the
reversed supply. Their purpose is to reduce the
current peaks to predetermined limits.
Transient conditions in dynamic or d.c. injection
braking are not likely to be very severe. Current
peaks of the order of 4 to 5 p.u. may be expected.
They are mainly caused by the injected direct voltage, rather than by the short-circuiting effect of the
d.c. [23]. The longer the duration of supply disconnection, the less significant the transients are likely
to be.
Capacitor braking - where the motor is disconnected from the supply and reconnected to a bank of
capacitors, is generally used to reduce speed to some
predetermined level. Where braking to standstill is
required, a combination of capacitor and d.c. injection braking may be employed. In either event, the
effect of the capacitors across the terminals is to
maintain the air gap flux, and the stator voltage may
build up to an overvoltage condition - in spite of the
drop in rotor speed. Current transients may be very
severe on reconnection to the supply since out-ofphase conditions are likely. The same effects obtain
when terminal capacitors are used for power-factor
correction [13].
Figure 5.23 [23] shows the typical transients which
occur when reconnecting a motor at full-load torque
and rated voltage. In this case the interruption
period was 93 ms.
Regenerative loads
When braking certain equipment, such as cranes
and lifts, mechanical energy may be fed back to the
power source in the form of electrical energy. This
energy may also be absorbed by other equipment
operating in the installation. The surplus (or reverse) power will cause any generator supplying the
equipment to act as a motor, tending to drive its
prime mover. The generating set's speed will
increase and the natural response of the prime
mover's governor will be to reduce the fuel supply.
If the speed continues to rise the governor may
become ineffective. A runaway condition is only
avoided if the reverse power is totally absorbed by
the prime mover's mechanical losses and by the
generator's electrical losses. Reverse power in excess of 10 % of a diesel-generator's rating constitutes
a dangerous condition. Runaway will almost certainly occur if fed back power exceeds 25 % - 30 %
of the rating. If regenerative load is connected to a
generator, the total of the other load elements
should, at least, be equal to the regenerated power.
Failing this, it may be necessary to connect a
continuously-rated resistive load to absorb the regenerated power [24].
Voltage unbalance
The performance of a 3-phase induction motor is
adversely affected if its stator phase voltages differ
in magnitude or phase displacement [13]. The limiting criterion for voltage imbalance in an installation
may well be the amount of unbalance that the
motors in the system can tolerate. BS 4999: Part 101
(IEC 34-1) requires motors to be designed to
operate under 'virtually balanced' voltage conditions. Any asymmetry exceeding 2 % could lead to
increased motor losses and overheating. Asymmetry
in this context means that the negative phase sequence (n.p.s.) component of voltage does not
exceed 2 % of the positive sequence component. See
Sub-section 10.4.6 of Chapter 10 for an explanation
of n.p.s. (The reader's attention is directed to the
review by Roper and Leedham (Bibliography [15])
for a detailed treatment of the topic of 3-phase
unbalanced systems.)
Harmonic voltages
A full discussion on the effects of harmonics on
induction motor performance is outside the scope of
this chapter. The increasing application of nonlinear loads makes it necessary to appreciate the
distorting effect they may have on the voltage
Load considerations
Load elements
waveform of high impedance generators (see Section 5.4.3). The effects can be reflected into the
performance of motors in the system. Figure 5.24
(adapted from [25]) shows the typical effect of
varying the amount of harmonic voltage in the
supply to a motor. Whilst the 2 nd harmonic is not
likely to be present in any significant amount, 5 th
and 7 th orders may be very evident on convertor
loads. Perhaps the most important effect on performance is the increase in harmonic copper losses
within the motor; resulting in increased local heating
5.4.2 Capacitive loads
In discussing the concept of synchronous reactance
(Sub-section 4.2.6 of Chapter 4), we established that
the effect of armature reaction was to demagnetize
the main field at zero power factor lagging (inductive) loads and to directly magnetize the main field at
zero power factor leading (capacitive) loads. There
will be a tendency to over-excite the generator, as
the proportion of capacitive load increases unless
the main field current can be 'reversed' by the action
of the machine's excitation control system. This is
not possible with an ordinary brushless generator.
There is also little chance of controlling the overexcitation on compounded machines (of the type
described in Sub-section 3.3.4 of Chapter 3), which
use the automatic voltage regulator to 'buck' the a.c.
exciter field through a separate winding; more so, if
the compounding circuit is not phase-compensated,
i.e. if it is not directly related to load power factor.
The self-exciting effect of purely capacitive loads
is, therefore, to produce a high terminal voltage,
limited only by the magnetic saturation of the
machine. The terminal voltage is determined by the
intersection of an impedance line (whose slope
corresponds to the amount of load capacitance),
with the open-circuit magnetization characteristic of
the generator. Figure 5.25 [26] shows typical terminal voltages to be expected when a generator is
self-excited by various amounts of capacitive load.
There must be a limit to the amount of capacitance that can be switched on to a given generator if
voltage stability is to be maintained. Modern 4-pole
and 6-pole machines, having synchronous reactance
values varying between 2 p.u. and 4 p.u. for ratings
between 20kVA and 1750kVA, can seldom be
safely exposed to more than about 25 %-30 %
purely capacitive load [26]. The presence of active
load (kW) will alleviate the situation. This is
because such loads will tend to increase the main
and exciter field currents and oppose the selfexcitation effects of the capacitive load element.
The limitation on capacitive load level is confirmed by the characteristic in Figure 5.26 [27] since
the lowest practical 'working' level of excitation
should be of the order of 0.75p.u .. The characteristic illustrates the effect of self-excitation. It shows
the excitation required to maintain constant voltage
as reactive load is varied from inductive (lagging) to
capacitive (leading). It should be noted that the field
excitation becomes negative at point A, and is
positive again at point B. Point A is that at which the
load's capacitive reactance is equal and opposite to
Load considerations
may include inductive chokes, high-reactance transformers, and electronic devices giving high-voltage
starting pulses. The lamps may draw high currents at
low power factor during run-up - which may be
anything up to 15 minutes in some cases. (See the
characteristics of Figure 5.27.) They also have the
habit of extinguishing if their supply is interrupted.
Unless special circuits are used, the pressure in the
lamp then remains too high for the arc to 'restrike'.
The lamp has to cool down before this happens and
it could take up to 3 minutes to cool.
Lamp control gear usually incorporates power
factor correction capacitors. Problems can arise with
small standby generators supplying banks of these
lamps. The generator is 'confronted' with a large
capacitive load at switch-on while the lamps are
It is advisable to consult both lamp manufacturer
and generator supplier if the principal role of a
standby generator is to supply banks of discharge
lamps. In practice, disconnection of as many as half
of the power factor capacitors is possible without
any significant deterioration in lamp performance
Figure 5.26 Excitation required to maintain
voltage, with increasing load [27]
the generator's direct-axis synchronous reactance
(Xsd)' B is that point where the load reactance is
equal and opposite to the generator's quadratureaxis synchronous reactance (Xsq)'
Two of the major problems encountered, in practice, are:
• the capacitance of long transmission lines, in
isolated areas;
• the capacitors of bulk power factor correction
schemes, applied to main distribution boards.
Not only are high steady-state voltages possible at
switch-on, but excessively high transient terminal
voltages will also occur if sufficient purely capacitive
load is present without any appreciable level of
active (kW) load. Dinsley and Glover [27] cite some
of the transient voltage problems associated with the
switching of capacitive loads. The power factor
correction capacitors of fluorescent lamp installations (such as those in departmental stores, where
this type of lighting may constitute the bulk of the
essential load fed from a standby generator) can
have the effect of imposing high transient stresses on
the rotating diodes of brush less generators. A noninductive and matched resistance in parallel with the
main field offers a simple and economic solution to
the problem [28]. See Figure 3.8 in Chapter 3.
Large installations of mercury or sodium metallized vapour discharge lamps are often found in the
industrial, horticultural, building construction and
public lighting sectors. These lamps require control
gear to provide a high voltage across them to cause
them to strike. Equipment used for this purpose
5.4.3 Non-linear loads
In recent years, industrial plants and domestic
equipment have made increasing use of solid-state
power devices such as thyristors, in industrial convertor systems, and triacs (bidirectional thyristors).
in household appliances. Loads using this form of
control are the major sources of harmonic distortion
in supply networks. The non-linear load currents
that characterize such equipment may well be
accommodated, within acceptable limits, where the
power source is a low impedance public utility
supply. But, where the source is a smaller capacity
higher impedance generator, deleterious distortion
of supply wave-form may be difficult to contain. In
the following sub-sections we shall briefly review
some applications of these control devices; and
examine the distortion effects likely to be induced in
network supplies.
Domestic loads
Domestic loads are mentioned because they may be
a significant portion of the load on generating plant
which supplies isolated communities and townships
in developing territories. The two basic systems of
power control applied to household appliances are
those using phase control, and those using burstfiring techniques.
In the phase-control method, power is regulated
by varying the point on the applied voltage waveform at which a thryistor's conduction begins. The
supply to the controlled equipment is then switchedoff every time the voltage passes through zero.
Load elements
Figure 5.27 Typical run-up characteristics for discharge
lamps (Courtesy: Crompton Parkinson Ltd)
In burst-firing control (also variously referred to
as integral-cycling, fast cycling, and cycle selection),
load current is switched on and off in blocks or
chains of complete cycles. Switching is arranged so
that each 'on' pulse is triggered at zero voltage - for
resistive loads, and at zero current - for inductive
loads. The required proportion of full-load power is
obtained by varying the number of cycles in the 'on'
and 'off' periods - for example, the on/off periods
must be equal for 0.5 p.u. load.
Analyses of the load voltage wave-forms of both
types of regulator would show that, for a given
degree of power transfer, there are more significant
harmonics present in integral-cycle control than in
phase-control [30].
The disturbance caused by any domestic appliance depends not only upon the characteristics of
the appliance itself, but also upon the characteristics
of the supply network and on the connection point
of the appliance to the network. Where the supply is
obtained from a local generator (of finite impedance
and producing, in itself, a distorted voltage waveform), the possibility exists for harmonic instability
or harmonic magnification, extending into the appliance controls, causing firing-pulse irregularity and
exacerbating the total system disturbance.
Joint studies were undertaken by electric utilities
and manufacturers in the Common Market countries to set limits to the disturbances originating
from domestic appliances. That work led to the
issue, by CENELEC (the European Committee for
Electrotechnical Standardization), of a harmonizing
document (European Standard EN 50 006 (BS 5406
and IEC 555-2) : Limitation of disturbances in electricity supply networks caused by domestic and similar appliances equipped with electronic devices,
1975). Its objectives are not only to limit the quantity of 'generated' harmonics, but also to lay down
requirements related to the use of phase-control
systems for thyristors [31]. Its scope is restricted to
non-professional equipment. (See also the following
sub-section dealing with industrial loads.)
Colour television receivers frequently use halfwave rectifiers. Where there are large populations
of sets, all simultaneously notching into the same
half-cycle of supply, their combined effect can be
significant [32]. Table 5.6 (taken from Kidd and
Duke [33]) shows the harmonic phase currents recorded at substations feeding housing estates in the
United Kingdom. The 2 nd and 3 rd harmonics were
largely attributed to television and radio receivers
operating at peak times.
It was found that distortion was mostly due to
2 nd, 3 rd, 4 th, and 5 th harmonics. Current magnitudes varied cyclically with television viewing times
and, to a lesser extent, with other household activities. Phase angles of the 2 nd, 3 rd, and 4 th harmonics often followed a regular pattern - similar to the
harmonic magnitude, but could differ appreciably at
distribution substations fed from the same primary
Load considerations
Table 5.6
Maximumharmoniccurrentsrecordedon a domesticestate
Time (May)
Fundamental (A)
Harmonic number
500k V A transformer
11/0.415kV 3ph
22.30-Fri, Sat, Sun
, =
less than O.4A
network. Most significantly, appreciable levels of
d.c. current were present in the neutral, and values
as high as 45 A were recorded [33].
equipment to float at a potential of a few volts
above earth.
3. Overheating of the neutral conductor, leading,
perhaps, to burn-out - especially on those older
installations where the neutral is not fully rated.
Industrial loads
Power distribution systems should be designed to
handle full-loads at a crest factor of 3: 1.
Where generators are used for standby to a mains
supply on UPS installations they are likely to be
confronted with non-linear loads with crest factors
as high as 2.8 when the the UPS system goes into a
bypass mode. See Suh-section 11.4.4 of Chapter 11.
Crest factor is defined as the ratio of peak current to
Lm.S. current. A linear load with a sine wave
current wave-form has a crest factor of 1.41 (see
Sub-section 4.2.4 of Chapter 4).
If a standby generator cannot supply the peak
current of a high crest factor (HCF) load, its output
voltage wave-form may become flat-topped. This
can have a marked effect on the SPMS load. The
capacitor of the switched mode supply is charged up
to the peak of the supply voltage wave-form. If this
is reduced by flat-topping, the d.c. voltage switched
in the SPMS is also reduced. This, in turn, will
reduce the low voltage operating margin of the
computer equipment, and may cause power supply
drop-out when transient load steps are applied [34].
A convertor is a source of harmonic currents
which produce voltage drops at harmonic frequencies across the supply system impedances. The harmonic currents generated will depend upon the type
of convertor used, whereas the resulting voltage
harmonics will relate to the properties of the supply
network. The harmonic voltage drops combine vectorially with the supply voltage to give a distorted
wave-form at all points in the network. The effect on
the converting equipment itself, and on other items
of plant in the network, depends upon the degree of
distortion, and the sensitivity of plant to it. The
effects on public supplies are limited in the UK by
application of the Electricity Council Engineering
Recommendation G 5/3 - Limits for harmonics in
the UK electricity supply system. This tabulates the
maximum permissible kVA ratings for singleconvertor and single-regulator installations, against
listed harmonic currents, for the point of common
coupling (PCC) with other users.
The effects of non-linear loads on generators feeding industrial sites are more significant and less
predictable. The common element in all industrial
solid-state power supplies is the convertor. It may be
defined as a device or circuit by which a.c. is
converted to d.c.; or vice versa. The term is usually
used to describe a rectifier, but the inverter, which
converts d.c. to a.c., is also grouped under this
generic description. Convertors are to be found in
variable speed motor drives, chemical and materials
processing plants, battery charging, switched mode
power supplies (SMPS) to computers and other
technical loads, and in the uninterrupted power
supplies (UPS) often associated with such loads.
Switched mode power supplies are to be found in
personal computers, computer terminals, small
main frames, video display units, and in printers and
other computer peripherals. Medium sized telecommunication loads are also becoming predominantly
SPMS. Switched mode supplies employ capacitors
connected to the mains supply through a singlephase diode bridge. The capacitors, being high
voltage, high capacitance units, require high peak
charging currents. SPMS current wave-forms contain very high third harmonics, and multiples of the
third harmonic (45 % third harmonics are not
unusual on high crest factor loads).
On 3-phase supplies SPMSs will be connected
between line and neutral on all three phases. This
means that the third harmonic currents generated in
each phase add together in the neutral conductor to
give a very high Lm.S. neutral current (it may be as
high as 1.7 times line current for crest factors of 3
[34]). The following problems may then arise [34]:
1. High voltage drop along the neutral conductoL
This effectively reduces the supply voltage at the
load, accentuating the SPMS drop-out problem
(described in more detail later).
2. Neutral-to-earth voltages are generated. This
would cause the neutral terminal at the load
Load elements
Reference was made in the previous Sub-section
to the CENELEC standard EN 5~6
(based on
IEC 555-2) which restricts the current harmonic
components taken from the mains supply by household electrical appliances. IEC Sub-Committee 77A
is now working on an extension of IEC 555-2 to
widen its scope to cover all equipments with load
currents up to 16A per phase, and proposes a
further extension to deal with equipment with load
currents above 16A per phase. When issued, the
new standards are likely to become mandatory
under the Electromagnetic Compatibility (EMC)
Directive of the EEC. Meanwhile, the European
Commission have drawn up a mandate for CENELEC to amend existing European Standards and to
initiate the other standards required by the Directive (89/336/EEC). This is to be on a phased timetable to 31 December 1992.
When a synchronous generator is loaded by a
convertor, severity of voltage wave-form distortion
will depend upon several factors, such as the ratio of
machine-to-convertor ratings, the characteristics of
the machine, and the performance parameters of the
convertor itself. We shall examine some of these
factors, and the salient characteristics of convertors.
Rectifiers. Bonwick and Jones [35, 36 and 37]
have considered the characteristics of three-phase
bridge rectifiers operating with synchronous generators, and against complex source impedances.
They have derived theoretical wave-forms for three
classes of generator coupled to both controlled and
uncontrolled bridge-rectifier loads. The composite,
Figure 5.28 [35], illustrates the comparison of the
(theoretical) wave-form distortions to be expected
for various combinations of generator class and
Load considerations
rectifier bridge. Generators were classified as follows:
Class 1 no damper windings; substantial transient saliency
Class 2 no
quadrature-axis damper with identical
properties to the field windings; no transient saliency
Class 3 direct-axis and quadrature-axis damper
windings; subtransient saliency
See Chapter 4 for an explanation of these machine
Oscillograms from experimental tests corresponded very closely with the theoretically derived
wave-forms. Non-sinusoidal and notched waveshapes of the type shown in Figure 5.28 can be
resolved by Fourier analysis into a system of sine
waves, each consisting of a fundamental frequency
and a number of higher-order harmonics.
Convertors may be considered to be ideal current
harmonic sources. The orders of the harmonics
injected into the supply are directly related to the
'pulse number' of the thyristor convertor; and are
given by the expression:
n = kp ± 1
where n
the order of the harmonic
i.e. the number of
thyristors which successively begin to
conduct during one cycle of the supply
frequency; alternatively, the ratio of the
ripple on the d.c. voltage to the fundamental frequency on the a.c. side of the
an integer: 1, 2, 3, etc.
p = the pulse number,
Typically, half-controlled and fully-controlled 3phase thyristor bridge convertors give six pulses per
cycle. The harmonics present on their a.c. side
would therefore be the 5 th, 7 th, 11th, 13 th, 17th,
19th, etc., orders of the fundamental load current.
Whilst this is true of the line currents, it is also
possible for 3rdharmonic currents to be present in
the phase current. The latter condition may be
avoided by making the thyristor commutations
unequally spaced in time, but this causes asymmetrical currents to flow and even-order harmonics
are introduced into the current wave-form [38].
In industrial installations it is usual to employ
filters to suppress the generated harmonics. Their
design requires careful consideration of the load
duty cycle and accurate knowledge of system impedance - to avoid them acting as 'sinks' for harmonic
currents generated elsewhere in the network [39].
With regard to system impedance, Ainsworth [40]
concludes that:
1. Controlled
fed from a highimpedance a.c. system (such as an independent
generator) may magnify abnormal harmonics originating in the a.c. system and also magnify
similar harmonics caused by thyristor controlsystem errors or impedance unbalance.
2. Where system harmonic impedance is above a
certain value, the convertor may exhibit harmonic instability.
3. Instability arising from high harmonic impedance
may occur in a.c supply systems of low shortcircuit level, and also in those of moderate shortcircuit level if there is resonance. Behaviour
largely depends on the type of control system
used. Short-circuit level is defined by the ratio:
Any harmonic order is likely under non-ideal conditions. Because resonance can also occur at any order
it becomes necessary to predict and calculate the
whole spectrum of frequencies possible in all circumstances. Yacamini and de Oliveira [41] describe
a computational method for finding the harmonics
in convertor systems. All interactive effects are
considered, at the design stage. These may include:
unbalanced a.c. system voltages and impedances;
any type of convertor/transformer arrangement;
virtually any pulse number;
a non-infinite a.c system representation;
system resonances in both the a.c. and the d.c.
• no assumption of smooth d.c side current;
• harmonic filters.
The effects of features within the convertor itself,
such as control methods, unequal commutation
overlaps, and firing balance, may also be modelled.
Harmonic suppression measures may include:
1. The use of input filter banks, tuned to the 5th
harmonic. They also have the advantage of improving the input power factor, thus reducing the
input kV A demand; the p.t. at no-load becomes
leading as a consequence. The no-load kVAr is
normally about 30 % of full-load kV A. It is
necessary to ensure that the filter input kVAr is
not excessively high, since some standby generators may over-excite with leading power factor
loads [44]. See Sub-section 5.4.2.
2. Increasing the pulse number of the convertor.
Several convertors may be arranged in groups to
form a single convertor unit. They are connected
in parallel on the a.c. side, and in series or
parallel on the d.c. side. Complementary 6-pulse
groups, arranged in this way, can give 'unit
systems' corresponding to 12- and 24-pulse convertors. Harmonics are of the order n = 12K ± 1
and n = 24K ± 1, respectively.
3. Phase-shifting - achieved with the use of special
rectifier transformers which alter the phasing of
Load elements
the secondary winding with respect to the primary. As a consequence, the angle at which the
harmonics are produced (and thus the wave
shape) is altered although the harmonic magnitudes are unchanged [44].
4. Reduction of the supply system impedance,
either by increasing the generator frame size or
by using a specially designed low-reactance
Inverters. The inverter form most widely used in
a.c. motor drive schemes and in fixed-frequency
changers is that known as the d.c. link inverter. The
a.c. supply to it is first rectified to give a direct
voltage. This may be either at variable magnitude
(see Figure 5.29(a» or at constant level (Figure
5.29(b». A three-phase inverter then converts this
d.c. into a.c ..
D.C. link inverters are classified by the output
wave-form generated. This may be quasi-square
shaped (Figure 5.29(a», or a synthesized sinusoidal
waveshape (Figure 5.29(b». The latter is obtained
from a constant d.c. link voltage by switching the
inverter section thyristors at high frequency. This
gives a constant-magnitude, variable-width, output
voltage pulse simulating a sine wave-shape. A constant output voltage/frequency ratio is then obtained
by varying the thyristors' switching speed and the
mark/space ratio of the voltage pulse. Equipment
giving this type of output is known as pulse-widthmodulated (PWM) inverters.
One of two configurations may be employed in
the quasi-square wave inverter. The first uses a
half-controlled, or fully-controlled, rectifier to give
a variable d.c link voltage. The output frequency is
then controlled by varying the timing of the inverterstage thyristor timing pulses. The firing circuits of
the rectifier thyristors may also be coupled with
those of the inverter section to give a constant
volts-per-cycle ratio. The second configuration
(b) Pulse-width-modulated inverter
Figure 5.29
of d.c. link inverters
Load considerations
employs an uncontrolled diode-bridge rectifier in
the input stage. The variable voltage for the d.c. link
is obtained by using a d.c. chopper.
As far as generated harmonics are concerned,
what applies to rectifiers also holds good for these
types of inverter. The independent generator, feeding this type of equipment will 'see' the harmonic
currents expected from a 6-pulse converting device.
The second type of inverter is a frequency
changer, which has no intermediate d.c. stage, and
is known as the cycloconvertor. The fixed supply
frequency is 'chopped' directly by means of thyristors, arranged to conduct alternately in two
inversely-connected convertors for each phase. An
approximately sinusoidal wave-form is attainable.
The main limitation with this type of inverter is that
its output is at low frequency - less than the supply
frequency. It is best employed for drive speeds
corresponding to about one-third supply frequency
and has inherent regenerative capabilities. In the
latter context, it should be appreciated that d.c. link
inverter drives of the type shown in Figure 5.29 are
often fitted with regenerative convertors when they
are used in reversing-drive applications. These regenerative capabilities are of particular significance to
consumers using independent generators as a power
supply source. The comments made in Sub-section
5.4.1 with regard to regenerative loads will equally
apply here.
Convertor power factor. We have established (in
Sub-section 4.2.3 of Chapter 4) that the expression:
power factor
true power (watts)/apparent power
always holds good as a definition for power factor.
The alternative definition:
is only valid when both the voltage and the current
wave-forms are sinusoidal.
As we have seen, the harmonic currents in the
input to a thyristor convertor distort the line current
wave-form. The distortion factor would be unity if
no harmonics existed. A further consideration is
that there is a phase displacement of the fundamental a.c. line current with respect to the a.c
supply line voltage (see Figure 5.28(c». The extent
of this displacement is affected by operating characteristics such as convertor phase control and 'system'
reactance. Because phase control is not possible in
an uncontrolled diode bridge rectifier, it would have
a fairly high and constant power factor. On the other
hand, one expects a thyristor convertor to have a
lower, and variable, power factor.
The power factor is a product of distortion and
displacement [41]. The displacement power factor,
which satisfies the ever-valid definition of power
factor (i.e. the ratio: input watts/input volt-amperes)
is approximately equal to the ratio of Eo/ Eoo [41].
Figure 5.30 (curve (a» illustrates the relationship
between power factor and phase control delay angle
0'. Curve (b) shows the effects of reduced power
factor due to com mutating reactance. The mode of
operation changes from rectification to inversion at
the point where the curves intersect the abscissa.
The symbols used in Figure 5.30 are defined as
ED = output voltage
Eoo = d.c. no-load voltage
= com mutating current
Xc = commutating reactance (which is the sum of
source reactance and transformer leakage
reactance - where the latter is applicable)
Es = source voltage (this could be the rated
secondary voltage of a transformer - where
= controlled-bridge delay angle, or firing angle
During commutation, current Ie is, effectively, a
2-phase short-circuit current. It is only limited by the
instantaneous voltage difference between source
phase voltages and source (commutating) reactance.
The following example [34] typifies the effect of
HCF loads with regard to system power factor.
The harmonic analysis of a 2.6 crest factor balanced
current wave-form gave these results:
Load elements
The total harmonic distortion
pendix 5.2) is given by:
This distortion
(see also Ap-
factor is given by fundamental/r.m.s.
= 100/120
= 0.83
Generally, the displacement factor of HCF loads is
less than 1 and, consequently,
very low power
factors obtain. McLennan [34] cites a typical value
of 0.63 for a computer load.
It should be appreciated
that the power taken
from the supply source is a function of the fundamental current only, i.e. harmonics have the effect
of increasing the r.m.s. current, without increasing
the power demand.
may therefore
arise between r.m.s. current readings and the wattmeters connected to the system. Also, harmonics
may cause errors in power factor meters, where
these only measure the displacement
factor and do
not compensate
for the distortion
factor. A p.f.
meter may indicate unity power factor when, in fact,
it may be as low as 0.7 [34].
Currents in the neutral conductor
In large lighting installations, as with HCF loads, the
distortion of the supply current wave-form can be
Contributory factors are the discharge
tubes themselves and the inductive chokes forming
part of their control gear. BS 2818 (Ballasts for
tubular fluorescent lamps) effectively permits a 30 %
third harmonic current to be 'generated'
by lamps.
This limit relates to a test source having a harmonic
voltage content not exceeding 3 % (this compares
with the 5 % to 10 % permitted by BS 4999 Part 101,
for a generator).
In practice, some generators are
likely to have open-circuit voltage wave-forms with
a harmonic content greater than 3 %. Some may
even generate significant third harmonic voltages
and their wave-form may worsen with load. See
Appendix 5.2 for the wave-form analysis of a typical
proprietary machine.
Even in well-designed lighting installations, where
the load may be equally balanced between phases,
third harmonic (or triplen) currents will be additive
in the neutral conductor. If each of the three phase
banks of the lighting gives a third harmonic current
of 20 % (which is not untypical), the result would be
60 % r.m.s. current flowing in the neutral. If the
load imbalance were such that there was full-load
current in two phases, and none in the third, the
20 % harmonic would produce over 100 % current in
the neutral [43].
Griffin [43] cites problems associated with standby generators. For example: the carefully balanced
250 kV A fluorescent lighting load in a warehouse
was said to have had 13 % of full-load line current in
its neutral when fed from the public electricity
supply. When the same load was fed from a
320 kV A standby generator, the neutral current rose
to 250 A (72 % f.l.c.). The presence of a high third
harmonic content in the generator's
output waveform was exacerbating
the situation. The solution
was to replace the generator's
11/12 pitch stator
winding with one having a 2/3 pitch. The latter coil
has the advantage of covering two harmonic polepitches and links zero third harmonic flux so that, in
theory, third harmonic e.m.fs disappear from the
phase e.m.f.. If a coil is short- or long-pitched by
-rrln, 3-rrln, etc., no harmonic of order n will survive
in the phase e.m.f. [13]. In the warehouse problem,
the replacement
winding resulted in the neutral
current falling 'to less than it had been on the
A similar situation arises on convertors
have fully-controlled
thyristor rectifiers fed from a
supply. As the firing angle of the thyristors approaches
90°, the r.m.s. current in the
neutral reaches a maximum value equal to the
full-load line current. It falls away towards zero at
full power as the firing angle approaches 180°, The
maximum neutral current consists entirely of 3 rd,
9th, 15th, ... , etc. harmonics [44].
The generator manufacturer
should be consulted
for those applications
where third harmonic load
currents are likely to exceed 15 %. Fortunately,
modern high-speed generators
in the range 50 to
500 k V A usually have 2/3 pitch stator windings.
These windings are also available at outputs up to
1500 kV A, but with, perhaps,
some sacrifice in
output or voltage range [43].
Sizing of generators for non-linear loads
It is difficult to generalize on the capabilities
standard generators,
as far as convertor loads are
concerned. Rules of thumb, which involve the oversizing of generating plant, are usually too rigid and
overly simplistic. Opinions differ as to the capabilities of 'standard' machines in coping with non-linear
loads. Some would place the limit as high as 70 % of
rating, whereas others are more conservative
suggesting a limit at nearer 50 %. In any event, by
far the best approach is to pass all available information on the proposed load to the generator manufacturers and give them the opportunity of selecting the
right machine for the application - consistent with
any limits imposed on wave-form distortion.
Machine manufacturers
use general factors, based
on experience,
to estimate the generator
size for
non-linear loads. The crux of the problem is one of
The current
Load considerations
produced by non-linear loads are constant, regardless
ofthe power source. The voltage distortion, however,
is a direct function of the generator
(Note: There is no one value of impedance which
can be used for all harmonic orders. The stator
winding, being of pitched configuration,
has varying
reactance for each harmonic order. Therefore,
true theoretical approach to evaluating the voltage
distortion would need to examine all orders individually. Whilst this may be a simple mathematical
analysis, evaluation of the generator reactance at all
orders of harmonics
is complex and subject to
significant inaccuracies [45]).
The main concern is the effect of the distorted
voltage wave-form on automatic voltage regulators
(A VRs), as this could influence stability. Those
A VRs which are fed from the main generator windings could be disturbed by heavy distortion interrupting their power circuits (and causing misfiring of
their output device) or by changing peak values in
their sensing circuits. The generator
system is just one of several 'closed loops' within a
power network supplying non-linear loads. Compatibility of these closed loop systems is essential for
stable network operation. (See Sub-section 5.4.4.)
The development
of excitation control systems
employing permanent magnet auxiliary exciters (see
Chapter 7) has done much to minimize the effects of
voltage distortion. The power of the A VR remains
constant, irrespective
of generator
output voltage
wave-form. Only the sensing circuits are subjected
to the voltage distortion. Improvements
in this area,
as a result of average (rather than peak) and 3-phase
(rather than single-phase)
sensing, gives the best
possible signal to the A VR, and better overall
regulation and stability. The isolation and filtration
techniques used in modern A VRs appear to provide
tolerable performance
in all but the most extreme
It is feasible to design generators
with specific
winding pitches and low reactance
in order to
minimize distortion levels. But, for sound commercial reasons, derating of standard industrial generators is the only practical solution. Fundamental
machine design changes for specific applications are
not viable options. Generator sizing is derived from
the manufacturer's
past practical experience,
the synchronous
(Xs) or subtransient
(X'd) reactance values are usually used as the yardstick [45].
(See Sub-sections
4.2.6 and 4.3.1 of Chapter 4.)
Newage International,
for example, have selected a
0.12 p.u. subtransient
reactance as a good practical
figure (on the basis of a 6-pulse, variable speed,
motor drive, having 26 % current distortion, giving
20 % voltage distortion at approximately 80 % of the
generator rating).
The following example [45] illustrates how a generator would then be sized to power a 6-pulse,
200 kV A input, uninterrupted
power supply (UPS)
system, plus a 30 kV A linear load. Because the UPS
rating should not exceed 80 % of generator rating, at
0.12 p. u. subtransient
reactance, the generator rating (at 0.12 X'd) must be (200/0.8)kVA
= 250kVA.
The maximum
total load is (200 + 30) kV A =
230 kV A. The generator
may then be sized on
250kVA - at 0.12p.u. X'd'
If the nearest standard generator has an industrial
rating of 275kVA
at O.13p.u. X"d, its rating at
0.12p.u. X'd would be:
275 x (12/13) = 254 kV A
It would therefore
be suitable for the proposed
Sensitive equipments
We shall now consider the effects that wave-form
distortion can have on equipment connected to the
On-site generators. We have established
wave-form distortion will affect, not only the sensing
and firing circuits of the A VRs associated with
but also (where a
machine is self-excited) the excitation power available from the main output winding. Furthermore,
the firing angle of the thyristor load approaches that
of any thyristor in the generator's
automatic controls, the two systems may interact to give overall
system instability. Separate pilot excitation will be
an advantage.
The presence of any 2 nd harmonics (and multiples thereof)
will induce pole-face currents in
generators. This could lead to local overheating.
Capacitors. BS 1650 (Capacitors for connection
to power frequency systems) limits the harmonic
overloads permissible on capacitors.
Because the
reactance of a capacitor is inversely proportional
it is particularly
sensitive to high frequency harmonics. The presence of any resonance
conditions in a power factor correction network will
not help either. A series inductor in the power factor
correction circuit will prevent this happening. The
inductor must have a reactance about 1/10 th that of
the capacitor bank at fundamental
Technical loads . These may include computers.
signals and radar installations, and
electronic office equipment.
All are sensitive, in
varying degrees, to wave-form distortion.
limits on total harmonic content are set at between
3 % and 5 %. Electronic office equipment (such as
and associated laser and ink-jet
printers) usually form a small, but critical, proportion of the total load in many commercial installations. The load attributable
to these machines is of
the order of 10 to 20 watts per square metre of office
area. It is customary to provide larger installations
with some form of standby supply, based on enginedriven
UPS systems.
Chapter 11.)
Load elements
Convertors operating under fully-controlled conditions tend to produce low power, high frequency,
signals in the range 150kHz to 30 MHz. These
signals are likely to cause disturbance to nearby
radio receiver and transmission equipment either by
direct radiation from the convertor housing itself, or
by being carried on the incoming and outgoing
power conductors to radiate elsewhere. Direct radiation from within the converter itself may be suppressed by simply earthing its housing cubicle.
Signals carried along the external cables require the
use of special filter, or suppression, networks to
prevent them radiating elsewhere. The technology is
described as radio interference suppression or radio
frequency interference (RFI) suppression.
Radio interference limits are defined in several
standards, notably:
• BS 800, for industrial and household equipment;
• BS 1597, and its associated Code of Practice
BS5260, for marine installations;
• the Federal German Republic's VDE 0875whose Grade G is very similar to BS 800 Part 3;
• military specifications and defence standards such
as those whose curves are compared in Figure
5.31, where the ordinate scale of dB(/l-V) is based
on 0 dB(/l-V) being equivalent to 1 /l-V. The
dotted curves show the RFI levels recorded on a
Figure 5.31 Limits of radio frequency interference
various standards (Courtesy: Newage International
set by
typical machine, in tests before and after suppression.
Many drive systems are required to operate in a fully
automated mode, controlled by telemetry links. It is
important that both convertor equipment and power
source are adequately suppressed against interference in the signal frequency range used for the
telemetry link.
Motors and transformers. Motors, like generators, are particularly affected by 2 nd order harmonics and multiples thereof. The effects are increased
losses and localized heating.
The presence of third harmonic currents in transformers will lead to increased iron loss and overheating of windings. Third harmonic voltages may
overstress insulation and could give rise to resonance, at third harmonic frequency, between transformer windings and network capacitance.
Cables. Cables should be capable of withstanding the effects of a total voltage distortion of 10 %
without damage. Harmonic currents may reduce
ratings due to: PR, eddy current, and skin and
proximity effects. A current distortion limit of 10 %
is recommended if reduction in rating is to be
avoided (33].
Control and protection devices. Unwanted harmonics are likely to affect generator excitation,
convertor firing controls, and electronic devices in
Load considerations
systems concerned with monitoring, alarm and protection functions. For example, 'notches' in voltage
wave-forms could cause incorrect operation of voltage and frequency monitors unless they are adequately filtered; overload protection relays fed by current transformers may trip at low current levels,
when harmonic currents are present.
5.4.4 Cyclic loads
Certain loads, such as radio identification beacons
or transmitters, vary cyclically at a regular rate, with
periodic times between 0.3 and 5 seconds. Where
regular load-pulse intervals interact with the prime
mover governor's response characteristics, speed
(and therefore frequency) oscillations are likely.
Improving engine speed response, by fitting, for
example, a fast-acting governor, may merely introduce instability and, possibly, excessive wear on
the governor's precision parts and linkages. Conversely, reducing governor response should alleviate a
hunting problem, but at the risk of putting normal
speed recovery outside acceptable limits. Therefore,
where regular load pulses are expected, a compromise must always be sought in selecting the correct
generator shaft system inertia and the correct governor response characteristics for the proposed duty
cycle. Sufficient rotating energy, in the form of
inertia mass, should be built into the generating
unit(s) to minimize the momentary speed changes
that will be outside the control response of governor
and turbo-charger (where one is fitted). To this end,
the largest standard flywheel capable of being fitted
to the diesel engine should be considered. The
additional rotor inertia of a larger frame generator
will also assist in this respect.
Manning [46] gives an example of typical cyclic
load conditions relating to a radio relay station. The
installation was for the External Radio Service of
the British Broadcasting Corporation (BBe). The
input power demand from the HV rectifiers feeding
each of the four anode-modulated transmitters varies rapidly with the degree of transmitter modulation and is dependent upon the programme material
being broadcast from the station.
(Koeppen [47] gives the following formula for
calculating the power required for a transmitter:
For the example cited [46], in worst-case conditions, all four transmitters are operated on the same
programme and their modulation level varies between zero and 100 %. The transmitter load fluctuates rapidly (at syllabic or tonic frequencies) between 1920kW and 3020kW (approximately 60%
load step changes). Generator voltage controls and
engine governors were designed to take account of
these load fluctuations in the frequency range 3 Hz
to 15 Hz.
The three basic interactive control loops that
would have had to be considered in a design exercise
of this type are shown, in simplified form, in Figure
5.32 [48]. The schematic shows the mutual interactions between the three automatic-control loops of
engine governor, generator control, and rectifier
control. The characteristics within the functional
blocks of each loop give an indication of their
individual step responses. The d.c. current from the
rectifier to the anode-modulated transmitter is the
final loop output in the schematic and has a direct
influence on the prime mover speed - the parameter controlled in the first of the three loops.
When the BBC broadcasts its signature tune and
time signals, using all four transmitters, rapid and
substantial load changes occur. Their nature is illustrated in Figure 5.33, which reproduces traces taken
during engine performance tests. These tests were
designed to simulate the hourly time-signal load
5.5 Related standards
Specific reference has been made to the following
British Standards in this chapter. Entries in brackets
contain reference to corresponding and/or technically equivalent international or world regional
BS5424: Part 1 (IEC 158-1 and lA) - Contactors
BS 2949 (lEC 92-5) - Rotating electrical machines
for use in ships
BS 5406 (EN 50006) - The limitation of disturbances
in electricity supply networks caused by
domestic and similar appliances equipped
with electronic devices
BS 2818 (IEC 82) - Specification for ballasts for
tubular fluorescent lamps
BS 4999 - General requirements for rotating
electrical machines Part 101 (lEC34-1) Specification for rating and performance
BS 1650 (IEC70) - Capacitors for connection to
power-frequency systems
BS 800 (VDE 0975 Grade N) - Specification for
radio interference limits and measuremenlS
for equipment embodying small motors,
contacts, control and other devices causing
similar interference
Related standards
Schematic legend
1. Engine speed control
1.1 Generator set inertia
1.2 Speed governor
1.3 Fuel injection system
1.4 Deviation produced by load torque
2. Generator
voltage control
Excitation circuit
Automatic voltage regulator
Deviation of frequency (rotary speed)
Deviation of load current Ie
3. Rectifier control
Thyristor circuit
Automatic voltage control linc. triggering circuits)
Deviation of a.c. input voltage
Time-delay of voltage increase (by filters)
4. D.c. load current response determined by load changes
(e.g. modulated-transmitter
as f'igure 5.33)
Figure 5.32 Typical schematic of the control loops in a
diesel generator/controlled convertor system (Courtesy:
Load changes 917 kW (baseload) ±443 kW
(Simulating G.M.T. signal)
Figure 5.33 Oscillograms showing variations of voltage and
speed/frequency during load changes imposed by hourly
transmitter time signals (Courtesy: NEI Allen Ltd)
Load considerations
BS 1597 - Radio interference suppression on marine
BS 5260 - Code of practice for radio interference
suppression on marine installations
5.6 References and bibliography
5.6.1 References
1. Valentine, A.C. Private power generation.
A.P.E. Engineering the technical journal of
Amalgamated Power Engineering Ltd., Bedford, England, Number 18 April (1975)
2. Windett, A.S. Reducing the cost of electricity
supply, 1st edn, Gower Press, (1973)
3. Bolton, D.J. A seasonal tariff for domestic
supplies. Proceedings of lEE, 112(2), 366--374
4. Freeman, P.J. Electric Power Transmission and
Distribution, 2nd edn, One of a Series of Engineering Science Monographs, George G. Harrap, (1977)
5. Smith, C.B. (Editor) Efficient Electricity Use,
Pergamon Press, (1976)
6. Rolls, T.B. Power Distribution in Industrial
Installations. No. 10 in a series of lEE Monographs, Peter Peregrinus on behalf of the I.E.E.
7. Dale, K.M. Dynamic-programming approach to
the selection and timing of generation-plant
additions. Proceedings lEE, 113(5), 803-811
8. West, Gerald S. The economic viability of private generation or total energy schemes. A
paper presented to lEE North Midland Centre
meeting Nov.1971 and based on one to the
lEE/ERA 'Distribution' Conference Oct. 1970
9. Perry, R.J. Isolated diesel power stations in
Western Australia. Electronics and Power
(lEE), October (1985) pp. 749-756
10. Sterling, M.J .H. Power System Control. No.6 in
Control Engineering series, Peter Peregrinus on
behalf of the lEE (1978)
11. Fitzgerald, A.E., Kingsley, C. Jr, and Kusko,
A. Electric Machinery, 3rd edn, McGraw-Hill,
12. Stein, R. and Hunt, W.T. Jr. Electric Power
System Components - Transformers and Rotating Machines, Van Nostrand Reinhold Company, (1979)
13. Say, M.G. Alternating Current Machines, 5th
edn, Pitman Publishing, (1983)
14. Stephen, D.D. Effect of system-voltage depressions on large a.c. motors. Proceedings lEE
113(3), 500-506 (1966)
15. Kayondo, G.E. Automatic load shedding and
reswitching. A paper to an lEE Colloquium on
'Design application and maintenance of large
No.170 pp 109 -116 (lEE) November (1978)
16. Barber, D.E. Effects of motor loads on generating sets. Electrical Export Review, JulySeptember pp. 55-59 (1973)
17. Escombe, F. Selection of on-site power plant
for starting motor loads. Power Generation Industrial, a series of three articles, January-May
18. Kendall, P.G. Light flicker in relation to powersystem voltage fluctuation. Proceedings lEE,
113(3),471--479 (1966)
19. Griffin, J. A.C. generators - construction and
performance. Paper D presented to lnpower
'86, a Conference jointly organized by the Institution of Diesel and Gas Turbine Engineers and
Fuel and Metallurgical Journals Ltd. September
( 1986)
20. Bibby, G.L. Electrical metrology and instrumentation. Chapter 6 of Electrical Engineer's
Reference Book 14th edn, Editors M.A.
Laughton and M.G. Say, Butterworth, (1985)
21. Reeves, A.E. (Editor) Newnes Electrical Pocket
Book, 18th edn, Butterworth, (1983)
22. Shoesmith, J. Power system disturbances due to
induction motor reswitching transients. A paper
presented to an lEE International Conference
Sources and effects of power system disturbances, Conference publication No. 110 pp. 1-5
(lEE) April (1974)
23. Smith, I.R. and Sriharan, S. Induction-motor
reswitching transients. Proceedings lEE, 114(4),
503-509 (1967)
24. Jones, D. and King, A.C. Standby power generation. A paper presented to the lEE's Second
Conference on Installation Engineering, Conference publication No. 251 pp. 10-14 (lEE) June
25. Wood, W.S., Flynn, F.P. and Poray, A. Effects
of supply voltage waveform distortion on motor
performance. A paper in Conference on Installation
No. 110
pp. 268-272 (lEE) June (1985)
26. Griffin, J. Why standby generators dislike capacitors. European Power News, May-June
27. Dinsley, J.L. and Glover, K.J. Voltage effects
of capacitive load on the synchronous gene~
rator. Proceedings lEE, 111(4),789-794 (1964)
28. Standby sets supplying fluorescent lighting. An
article by Petbow Ltd. in Electrical Times, April
29. Shepherd, W. and Gallagher, P.J. Some properties of thyristor controlled resistive load with
integral cycle switching. A paper in conference
on Installation Engineering, publication No. 110
pp. 188-192 (lEE) June (1985)
30. Lingard, B.W.; Johnson, R.W. and Shepherd.
W. Analysis of thyristor-controlled single-phase
Referencesand bibliograpy
loads with integral-cycle triggering. In correspondence to Proceedings lEE, 117(3),
607-608 (1970)
31. Deloux, G. International standardization and
electrical power supply network disturbances. A
paper in Conference on Installation Engineering, publication No. 110 pp. 223-227 (lEE)
32. Corbyn, D.B. Abnormal harmonic disturbances; assessment and mitigation. A paper in
Conference on Installation Engineering, publication No. 110 pp. 235-240 (lEE) April 1974
33. Kidd, W.L. and Duke, K.J. Harmonic voltage
distortion and harmonic currents in the British
distribution network, their effects and limitation. A paper in Conference on Installation
Engineering, publication No. 110 pp. 228-234
(lEE) April 1974
34. McLennan, AA High crest factors on UPS
systems - comparative technologies. A paper
read at a Seminar presented jointly by: Liebert
Ltd., Emerson Electric Ltd., Dale Electric of
Great Britain Ltd, and Newage International
Ltd., at Croydon in March (1990)
35. Bonwick, M.E. Characteristics of a diodebridge-loaded synchronous generator without
damper windings. Proceedings lEE, 122(6),
637-642 (1975)
36. Bonwick, M.E. and Jones, V.H. Rectifierloaded synchronous generators with damper
windings. Proceedings lEE, 120(6), 659-666
37. Jones, V.H. and Bonwick, W.J. Three-phase
bridge rectifiers with complex source impedance. Proceedings lEE, 122(6),630-636 (1975)
38. Bird, B.M., Marsh, J.F. and McLellan, P.R.
Harmonic reduction in multiplex convertors by
triple-frequency current injection. Proceedings
lEE, 116(10), 1730-1734 (1969)
39. Johnson, P.A Reducing the effects of harmonic
distortion. Electrical Review (1977)
40. Ainsworth, J.D. Harmonic instability between
controlled static convertors and a.c. networks.
Proceedings lEE, 114(7), 949-957 (1967)
41. Yacamini, R. and de Oliveira, J.c. Comprehensive calculation of convertor harmonics with
system impedances and control representations.
Proceedings lEE, 133 Part B (2), 95-102 (1986)
42. Forrester, W. Harmonic reduction of rectifiers.
A paper read at a Seminar presented jointly by:
Liebert Ltd., Emerson Electric Ltd., Dale Electric of Great Britain Ltd, and Newage International Ltd., at Croydon in March (1990)
43. Griffin, J. Fluorescent lamps, neutral currents
and standby generators. Electrical Review,
January (1979)
44. Nickels, L.E. Variable-speed solid-state d.c.
drives. Chapter 26 of Switchgear and Control
Editor Robert W. Smeaton,
McGraw-HilI, (1977).
45. Moody, R.L. and Rounce, W.J. UPS and diesel
generating set compatibility. A paper read at a
Seminar presented jointly by: Liebert Ltd.,
Emerson Electric Ltd., Dale Electric of Great
Britain Ltd, and Newage International Ltd., at
Croydon in March (1990).
46. Manning, J. Allen engines on BBC World Service. Product Information, 1, W.H. Allen/NEI
Allen. Ltd., Bedford, England, (1984)
47. Koeppen, B. Diesel-generator power supply for
transmitter stations. Brown Boveri Review, 73,
723-727, (1986)
48. Boettger, K. and Schmidt, J. Operating experience with secured power supply systems comprising standby diesel driven generating sets and
rectifier/charger units with batteries. A paper
presented to the 3rd International Telecommunications Energy Conference (Intelec 81). Conference publication No. 196 pp. 338-345 (lEE)
5.6.2 Bibliography
The references listed above contain, within themselves, useful cross-references to related papers on
the topics covered. Additionally, the lEE Conference publications mentioned in [15], [22], [24] and
[48] also contain a significant number of contributions which may be of help to the reader who is
researching this chapter's subject matter.
In this context, the lEE Conference publication
No. 272 covering the first International Conference
on Industrial Power Engineering, organized by the
lEE in London in December 1986, also includes
helpful and informative papers on electrical power
system design and the nature of loads in industrial
The following literature, reports and papers also
cover the various topics of this chapter, in some
1. ANSI/IEEE Standard 241-1983 the American
National Standard prepared by the Institute of
Electrical and Electronic Engineers, Inc. IEEE
Recommended Practice for Electric Power
Systems in Commercial Buildings, the Gray
Book 3rd Edition, 1983 published by the IEEE
and distributed by Wiley-Interscience.
2. Sutton, K. W. A Consultant's approach to total
energy. Electronics and Power, pp 614-618,
May 1975 (lEE).
3. Roberts, C.A and Perry, E.J. Statistical Analysis of Integrated Energy Systems Pre-Print 363,
Institution of Diesel and Gas Turbine Engineers
4. Total energy for Leeds Infirmary. A report in
Diesel Engineering, Spring (1977)
5. Was, G.S. and Golay, M.W. Cogeneration - an
energy alternative for the U.S.? Energy, 4,
1023-1031 (1979)
Load considerations
6. Goldman, S.B. and Golay, M.W. Strategies for
scheduling power plants producing both heat
and electric energy. Energy, 4, 1033-1051
7. Dettmer, R. Combined heat and power - the
cogeneration game. Electronics and Power, October, 735-739 (1985)
8. Norris Economic comparisons in planning for
electricity supply. Proceedings lEE, 117(3),
593-605 (1970)
9. Snider, L.A. and Smith, I.R. Prediction of
performance of induction-motor/synchronous
Proceedings lEE,
120(1), 79-86 (1973)
10. Payne, M.G. Motor starting on diesel generators. Electronics and Power, June, 479-483,
11. Duguid, J.N. Starting 3-phase induction motors. Electrical Times, December (1975)
12. Smith, I.R. and Sriharan, S. Transient performance of the induction motor. Proceedings
lEE, 113(7), 1173-1181 (1966)
13. Sriharan, S. Digital simulation of a group of
induction motors. Proceedings lEE, 122(12),
1399-1401 (1975) (a computer-program description)
14. Harrison, D. The dynamic braking of induction
motors. Proceedings lEE, Part A, 103121-133,
( 1955)
15. Roper, R.D. and Leedham, P.J. A review of
the causes and effects of distribution system
3-phase unbalance. A paper presented to an
lEE International Conference Sources and effects of power system disturbances, Conference
publication No. 110 pp 83-88 (lEE)
16. Gallagher, P.J., Barrett, A. and Shepherd, W.
Analysis of single-phase rectified thyristorcontrolled load with integral-cycle triggering.
Proceedings lEE, 117(2),409-414 (1970)
17. Freris, L.L. Multigroup convertors with series
a.c. connection. Proceedings lEE, 118(9),
1247-1252 (1971)
18. Williams, S. and Smith, I.R. Prediction of performance of synchronous-generator/transformer/bridge-convertor installations. A summary
paper in Proceedings lEE, 122(11), 1286--1288
19. Clark, G.R. and Ferry, G.A. Modern power
electronic devices and their application to the
control of power transmission and industrial
supplies. GEC Review, 5(1), 4-17 (1989)
20. Barnes, R. Harmonics in power systems. Power
Engineering Journal, 3(1), 11-15 (1989)
21. Brown, J. and Maylin, R.F. Electricity supplies
for Addu Atoll. Power Engineering Journal,
3(1), (1989)
22. Yacamini, R., Hu, L. and Fallaize, R.A. Calculation of commutation spikes and harmonics on
offshore platforms.
Part B, (1) (1990)
Proceedings, 137,
Appendix 5.1 - Selection of power
plant for load conditions
The following example has been devised to demonstrate the practical methods used for assessing the
sizes of a.c. generator and prime mover required for
a given installation.
A standby generator is required to supply the following essential loads within a manufacturer's premises with 415 volts 3-phase 50 Hz, in the event of a
utility supply failure.
Load details
Single phase - 240 V
1. Fifteen high-pressure mercury vapour, highbay
luminaires (400 W nominal each) and 60 fluorescent fittings (100 W nominal each), all corrected
to 0.95 power factor lagging.
2. 10 space heaters (5 kW each) at unity power
Single phase - 415 V
3. Seven welding machines each rated at 12A primary current at 0.4 power factor lagging
Three-phase - 415 V
4. A process control system rated at 10kW at 0.85
power factor lagging.
5. A series of motors driving process line equipment, as follows:
1kW (7 number)
3kW (3 number), all d.o.!. started
6. One motor, 30kW d.o.!. started;
One motor, 80kW star/delta started.
Governing conditions
1. For safety and operational reasons, the essential
lighting (load element 1), the process control
system (load element 4), five 1 kW motors and
one 3 kW motor (load element 5) must be restored as soon as possible, after utility supply
2. The 80 kW motor (load element 6) may not be
started until all other motors are on-line.
Selection of power plant for load conditions
3. The process control equipment is sensitive to any
voltage dips of more than 10 % magnitude.
4. The remainder of the site's plant will tolerate a
20 % voltage dip.
Load application stages
Based on these governing conditions, the following
stages of load restoration were agreed with site
Full load
75 % f.l.
50 % f.l.
estimate of the peak starting power requirement
can be obtained from the expression [17]:
peak power
Load element
Process controller
1 No. - 3 kW motor d.D.!.
5 No. - 1 kW motors d.o.!.
1 No. - 30kW motor d.o.!.
2 No. - 3 kW motors d.o.!.
2 No. - 1 kW motors d.o.!.
1 No. - 80 kW motor star/delta
7 No. - welding machines
10 No. - space heaters
The running (steady-state) and starting (transient)
kVA and kW requirements of each load element
must first be analysed, before being tabulated for
each of the load-application stages proposed.
Motor nameplates, or manufacturers' literature
will provide details such as:
full-load power
power factor
starting current
starting torque
peak (or pull-out) torque
Given such data it is possible to calculate:
1. the starting kVA, which may be obtained directly
from the starting current;
2. the running kW requirement: this from the ratio
of operating load divided by the motor efficiency.
If the operating load is not known, assume the
worst case, i.e. the motor is running on full load;
3. the running kV A: this is given by the running
power, as at 2 above, divided by the power
factor. It should be appreciated that power factor
will vary considerably with the operating load.
Typical values for the sizes of motor considered
in this example are tabled below;
4. the maximum kW demand at starting: since
information on the maximum power factor during starting is seldom available, a reasonable
[rated power x peak torque
(p.u.)]/motor efficiency
Peak or pull-out torque is usually between 2 and
2.8 times rated torque. For this example, we shall
assume 2.S p.u.
Load analysis
Single phase loads
Lighting (load element 1) The IS mercury vapour and 60 fluorescent lamps may be balanced over
the three phases. The load per phase would then be
as indicated in Table SA.
Space heating (load element 2) Since there ate ten
space heaters over three phases, one phase must
supply four heaters; the loads are therefore as
detailed in Table SB.
Welding machines (load element 3) With each of
the seven welding sets across one pair of 41SV line
conductors there is a 'worst-case' condition of three
sets across a particular pair of lines.
This gives the loading conditions shown in Table sc.
Three-phase loads
Process controls
The load is:
Induction motors
Typically, 47 kYA
47 x 0.5 = 23.5 kW
The loads are listed in Table SD.
Load demand at each stage
Strictly speaking (as discussed in Section S.4.1 and
illustrated in Figure S.21), the effect of power factor
in transient load switching should be considered by a
vectoral summation of existing and incoming loads
at each stage. However, for the purpose of this
exercise, we shall ignore that requirement and make
an arithmetical addition instead. This is normally
Load considerations
Table 5A
Mercury vapour
5 x 400 x 1.8 = 3.6kVA
and 3.6 x 0.95 = 3.4 kW
20 x 100 x 1.8 = 3.6kVA
and 3.6 x 0.95 = 3.4 kW
1.7 kW
(see Figure 5.27)
and 2.3 kW
Total equivalent 3-phase load
20.4 kW
21.6 kV A
31.5 kVA
Table 58
Maximum single-phase load
as running
Equivalent 3-phase load
Table 5C
Maximum single-phase load
Equivalent 3-phase load
(3 x 415 x 12)/
= 14.9kVA
14.9 x 0.4 = 6kW
55.8 kW
Table 50
Rating (kW)
at rated power
Power factor
at rated load
Power requirements
80 (Y-.i)
acceptable for sizing exercises of this type as the
arithmetic addition always gives a larger kYA load
demand figure than the vectorial addition.
The three-phase starting and running (kY A and
kW) requirements are summarized in the following
tables and are based on the stage load elements
listed earlier in this Appendix.
The voltage dip experienced by the generator is
related to the highest applied kYA at any of the
three load stages (column 1 of Table 5 E). From
Table 5.E it will be seen that the maximum is
390.9 kY A; and occurs at stage 3.
Referring back to the governing conditions, we
see that a critical factor is the voltage dip limit of
10 % imposed by the process control equipment.
Assuming that the generator voltage dip characteristic of Figure 5.17 applies in this case, it becomes
necessary to restrict the applied kY A to no more
than about 0.5 times the maximum rating of a
generator frame if the voltage dip is to be limited to
Selection of power plant for load conditions
Table 5E
Starting and running kVA requirements
kV A applied
this stage
Starting k VA
Running kV A
Peak kV A
kV A this
Running kV A
kV A from
10 %. This means that we would need to select a
generator rated for 782kVA (= 390.9 -:- 0.5), i.e.
2.7 times the steady-state kV A requirement
(294.4 kV A) of our essential load. Obviously, this is
not an economical proposition. We are faced with
two practical alternatives:
1. The first is to 'remove' the performance restriction imposed by the process controller, by
supplying it from a dedicated inverter, thus isolating
it from all disturbances in the standby generator
supply (see Chapter 11). This means that we could
work to the wider (20 %) voltage dip tolerance
applicable to the remainder of the consumer's essentialloads (governing condition 4). This being so, we
can select a generator frame with a maximum rating
of 390.9 kV A to meet the less arduous 20 % voltage
dip tolerance (see Figure 5.17). The selected generator is now only some 1.3 times the capacity
required by our steady-state load. This is not unreasonable, economically, and it does allow for the
'accommodation' of further essential loads, either
now or at some future date when the site is expanded.
2. The alternative is to rearrange the load switching sequences - provided, this is acceptable to the
user. So that, for example, an intermediate loadrestoration stage is introduced between stages 2 and
3. The revised load-switching pattern could then be:
The revised start/run kV A requirements would
then be as shown in Table 5 F. This now gives us a
maximum applied kVA, at any stage, of 361.8 (as
opposed to the 390.9 kVA of our original loading
pattern). This means that, if we are to meet the 10%
voltage dip limitation imposed by the process controller, the selected generator frame now needs to
Total kVA
be rated for 723.6kVA (= 361.8 -:- 0.5) i.e. 2.5
times the required steady-state rating - still not an
economical proposition .•
Perhaps the best compromise would be to provide
the process controller with a separate and 'buffering'
inverter and, at the same time, revise the loading
stages as Table 5 F above. If this is done, the
maximum rating of the generator frame needs to be
at least 361.8 kV A, in order to meet the 20 %
voltage dip limitation.
As explained in Section 5.4.1 the selected
machine must have an overload capability greater
than the largest peak kV A figure given in column 4
of Tables 5 E and 5 F. In this example these maxima
are 486 and 407.7kVA; and equate to per-unit
overload requirements of 1.24 (486/390.9) and 1.38
(407.7/294.4), respectively. Both are well within the
2.0p.u. capacity, for 10 seconds, associated with
generators of this size. Generally speaking, in order
to achieve the minimum size of generator based on
overload performance, it is desirable to apply starting loads (column 2) in descending order of magnitude, i.e. the largest starting load should be
switched-on first, and the smallest last. Whilst this is
the ideal, it is seldom achieved in practice.
Having considered the optimum sizing of the
generator (against kV A requirements), it is now
necessary to analyse the kW requirements. It is
these that will determine the rating of the prime
mover to be used. Table 5 G gives the starting and
running kW demands at the three stages corresponding to those of Table 5 E.
The first conclusion to be drawn, when comparing
the total running kW (249.2) from this table with the
corresponding total kVA (294.4) in Table 5 E, is
that the essential load's power factor is of the order
of 0.85 (249.2/294.4). This implies that the selected
generator should not have to be derated for power
factor (see Chapter 3, Section 3.7.2).
The next point to be noted is that the peak
starting demand is 417.4kW. When the generator
losses are taken into account (the efficiency of a
machine of this size is approximately 93 %), this
demand equates to 449kW, and is the output required from the prime mover during the load
Load considerations
Table 5F Starting
and running
kV A applied
this stage
Table 5G Starting
and running
this stage
kVA requirements
Starting k VA
Running kVA
for revised stages
Peak kVA
kVA this
Running kVA
kVA from
Total kV A
k W this
Running kW
k W from
Total kW
kW requirements
Starting kW
Running kW
Peak k W
changes in stage 3. This means that the prime mover
need only be continuously rated for about 408 kW,
at site conditions since advantage can be taken of its
10 % overload capacity to meet the momentary
449 kW demand. A prime mover of this capacity
would normally be coupled to a generator of some
475 kVA rating. Were our generating set to be so
rated, we would expect a voltage dip of the order of
17 % when switching the 390.9 kV A required in
stage 3 of Table 5 E. This is still an unacceptable
performance as far as the process controller is
concerned. Furthermore, a 475 kV A machine is 1.6
times the capacity required by the total steady-state
We shall now analyse the kW requirements of the
4-stage load restoration pattern of Table 5 F to
determine if any reduction in peak demand results.
The analysis is presented in Table 5 H.
Clearly, this has had the desired effect of reducing
the peak starting requirement (from the 417.4 kW of
Table 5H Starting
and running
this stage
kW requirements
Starting kW
Running kW
Table 5 G to 301.6 kW). Taking generator efficiency
into account, this revised demand means that the
prime mover need only have a momentary capacity
of about 325 kW with a corresponding continuous
rating of about 295 kW. This size of prime mover
would normally be equipped with a generator of
some 340 kVA. This rating is only 15 % higher than
the total steady-state
load requirement
294.4 kVA and would be quite acceptable in our
case. Unfortunately, unless the selected generator
frame is capable of a maximum rating of 361.8 kVA
we shall have voltage dip problems - the dip will
exceed our permissible limit of 20 %.
To meet the conditions set by the consumer, it is
necessary to use a generator with a minimum capacity of 361.8 kVA; and a prime mover with a minimum continuous output of295 kW.
for 4-stage switching
Running kW
kW from
Peak k W
k W this
Total kW
Selection of power plant for load conditions
One option is to install a generating unit employing a 295 kW prime mover coupled to a generator,
whose frame is capable of a maximum rating of
362 kVA, but would be nameplate-rated
340kVA (to match the prime mover output). This
would not only result in a 'non-standard' generating
set, but it would also restrict the potential to accommodate non-essential, but desirable, loads, and any
future essential loads resulting from expansion of
the consumer's site operations.
Provided adequate funds are available, it would
be prudent to install, initially, a 'standard' unit using
a prime mover rated 20 % higher than that now
required. This would give a generator output of
about 360 kV A. Much will depend upon where this
rating falls within a machine manufacturer's range.
For example, it is possible that the 360kV A rating is
obtainable from a generator frame whose maximum
output is 420 kVA, giving increased potential for
coping with inrush kVA loads.
It should be appreciated that, unless some form of
auto-recloser is used on the contactor switchgear
controlling the lighting, heating, process controller
and d.o.!. motors, it is unlikely that the standby
generator will experience simultaneous switching of
the load elements at each stage. The load steps
calculated in this example, therefore, represent
'worst-case' emergency conditions.
These conclusions are, of course, based on the
assumption that an inverter will be installed to give a
dedicated supply to the process control equipment.
Again, capital permitting, it would be prudent to
incorporate a battery energy store in the inverter
equipment to provide the processor with an uninterruptible power supply. See Chapter 11.
Load considerations
Appendix 5.2 - Wave-form analysis of
a typical generator
The table below gives the wave-form analysis of one
of a range of generators manufactured by Newage
International of Stamford, England. The machine
has a 3-phase, 50Hz continuous rating of 160kW.
Results were obtained using a Camille-Bauer wave
Basic terminology
Governor selection
Fundamental operating principles
Typical governors
Recent trends in advanced control systems
6.1 Introduction
In this chapter we shall consider the principles
behind the many forms of speed governor used in
the control of diesel engines in generating plant.
First, the meanings of commonly used governing
terms will be explained and, then, the factors that
influence the selection of governors will be examined. This will be followed by a discussion of the
fundamental operating principles of the various governor types, before a more detailed study is made of
the constructions and actions of some commercial
The reason for the complex mechanical solutions
offered is that any engine system is inherently
unstable. This is due to the time lag between modification of the fuel position and a resulting new value
of torque produced at the crankshaft. For example,
if a 4-stroke engine runs at 1200 rpm it must
complete two revolutions before any change in
fuelling rate has an effect. This will take 60/1200 s or
50 ms. If the natural acceleration of the engine is
60 % per second (a fairly normal figure), the result
would be a 3 % excursion from set speed, and the
fuel pump/governor/engine
combination would
oscillate continuously to correct this condition - a
phenomenon known as hunting. Therefore, in addition to providing a precise regulation of speed, a
governor must include some compensating system to
control this instability.
6.2 Basic terminology
The definitions that follow are not necessarily precisely those given in any national standards, rather,
they are largely based on glossaries of terms contained in [1] to [7], and in the British Standard
Specification BS 5514 - Reciprocating internal combustion engines: Part 4. Speed governing. The intention here is to convey, in as simple a form as
possible, the meanings of those terms that are in
most common use.
Actuator: A device in a governing system which
provides the final control output, usually to the fuel
pump rack(s)
All-speed governor: Is one which may be adjusted to control an engine's speed at any selected
value within a predetermined range. It is also referred to as a variable speed governor. It differs
therefore from a single speed governor, which is
intended to control an engine required to operate at
one predetermined speed, and from a multiple speed
governor, which may be adjusted to control the
speed of a prime mover at two, or more, predetermined speeds.
Control action: This is usually defined under
three headings:
1. Proportional, which implies engine control proportional to the difference between actual speed
and the reference speed setting (this difference is
sometimes referred to as the 'offspeed' value).
Proportional control gives a continuous linear
relationship between the output from, and the
input to, the governor.
2. Integral control wherein the governor's output is
proportional to the time integral of the input, i.e.
the time of the offspeed.
3. Derivative control, where the output is proportional to the rate of change of the input or, in
other words, the rate of change of the offspeed.
Dead band or dead zone: Is the range of speed
(usually very small), expressed as a percentage of
the nominal rotational speed, through which the
governor will operate before it responds to an input
and makes a corrective movement of the fuel rack
actuator. The related term is sensitivity, which is
defined as the smallest speed change for which a
governor will affect a movement of the fuel controlling actuator.
Dead time: Is the time interval between initiation of an input change signal to a governor and the
response of the engine to this signal. The delay in
response is due to several factors, which are discussed in Section 6.3.
Governor: A device or mechanism that senses a
parameter and automatically controls and maintains
it at a desired level. In the context of a.c. generation, the regulated parameter is engine speed, and
the governor's purpose is to keep the speed constant
under all conditions of loading. It does this by
controlling the fuel consumption, so that a rise in
speed is arranged to reduce the fuel intake and a fall
in speed to increase it. Definitions have been given
for all-speed, single speed and multiple speed governors. We shall be considering other types, such as
load-sensing governors, later in the chapter.
Hunting: Also called 'cycling' is a rhythmic and
repeated variation of speed, sometimes due to overcontrol by the governor but not always necessarily
so. If the governor is responsible, manually and
momentarily 'blocking' the fuel control actuator will
eliminate the condition. If speed surging persists,
the governor is not the cause. Also see 'stability'
Hysteresis: The speed 'difference', at a given
load, between conditions of increasing and decreasing load. It is expressed as a percentage of rated
Isochronous speed governing: Implies constant
speed operation, regardless of load, within the rated
capacity of the engine controlled. It equates to zero
percentage speed regulation or zero permanent
speed droop (see later definitions).
Pilot valve: One which is used to regulate hydraulic pressure to the piston and cylinder of a
Engine governing
servomotor, which, in turn, may be employed to act
directly into the fuel control mechanism.
Recovery time: Also referred to as 'settling
time', is the time interval (expressed in seconds)
between the instant when the speed departs from its
steady-state band value after a load change, and the
instant of return to speed control within the specified steady-state band associated with the new load.
See Figure 6.1(a).
Response time: In governor parlance, this refers
to the time required for its output to affect a speed
change from its initial value to a large specified
percentage of the steady-state. See Figure 6.1(b).
Kates and Luck [3] use the term promptness to
convey this idea of speed of action by the governor.
Much will depend upon the power of the governor
relative to the work it must do. The greater the
power the shorter the response time. See also work
capacity, explained below. The governor system's
dead band will determine the time delay before
corrective action is taken. Its promptness will determine how rapidly the corrective action is completed.
Sensitivity: See 'dead band'.
Speed band: The width of the speed/time envelope (i.e. the difference between upper and lower
speed instantaneous values) expressed as a percentage of the declared speed. Thus the width of the
envelope under steady-state conditions would be
called the steady-state speed band. See Figure 6.1(c)
Speed drift: Gradual deviation from the desired
governed speed.
Speed droop: Is the percentage change in speed
corresponding to full travel of the governor output
shaft (i.e from its full speed to its zero speed
position). It is a characteristic of governor operation
that provides 'stability', and one that is necessary
when two or more engines fitted with speed-sensing
governors are operating in parallel, mechanically or
electrically, to share load proportionally [2]. See
also Chapter 8: Parallel operation of generator sets.
Speed regulation or steady-state speed regulation:
Is the speed increase from full to zero power output
of the engine, without adjusting the governor. It is
expressed as a percentage of the declared speed at
rated power.
Given the speed regulation, it is possible to compute
the change of speed that will occur for partial
changes of load. See Figure 6.2.
The terms speed droop and steady-state speed
regulation are often used synonymously, but they
have different connotations. The speed droop is a
measure of the speed change necessary to cause the
governor's control shaft (or servo) to travel through
its full working range. See Figure 6.3. It is simply the
Governor selection
Figure 6.3 The relationship
given governor
of 'droop'
to 'regulation'
for a
governor characteristic which requires a decrease in
speed to produce an increase in fuel supply to the
engine. Since an increase in fuel is required if the
engine is to carry more load, it follows that increased load means decreased speed. Speed droop is
a function of governor design and adjustment only
Regulation, on the other hand, is dependent not
only upon the speed droop setting, but also upon the
percentage of governor output shaft movement (or
servo stroke) which is required to move the fuel rack
between its no-load and rated-load positions. If the
governor, whose speed-droop characteristic is
shown in Figure 6.3, were so connected to the
engine's fuel rack that only 50 % of its servo travel
was required to move the rack between no-load and
rated-load, the regulation would be 2 % although
the speed droop is 4 % [5].
Unlike speed regulation, the engine's speed droop
may be either permanent or temporary. If the speed
droop is permanent, the governor's output shaft will
come to rest in a different position for each speed.
This would affect the speed regulation because the
engine's final speed (as we see from Figure 6.2) is
different for each level of loading. If the speed
droop is temporary, the governor's output shaft will
always come to rest at the same speed, so that the
engine's final steady speed remains constant regardless of load [3]. The temporary speed-droop feature
(then described as compensation) is normally designed into the servo movements of isochronous
hydraulic governors to give stability in operation.
Stability: The capacity of a governor to maintain
system speed equilibrium within specified limits
(with constant or varying loads) without hunting. It
represents the governor's ability to attain a pre-
Engine governing
selected running speed and to settle down quickly.
Every governor has a definite response time. One
with poor stability will oscillate above and below the
required speed (i.e. it will hunt) or take an excessive
time to cease oscillating .
Surging: This is the same as hunting but the
periodic speed swings are always of large magnitude.
Time constant: Gives a measure of the response
of the 'governing system' to a step change. It
determines the speed of the system reponse to the
change. The typical step response of Figure 6.1(b) is
reproduced in Figure 6.4 with the tangential speed
decay curve shown by a dotted line. The time
constant of this exponentially decaying characteristic is defined as the time taken (T seconds) to
complete 63.2 % of the total decay.
Transfer function: Is the mathematical (or graphical) expression for the relationship between the
input and the output of an electrical network, or of a
mechanical element, in any closed loop control
system. Expressed as the ratio of output/input for
the particular element, it may be considered as 'that
property of the circuit or element by which the input
must be multiplied to give the output' [6]. Or again,
it is defined as the ratio of the response produced by
a system to a disturbance [7]. In its simplest form,
the transfer function is a constant and is independent of time and, therefore, frequency. The 'form'
of the output is identical to that of the input. It may
be of greater or smaller amplitude depending on the
value of the constant [6].
Work capacity; Of a governor, or an associated
actuator, is a statement of the amount of work it can
do at its output lever or shaft. It is expressed in
joules (or in foot-pounds, inch-pounds). Selection
for a particular application will also be influenced by
available torque, given in Newton-metres (poundfeet) and by lever or shaft angular travel (in mechanical degrees).
6.3 Governor selection
Certain data are required if correct governor selection is to be made for the generator duty envisaged.
National standards exist in most countries for
engines applied to generators. Governor type will
depend upon:
1. the specified standard of governing; and
2. the inertia of the combined engine-generator
The applicable British Standard for the speed gov·
erning of reciprocating i.c. engines is BS 5514: Part
4. Its equivalent International Standard is ISO 3046:
Part 4.
BS 5514 defines five classes of governing accuracy
in its Table 2, and the parameters for the governing
systems are given in Table 3. Those applicable to
single speed operation, as on generator duty, are in
the Class A category, and are defined as follows:
Accuracy requirement
Highest requirements of governing
High requirements of governing accuracy
Normal requirements of governing
The British Standard does not give examples of
installations for which each class is suited. Also,
parameter values are not given in Table 3 for the
highest accuracy (Ao) class. Those 'which are important in the context of a particular application should
be subject to agreement between the manufacturer
and the customer. Remaining values should conform to the next lower class.' Acceptable limits for
the following parameters should be quoted by the
customer, in the first instance:
Speed droop.
Steady-state speed band.
Momentary overspeed and underspeed.
Recovery time.
The customer should also declare any other special
requirements such as parallel operation, etc. The
importance of the combined inertia of engine, flywheel and generator in achieving specified governing will be discussed later.
It can be assumed that, for those applications
where Class Ao governing or parallel operation is
specified, some form of external (remote-mounted)
governor, such as a mechanical-hydraulic or an
electrically powered unit, will be required. This is in
contrast to, say, the combined all-speed mechanical
governor and block type fuel injection pump, used
on high speed engines for Class Az governing (and,
possibly, Class Al also). See Section 6.5.1 later in
the chapter.
Fundamentaloperating principles
The moment of inertia (mk2) and the flywheel
effect (GD2) of the combined engine and generator
will have an important bearing on the choice of
governor type.
When a load change takes place, the difference
between the torque developed by the engine and
that required by the load goes into accelerating or
decelerating the engine. The amount of the acceleration or deceleration depends upon the inertia of the
rotating parts. The basic laws of dynamics apply.
Whenload is suddenly removed, the surplus power
causes the machine to speed up. This power is
converted into kinetic energy, which is directly
proportional to the inertia of the machinery and the
square of its rotating speed. It follows, therefore,
that a generating set with high inertia will experience less increase in speed than one with low
inertia. Conversely, as load is suddenly increased
beyond the power developed by the engine, the
rotating machinery gives out kinetic energy and
slows down. The greater 'the flywheel effect' the
easier it becomes to maintain steady running with
smaller transient speed deviations.
Flywheels of the highest practicable inertia should
therefore be used on generating sets where closetolerance governing is required. It may then be
possible to employ a relatively simple mechanical
governor (in conjunction with a high inertia flywheel) to meet a tight governing specification in an
economical manner without recourse to a more
sophisticated and expensive governing system.
A flywheel then has the effect of supplementing
the governor action by reducing engine speed deviations. The time required to correct the rate of fuel
injection on sudden application or removal of load is
a function of the governor. The governor's internal
time intervals are so small as to be meaningless, and
they are included in the total time constant of the
governor (see the definition in Section 6.2). There
will also be a further discrete interval in time before
the engine will respond to any change in the rate of
fuel injection. This dead time has at least two
components in the case of diesel engines. The first is
the time delay between the charging of a cylinder
with fuel and its conversion to torque. The second is
the interval then required for all the cylinders which
simultaneously contribute to the developed torque
to be firing at the new level. The latter time lag may
amount to one-third of an engine revolution [5]. The
extent of the dead time within the engine itself is
influenced by the number of cylinders, the engine
speed, and the type of fuel injection system used.
The instantaneous torque developed by an engine
depends on the position of each piston and its
connecting rod in relation to the crankshaft angular
position. It will vary about a 'mean value', causing
the crankshaft speed to fluctuate about a mean
speed. This change of angular velocity gives rise to
the effect known as cyclic irregularity. It is defined
as the ratio of the total variation in speed during one
engine cycle, to the mean speed when the engine is
running at any load up to, and including, rated load
and rated speed.
It is important that the cyclic irregularity is not
excessive on diesel generating plant, particularly on
those required to operate in parallel. This topic is
discussed in some detail in Chapter 8: Parallel
operation of a.c. generating sets. Suffice it to say, at
this point, that another purpose of a flywheel is to
smooth out any cyclic speed variations.
In summary: the provision of higher inertia will
allow the governor more time to take corrective
action on sudden load change and will permit steadier running with less speed deviation.
Adding inertia to a generator rotor, or increasing
the mass and diameter of an engine flywheel, costs
money. Also, it should be appreciated that when
large percentages of load are removed from an
engine and the fuel injection rate is correspondingly
reduced, the only retarding torque available for
deceleration comes from that provided by windage
and friction within the rotating machines themselves. If the inertia is high, it may take an appreciable (and unacceptable) time for the speed to return
to specified limits. Some compromise is, therefore,
always necessary between the limits specified for
speed deviation and the time of return to rated
speed. Choice of inertia may frequently be determined by which of these is more important in the
given application [5].
6.4 Fundamental operating principles
6.4.1 The mechanical governor
Every speed governor must include at least two
components: the first, to measure or sense the
speed; and the second, to translate this measurement into a movement of the governor output shaft
to control the device(s) that regulate the amount of
fuel injected into the engine cylinders. In the simplest governors both functions may be combined in
the one mechanism. Where any significanit force is
required it is necessary to employ a servomechanism
to amplify the low energy available from the speed
sensing elemc;:nt. 'Servo' is derived from the Latin
servus (a slave) thus expressing the idea that the
output or controlled quantity is made to slavishly
follow the input (or controlling) quantity [7].
The speed sensing element customarily applied to
and mechanical-hydraulic
Engine governing
governors is the centrifugal bal/head. The principle of
its operation is illustrated by Figures 6.5 and 6.6 [5].
In the mechanism of Figure 6.5 spherical weights
are fixed to the ends of arms pivoted near the axis of
rotation of the governor spindle. The flyweights are
able to move radially in a plane through the spindle's axis of rotation. Links are attached to the two
arms and to a collar which forms a sleeve mounted
on the governor spindle. When the weights are
moved outwards, the collar travels up the spindle.
The spindle is rotated by the engine, possibly
through a gear train. The spinning weights produce
a centrifugal force, always acting at right angles to
the axis of rotation. As speed increases, this force
increases. The weights move further apart and raise
the collar, which is connected, in the case of the
direct-acting mechanical governor, to the throttle.
(Note: This latter term is a convenient way of
describing the fuel-control mechanism, whatever
fuel injection system is used). The reverse action
occurs when the speed decreases.
Because the centrifugal force is at right angles to
the axis of rotation, it creates a torque about its
pivot (see Figure 6.6). This torque is equal to the
product of the force and the vertical distance (A) of
the ball below the pivot. If no other forces are
present, the torque is balanced, at equilibrium, by
an opposing gravity torque. The latter is the product
of the weight of the flyball and the horizontal
distance (B) to the pivot.
As speed increases, centrifugal force increases
and the flyball moves outward. This movement
reduces the torque arm A whilst increasing the arm
B associated with the gravity torque until a point of
equilibrium is reached when
There is a singular equilibrium position of weight
and collar for each speed of rotation. This relationship will not exist if friction is added to the system.
The torque equation on speed increase from an
equilibrium position then becomes:
centrifugal force torque
gravity torque
+ friction
before any further movement occurs.
Conversely, on a speed reduction the equation must
centrifugal force torque = gravity torque - friction
before any further movement is possible.
There is therefore a speed range within which no
corrective throttle motion (via the collar) is possible.
This is the dead band, defined in Section 6.2.
To obtain the greatest sensitivity from a governor
mechanism of this type, the ratio:
should be as large as possible.
One has two options in this regard:
1. To minimize both the friction and the magnitude
of the force required for throttle control. This
would entail careful construction to reduce friction and the use of a servomechanism to amplify
the low-power ballhead output.
2. To increase the force available from the ballhead. This may be achieved either by increasing
the size of the ballhead or by running a small
ballhead at increased speed.
6.4.2 The mechanical-hydraulic
Otherwise referred to as the 'servo governor', this
type still employs a ballhead as the speed sensing
element, but, in order to provide a higher power
level output to the engine's throttle control device,
use is made of a servomechanism. Vane type hydraulic servos and hydraulic motors have been used,
but the commonest form of servomechanism is the
hydraulic power piston. The fluid medium is usually
oil - supplied under pressure from a pump driven
from the controlled engine. Lubrication is therefore
Fundamentaloperating principles
inherent within the governing system. The possibilities of oil leaks, difficulties with flow at extremely
low temperatures, and the necessity to maintain a
clean oil supply may be considered disadvantages.
These are far outweighed by the advantages of
speed in operation, and in output power offered by
the servomotors. By selecting the appropriate piston
size and working oil pressure the power of the
governor (or its work capacity) can be customed to
suit the largest engines.
The servomotor most commonly applied is the reciprocating piston type. Three variants are used and
these are illustrated diagrammatically in Figure 6.7
The double-acting piston (Figure 6.7(a» requires
a pilot valve, which has two control lands simultaneously regulating oil flow to and from opposite
sides of the piston. The piston rod passes through a
seal in the body and is connected to the throttle
device. (Note: the land is defined as the enlarged
diameter of the value plunger and is used to guide its
movement or to control the flow of fluid through its
ports. )
Figure 6.7
Variants of reciprocating
piston servos (a)
The single acting, spring loaded piston (Figure
6.7(b» uses a single land pilot valve to feed pressurised oil to it. In one direction of piston travel the
hydraulic force overcomes spring and external load,
and, in the other direction, the spring discharges the
oil from the cylinder and overcomes the external
load imposed by the throttle controls.
In the differential piston, the piston rod (connected to the external load) is sized to reduce the
effective area on its side of the piston to about
one-half of that of the other side. It will be seen
from Figure 6.7(c) that oil is maintained at supply
pressure in the piston rod area of the cylinder, and
that oil flow in and out of the other side of the
cylinder is controlled by a single land pilot valve.
A ballhead configuration
Before we go on to consider elementary arrangements of hydraulic governors, we shall look at the
ballhead design configuration used in the governors
described in Figures 6.9 to 6.12. The weights (no
longer 'balls') are arranged to swing about pivots
located on opposite sides of a ballhead, which is
rotated by the engine through gears (see Figure 6.8).
The 'toes' on the ballarms, which are at right angles
to the flyweights, convert the inwards and outwards
movement of the flyweights (caused by changing
centrifugal forces) into an axial movement of the
speeder rod, through a thrust bearing. The centrifugal force is opposed and balanced by the force
exerted by a compressed speeder spring.
The speed adjusting device consists of a plug into
which the upper end of the speeder spring is fitted,
and which is adjusted up or down to change the
speeder spring force. Thus, raising the plug reduces
the spring force and allows a lower engine speed for
the same load condition. Conversely, lowering the
plug compresses the spring, increasing its force and
resulting in a higher engine speed for the same load.
The conical spring gives better speed control
Figure 6.8 A ballhead configuration
Governor Company)
Engine governing
through a wide range than an ordinary cylindrical
spring would.
The speed adjusting screw on governors fitted to
parallel running engine-generators may be operated
through a small reversible, fractional horsepower,
electric motor. The drive is through a friction clutch
permitting overtravel without damage. The motor is
controlled by speed raise and lower push buttons,
located within the generator switchboard. (See also
Chapter 8.)
opens the pilot valve in the opposite direction and
the throttle operates to decrease the fuel supply.
Once again, by the time engine speed reaches its
preset value, the throttle has overtravelled. The
engine underspeeds and the cycle repeats itself. In
other words, the throttle is continually moving in an
endeavour to take corrective action - it hunts. Stabilizing controls must be added to the system, in
order to ensure satisfactory operation.
The speed droop governor
The elementary hydraulic governor
We shall now consider the very elementary hydraulic governor illustrated diagrammatically in Figure
6.9 [5]. A ballhead of the type just described
operates in conjunction with a pilot valve to control
a reciprocating piston servo. For this description we
shall assume that an adequate oil supply is available.
The ballhead, as we have established, is driven at a
speed proportional to that of the controlled engine.
The piston servo is connected to the engine throttle
device. The ballhead and directly-connected pilot
valve have only one equiblibrium position. This
occurs when the valve is closed, and oil is neither
admitted to, nor discharged from, the servo cylinder. For a given setting of the speeder plug (not
shown at the top end of the speeder spring), the
ballhead and pilot valve will be in this position at
only one speed. The speed sensing element is,
therefore, inherently 'isochronous'.
The one big disadvantage with this rudimentary
system is its instability. Because of the inertia
masses of engine and generator, engine speed does
not instantly assume a value proportional to the fuel
rack position. This means that if the engine speed is
below the governor speed setting, the pilot valve
takes up a position that allows the servo piston to
move and 'open up' the throttle. Before the engine
speed can increase to the set level (so that the pilot
valve may centre and stop the servo), the fuel has
already been increased too much and the engine
continues to speed up past the setting point. This
The simplest method of securing stability in the
system just described is to provide speed droop
within the governor [2]. It should be appreciated
that the simple flyball mechanical governor (Figure
6.5) has inherent speed droop characteristics. This is
because the only way in which the throttle opening
can be increased, for instance, is by inwards movement of the flyweights, requiring a decrease in
Speed droop gives stability because the throttle
can only take up one position for anyone speed (see
Figure 6.3). This means that when a load change
calls for a speed change, the resulting governor
action ceases at that particular point which gives the
amount of fuel required for the new load. In this
way, unnecessary governor movement and any tendency to over-correction (hunting) is avoided. It is
of course important to ensure that the speed droop
is large enough to cover any time delay while the
engine itself is responding to the governor action. If
the droop is insufficient, some hunting may still
occur while the engine is returning to its steady
speed after the first momentary speed change [3].
Permanent speed droop can be incorporated into
the system illustrated in Figure 6.9 by employing a
mechanical interconnection between the power servo and the governor speeder spring. Then, as fuel is
increased the speed setting is decreased. As we have
seen (when discussing the speed adjusting device of
Figure 6.8), reducing the spring force reduces the
engine speed and, conversely, increasing the spring
force increases the engine speed. One such speed
droop device consists of a lever of suitable ratio
between servo rod and the top end of the speeder
spring (see Figure 6.10, and Figure 6.21 later in the
text) .
The equilibrium relationship between speed setting and servo position may then be represented by a
line which slopes (or 'droops') downward to indicate
decreased speed setting against movement of the
servo in the 'increase fuel' position [5]. (See Figure
Obtaining speed droop by action on the speeder
spring is not the only method used in hydraulic
governors. An alternative is to change the flyweight
position required to centre the pilot valve at its
equilibrium position. This effectively changes the
Fundamental operating principles
Figure 6.10
Principle of speed droop lever
speed of the flyweights for a given speeder spring
setting. One method uses a 'floating' lever connection between speeder rod, servo and pilot valve (see
Figure 6.11, and Figure 6.27 later in the text).
The compensated isochronous governor
While hydraulic governors of the type described
have several commendable features (accuracy and
sensitivity, and more power than mechanical governors of similar dimensions), they are not isochronous in operation and their speed droop setting is
not conveniently adjustable. Since it is sometimes
desirable to have an isolated engine-generator run
isochronously, it is necessary to introduce transient
or temporary speed droop (generally called compensation) to prevent any tendency of the governor to
over-correct. This may be accomplished in a number
of ways.
One system (see Figure 6.12) employing so-called
pressure compensation consists of a buffer piston,
two buffer springs, a needle valve, and a compensating land on the pilot valve plunger. It works on the
principle of applying pressure to the pilot valve
plunger, thereby adding to or subtracting from the
speeder spring force to effect a temporary change in
speed setting.
Pressurized oil actuating the servo piston is directed through the control port into a 'buffer' system
which consists of a piston centred between two
springs. Deflection of this piston produces a pressure differential across it. This, in turn, is transmitted
to the area above and below the compensating land
(or receiving piston) on the pilot valve plunger. The
needle valve permits equalization of the pressure
across this land and acts to restore the initial speed
As oil flows to the servo the buffer piston is
moved against the force of its centring spring. This
gives a higher pressure on the lower side of the
compensating land and produces an upward force on
the pilot valve. The effect is a decrease in the force,
which the flyweights must balance. The pilot valve
then centres before the original speed is regained
(i.e. at a lower speed) and provides speed droop.
The differential pressure disappears as the displaced
oil is permitted to leak across the needle valve. The
temporary droop signal is, therefore, dissipated.
The buffer piston returns to its equilibrium position
and the speed setting reverts to its original value.
The precise recentring of the buffer piston is not
essential for exact return of the system to normal
speed. Equalization of pressure is all that is required
for this to happen.
Engine governing
Applications summary
The basic arrangement of Figure 6.9 is not used in
generator applications because of its inherent instability. The speed droop types described under Figures 6.10 and 6.11 are extensively used on engine
driven generators and are suitable for parallel operating units. The compensated isochronous type of
Figure 6.12 has many applications. These include
single running standby generators feeding 'technical'
loads, where stability and constant frequency,
regardless of load, are primary factors. It is possible
to combine temporary and permanent droop in a
single governor and this is a useful feature where
parallel operation, requiring the proper division of
load, is contemplated. The degree of accuracy of
load sharing is improved by increasing droop. In
pressure compensated isochronous governors the
addition of droop provides only minimal improvement in stability. It should be appreciated that with
5 % droop, a speed setting error of 1 % between
generating sets will give a load imbalance of 20 %. If
speed droop is reduced to 2 %, the discrepancy
would be 50 %, and with only 1 % droop, the
imbalance could well be 100 %. (see Sub-section
8.6.4 of Chapter 8.) A 'compensated' isochronous
governor may be set for 4 % permanent droop to
give good load division, and with that amount of
transient droop which is required for stable governing [5]. In most proprietary hydraulic governors
speed droop adjustment is provided inside the governor, and is achieved by varying the ratio of the
interconnecting lever between the throttle and
speed setting elements.
6.4.3 The electric governor
In its very basic form (Figure 6.13) the electrically
powered governor is a direct alternative to the
simple mechanical governor. It comprises a speed
sensing device, an electronic speed controller, and
an actuator connected to the fuel control linkage.
Only the actuator and the speed sensor need be
installed on the engine. (Note: Some manufacturers offer integral injector pump mounted actuators which eliminate the need for an external fuel
system control linkage and actuator support brackets.) The basic system can be expanded for a variety
of control applications.
Speed sensing
It is possible to use the output power frequency as
the speed signal in generating sets provided it is
within the frequency range of the governor. Clearly
this is only feasible for single unit, non-parallel,
operation. Almost all proprietary electric governor
systems use a magnetic pick-up (Figure 6.14),
mounted in close proximity to a ferrous projec-
tion(s) driven by the engine, to detect prime mover
speed. The flywheel starter ring gear is normally
used because of ease of installation and the high
frequency signal obtained, and because it is
mounted on a nodal point of the shaft. See Figure
6.13. The magnetic pick-up (MPU) signals the
number of gear teeth which passes its tip, or pole
piece, per second - the signal being directly proportional to the engine's speed. The output from the
sensor is determined by the peripheral speed of the
teeth, their size, and their distance from the pickup's tip. It is important that the governor manufacturer's installation instructions are closely followed.
These would extend to the wire leads to be used.
These may need to be of the shielded (and/or twisted
pairs) type over their entire length up to the speed
control unit.
Speed control
Basic control units compare the actual electrically
measured speed from the sensor with the adjustable
and preset desired (or reference) speed. The resultant error signal is amplified and fed to the actuator
(Figure 6.15).
Failsafe circuitry in the controller senses MPU
frequency and operates to move the actuator into
the fuel shutoff (or minimum fuel) position.
Fuel control actuators
The actuator's purpose is to position the throttle
controls which adjust the fuel supply to the engine.
Actuators may be either electrohydraulic or electromechanical. Some electro hydraulic types may require an engine drive for an internal oil pump. Others
may be driven by an electric motor (or an oil motor)
if no drive pad is available on the engine.
Figure 6.16 [8] is the schematic diagram for a
Woodward type EG-3P electrohydraulic actuator.
Its essential element is a transducer which controls
oil flow to and from a power piston through the
action of a polarized solenoid. The position of the
terminal (or output) shaft is proportional to the
input current to this solenoid, which controls the
movement of a hydraulic pilot valve plunger. The
relationship between actuator position and input
voltage is linear.
Very briefly, the mode of operation is as follows:
1. The movement of two opposing pistons (the
loading and the power pistons) rotates the actuator terminal shaft to which the engine fuel
linkage is attached.
2. Pressure oil from the pump is supplied directly to
the underside of the loading piston and tends to
turn the terminal shaft in the 'decrease fuel'
Fundamentaloperating principles
Figure 6.13 Basic electrically powered governor system
(Courtesy: Woodward Governor Company)
3. The power piston can only move downwards
when the oil trapped beneath it escapes to the
sump. Oil flow to and from this piston is controlled by the pilot valve plunger.
4. The greater of two forces moves the pilot valve
plunger up or down. The first, a downward force,
is provided from a two-coil solenoid which is
energized by the output signal from the governor
speed controller. The pilot valve plunger is connected to a permanent magnet spring-suspended
in the solenoid field. The second force is the
resultant of (a) an upward force on the pilot valve
plunger due to the action of a centring spring
fitted above the solenoid core case, and (b) an
opposing downward force due to the restoring
spring. The magnitude of force (b) depends upon
the position of the restoring lever. It reduces as
the terminal shaft rotates in the 'increase fuel'
direction and increases with rotation in the
'decrease fuel' direction.
5. The resultant force from the combined output of
centring and restoring springs always acts to urge
Prime mover protection
Figure 10.7 The flameproof model of a combined oil level
regulator and safety switch (Courtesy: Frank W. Murphy
Mfr. Inc.)
oil at high-pressure between the pump and the
injector. It is subjected to severe vibration and
unless adequately supported and clamped along its
route length, is susceptible to failure. Such failure
may range from a minor leak to a complete pipe
fracture. If the consequent oil spray reaches the
exhaust manifold or a turbo-charger body a severe
fire hazard results.
In the early 1970s engine manufacturers W.H.
Allen designed and fitted fuel-pipe fracture protection equipment to engines installed in unattended
machinery spaces in ships. This is still a standard
feature of their S37-G heavy-fuel engines. Each
high-pressure fuel pipe (and its end connections) is
completely encased within a strong, sealed, flexible
sheath which is longitudinally compressible to facilitate access to the connections (see arrangement
diagram (a) of Figure 10.8). Leakage from any of
the high-pressure fuel pipes is contained within the
sheath and is drained to a common manifold from
where it is fed to a drain tank equipped with a level
switch. This arrangement considerably reduces the
fire risk by giving warning of any fault in the fuel
injection piping. For less stringent requirements the
seal at the fuel pump end of the injector pipe may be
omitted. Any fuel contained within the outer sheath
is then allowed to drain down over the fuel pump
into a gully in the engine housing, and is collected in
a drain tank [4].
Figure 10.8 Fuel pipe fracture protection equipment (a) Full
injection pipe sheath arrangement (b) Close up of cylinder
head showing fuel injection pump and injector with fuel
pipe fracture protection sheath and piping to drain tank
(Courtesy: NEI Allen Ltd)
Prime mover and generator protection
The Southampton-based company Giro Engineering Ltd markets a range of rigid sheathed fuel
pipe kits for some 30 manufacturers' engines. Each
of the Duoline fuel pipes consists of a high pressure
fuel injector pipe contained within a rigid, tubular
steel, outer sheath firmly locked onto special end
fittings. In the event of failure fuel is retained in the
sheath and piped away along the gap between the
concentric pipes to an alarm tank. Since each pipe is
only marginally bulkier than a single-pipe system
existing pipe routes and clamping points are usually
used. Special indicators may also be fitted to show
the exact location of failure. Some engine manufacturers provide the Duoline system as standard
equipment. Others offer it as an optional item.
10.2.5 Flow monitoring
On the larger water-cooled engines it is useful to
have some means of verifying that the coolant
circuits are operating satisfactorily. We have already
seen how use may be made of a pressure switch in
the engine's inlet to detect coolant circulation failure. Other devicei commonly applied are visual flow
indicators, flow alarm switches, and instruments
that combine both functions.
pOSItiOning for the most effective protection on
generating plant. The switch is 'armed' by pressing
the reset push button which moves the tripping latch
to a position where the holding power of the permanent magnet is sufficient to keep it in a 'latched'
position. The air-gap is factory preset to 0.01 inches
(0.25 mm) for minimal shock. Adjustments are
made by moving the armature linearly toward or
away from the holding magnet. This is done by using
the slotted sensitivity adjustment screw. A malfunction in the reciprocating or rotating parts of the
protected plant (see also generator rotor faults in
Section 10.3.2 of this chapter) may result in abnormal shocks or vibrations. Should this happen,
the sliding inertia mass in the instrument causes the
armature to separate from the adjustment rod and
break the magnetic coupling. The switch contacts
are then actuated.
The switch is adjusted so that its contacts are not
actuated during start-up and normal operating conditions. This is most easily done by turning the
adjusting screw in its 'more sensitive' direction until
the switch trips. The screw is then turned in the
opposite direction (in about 45" increments) until
the switch does not trip.
10.2.7 Engine shut down
10.2.6 Engine-generator
Detection of excessive vibration is particularly desirable in unattended or partially manned machinery
spaces. Switches sensitive to abnormal oscillation in
any plane of motion may be installed and used to
stop the prime mover when excessive vibration
occurs. The Murphy series VS-2 switch shown in
Figure 10.9 is an example of one such device.
The unit is rigidly mounted to the monitored
equipment. The inset diagram shows location and
Mention has already been made of the methods used
to effect shut down. We shall now consider some of
these in more detail.
Electrically operated devices
One of these is the solenoid-operated fuel shutoff
valve. A typical example is the Murphy series SV
valve illustrated in Figure 10.10, which operates on
12, 24 or 32 volt d.c. It should be placed in the fuel
Prime mover protection
Figure 10.10 Solenoid-operated fuel shut-off valve
(Courtesy: Frank W. Murphy Mfr. Inc.)
supply line as close as possible to the fuel injection
pump and after the final fuel filter. It may be
arranged for either 'energized-to-stop' or 'energized-to-run' configurations. In the latter arrangement activation of engine-mounted safety switches
would cause the supply to the solenoid to be interrupted - by contacts in the shut-down relay of the
protection circuit. The valve closes and cuts-off the
fuel to the engine. A check valve may be fitted in the
excess fuel line to the service tank to prevent the
injection pump siphoning fuel back through the line.
This ensures that the engine shuts down quickly.
Not all engine makers will permit shut down by
fuel cut-off. Where the method is sanctioned, use is
usually restricted to mechanically governed engines.
The following points should be noted [5]:
1. The valve should only be used as a means of
emergency shut down. If used as a regular means
of shut down, fuel aeration could lead to starting
problems ..
2. The engine is likely to run-on for a considerable
period after valve operation.
3. Care must be taken when fitting the valve to
ensure that contaminants are not introduced into
the fuel system.
This form of shut down must not be used on
hydraulically governed engines. Operation of the
valve will result in loss of governing and a dangerous
run-away condition. We have established (in Chapter 6) that external solenoid shut-down valves are
used for the smaller hydraulic governors. Their sole
purpose is to drain the governor's control oil into the
engine sump. They are not used as fuel shut-off
The preferred method of stopping mechanically
governed engines is to operate the stop lever on the
fuel injection pump. This is done by electrical,
hydraulic, or mechanical type actuators. The shutdown arrangement should be of the fail-safe type,
i.e. engine shut-down should take place automatically in the event of actuator power supply failure.
The usual arrangement is for the actuator to hold
the stop lever in the run position against a return
The term fail-safe can have different connotations. For example, when it is applied to emergency
generators in ships the vital consideration is not the
generator itself but the vessel it safeguards. In these
circumstances, failure of the control supply to a
shut-down device should not be allowed to jeopardize the ship's safety, especially when it is manoeuvring in close waters. Engine protection in such cases
is arranged to operate in an energized-to-shut-down
The normal shut-down device on mechanicallygoverned engines is the d.c. solenoid actuator. The
cut-away illustration in Figure 10.11 is of a unit
which has two working coils. Both are initially
energized to give the high pull-in force required.
The self-contained, heavy duty, double break switch
disconnects the lower resistance pull-in winding, on
completion of the plunger stroke. This leaves the
higher resistance, continuously rated hold-on coil in
circuit. This coil provides the necessary force to hold
the injection pump's stop mechanism in the run
position. Typical currents and forces for a 24 V d.c.
solenoid of this construction are:
• a pull-in current of 23 A;
• a hold-on current of 0.3 A; and
• pull and hold force ratings of 75 Nand
25 mm, respectively.
180N at
The upper illustration of Figure 10.11 shows the
application of this type of shut-down solenoid to a
proprietary engine.
It is important that any linkage used with dual-coil
solenoids is free of binding during the plunger's
movement. The plunger must be allowed to complete its travel into the solenoid body. This ensures
that the internal switch automatically disconnects
the low resistance (short-time-rated) pull-in coil.
Otherwise excessive heat is generated within the
solenoid body and this may cause irreparable damage within a very short time (over 30s). The solenoid's mounting must be sufficiently strong to withstand the pull forces and all likely vibrations and
shocks. Well designed systems will include an overload device (such as a miniature circuit breaker), to
protect the solenoid.
The auxiliary terminal shown in Figure 10.11 may
be used for a number of purposes, for example, to
Prime mover and generator protection
Non-electric systems
A typical example of a hydromechanical system is
that shown in Figure 1O.12(a). It uses engine lubricating oil to 'arm' a snap-action, spring-actuated
shut-down device. This is directly connected to the
fuel pump control lever or to an intake air shut-off
valve. The trip actuates on falling oil pressure. A
third option is a fuel shut-off valve, which gives
immediate shut down when installed in 'pressuretime' fuel injection systems of the type used on
Cummins engines. If the valve is fitted to engines
with conventional jerk pump fuel systems overspeed
sensing must not be used.
The handle of the 2800 safety device is connected
to the fuel pump control lever or to an air butterfly
in the intake air system. Before starting the engine
the 2800 arm must be cocked and latched in the
'fuel-on' position. When engine lubricating oil pressure rises to a preset level the latch drops out and
the device is 'armed'. When engine oil pressure falls
to the preset point the 2800 trips and stops the
The other protective sensors shown in Figure
1O.12(a) are essentially vent valves which drop oil
pressure at the 2800 when dangerous conditions
occur. They are arranged to vent oil back to the
engine sump when activated. The system is 'fail safe'
because breakage of an oil pressure line or excessive
oil leakage or a plugged orifice will cause engine
shut down.
The 4143 differential pressure valve is designed to
protect the engine against a blown radiator hose or
excessive cavitation in the jacket coolant system.
When coolant drops to dangerous levels most temperature sensors will not detect the condition until it
is too late. This is because the sensing element does
not respond in air or steam pockets as fast as it
would with the scrubbing action of jacket coolant
flow. The 4143 should be adjusted to vent oil in the
safety shut-down system at a pressure differential
slightly below that which occurs at low-idle engine
speed. There are right and wrong ways of piping the
sensing devices into the system, and the manufacturer's literature should always be consulted.
Another proprietary unit often used on the
smaller engines is the type OV equipment marketed
by Teddington Industrial Equipment (see Figure
10.12(b)). It incorporates a latch which operates
through linkwork to lock the injection pump's stop
lever in the 'run' position. The latch may be released
in anyone of three ways:
Figure 10.11 A dual-coil construction solenoid actuator
(Courtesy: Synchro-Start Products Inc.)
indicate that the plunger is properly seated after
pull-in or to give a sequencing signal to the engine's
control logic.
1. manually, by means of a push button on the unit;
2. by a pressure-sensitive capsule housed within the
unit and influenced by engine lubricating oil
3. by a temperature-sensitive capsule, also incorporated within the unit and influenced by coolant
Prime mover protection
Figure 10.12
Examples of non-electric shutdown systems
(a) A hydro-mechanical
safety shutdown system (Courtesy:
Amot Controls Ltd) (b) A mechanical
(Courtesy: Teddington Industrial Equipment Ltd).
Prime mover and generator protection
The capsules are set to release the latch at specific
pressures and temperatures. When the latch is released an external spring on the unit returns the
control lever to the 'stop' position and closes the fuel
rack. The cause of shutdown is indicated on one of
two dials mounted on the unit.
Mention has already been made of the use of an
intake air butterfly valve as an effective shut-down
device. An engine will stop if it does not receive
sufficient oxygen to burn the fuel. One of the worst
hazards on engines operating in areas where
inflammable gases are likely to be present (e.g. oil
refineries and off-shore oil rigs) is engine overspeeding due to ingestion of these gases. In such conditions, even when the diesel fuel supply has been
cut-off, the engine will continue to run and the
governor is unable to control the speed. The
inflammable gases passing along the air inlet manifolds act as a fuel. There is also the real danger of
the ambient gas cloud surrounding the engine being
ignited by the passage of flame from the engine
cylinders (through the air inlet) as the valves
'bounce' on an overspeeding engine. Another cause
of engine 'runaway' is overfilling of oil-bath intake
air cleaners.
A device that may be used to provide automatic
overspeed protection in such circumstances is the
shut-down valve marketed by Chalwyn Equipment.
The principle of its operation may be explained by
reference to Figure 10.13. The unit, which contains
a spring-loaded poppet valve, is fitted upstream of
the air inlet manifold and is actuated by the pressure
Figure 10.13 Air intake shut-off valve (Courtesy: Chalwyn
Equipment Ltd)
differential across it while the engine is running.
This gives a valve closing force which is restricted by
springs. The tension of these springs can be adjusted
to give closure only when engine speed exceeds the
desired level. Closure is instant and automatic.
The valve is held in its closed position by the
partial vacuum created in the air inlet manifold as
the engine stops. Once the engine has stopped the
vacuum is dissipated through normal engine clearances and the springs reset the valve automatically.
The engine is then ready for restarting.
An anti-flutter mechanism is fitted to the smaller
versions of the valve. This consists of a spiral
adaptor on which the valve moves. It ensures valve
stability and prevents premature closure due to
valve fluctuation when used on single and oddcylinder engines.
Dual-fuel and gas engines
Dual-fuel engines are stopped by cuting-off both gas
and pilot diesel fuel supplies. Spark ignition engines
are stopped by earthing or interrupting the ignition
circuit, and shutting-off the gas supply.
10.2.8 Some examples of monitoring and
annunciator equipment
Where multiple alarms and/or shut-down protection
are to be provided it is well worth considering the
use of proprietary systems. A 'buyer's guide' is quite
out of the question here and the examples that
follow are merely intended to illustrate what can be
The ideal installation would provide for the control and condition monitoring of the power plant
from one central location. Figure 10.14 illustrates
this point. The spacious control room accommodates suites of panels which allow control of
engines and generators. Facilities include synchronizing, remote switching, generator protection and
comprehensive alarm and fault indications.
Use has been made of multi-channel alarm annunciators for each of the four generators installed.
Annunciators of this type are available from a
number of manufacturers. Typical examples are
illustrated in Figure 10.15. That shown at (a) is a
Highland Electronics type MPAS 40 microprocessor
controlled system which can cater for up to 26
alarm channels. It is possible to include both prime
mover and generator alarm functions in the one
annunciator panel.
The Murphy Series ST alarm panel (illustration at
1O.15(b) is also designed to cater for all applications
where normally-open or normally-closed sensor
switch, or protection relay, contacts are used. Units
in the range can cater for 5 or 10 alarm points.
The Modex Automation engine control/alarm
unit illustrated in 10.15(c) is available in manual and
Protective relaying
Protective relaying
Figure 10.14
Engine instrument and alarm panels in the
control room of a Pacific island power station (Courtesy: NEI
Allen Ltd)
automatic start versions, catering for low oil pressure, high engine temperature, failed to start, charge
fail, and overspeed alarm conditions.
The Autostop Unit (Figure 1O.15(d» is one of
several modules in a diesel engine control system
marketed by Selco.
At the other end of the scale, the Pyropress
MU2000 series system provides a comprehensive
automatic condition monitoring and status indication facility. Continuous monitoring of up to 21
variables is possible. Inputs may be derived from
temperature, pressure, speed, current, power and
phase angle transducers. Each input channel can
have one or two independently adjustable setpoints
whose status is indicated by LEDs (light emitting
diodes). These, and the channel levels, may be
digitally displayed at any time. Analogue retransmission of any or all channels (to drive meters
and/or recorders at remote locations) is an optional
Each of the systems in the Pyropress range requires a common control module. This houses
power supplies, digital meter readout, and other
common functions (see the top left-hand section of
the unit illustrated in Figure 10.16). A module is
required for each input measured. Unused spaces
are blanked-off. Alarm and shut-down systems
require an additional common circuits module. A
group alarm module provides the user with additional relay outputs for any setpoint or group of
setpoints. A maximum of eight contact outputs is
available from each group alarm module.
10.3 Protective relaying
BS 142: Section 1.1 Glossary of protection relay
terms defines a protection relay as 'an electrical
relay designed to initiate disconnection of a part of
an electrical installation, and/or to operate a warning signal, in the case of a fault or other abnormal
condition in the installation'. Protective relaying is
concerned with the application of these relays to
elements of systems in such a manner as to effect
their prompt disconnection when undesirable conditions (such as short-circuits) occur in them, or when
they begin to operate in an abnormal manner, i.e.
when their dynamic functions exceed design parameters.
10.3.1 Philosophy of application
A protective relaying system should be capable of
detecting a network element failure or an abnormal
condition within the element with a speed commensurate with the particular power system's requirements, and then operate the minimum number of
circuit-breaking devices to isolate the defective element from the network. The advantages of speedy
isolation are:
1. Prevention, or at least minimization, of damage
to the isolated element to reduce the downtime
and expense related to its repair and so permit its
restoration to normal service in the shortest
possible time.
Prime mover and generator protection
Figure 10.15
Examples of monitoring and annunciator
equipments ((a) Courtesy: Highland Electronics Ltd (bl
Courtesy: Modex Automation (c) Courtesy: Frank W.
Murphy Mfr. (d) Courtesy: Selcol
Protective relaying
Wherever possible, protective relaying should
be so arranged that one zone's protection takes
over from another's, should the latter's devices
fail for any reason. This is known as overall
back-up protection and is discussed in more
detail later in this Section.
3. Speed, which as we have just seen is essential
when rel110vinga failed element from a network.
In this context, the characteristic embraces both
the initiating and the corrective action taken; that
is, the time for a breaker to operate and to isolate
the circuit must be included.
4. Reliability. The capability of a protective device
to perform repeatedly and accurately within its
design limits. It implies certainty of operation
and that all extraneous causes will not jeopardize
this assurance.
Figure 10.16 Plant condition-monitoring
The Pyropress Engineering Co. Ltd)
system (Courtesy:
2. Minimizing interference with normal operation
of the overall power system or of any dedicated
end-process that may be associated with the loss
of the isolated element.
This latter aspect of protective relaying is of special
significance where reliability of service is of paramount concern to power system users. The speed
with which protective devices operate therefore has
a direct bearing on continuity of service to the
otherwise healthy parts of the power system. Rapid
fault-clearing will keep generators and motors stable
and help them 'ride through' any momentary system
disturbances [6]. During a short-circuit paralleled
generators will tend to lose synchronism if a fault is
allowed to persist.
The essential attributes of any well-designed protective relaying scheme must therefore be:
1. Sensitivity.
The ability to detect and initiate
operation against the minimal variation in a
dynamic function.
2. Selectivity, which relates to a relay's ability to
recognize a fault condition in the zone or in the
item of plant which it is intended to protect, and
to operate only into those zones, and in that
sequence, which the protective system design
envisaged. It should be capable of ignoring faults
arising in other zones. Some devices are inherently selective and others are not. Selectivity is
achieved [7] by:
• comparison of measured values on both sides
of a protected item;
• current and impedance grading; and
• time grading.
More of this later, when we discuss particular
10.3.2 Relay types
These fall into two categories:
1. the electromechanical; and
2. the solid-state (electronic) type, using either
digital logic or microprocessors.
Electromechanical relays take many forms, of which
the most common are the attraction and the induction types. Attraction type units work on the solenoid principle with a plunger being drawn into a
magnetic coil. They work satisfactorily on both d.c.
and a.c. Therein lies one of their limitations when
used for overcurrent protection. They act so fast
that they trip on the d.c. offset or transient peak
rather than on the actual setting which is much
lower. This characteristic, known as 'transient overreach', should be taken into consideration when
setting relays of this type [8].
The more commonly used relay is the induction
disc type which is of similar construction to an
integrating a.c. energy meter. It works on the same
principle as an induction motor. Magnetic fluxes
cause eddy currents to flow in a metallic, nonmagnetic rotating disc. The torque so produced
causes the disc to turn - in an identical principle to
that used in a conventional domestic meter. After a
predetermined angle of rotation the disc is arranged
to trigger a trip mechanism. Variations of current
and/or voltage signals, and the use of varied flux
distribution patterns and damping arrangements,
offer many permutations in operational characteristics. It hasn't the application limitations of the
attracted-armature relay and its versatility accounts
for its wider use in relay schemes.
The advent of semi-conductor devices such as the
transistor and the thyristor resulted in the development of solid-state relays. They offer alternatives to
electromechanical relays for most relaying applications. Not only are they physically smaller than their
electromechanical counterparts but they are also
Prime mover and generator protection
faster-acting and give improved performance. They
may be used in conjunction with electromechanical
devices provided the design of such composite
schemes makes due allowance for their disparate
time-delays. A further advantage of static relays is
the lower burden imposed on current transformers typically, 0.5 VA at setting current, for an overcurrent relay, compared with the 3 VA of the induction
disc types. Because of their resilience and very low
mass they are also able to withstand far greater
levels of mechanical shock and vibration - a very
considerable advantage in respect of transportation
and site erection.
Because they are static devices (apart from the
final output elements, which are electromechanical
relays) they require less maintenance. However,
when considered in the context of diesel generating
plant, fault diagnosis and repairs are probably
beyond the capability of on-site maintenance personnel, particularly in the least-developed countries
where the technology associated with electromechanical relays is more likely to be understood by the
competent electrical technician.
On the debit side static relay components are
prone to damage in high-humidity environments
unless they are encapsulated. Also, transistors and
thyristors are sensitive to transient voltages (arising
from switching operations) which may be injected
into the protection equipment either directly,
through metering transformers, or by electric or
magnetic field induction. This is also a significant
consideration where a communications link is
employed in a protective scheme. Any electrical
noise and extraneous signals are a potential hazard
to the security and correct operation of the system
Protection relays usually operate directly into the
trip circuits of circuit-breakers. Use is often made of
auxiliary or intermediate electromechanical relays
to provide higher switching power on the tripping
contacts or to multiply the number of circuits affected by the operating signal.
The design of static protection initially developed
as single or dual quantity measuring relays which
could be used as direct substitutes for electromechanical units. The earliest relays employed semiconductor devices such as silicon planar transistors.
The modem trend is to use digital gate and logic
techniques, custom-built medium scaled integrating
(MSI) digital circuits, and memories and microprocessors, to give widely programmable multi-function
relays that meet the requirements of complex power
systems [10]. Examples of proprietary relays are
given in Section 10.5.
Analogue input signals from measurable power
system quantities (such as current, voltage, phase
angle and power) are converted to digital form and
are measured and compared with reference 'setting'
values. The resultant signals are processed either by
digital logic or by a microprocessor (see the simplified block schematic of Figure 10.17). Discrete
logic can perform a number of relatively simple
operations in parallel. A microprocessor can handle
highly complex logic tasks but is not ideally suited to
parallel operation [10].
10.3.3 Transformers in protective relaying
Current transformers (c.t.s) and voltage or potential
transformers (v.t.s) are combined with relays and
circuit-breakers to provide the blocks upon which a
protection scheme is structured. Their purpose is to
Protective relaying
isolate the protection circuits from the higher voltages of the main power networks and to change the
network working currents and voltages to values
suited to the protection relays. Primary and secondary windings are electrically separated and the only
coupling between them is that of magnetic flux.
We can only deal here with those features and
that terminology which are of significance in our
present discussion ..
Current transformers
These may take several forms. Those of direct
interest to us are of the air-cooled bar primary and
the wound primary type. The latter have particular
application to lower primary currents. As the
description implies bar primary c.t.s use the cable or
busbar of the power network circuit as the primary
winding. The secondary winding is completely insulated and is permanently assembled on a core of
rectangular or circular shape. This format is also
referred to as a window-type or a doughnut C.t. in
American practice where a bar-type or in-line transformer is taken to be one which has a primary bar
conductor permanently mounted in the core window, and to which the network conductor or busbar
is connected [6, 8].
The wound primary type c.t. has primary and
secondary windings insulated from each other and
assembled on the one core. By definition they are
classed as high reactance. They are less accurate
than the bar primary type C.t.S and are not allowed
for high accuracy (Class X) duty.
British Standard BS 3938: Current transformers,
which is related to IEC 185, specifies c.t.s for use
with electrical measuring instruments and electrical
protective devices and gives guidance on their application and use. The standard defines classes of
accuracy for protective C.t.S. Each class is associated
with a limit of error. In all practical transformers
their current ratios differ from their theoretical
(turns) ratios and their phase angle (i.e the angle
beween primary and reversed secondary-current
phasors) diverges from the theoretical 180 due to
the iron losses within them. These errors are referred to as the ratio and phase angle errors, respectively. For Class 5P and lOP c.t.s the limit of error is
5 % and 10 %, and takes account of both phase
angle and ratio error. The 'P' refers to 'protection'
as opposed to 'measuring'. 5P is employed for
sensitive protection (such as differential schemes)
and also where protection co-ordination requires
relays to be set close to each other. lOP is used
where a 10% error is acceptable. Limits for Class X
devices are not explicitly stated [11].
Two other terms commonly associated with protective transformers are burden and output. The
burden is the load connected to the secondary
winding and is expressed as ohmic impedance at a
certain power factor. It should be noted that this
power factor is not the power factor of the system
network being protected. Rated burden is thus the
impedance of the secondary circuit when carrying
the rated current at a stated power factor. The
output, on the other hand is the J2Z product expressed in volt-amperes (VA).
The smallest operational errors are to be expected
when transformers are loaded to between 50 and
70 % of their rated burden. Rating plate burdens
may be exceeded in exceptional cases (e.g. 125 %
indefinitely and 150 % for about 15 minutes) without damage to the transformer, but accuracy will be
affected when this happens. The error is doubled
when the load is doubled.
The accuracy classification should be just high
enough to provide an adequate supply to the connected load. The higher the class, the greater the
cost of the transformers. Where there are many
circuits on one project the cost of over-specifying
may become significant. The main consideration for
LV and HV switchgear is space, which is always
limited. Accommodation difficulties can result if
c.t.s are over-specified.
Preferred secondary current ratings are 5 A, 2 A
and 1 A. Where long circuit leads cannot be avoided
voltage drop considerations will dictate the use of
1A rated secondaries.
A current transformer operates as a lowimpedance device and its secondary circuit should
always be closed through series-connected relays
and/or instruments. If the load circuit is opened
high-voltage peaks can occur in the secondary winding. This will put insulation at risk and overheat the
transformer core. In theory, a short-circuiting switch
should be provided for the winding. It should be
arranged to be closed when no low-impedance loads
are connected in the circuit. In practice, it is never
fitted since any withdrawable component (such as a
protection relay fitted in a drawout case) has automatic shorting at the case terminals, and on hardwired circuits it is assumed that wires will not be
disconneceted deliberately while on-load.
While BS 3938 offers guidance on the use of c.t.s
for the dual purpose of measurement and protection, it is extremely difficult to produce a C.t.
capable of efficient use for both purposes. It is only
in those instances where the simplest forms of
protection and instrumentation are used (e.g. a
direct-acting overload device in series with an ammeter) that the dual-purpose role can really be
justified [11].
Voltage transformers
These are made for specific primary voltages. For
voltages up to 15kV they are usually of the cast-inresin, dry type.
Prime mover and generator protection
British Standard BS 3941: Voltage transformers
(IEC 186) defines performance requirements and
gives guidance on applications. Preferred secondary
voltages range from 1l0/v'"3, for earth-connected
v.1.s, to 220 V/phase. Operation up to 120% rated
voltage is possible while still maintaining accuracy.
Above this level overheating could occur. As in the
case of c.1.s, the permissible error limits determine
the accuracy classification. Voltage transformers for
protective service are classified as either 3P or 6P,
having voltage ratio errors limits of ±3 % and
±6 %, respectively.
Practice differs on the use of fuses in protective
V.t. circuits. Where they are employed on the primary side of transformers they are only intended to
clear a potential primary winding short-circuit
caused by insulation failure. They will not protect
against overload on the secondary side. Even a short
in the secondary will not give sufficient primary
current to blow the fuses. Low voltage fuses should
be fitted in all secondary circuits which are not
Readers will find Section 15 in [6]; Section 17 in
Item [3] of the Bibliography; and Section 4 in Item
[4] of the Bibliography of particular value in obtaining an understanding of American practice on
appropriate American Standard for instrument transformers is ANSI/IEEE C57.13 - 1987.
10.3.4 Co-ordination
We have discussed in Chapter 9 the need for incorporating selective tripping (discrimination) in distribution networks using circuit-breakers and fuses.
The aim must be to ensure that only the protective
device before the network fault trips. Branch circuits
of the same or higher level should not be affected.
This is achieved by arranging for devices nearer the
power source to have progressively higher trip current settings and longer fault-clearance times. The
same principle applies to relay protective schemes.
The time-current characteristics of the protective
devices to be co-ordinated should be plotted and
compared on a single sheet of log-log graph paper.
This systematic approach will help in the choice of
the right devices to achieve the desired selectivity.
Computers are increasingly used in the analysis of
power systems. Software for overcurrent grading
and the programming of relay curve plotting is
becoming more widely available for unit protection
schemes [12].
Co-ordination margins between breakers and
fuses have already been dealt with in Chapter 9.
When overcurrent relays are used, either with other
relays or with breakers and/or fuses, clear timemargins should be provided between all protective
devices in a co-ordinated system. Breakers and fuses
have band characteristics to allow for manufacturing
tolerances whereas overcurrent relays have line characteristics. When selecting for co-ordination it is
therefore necessary to ensure that the correct edge
of a band characteristic is used. If a relay is downstream of a banded device, the lower limit of the
band is used. The converse applies for relays upstream and nearer the power source.
The drawback with fuses, miniature circuit
breakers, and some moulded-case breakers is that
their short-circuit trip characteristics are not adjustable. This can be a serious drawback on installations
where prospective short-circuit fault levels are likely
to change. For example, where power is provided by
mutiple generators the value of fault current will
depend upon the number of generators in service at
the time the fault occurs. Ideal selectivity is only
obtained by using protective devices that give adjustable sensitivity or variable operating times or,
better still, a combination of both.
10.3.5 Overcurrent relays
The overcurrent relay is by far the simplest type of
functional protective device. Its purpose, as its
description implies, is to operate when more than a
predetermined amount of current flows in the portion of the power network which it protects. There
are two basic types:
• instantaneous; and
• time-delayed.
Theoretically, the instantaneous relay is designed to
operate with no time delay, immediately measured
current exceeds the relay's current setting. In practice, operating times may vary from about 1 cycle to
about 10 cycles of power frequency, depending on
the relay construction, i.e. whether it is of the
solid-state or the electromechanical type.
Where overcurrent relays are used in several
series steps it is necessary to maintain a time margin
between them. A margin of 0.25 to O.4s, at maximum fault current is typical. This allows for breaker
operating time, relay overtravel, variations in the
manufacturing tolerances of c. 1.s, and a safety margin [6]. See Figure 10.18 for a network fed from one
end. Note that the final-circuit relays are set for
instantaneous operation. Stouppe [6] suggests that
all overcurrent relays other than those applied to
network incomers should incorporate instantaneous
attachments and time-delay elements. This will help
achieve overall selectivity where several in-series
protection steps are used and still maintain reasonable operating times at the circuit-interrupting
devices nearest the power source.
Before going on to discuss the various timedelayed characteristics we shall consider the setting
adjustments provided on the relays. The first is a
current (or plug) setting. The term plug derives from
the fact that on an electromechanical relay the
Protective relaying
Prime mover and generator protection
Protective relaying
For time setting of definite time
relays, the time adjuurnenl
is calibrated
in seconds
- - - - -.-
I nverse
time relay
= Time multiplier
Figure 10.22
Comparative characteristics of definite time
relay and inverse definite minimum time relay (Courtesy:
GEC Alsthom Measurements Ltd)
more sophisticated designs of magnetic and solidstate direct-acting devices. While they may not quite
match-up to the accuracy and close calibration tolerance of the inverse-time relay with definite minimum time (IDMTL), they provide similar characteristics and allow a considerable measure of coordination and discrimination. Their use has been
largely linked with the modern low-voltage, airbreak circuit-breaker [11].
Figure 1O.24(a)) to match the generator's thermal
characteristic. Under fault conditions, when the
busbar voltage falls, the relay operates on a standard
IDMT characteristic, permitting it to be time-graded
with other IDMTL relays in the system.
An alternative technique is employed on voltage
restrained overcurrent protection, where the IDMT
characteristic of the relay is modified continuously
according to the voltage at the machine's terminals.
See Figure 1O.24(b).
Voltage controlled relays
Unfortunately, one of the difficult areas of overcurrent relay application is in generator protection.
This is because of the tendency of a generator's
sustained fault current to be low if its automatic
regulator is out of service and if no current-forcing
circuitry is employed (see Chapter 4). Also, under
fault conditions busbar voltage falls to a level lower
than that obtaining on normal overload.
Voltage controlled overcurrent relays which incorporate an instantaneous undervoltage unit are designed to cater for these conditions. They do this by
employing two operating characteristics, either of
which may be 'selected' by the undervoltage relay
which monitors the bus bar voltage. In overload
conditions when busbar voltage is near normal the
relay operates on a long IDMT characteristic (see
Protection zoning
On the larger power systems each item of equipment
forms a zone. These zones are arranged so that they
overlap as shown in Figure 10.25. Each is protected
by what is termed a unit system of protection, which
operates only on the occurrence of a fault within its
own zone. It must remain stable under conditions
arising from through-faults.
Non-unit systems (e.g. those where overcurrent
relays are used in a time-graded discrimination
scheme) will respond to faults beyond the actual
zone of protection. Normally, relay(s) in the faulty
zone will clear the fault before those in other zones
operate. Should the relay(s) in the faulty zone fail to
operate, those in the non-unit-system zones will.
This is referred to as overall back-up protection [14].
Prime mover and generator protection
Although primary protection is mainly concerned
with complete or partial short-circuits (between
phases and to earth) it also covers abnormal conditions such as persistent overload on thermally rated
plant, rotating
machines, transformers,
reactors. The response in such applications need not
be as fast as that required to remove a short-circuit.
Nevertheless, in most applications the fastest
possible primary protection is beneficial.
Standby relaying is usually provided for added
protection against short-circuits only since they are
the commonest cause of failure in power systems.
Indeed, experience indicates that standby protection
for other than short circuits is not economically
justifiable [15].
Readers will also encounter the terms restricted
and unrestricted zones. These are defined later in
sub-section 10.4.1.
10.4 Generator protection
We shall now examine those fault conditions that
can occur on a generator in service, and discuss the
relay protection equipment that may be applied in
each case.
Figure 10.23 Operation time characteristics of a
time-delayed overcurrent element in a static modular relay
(Courtesy: GECAlsthom Ltd)
Within the zones themselves two categories of
protection may be applied:
1. Primary protection, which is the 'first line of
defence', and operates to remove only the
faulted element from the total network.
2. Standby or 'last resort' protection, which uses
relays to isolate network elements by either
operating the same circuit-breakers as the primary protection or by tripping breakers in adjacent
zones to disconnect a greater portion of the
power system.
It will be noted from Figure 10.25 that each power
network item (generator, generator-transformer,
transmission feeders, etc.) is interconnected to the
network through circuit-breakers. Because the
zones overlap any failure in a region where overlap
occurs will result in the operation of more breakers
than the minimum necessary to clear the fault. If
there was no overlap, failure in the regions between
zones would not be covered and no breakers would
trip. Overlap is the lesser of two evils [15]. In
practice, the extent of overlap is limited so that the
operation of too many breakers is avoided.
Stator winding faults
Rotor faults
Over- and under-voltage
Over- and under-frequency
Unbalanced loading
Loss of excitation
Failure of prime mover
Loss of synchronism
Before we do this it is as well to mention, for the
benefit of a broad-based readership, that the
diagrams which follow will not use the device function numbering system commonly applied to protection engineering schemes. The reader needs to be
aware of the system, however, and an abridged list
of those functional numbers which may relate to
engine-generator protection schemes is given in
Appendix D.
Device numbers indicate the operational function
of the device. They had their origin in American
practice, where they are mainly used in the context
of control equipment associated with switchgear.
They are also applied to other electrical diagrams
and specifications. BS 3939: Graphical symbols for
electrical and electronics diagrams - Guiding principles has adopted the numbering scheme originally
introduced in America by the National Electrical
Manufacturers Association (NEMA) in its standard
SG-5 in 1959 (Power Switchgear Assemblies Part 8)
and later issued by the American National Standards Institute (ANSI) in their document ANSI
Generator protection
Figure 10.24 Typical characteristics for voltage controlled
and voltage restrained overcurrent
relays (a) Voltage
controlled relay - set at 100% (5A), with time multiplier
setting (TMS) of 1.0. (b) Operating characteristic of voltage
restrained relay, with TMS = 1.0 (Courtesy: GEC Alsthom
Prime mover and generator protection
C37.2 -1979, Electrical Power System Device Function Numbers. There are minor differences between
the British and American documents in certain
functional titles and definitions but in each case the
'intent' of the device is the same.
10.4.1 Stator winding faults
Any fault caused by insulation failure within the
windings of a machine will only exist within the
machine itself. Opening the main output circuitbreaker will not protect the windings. Of course it is
important that any breaker between generator output terminals and system busbars is opened in order
to prevent other generating plant in the network
feeding into the faulty machine. If the generator's
main field continues to be energized the e.m.f.
generated internally will maintain the fault current.
One of the first actions must therefore be to quickly
remove the excitation supply to the main field. This
is a relatively straightforward
procedure on
machines employing d.c. exciters, or on certain
forms of static excitation where the main field is fed
through sliprings on the rotor shaft (see Chapter 3
Figure 3.5). The rate at which the fault current is
reduced is then dictated by the speed with which the
magnetizing flux decays within the machine.
Where 'brushless' excitation is employed one can
only act to remove the a.v.r. fed supply to the a.c.
exciter's field winding. The aggregrate timeconstants of the excitation system in such cases
means that the flux decay time is longer and,
consequently, so also is that of the fault current
decay. It is possible on machines of the type
described in Figure 3.34 of Chapter 3, which have
'buck and boost' a.C. exciter field windings, to use
the regulation field winding (that providing the
subtractive or bucking effect) to help speed up the
process but its influence can only be marginal.
Stator winding faults may develop through failure
of insulation between turns of a winding, between
phases, or between phases and earth. Interturn
faults are difficult to detect and protect against. Roe
[8] describes a scheme that may be employed where
generators use a parallel-circuit winding (i.e each
phase has two windings in parallel). With this form
of construction a 'jumper' connects the parallel
windings together at the halfway point. Its purpose
is to carry the small circulating current that flows
due to the inequality that exists between the windings. The method described detects any increase in
the circulating current caused by short-circuits between turns on one of the paralleled windings of a
phase. Because the circulating currents are only of
the order of 0.005 times rated current, low-burden
extremely inverse overcurrent relays with specially
designed (split-phase) c.t.s are required. Close liaison is necessary with the generator manufacturer and
particular attention must be paid to winding connec-
tions and external leads. The engineering effort and
expense involved in such measures will rarely be
justified, especially when it is realized that most
interturn faults will quickly develop into earth-faults
and may be cleared by the stator earth-fault protection. There remains the possibility, however, that
extensive damage could arise if the interturn fault
occurs at the winding ends before it develops into
one detectable by other protection [16].
While phase-to-phase faults usually evolve from
interturn faults, they may occur without any such
preliminary. Once breakdown in interphase insulation begins it inevitably leads to an immediate
earth-fault. The damage caused to the stator by the
short-circuit current can be considerable. It is very
unusual to have a phase-to-phase fault clear of
Forms of earth-fault protection
The basic forms of earth leakage or earth-fault
protection using single relays are shown in Figures
10.26, 10.27 and 10.28. Diagram (a) of Figure 10.26
illustrates the use of a core-balance transformer
(c.b.t.). The three phase conductors are passed
through the ring core of the transformer which
carries a detector winding (or sensing coil). Under
balanced conditions the currents in the three phases
vectorially sum to zero. Any out-of-balance or residual current resulting from the earthing of a line
conductor sets up a flux in the ring core. This flux
links with the detector winding and induces a voltage within it which energizes a fault relay preset to
pick-up at a certain level. In another form of c.b.t.
the ring core is provided with three primary windings to which the line conductors are connected.
Diagrams (b) and (c) illustrate alternative methods of applying the residual current detection
principle in earth-fault protection. They use individual c.t.s in 3-wire and 4-wire circuits, respectively.
Current balance under normal circuit conditions is
between the secondary currents in the c.t.s. In arrangement (a) it is the fluxes in the ring core of the
c.b.t. that are balanced.
Diagram (d) illustrates how a single relay may be
used for earth-fault detection. The relay is operated
from a C.t. inserted in the neutral-earth connection
of the generator. It is important to realize that all
these forms of protection do not protect against
open circuit in a line, i.e. single-phasing [11].
Neutral earthing
The methods most likely to be used in the installations under discussion are:
1. Solid earthing, which is normal British practice
for low voltage systems below 600 V.
Generator protection
Prime mover and generator protection
Figure 10.28 The principle of restricted earth-fault
protection (a) Fault within the zone (b) Fault outside the
2. Low-resistance earthing for systems between
1 kV and 15 kV, employing either metallic grid or
liquid type earthing resistors. In European practice reactors of corresponding ohmic value are
often employed.
The intention in both methods is to provide a
solidly-earthed system. The purpose of the low
resistance in method 2 is to limit the earth fault
current either:
1. to full load current, for convenience of c.t. ratios
and co-ordination of protection; or
2. to approximately 100 A - being the minimum
order of current that will still allow discrimination using core balance c. t.s to achieve the required sensitivity; or
3. to approximately 5 A - being the minimum that
can still be detectable although not allowing
discrimination between zones.
Occasionally, HV systems involving generators are
solidly earthed. The main reasons for earthing low
voltage systems are:
• to stabilize the voltage of the system with respect
to earth;
• to ensure that the voltage between any phase and
earth does not normally exceed the phase voltage
of the system;
• to anchor the neutral point of the system so that
its voltage potential does not fluctuate;
• to allow the system protection to operate in case
of fault between any phase and earth;
• to decrease hazard to human life.
Earth resistances of less than 1 0 may be obtained in
favourable soil conditions (using earth plates, mats
or driven electrodes in parallel arrangements). BS
Code of Practice CP 1013 (Earthing) recommends a
value of 10. Even in relatively high-resistivity soils
it is possible to achieve values of the order of 2 to
In this discussion we are concerned only with the
engineering of protective systems. The objective is
to insert sufficient resistance in the neutral earth
path to limit fault currents to levels that minimize
damage to plant while ensuring that current magnitudes are adequate for detection by, and operation
of, protective relaying; but these are not the sole
Where on-site generators are employed the power
source's earthing impedance is just one element of
the earth fault loop for final circuits at the most
distant points of the consumer's installation. The
lEE Wiring Regulations give values of the maximum permissible earth-fault loop impedance (Zs)
for all the conventional types of protective devices
used in such final circuits. Due regard should be
paid to the needs of the installation, as a whole
before finalizing the power source's earthing arrangements.
Readers will encounter references to other forms
of earthing such as high-resistance earthing, and
earthing through arc suppression coils and through
voltage transformers. These are special methods
which are not associated with the size of generator
or the operating voltage levels that concern us.
Briefly, they may be considered to give an insulated
neutral and offer the advantage that a system may
continue in service after a first earth-fault occurs.
Locating such a fault can be difficult. If a second
earth-fault (i.e. on another phase) occurs, one has a
phase-phase-earth fault. This can cause considerable
damage if one of the earth faults is not cleared
Choice of earthing method is dictated by the
network's operational requirements such as those of
power and lighting loads, of continuity of service,
and of safety and cost. International marine regulations, for example, permit the use of both isolated
and earthed neutral systems for all vessels except
tankers. Earthed systems are forbidden in tankers.
Choice, however, is almost inevitably for isolated
systems - in order to avoid loss of vital services
should a single earth-fault occur [17].
Generator earth-fault protection
When earth-fault protection is applied to a generator circuit it should be of the type known as
restricted earth-fault protection. The term implies
that there is full discrimination between earth-faults
external to the protected zone and those that occur
within the zone. The alternative configuration
known as unrestricted protection implies that no
such discrimination is given. Any earth-faults
occurring well outside the immediate circuit to
which the protection is applied may cause unnecessary opening of generator circuit-breakers. All of the
schemes illustrated in Figure 10.26 are classified in
the unrestricted protection category.
Figure 10.27 compares basic forms of both types
of protection applied to 4-wire generator systems.
Diagram (a) shows the simplest arrangement and
covers all zones from the generator windings to the
final circuits in the load network. Regardless of
where the earth-fault occurs, the return current path
to the generator neutral must always pass through
the c.t. and therefore operate the earth-fault relay
(e.f.r.). Diagram (b) is for an arrangement that
monitors faults in the load network only and does
not detect earth-faults in the generator windings.
Note the relationship between the earthing point in
this diagram and that in diagram (d). Diagrams (c)
and (d) indicate the different arrangements necessary for restricted earth-fault protection. The location of the neutral earthing point in relation to the
protection c.t. in the neutral conductor determines
whether 4 or 5 c.t.s are employed.
The principle of restricted earth-fault protection is
explained in Figure 10.28. The secondary current of
the c.t. placed in the star-point-to-earth connection
is shown (in diagram (a)) as circulating entirely
through the e.f.r.'s sensing coil. This is not strictly
true as a proportion of the current will divide
through the secondaries of the four line c. t. s. The
current drawn by these four 'idling' c.t.s is excitation
current. It is related to the voltage generated by the
working (the fifth) c.t. at the point of common
connection and the c.t. magnetization curve. The
magnitude of this current is very small compared
with the earth-fault relay's setting. The overall sensitivity will be decreased, but only by a small amount
- e.g 10%. High impedance relays are used. A
stabilizing resistor may be necessary in order to
prevent relay malfunction on heavy through faults.
We shall now consider the application of earthfault protection to single and multiple generating
plants operating in various modes. Combining the
arrangements of Figures 1O.27(b) and 1O.27(d) we
can derive a scheme that provides earth-fault protection in all zones for a single base-load generator
installation [18]. The operation of relay e.f.r.a. in
Figure 10.29 indicates the presence of a fault in the
machine's windings. Relay e.f.r.l. covers for faults
developing in the load network. Any signal from
e.f.r.a. should be used to shut down the prime mover
and trip the output circuit-breaker or contactor.
Action should also be taken (especially on the larger
generators) to suppress the excitation circuit. For
faults in the load network it is only necessary to trip
the circuit breaker. The generator may continue to
run if deemed desirable.
Prime mover and generator protection
Figure 10.29
in all zones for a single
Schemes fqr multi-power source systems may use
the arrangement shown in Figure 1O.27(d). It is
equally applicable to installations where a single
generator operates in standby to a utility supply and
to installations which contain multiple generators.
Anyone, or a combination of all, of the arrangements in (a), (b) and (d) of Figure 10.27 may be
used for non-paralleled multiple generators utilizing
4-pole breakers or contactors.
Special precautions are necessary, however, where
multiple power sources operate in parallel or
employ 3-pole output switching devices. Simultaneous earthing of the neutral conductor at more
than one point with respect to the c.t.s is proscribed.
This is because there can be no guarantee as to
which return path the fault current will take.
A scheme that meets all the requirements for
parallel or non-parallel running generators is shown
in Figure 10.30 (18]. It uses 3-pole contactors or
circuit-breakers. Relays e.f.r.a. and e.f.r.b. give restricted earth-fault protection for the generator windings. Relay eJ.r.1. protects the load network zone
regardless of which generator is operating. A signal
from this relay should be arranged to trip outgoing
load feeder contactors or circuit-breakers.
Operators of private standby generating plant in
the United Kingdom need to be aware of the
Electricity Supply Company's regulations relating to
the use of multi-pole changeover switching devices.
Much will depend upon the method of earthing
applied to the installation. Until the relaxation of
regulations in 1965 all Area Boards in the United
Kingdom provided a single neutral earth at their
substation distribution transformer (many still do to
this day). This requires 4-pole switching to be used
(see also Figure 10.31). Where a relaxation has been
applied and the use of protective multiple earthing
(p.m.e.) or multiple earthing of the neutral (m.e.n.)
has been permitted, 3-pole switching is usually satisfactory. It is a requirement of p.m.e. that the neutral
is earthed not only at the Supply Authority's end but
also at the consumer's end of the distribution network. Also, that all extraneous and exposed conductive parts (e.g. metalwork) within a consumer's
premises shall be bonded together and to the
Figure 10.31 shows typical protection configurations using 3- and 4-pole switching arrangements. In
both cases relay e.f.r.a. provides restricted protection for the standby generator windings and eJ.r.1.
gives unrestricted protection to the load distribution
network - regardless of supply source. The latter
relay should be arranged to trip both supply source
Figure 10.32 shows a simplified arrangement for
4-pole switching using one earth-fault relay to give
unrestricted protection in all zones, but only when
the standby generator is supplying the essential
loads. No protection is given when the utility supply
feeds the site loads (see also Figure 10.33).
Anxious to standardize on a scheme that would
cater for both non-p.m.e. and p.m.e. systems (since
non-p.m.e. systems were being increasingly converted to p.m.e. in its area) the United Kingdom's
East Midlands Electricity Board produced a code of
practice in the mid-1970s. This devised a common
arrangement for both systems permitting the use of
3-pole, electrically and mechanically interlocked
changeover contactors.
The salient features of the scheme, reproduced in
Figure 10.33, are that the installation's bonding is in
accordance with the lEE Wiring Regulations and
Standby-to-mains generator supply with
3-pole and 4-pole switching, giving restricted and
unrestricted earth-fault protection
Figure 10.31
Prime mover and generator protection
Figure 10.32 Unrestricted earth-fault protection in all zones
only when the standby generator feeds the load.
restricted earth-fault protection is provided for the
standby generator zone. The generator's neutral is
connected to the consumer's earth bar through a
core balance c.t. All the phase conductors are
threaded through the transformer. Restricted protection is therefore given for all earth-faults on the
generator and between it and the core balance
transformer (c.b.t.). (Note: for the larger sizes of
generator with large diameter cables it is more
convenient to use four c. t.s as in the arrangements of
Figure 10.31.)
The relay operated by the c.b.t. performs the
following functions:
1. trips the controlling contactor to the essential
load network;
2. stops the generator prime mover.
If prime mover shutdown is not provided the standby generator's neutral must be switched along with
the three phases - i.e. a 4-pole contactor must be
fitted. This is because a 3-pole contactor would not
interrupt the fault-current circuit which would con-
tinue to be maintained through the fault-to-earth
and earth paths.
A standby earth-fault relay designed to give backup protection for earth faults on the generator is
shown as an optional extra. It is necessary to incorporate a time delay into this relay in order to give
discrimination with fuses or other earth-fault protection devices in the load distribution sub-circuits.
The schemes shown in Figure 10.31 are equally
applicable to plants operating in parallel with a
utility supply - typically, in peak lopping installations (see Chapter 11). Since it is usual to interconnect the neutrals of utility supply and generator(s)
for the parallel-running mode, 4-pole switching
is unnecessary. Three-pole arrangements are the
Three-pole switching can result in malfunctions of
protection if both of two sources of supply (e.g.
utility transformer and local generator circuits) have
a neutral-to-earth connection. In such cases, if the
parallel neutral/earth paths could cause malfunction
of the standby (or back-up) earth fault protection,
4-pole switching should be employed.
Generator protection
Figure 10.33 Standby generator scheme suitable for p.rn.e.
and non-p.rn.e. installations using 3-pole changeover
The overcurrent relay has been discussed in some
detail in Section 10.3.5. It is fairly common practice,
where the minimum earth-fault current is less than
the normal full-load current, to combine two overcurrent and one earth-fault element in a single
triple-pole relay. The connections are as those
shown in Figure 10.34. The centre element is usually
the earth-fault one. Elements may have combinations of time-delayed and instantaneous operation.
It is very important when using this arrangement
to ensure that the current transformers operate well
below their saturation levels under all fault conditions. That is, they should have sufficiently high
knee-point voltages [14]. (The knee-point voltage is
defined in BS 3938 as the point on the magnetization
curve at which a 10% increase in excitation voltage
produces a 50 % increase in excitation current.)
The reader will come across the term directional
relays. In their overcurrent form they are not
applied to generator zones and are more suitable for
the protection of ring mains, parallel feeders, parallel transformers, and transformer feeders. They
should not be confused with 'power directional'
relays (see Section 10.4.8). Directional relays res-
pond to the phase relationship between the current
flow and a reference voltage, which is fed to an
additional voltage element within the relay. They
will only operate when an overload or short-circuit
current flows in the desired direction.
Prime mover and generator protection
The schemes we have just discussed give technically sound and economically viable solutions to
the problem of internal earth-fault protection for
small low-voltage machines (and indeed are basic to
the protection of larger machines). Their fault clearance times tend to be too slow for the higher
voltage, larger output, and more expensive
machines. Serious consideration should be given to
the use of a truly unit-type protection (described in
the next sub-section) for machines:
• above 1 MV A;
• above 5 kV; or
• above 500kVA at 2.2kV and higher.
Differential protection
Often called balanced or Merz-Price protection, it
uses the most selective of relaying principles. It
works on the basic proposition that, under healthy
conditions, the current entering into a protected
section (a winding, in our case) is identical in
magnitude and phase angle to that leaving it. Current transformers of equal ratio are located in the
input and output lines of the protected section and
are connected by pilot wires to a current sensitive
(differential) relay.
The single-phase diagram of Figure 10.35 depicts
the basic principle of the circulating-current system.
During normal conditions in the protected zone
Figure 10.35 The principle of circulating current protection
(a) In normal conditions or for an external fault (b) For
internal fault condition
(diagram (a» no current flows through the relay's
operating coil which is connected across the equipotential points a and b in the pilot-wire loop. A
similar situation applies when a fault occurs outside
the zone. The direction of the circulating current in
the 'differential' loop is unchanged since the primary
current entering and leaving the c. Ls is of the same
magnitude and in the same direction. The system is
therefore insensitive to load current changes and to
any through-fault currents (i.e. it is stable for all
conditions outside the zone).
If an 'internal' fault occurs within the protected
zone between the two c. Ls the equality of input/
output currents (i.e. the current balance) is disturbed. See diagram (b) in Figure 10.35. The
secondary current flow in the right-hand c.L is zero.
The only current flowing in the relay coil is that from
the left-hand C.t. The relay operates when the
magnitude of this current exceeds the relay's pick-up
Figure 10.36 shows the connections for a generator's differential protection scheme. It consists of
two sets of high quality c.Ls with similar characteristics, connected to the three elements of a
circulating-current relay. Note the need to have all
six ends of the stator windings brought out to the
machine's terminal box. The c.Ls should not be used
for any function other than that of differential
protection, because unequal secondary loading may
occur [11]. Despite this precaution it is quite
possible for supposedly identical sets of c. Ls to give
slightly differing ratios when subjected to high asymmetrical through-fault currents. Such currents,
which have a marked d.c. component (see Figures
4.33 to 4.36 of Chapter 4), can cause the transformers to reach saturation level. Any variations in
the magnetizing characteristics of the c.Ls will then
result in some unbalance current (or spill-current)
flowing through the relay winding, even though the
fault may be external to the protected zone.
There are ways of ensuring stability under these
conditions. One method is to use a so-called high
stability circulating current relay. This is a voltageoperated high impedance relay set to operate at a
voltage slightly higher than that developed by the
current transformers under maximum external fault
Because of its high speed and immunity to a.c.
transient currents the high stability relay is more
usually applied to small and medium sized generators. Larger machines (and directly connected
generator-transformer units) may employ another
method known as biased differential protection. It is
particularly suited to those circumstances where c.Ls
on the line side of a generator are likely to have
dissimilar characteristics to those on the neutral side
of the windings.
The bias slope (see Figure 10.37) of a relay is
defined as the ratio:
Generator protection
Figure 10.36 Circulating current (differential) protection
applied to a 3-phase generator
The contacts of 'biased' relays are arranged to
close only if the differential current for a corresponding through current falls within the relay's
operating region, i.e above the bias line setting. The
setting of the relay is the minimum current (expressed as a percentage of generator rated current)
which will trigger the relay.
The percentage bias is a fixed parameter defined
at the minimum current setting as:
The characteristics depicted in Figure 10.37 include
that for a relay with a 10 % bias. A generator,
typically, requires 10 % values for both setting and
bias. The upper curve in Figure 10.37 therefore
gives a very satisfactory relay operating characteristic for all but the smallest generators.
Solid state relays are widely applied in bias differential schemes. Restraining and differential operating voltages are compared in transistorized comparator stages to operate into output units of the
attracted armature type. Such relays (e.g. GEC
Alsthom Measurements type DTH 31 and P & B
Engineering type PBSD 3A5 J6) may also incorporate second harmonic restraint features to inhibit
operation by the normal inrush currents produced
when transformers directly connected to generators
are first energized. Descriptions are outside the
scope of the present discussion, which must be
confined to system principles only.
When a generator is directly connected to a
step-up transformer without any intervening switchgear it is normal practice to include both generator
and transformer in the one differential protection
Figure 10.37 Operating characteristic of a percentage
biased differential relay (Courtesy: GECAlsthom
Measurements Ltd)
scheme. One set of C.t.s is then placed at the
generator neutral end of the zone and the other in
the transformer HV lines before the circuit breaker
- as shown in Figure 10.38. Notice that the c.t.s are
connected in star at the generator neutral end and in
delta at the transformer HV side. This is done in
order to compensate for the 30° phase shift introduced by the delta-star step-up transformer.
Because the power transformer usually has tapchanging facilities it is extremely difficult to match
both sets of protective c.t.s for all voltage ratios. The
generator-transformer units should therefore always
be of the percentage biased type.
The star-point of the HV side on the power
transformer is likely to be earthed through a neutral
resistance. It is therefore normal practice to include
Prime mover and generator protection
Figure 10.39
Combined overall biased differential and
restricted earth-fault protection for a directly connected
generator-tra nsfo rmer.
a restricted earth-fault relay (as shown in Figure
10.39). It is also necessary to time-delay the operation of the differential protection in order to cope
with the high inrush magnetizing currents of the
power transformer. Where this is unacceptable use
may be made of a percentage biased relay with
second harmonic restraint features - as discussed
None of the schemes described in this and the
previous sub-section give 100 % protection of the
stator winding against internal faults. There is no
difficulty in detecting an earth-fault near the generator output terminals. The earth-fault voltage and
hence the current flowing to earth is sufficient to
ensure reliable operation of protection equipment.
This is especially true when the neutral-to-earth
resistance ensures that the magnitude of the fault
current is greater than that of the machine's full-load
current (i.e greater than 1 p.u.). The situation is
markedly different if an earth-fault occurs near the
star-point of the stator winding. The voltage behind
the fault is then usually insufficient to create a
relay-energizing current greater than its setting
The percentage of generator winding protected
depends both upon the magnitude of fault current
Generator protection
that the neutral resistance (more strictly, the impedance) is designed to pass, and upon the setting of
the relay. The two must be correlated to obtain the
optimum stator protection. Where a generator is
connected directly to a load the value of the earthing
resistance RN is usually chosen to be about 1 p.u.
This means that 1 p. u. phase voltage is measured
across it when rated full-load current flows through
it. Given a resistance of 1p.u. and a differential
protection relay setting of 20 % of the generator's
full-load current, an earth-fault 20 % distant from
the star-point would create just sufficient fault current to trip the relay. In other words, 80 % of the
stator winding is protected. Reducing the relay
setting to S% would give 9S % of the winding
protected; and so on.
If the neutral resistance (RN) were reduced to,
say, 0.9p.u. with a relay setting at 20%, an earthfault 18 % distant from the star point (i.e. 0.9 x 20)
would just be detectable. Protection would therefore be afforded to 82 % of the winding. Figure
10.40 [14] illustrates the argument.
It can be shown [14] that x, the position of the
fault from the star-point, expressed as a fraction of
the whole phase length (i.e the unprotected fraction
of the phase winding) is given by;
RN is earth resistance value, in ohms;
kIs is the setting of the protective system
measured on the primary side of the c.t.
(i.e the primary fault setting).
Furthermore, if all these quantities are expressed in
per-unit terms (of the generator's rated current), the
relationship for x becomes;
10.4.2 Rotor faults
In modern practice it is usual for the generator's
main (rotor) field circuit to be fully isolated from
earth. The presence of a single earth-fault due to
insulation failure does not jeopardize the machine.
A second earth-fault some physical distance from
the first would, however, short-circuit a major part
of the field winding. If this does not take the
machine out of service it will certainly cause magnetic unbalance in the rotor and possible bearing
On machines in which excitation is affected
through shaft-mounted brushgear it is possible to
employ direct detection methods using, for
example, a potentiometer across the field winding,
with a centre tap connected through a sensitive
moving-coil relay to earth. Alternatively, d.c injection techniques are employed where, in the presence
of a fault, current flowing in a bias circuit is detected
by a sensitive balanced-armature relay or by a
moving-coil relay. a.c. injection methods are also
applied, but any schemes that introduce currents in
the main shaft should not be entertained as they are
likely to accelerate bearing failure [8].
Generators in the range we are considering are
more likely to be of the 'brushless' type. Direct
access to the main field will not be possible so that
any of the fault detection techniques just described
are out of the question. We have considered in
Chapter 3 (Section 3.3.2) some of the methods that
may be used to detect rotating diode failures. An
open-circuit on anyone of these devices (or a single
earth-fault on the winding, for that matter) would
produce little change in excitation level. Alarm
indication only would suffice in such circumstances.
A short-circuited diode is more serious. The resulting magnetic imbalance will tend to cause abnormal
vibration in the machine. This can be detected by a
vibration monitor. Action should then be taken to
unload the machine or to take it out of service.
10.4.3 Over and under-voltage
Protection against abnormal voltage conditions is
not often applied to generators. Over-voltage can be
produced by overspeed or by automatic voltage
Prime mover and generator protection
regulator failure. It may overstress generator winding insulation but is more likely to endanger connected sensitive loads such as computers and communication equipment.
The most commonly applied relays are of the
solid-state type in which the input signal is rectified
to give a proportional d.c. voltage which is compared with a preset reference voltage. The resultant
of this comparison is fed, via a transistorised amplifier, to an electromechanical output relay. Instantaneous and time-delayed versions are available.
Settings must be selected to allow for network
transients (e.g. due to load switching) but to trip the
machine before dangerous voltage levels are
reached. Where generators are operating in parallel
it is necessary to discriminate between healthy and
faulty units. Because over-voltage results from excessive excitation a protection system operating on
combined over-voltage and high exciter field current
may be used [19]. Due allowance must be made for
the range of excitation in normal operation (i.e.
from zero, on load removal, to maximum when
field-forcing on load switching or short-circuit).
Under-voltage settings should similarly take account
of the voltage dips that are expected during load
Where voltage surges are likely to be induced by
lightning discharges, on lines external to the generating plant surge protection should be considered
- especially where vacuum breakers or contactors
are used with the larger and higher voltage
10.4.4 Over- and under-frequency
In most instances the prime mover will be fitted with
an independently-acting overspeed device (see Section 10.2.3). This is usually set at about 15 % above
normal speed. Over-frequency protection is then
quite unnecessary as far as the generator itself is
concerned. However, some specialized loads are
prone to damage if operated at other than design
frequency. Frequency relays are arranged to trip the
load breakers when the supply frequency drifts
outside predetermined limits.
On single generator installations overloading will
result in frequency drop. The problem is that if no
action is taken to alleviate the situation, the machine's a.v.r. continues to boost excitation to compensate for loss of speed. If the condition is allowed to
persist rotor overheating can occur. Most modern
a. v.r.s are fitted with under-frequency protection but
these circuits are designed to operate at low speeds
and are only meant to protect the generator against
engine idling conditions. Under-frequency relays
may be used to initiate load-shedding by tripping
selected load breakers in multiple steps with a
time-delay between steps. A stability study would
determine the frequency settings required at each
step [8].
Frequency relays may be of the solid-state type or
of the polarized sensitive balanced-armature type
(such as the GEC Alsthom Measurements type
FMG), where the relay has three separate windings.
The operating and restraint windings are energized
through series circuits tuned to resonate at frequencies on either side of the setting. The third winding
provides a small bias in the restraint direction to
prevent transient operation if the supply is switched
on at a frequency close to the relay setting. On the
under-frequency relay the operating winding is
tuned to a frequency below the setting, and the
restraint winding to a frequency above the setting.
As supply frequency falls the current in the restraint
winding decreases whilst that in the operating winding increases - as its tuned circuit approaches resonance. The over-frequency relay is similar in operation except that the tuned circuits are reversed.
The solid-state versions give greater accuracy but,
because they use square-wave pulsing controls, their
measuring logic can be adversely affected by harmonics in the input wave-form. Low pass filters fitted
between the protection v. t. secondary and the relay
are therefore required when the total harmonic
distortion exceeds 5 % of the fundamental frequency.
10.4.5 Overload
On partially-manned or unattended sites it is necessary to include protection against prolonged overloading. We have established elsewhere (Chapters 2
and 3) that diesel generating sets rated to ISO
standards are capable of meeting 10 % overload
conditions for periods of 1 hour in any 12 hours of
continuous operation. Fuel systems are so arranged
that once an engine operates beyond the 10 %
overload region supply is restricted by preset control
or fuel-limiting stops. Load increases that call for
fuel amounts beyond this limit merely result in the
engine's speed falling below the governor control
setting. This not only results in engine overheating
and high pressures on its bearings but it also raises
generator temperature. In the previous section we
discussed the use of frequency relays to detect the
reduced speed condition and to initiate the
appropriate load-shedding actions.
In Section 10.3.5, we considered the merits of the
modern direct-acting overload devices linked with
low-voltage switchgear and suggested that they afforded adequate protection on the smaller installations. On larger base-load installations serious consideration should be given to the use of frequency
and/or IDMTL overcurrent relays. The latter may
be voltage restrained to give better discrimination in
overall system protection schemes.
Generator protection
Other devices which may be used on the larger
machine installations are:
Power relays
• temperature relays; and
• power relays.
Temperature relays
A number of resistance temperature detector
(RTD) elements may be embedded into the
machine windings during manufacture. The RTD
elements may be of platinum or copper - 100 nand
10n resistance, respectively. They are strategically
positioned by the machine manufacturer to detect
hot-spots and localized heating, such as that caused
by short-circuited laminations within the stator.
Each RTD element forms one limb of a balanced
Wheatstone bridge. Change in the resistance of any
element as a result of its temperature rise unbalances its bridge circuit. The resultant voltage is
amplified and is compared with a common voltage
reference corresponding to the preset temperature
trip level of the relay.
In the six-point relay illustrated in Figure 10.41 all
the comparator outputs are fed via a relay driver to a
multi-input OR gate which energizes the output
relay if one or more RTD elements exceeds the trip
setting. The output relay is also energized if any of
the six input channels are open-circuited. Light
emitting diode (LED) indicators show the 'normal'
and 'tripped' conditions for the relay and also which
of the inputs are exceeding the trip level. Adjustment of the trip point is by a multi-turn potentiometer accessed from the front panel. Temperature is
adjusted by SOCper turn. Since no visual readout is
given it is prudent to seal the setting and mount the
relay within the control gear enclosure. This will
militate against the operator who may be tempted to
raise settings when under pressure to maintain operations.
The relationships between the three forms of power
is dealt with more fully in Chapter 4. It will help in
our present discussion to think of watts being put
into the system (in the form of torque) by the prime
mover, and that VArs required to support the
voltage and transport the watts through the system
are put in by the generator's a.v.r. Power relays may
be connected for either watts or VAr measurement
by a simple rearrangement of the voltage and current connections across (and in) the power system
lines. By virtue of its sensitivity to phase angle the
power relay is inherently directional in operation.
This characteristic is very useful when it comes to
sensing loss of prime mover power - as we shall see
in a later sub-section (10.4.8)
For high speed applications electromechanical
induction-cup and solid-state relays are suitable
alternatives. A single relay is sufficient for balanced
conditions on three-phase, 3-wire or 4-wire systems.
Where unbalanced loads are likely a relay must be
employed on each phase or a triple pole unit must be
10.4.6 Unbalanced loading
All faults other than those involving all three phases
of the system will give rise to unbalanced currents.
These may be resolved into positive-, negative-, and
zero-sequence components. The relative values and
the relative phase angles of these fictitious components of current will vary from one asymmetrical
fault condition to another.
While faults external to the generator should be
quickly cleared by circuit protection, any failure of
remote protection to operate or of related circuit
breakers to trip would result in the faulted circuit(s)
remaining connected to the generator. We are concerned here with the effects that sustained external
faults which have not been cleared by the
appropriate protection will have on the generator.
The worst condition for the machine occurs on
those faults that give rise to negative phase sequence
(n.p.s.) currents in the stator windings. These are
phase-to-phase faults (between lines) and phase-toneutral faults (or line-to-earth, where the neutral is
Prime mover and generator protection
The n.p.s. components of the current rotate in a
direction counter to the d.c. field system of the
generator. This produces a stator flux which cuts the
rotor at twice its rotational velocity and induces
double-frequency currents in the field system and in
the rotor body. The resulting eddy currents may be
very large and cause severe overheating of the
rotor - especially at those points where the circulating eddy current is concentrated by the winding slots
of cylindrical rotor machines. Heat concentrates in
parts of the coil binding rings and surface fusion has
been known to result [10]. The heating effect is less
marked in salient pole constructions, particularly
those employing laminated poles and heavy amortisseur windings (see Chapter 3). The presence of the
damper windings has a powerful limiting effect on
the double-frequency currents in the field system.
The measure of the machine's tolerable thermal
overload is given by a constant factor (A. Allen,
1987, personal communication) which is the product:
Generator protection
Rgure 10.42 Block schematic diagram of a negative phase
sequence current relay (Courtesy: GEC Alsthom
Measurements Ltd)
10.4.7 Loss of excitation
Loss of excitation in a generator can be caused by:
• a failure within the main field system (as discussed in Section 10.4.2);
• a failure in the a.V.L; or
• failure within main and/or auxiliary exciters in the
excitation chain.
The condition is readily detected by an undervoltage relay on a single machine feeding a load
network. When the generator is one of several
running in parallel, or if it is connected into an
'infinite system' such as a utility supply, loss of
excitation may result in the machine losing
synchronism. It then runs asynchronously as an
induction generator deriving its excitation from the
reactive stator current provided by the power
system. The magnitude of this reactive current may
well exceed the rating of the generator so that its
stator windings are overloaded. Additionally, the
slip frequency currents induced in the rotor damping windings will cause abnormal heating of the
rotor [16]. Asynchronous operation will not damage
the machine immediately. Conventionally cooled
machines should be capable of operating at full load
in this mode for about 5 minutes.
Failure of a.V.L supply to the exciter field may be
easy enough to detect but main rotor field circuits
Prime mover and generator protection
the machine's terminal voltage falling while the
current begins to increase. The apparent impedance
of the machine therefore decreases and its power
factor changes. An impedance relay may be used to
detect this change from normal load value. These
relays can be constructed to give various R-X characteristics. Application theory is fairly complex and
must be outside the scope of our discussion. Attention is directed to [9] and [16] for further reading on
the subject.
The most common relay characteristic is the circular offset mho shown in Figure 10.45. The offset and
the diameter of the circle of the relay characteristic
are both adjustable. The relay is usually set so that
the offset is equal to half the generator's transient
reactance (X'd) and the diameter of the characteristic is equal to the direct axis synchronous reactance
(Xd). When the machine experiences a loss of field
its terminal impedance moves into the negative
reactance area. The relay operates as soon as the
operating-point locus moves inside the circle - the
relay characteristic setting. The impedance phasor is
drawn for lagging currents at heavy and light loads
(Ia and lb, respectively).
Loss-of-excitation protection schemes are usually
arranged to shut down the failed generator if adverse field conditions persist for a predetermined
10.4.8 Failure of prime mover
Should the prime mover fail (say, due to fuel
problems) while a generator is running in parallel
with others or with a utility supply, the generator
will 'motor' and will draw power from the system
into which it is connected. Diesel engines are less
Generator protection
susceptible to damage caused by generator moto:ing
than some other types of prime mover. Nevert'1eless, some means of detecting the reverse power
flow is necessary, especially to protect against 'inverted' running should mechanical seizure occur on
the prime mover (e.g. bearing heating leading to
closure of fuel valves).
Selection of the proper protection depends upon
the amount of power that is required to drive the
prime mover when no fuel is supplied. For diesel
engines this, when expressed in terms of a percentage of the full-load nameplate kW rating, is of the
order of 25 % for 4-stroke engines and between 15
and 20 % for 2-stroke engines. These figures may be
very much lower for well run-in engines. In all cases,
the engine manufacturer should be consulted for
actual values.
Protection is provided by a directional (reverse
power) relay set to operate at a level lower than the
predetermined motoring power of the prime mover.
Typically, this is below 10 % of machine rating. The
sensitivity of the chosen device depends upon the
generator's application. In some applications, such
as on marine power systems, maintenance of supply
is such a critical requirement that only large (but
safe) amounts of reverse power would be sufficient
cause for tripping a generator. On others the generator may be tripped at very much lower levels.
Relay sensitivity may then be limited only by the
need to avoid nuisance tripping during synchronizing operations or during transient load changes.
The true power taken by the generator on inversed running should be monitored. Due account
must be taken of the power factor of the current
drawn from the system. This may be either lagging
or leading depending upon the excitation level of the
failed machine. A true 'wattmetric' relay having
accurate quadrature adjustment must be used. Unlike directional relays for fault power applications
the reverse power relay should not be fitted with
phase angle or low voltage compensation [16].
In most cases, single phase power measurement
suffices because loss of prime mover torque affects
all three phases equally. True power (kW) is measured by voltage taken from line to neutral and
current is measured from the same line. Where the
neutral connection is not available three phase voltage sensing, or a power relay with an in-built 30°
voltage phase shift, may be used.
Suitable protection relays may be of the electromechanical (induction cup) type or of the static
type. An example of the latter is shown diagrammatically in Figure 10.46. It measures I cos<f,>,on the
not unreasonable assumption that voltage may be
regarded as constant and balanced. The relay has a
definite time characteristic, ensuring that maloperation does not occur due to momentary power
reversals during synchronizing or during power
swings on the system. The unit uses a phase sensitive
bridge to produce an output which is amplified and
compared with an adjustable d.c. level to give the
settihg range (Ps), in multiple steps - between 3 %
Prime mover and generator protection
and 30 % of the nominal power. The output from
this level detector controls a timer with a continuously adjustable range up to 10s. The relay is
self-powered from the measuring voltage transformer. An alternative version is available for
phase-to-phase v.t. connection.
10.4.9 Loss of synchronism
Severe system faults or operation at high load levels
with leading power factor (and hence a relatively
weak field) may cause a machine to pull out of
synchronism. The generator and its prime mover are
subjected to violent torque oscillations. Synchronism can be regained if the load is quickly reduced. If
this fails it would be necessary to take the machine
out of service. The alternative is to remove the
machine's excitation to make it run asynchronously.
This immediately removes the power oscillations.
Reclosing the excitation at a low value will then
allow the machine to resynchronize smoothly [9].
The pole-slipping that can occur with such system
disturbances results in a flow of synchronizing power
(see Chapter 8) which reverses in direction twice
every slip cycle. This is the cause of the violent
One convenient and reliable method of poleslipping protection uses measurement of the magnitude, direction, and rate of change of load impedance relative to the generator's terminals. If satisfactory unit discrimination is to be achieved a great
deal of care needs to be exercised in designing and
setting pole-slipping protection which is based on
load impedance measurement. The application of
offset mho-type relays to this form of protection
requires fairly complex analysis. The techniques
involved are outside our present scope. Pole slipping
protection is seldom applied to diesel generating
plants. If used, it is usually only on the very largest
10.5 Typical protective relaying
Having described the protection equipment that
may be applied for specific fault conditions, we shall
now examine typical schemes for single and multiple
generator installations. These are not meant to be
definitive as system needs will vary. Experienced
guidance should be sought in obtaining the best and
most economical packages for any proposed installation. In some cases a client's specification will
dictate the relaying requirements. This should never
imply finality in design terms. A contractor's own
study, based on his specialist equipment manufacturers' recommendations, may suggest a different
approach which could well satisfy the 'sense' of a
client's overall requirements. If this is so, the alter-
native proposals should be defined and the information on which they are based should be declared.
Where plant is to be unattended the aim should be
to provide automatic protection of both generator
and prime mover against all harmful conditions.
Where attendants are present automatic shut-down
could result in the unnecessary removal of generator(s) from service for those abnormal conditions
which may be corrected by attendants. Opinions will
differ as to which conditions should be left to the
action of attendants.
Table 10.1 (in Sub-section 10.5.4) offers guidance
on the monitoring and protective features to be
provided for generators. Consideration must always
be given to the cost and functional importance of the
plant when determining needs. The objective must
be to provide the minimum of protection consistent
with adequate coverage of all conditions that are
liable to cause damage or affect the continuity of
supply [6, 10].
In the discussions that follow we shall start with
the very basic standby-to-utility generator, then
expand this to consider parallel running machines
feeding isolated loads, and finally consider the protection of peak-lopping and cogeneration plants. In
each case the aim is to recommend the basic features
that should be applied and those which should be
10.5.1 Single generator protection
The single line diagram of Figure 10.47 defines the
minimum protection for a single standby unit which
is isolated from the utility supply by some form of
changeover switching (see Chapter 11, Section
11.3). The protection consists of a three-element,
voltage controlled overcurrent relay for phase-tophase and phase-to-earth protection. A separate
earth-fault relay provides back-up phase and earthfault protection for systems where the machine's
neutral is solidly earthed. The overcurrent relay will
also provide back-up to the distribution network.
That is, it will detect faults external to the generator
and trip its breaker should other protective devices
in the network fail to operate or fail to clear the
faults quickly enough.
Multi-purpose protection equipment based on
microprocessors is now widely available. They are
easily activated and set by the user for his specific
application. The Asea Brown Boven (ABB) Series
SPAG units are typical examples. Designed for
stand-alone, small and medium sized stationary and
mobile generating plants, they offer a number of
functional modules giving flexibility in the design of
operational schemes for various applications.
System reliability is enhanced by self-supervision
features, including self-diagnostics. Functional modules include:
Typical protective relaying schemes
Figure 10.47 Single-line diagram of basic protection for a
standby generator
overvoltage and undervoltage;
reverse power; and
earth faults.
Direct numerical display is provided for setting
values, voltages, load currents, memorized fault
data, indications and status information.
10.5.2 Multiple generator plant
When two or more generators are operated in
parallel the protection requirements are enhanced
to include a reverse power relay for each machine
and some means of synchronizing and paralleling
the generators to the busbars. Figure 10.48 shows
the basic protection required for a two-unit configuration.
Prime mover and generator protection
The differential protection scheme has been
shown as an option but one which should be
seriously considered for those installations where
larger unit-size machines are employed (see Subsection 10.4.1). Choice of the synchronizing device
will be determined by the operational needs of the
plant. If the installation is permanently manned and
generators are started and stopped infrequently the
synchronizing of incoming machines may be' supervised by check relays. These verify that the proper
system conditions exist before breakers can be
closed. Where frequent synchronizing operations
are the norm or where remote control is provided, it
may be necessary to use automatic synchronizing
relays for each machine. These are then arranged to
interact with the generator's excitation controls and
the engine's governor to give voltage and frequency
matching. They also automatically compute the correct advance angle before giving a closing signal to
the generator circuit breaker. See Chapter 8 for
more details of check- and auto-synchronizing
Multi-functional relays specifically designed for
the protection of parallel-running generators are
also widely available. One such equipment is the P
Figure 10.49 An example of a multi-function generator
protection relay (Courtesy: P & B Engineering Co. Ltd)
& B Engineering type PBSG 310 relay shown in
Figure 10.49. The relay package includes three
plug-in measuring units (shown on the right of
Figure 10.49). Unit PBCP 3B2 caters for reverse
power and dual-stage over-voltage protection. The
reverse power function includes a time-lag feature.
Units PBCJ 3B3 are used for high- and low-set
overcurrent protection and for load shedding. The
overcurrent units can be set to give either definite or
inverse time characteristics.
10.5.3 Parallel operation with a utility supply
We shall now consider the protection requirements
for in-plant generators and utility tie lines in peaklopping and co-generation schemes .
Additional protection to that already defined for a
standby generator scheme (discussed in Section
10.5.1) will be required. It is essential that the
generator is removed from service as quickly as
possible for any faults developing within the plant
itself or within the utility's feeder to the plant.
Should a fault develop within that portion of the
utility's network which is associated with the consumer's plant the feeder must be the first to be
Typical protective relaying schemes
interrupted and the last to be restored. Most utilities
will expect the consumer to pay for the additional
protection required to remove in-plant generators
from the affected line.
The majority of private generation systems are
connected to a radial feeder. Any fault in that feeder
or in the substation associated with it is effectively a
step change in the generator's load. The generator
will attempt to support the connected load and also
supply the fault current. The machine overloads and
prolonged voltage and/or frequency dips will result.
The under-voltage and under-frequency relays of
Figure 10.50 are included for this purpose. They
should be arranged to disconnect the generator from
the faulted zone. Over-frequency detection guards
against engine governor failure, and under-voltage
against generator excitation control failure.
It is a requirement of many Supply Authorities
that in-plant generators are totally isolated before
the tie-line breaker is closed. It is then usual to
inhibit closure of that breaker by placing the voltage
relay contacts in series with the breaker's closing coil
contacts. Automatic synchronizing facilities capable
of adjusting generator speed and voltage will reduce
the time required to re-establish the utility supply to
the plant. '
Any protective devices included in the plant's
generating system will need to be approved by the
Supply Authority. Also, its personnel will require
right of access to the consumer's protective equipment for periodic calibration and tests.
Those co-generation schemes which employ induction generators will not require synchronizing
equipment since the generators usually draw their
excitation VArs from the utility. Voltage and frequency are controlled by that source. When the
utility supply is disconnected the induction generator's terminal voltage decays rapidly. In other
words, it needs the utility system in order to generate an output (see Chapter 11, Section 11.3.4).
Operation of the under-voltage relay is delayed so
as to cater for system voltage dips due, for example,
to the starting of large motors elsewhere in the
utility-fed network. The time-delayed output of the
relay trips the in-plant generator breaker. Overvoltage relays may act to instantaneously trip the
generator circuit breaker. The frequency relays
serve to isolate the plant generator when the utility
supply fails.
A point worth noting is that capacitor banks are
commonly employed by industrial consumers for
power factor correction on their premises. These are
shown in the single line diagram of Figure 10.51.
They will provide continued excitation (VArs) to
the induction generator. This being the case, the
generator's output voltage and frequency will be a
Prime mover and generator protection
function of prime mover torque and of the total load
imposed by the consumer's installation plus that of
other consumers tied to the utility's radial feeder.
The condition will be detected by frequency and
under-voltage relaying and the induction generator
will be tripped.
Reference is made in the next chapter (Section
11.4.4) to the Electricity Council's Engineering Recommendation G 59 (Recommendations for the connection of private generating plant to the Electricity
Boards' distribution systems). In our context, it lists
the objectives for the protective equipment to be
provided by the private generator. These are summarized as:
1. to inhibit connection to the Board's supply unless
all the latter's phases are energized and within
agreed protection settings;
2. to disconnect the generator when a system abnormality results in deviation of voltage or frequency at the point of supply;
3. to disconnect the generator on loss of Board's
4. to ensure automatic disconnection (or alarm indication where competent attendance is available)
of the in-plant generator when failure of any
supplies to the protective equipment inhibits its
correct operation.
G 59 goes on to indicate the minimum protection
necessary for both HV and LV supply arrangements. This includes the detection of:
loss of mains
It acknowledges that other protection may be required, and may include the detection of:
neutral voltage displacement
earth fault
reverse power
The rider is added that if automatic resetting of the
protective equipment is used there must be a built-in
time delay in order to ensure existence of a healthy
supply for a continuous period of 60 s. The automatic reset must also be inhibited for faults on the
Private Generator's installation. Staged timing may
be necessary where more than one generator is
plant when determining the protective measures
required for lower rated machines.
The selection column classifies the protection
systems as follows:
recommended for general application
where required
In practice, one relay can be used for several
functions. For example, one triple pole overcurrent
relay may be used for both phase and earth fault
(1) With HV generators.
(2) Only when continuous operation on unbalanced load is likely.
(3) For parallel running generators.
(4) Not usual below 100 kW, or in non-critical
(5) Usually applied: (a) above 1 MVA;
(b) above S kV;
(c) above SOOkVA at more
than 2.2 kV.
(6) On generators equipped with static excitation
but not on brushless machines.
(7) For protection against external overload and
reduced cooling, e.g. high ambient temperatures or dirty air filters. Usually fitted to
machines above 5 MV A.
10.6 Related standards
Reference has been made in this chapter to the
following British Standards. Entries in parentheses
are the corresponding and/or technically equivalent
international standards.
BS 142: Part 1: Section 1.1 - Glossary of protection
(IEC SO(446)
relay terms
BS 3938 (IEC 185) - Current transformers
BS3941 (IEC 186,IEC 186A) - Voltage transformers
BS 3939 (IEC 617) - Guide for graphical symbols
for electrical power, telecommunications, and electronics diagrams
CP 1013 - Earthing
10.7 References and bibliography
10.7.1. References
10.5.4 Relay application guide
Table 10.1 lists the recommended protective features to be applied to generators. It is not practicable to cover every application. Consideration must
be given to the cost and functional importance of the
1. The Lubrication of Oil Engines, Shell International Petroleum Company Ltd
2. Kates, E.J. and Luck, W.E. Diesel and High
Compression Gas Engines, 3rd edn, American
Technical Society, (1.974)
Referencesand bibliography
10.1 Choice of a.c. generator protection systems
Protection systems
Above 1 MV A Below 1 MV A
Short -circuit
Stator earth fault
Restricted earth fault
Unbalanced load
Reverse power
Over-/under voltage
Over-/under frequency
Loss of excitation
Rotor earth fault
Winding over temperature
3. Davies, M. Governors and governor gear.
Chapter 13 in Diesel Engine Reference Book, 1st
edn, edited by L.C.R. Lilly, Butterworth, (1984)
4. Allen fuel-pipe fracture protection equipment.
Apex, No 6 March, NEI Allen Ltd (1971)
5. General installation manual, 2nd edn, Publication No. 801SER02801164, Perkins Engines Ltd
6. Stouppe, R.L. Jr. System protection and coordination. Chapter 12 of Switchgear and Control
Handbook, edited by Robert W. Smeaton,
McGraw-Hill, (1977)
7. Fettback-Heyden, K.H., Ilar,F. and Ungrad
Dr. H. A computer programme for testing
protection selectivity in medium-voltage and
low-voltage systems, Brown Boveri Review, No.
11(72), November, BBC Brown Boveri & Company Ltd (1985)
8. Roe, L.B. Practical Electrical Project Engineering, Mc Graw-Hill, (1978)
9. Cliff, J .E. Switchgear and protection. Chapter
15 in Electrical Engineer's Reference Book, 14th
edn, edited by M.A. Laughton and M.G. Say,
Butterworth, (1985)
10. Protective Relays - Application Guide, 3rd
edn, GEC Measurements, GEC pic (1987)
11. Lythall, R.T. The J&P Switchgear Book, 7th
edn, General editor C.A. Worth, Butterworth,
second revised impression (1976)
12. McLeay, K. Analysing power systems using
computers. Power Engineering Journal, 2(4),
July (1988)
l3. Newcombe, R.W. Application of inverse time
overcurrent relays. Publication R-5226B, GEC
Alsthom Measurements Ltd
Section reference
and (Notes)
(1 &
(1 &
(1 &
(1 &
14. Guile, A.E. and Paterson, W. Electrical Power
Systems, Vol. 1, 2nd edn, Pergamon Press,
reprinted (1978)
15. Protective relays and automatic synchronisers.
Proceedings from a Seminar held at Staveley
Electrotechnic Services Ltd (Basler Electric)
January (1983)
16. Rushton, J. and Mewes, K.G.M. Protection
generatortransformer units and transformer feeders.
Chapter 12 in Power System Protection, Vol.
3 - 'Application', 2nd edn, edited by The Electricity Council (Peter Peregrinus Ltd.) (1981)
17. Bailey, A.C. Transportation: ships. Chapter 31
of Electrical Engineer's Reference Book, 14th
edn, edited by M.A. Laughton and M.G. Say,
Butterworth, (1985)
18. Moody, R.L. Generating plant fault protection.
Electrical Equipment, January (1975)
19. Ames, R.L. Marine a.c. generating systems.
LSE Engineering Bulletin, 12(1), January (1972)
10.7.2 Bibliography
The reader is particularly referred to:
1. The J&P Switchgear Book listed under Reference [11]
2. The three volumes which make up the Electricity
Council's publication listed in Reference [16]
3. IEEE recommended practice for electrical power
distribution for industrial plants - the so-called
'Red Book', IEEE Std 141-1987, The Institute of
Electrical and Electronics Engineers Inc. and
distributed by Wiley-Interscience, a division of
John Wiley & Sons Ltd
Prime mover and generator protection
4. IEEE recommended practice for the protection
and coordination of industrial and commercial
power systems - the so-caJled 'Buff Book', IEEE
Std 242-1975, The Institute of Electrical and
Electronics Engineers Inc.
5. Recommendations for the connection of private
generating plant to the Electricity Boards' distribution systems. Electricity Council Engineering
Recommendation G 59.
6. The installation and operational aspects of private generating plant. Electricity Council Engineering Recommendation G 26.
7. Notes of guidance for the paraJlel operation of
private generators with Electricity Boards' low
voltage networks. Electricity Council Engineering Recommendation G 47/1.
Also, useful information on related topics is
obtainable from equipment manufacturers, GEC
Alsthom Measurements brief guide Generator pro-
tection (publication R-5061A), and Relay application
guide (publication R-5582), are typical examples.
The journals and specialist-conference publications of learned societies report on the latest advances in design and engineering practices. For
example, the lEE's Conference Publication 272
December 1986 includes the following papers:
(a) Enhancing control and protection for improved
plant reliability.
(b) Combined solid state system for protection,
automation, measurement and tele-transmission
in MY switchgear.
(c) Protection against reverse and under-power
conditions by modern electronic relays.
(d) Computer-aided design of industrial protection
coordination with assessment of relay settings
under transient conditions.
(e) Microprocessor-based power system protection
Emergency and
A.C. power
Short-break systems
11.1 Introduction
Iii this chapter we shall examine the types of emergency and standby power systems that are available
to the user who has a need for an alternative or
additional electrical power source to that which is
normally obtained from, say, a public utility. Reserve sources of power may be needed for critical
devices and equipment in installations such as
hospitals, communication and computer centres,
continuous-process manufacturing, military security, airports, and marine navigational systems.
An emergency power supply is defined as an
independent reserve source of electrical energy for
critical equipment whose failure would jeopardize
the health and safety of personnel or result in
damage to property. There are mandatory requirements in most countries for the provision of emergency power supplies where unacceptable risk to
human life or health results from the loss of the
primary power source.
A standby power supply is an independent reserve
source of energy which provides power of acceptable quality and quantity so that a user's facilities
may continue in satisfactory operation on failure or
outage of the normal power supply source.
Emergency power supplies (EPS) are specifically
associated with life safety and life support systems.
Standby power supplies (SPS) are usually associated
with production processes, data and communication
systems, and building 'utility' systems (such as heating, refrigeration,
etc.). The power systems
described in this chapter are applicable to both
fields. The terms 'emergency' and 'standby' are used
fairly loosely here to describe systems which may
serve multiple purposes. Either term may be
appropriate to the needs of a particular user's application.
11.2 Plant categories
Reserve plants are normally categorized according
to the duration of the supply interruption tolerable
to the installations they serve. They are identified in
one of two groups:
1. short-break systems; and
2. no-break systems.
The first relates to installations able to tolerate a
break in supply from about one-half second to 60
seconds. The second category infers that no discernible interruption in supply is acceptable to the
serviced load. Plant designed to meet this requirement must be capable of providing enough stored
energy to cover the time-gap between primary
power failure and the availability of a standby power
source. The most commonly used energy stores are
secondary batteries and kinetic devices such as
flywheels. Plant in this category are generically
referred to as uninterruptible power supply (UPS)
11.3 Short-break systems
Where a supply break of 15 seconds or more is
tolerable a fully automatic standby generator will
usually fall into this category. For shorter break
times a continuously running plant or a UPS may be
necessary. An alternative system using the kinetic
energy stored in a flywheel and described in Section
11.3.1 has been successfully applied on Category II
international airports, where supplies to critical airfield lighting may not be interrupted for more than 1
second. A feature of all standby power systems of
this type is that they incorporate some form of
power transfer switching and there must, therefore,
always be a period of power interruption to the load.
This may be anything from a fraction of a second to
10 or 15 seconds, depending on the system used.
11.3.1 Single generator schemes
The majority of standby generator installations are
of this type. Size of plant is determined by the
magnitude and nature of the essential loads which
have to be supplied during the period of the emergency. Special consideration should be given (as we
have seen in Chapter 5) to those load elements
which particularly affect the sizing of the prime
mover and/or its coupled generator. Such elements
would include induction motor starting, and capacitive, cyclic and non-linear loads. The generating unit
may be brought into operation either manually (e.g.
push-button started) or automatically, on primary
power failure. Choice is largely determined by the
permissible period of power supply interruption and
by the practicalities of effectively isolating essential
from non-essential loads within the installation. In
some circumstances it may be impossible to achieve
a satisfactory split between these two load categories
on an existing installation. Recourse may then have
to be made to manual supply-changeover switching
followed by a routine switching-off of all nonessential loads before the standby generator is
No attempt will be made in this discussion to
cover every eventuality. Typical arrangements only
are considered. Figure 11.1 shows the simplified
control circuit for a single standby unit started
automatically on primary supply failure. It is often
referred to as an automatic mains failure (a.m.f.)
unit. The mains failure sensing equipment provides
a signal to the prime mover's starter system. It
monitors the normal supply source for under- and
over-voltage faults, for phase failure and phase
reversal, and for frequency deviations beyond preset
Emergency and standby A.C. power supplies
of the transfer switch should be employed for electric starting sets. It should have a selectable boost
charging facility. This is highly desirable for commissioning and for maintenance operations. Some
generating sets are fitted with small engine-driven
battery charging alternators. Since they are only
effective whilst the engine is running they must not
be employed as the sole charging devices for standby
generator batteries.
Starter batteries are either of the heavy-duty
lead-acid or alkaline type. The latter, in its nickelcadmium form, is favoured for standby generators
because it can be left for long periods in any state of
charge. Its disadvantages are that it is bulkier than
the lead-acid battery of similar capacity, it is more
expensive, and it tends to have a larger terminal
voltage drop with heavy current starting drain [3].
See Appendix B of this handbook for further details
on batteries.
Starting aids
It will be appreciated that on short-break standby
sets of the type illustrated in Figure 11.2 there is
only one chance for the engine to start before the
output frequency drops below an acceptable level.
The flywheel is designed to sustain the impacting
load for only a given period of time before the
frequency limit is reached. It is essential that the
engine is successfully started and clutched onto the
spinning flywheel/generator assembly within this
period. Engines using conventional electric motor
starters can usually be up to full speed in about 6 to
8 s given some form of preheating of engine coolant
and/or lubricating oil.
For an engine to start successfully the heat generated within its cylinders during the compression
stroke must be higher than the self-ignition temperature of the fuel used. Higher cranking speeds reduce
both the leakage of air and the loss of heat from the
air during the compression stroke. Engine starting
capability is significantly improved by employing a
starting gear which gives an increased cranking
speed [4]. One successful method used on shortbreak (and no-break) engine/flywheel sets is to
arrange for the control system to immediately energize the engine output shaft clutch (clutch (2) in
Figure 11.2) on mains failure. This uses some of the
flywheel's kinetic energy to accelerate the engine
from standstill up to firing and self-sustaining speed.
This obviously calls for a clutch capable not only of
transmitting the required torque but also of handling
the heat generated by the resulting friction energy.
One must bear in mind that the clutch's slip speed
(i.e. the speed difference across it) is approximately
generator synchronous speed. Much depends on the
magnitudes of the rotational inertias involved. Generally speaking, elastic clutches relying on the use of
helical or spiral springs to store energy are preferred
to the friction types. The problem of heat dissipation
does not arise to any significant degree with elastic
clutches. In friction type clutches the energy released in the gradual pick-up of the driven friction
linings (a function of the total transmitted torque)
has to be discarded [5]. The torque capacity of the
clutch selected may have to be uprated by a service
factor of 2.5 to 3.0 (over the nominal required
value) for this type of starting duty [6].
Starting aids include:
1. pre-heating of engine coolant and lubricating oil
by means of thermostatically controlled immersion heaters energized from the primary power
2. electric glow plugs or rapid warm-up heater plugs
on the smaller high-speed engines;
3. proprietary automatic systems, such as those
marketed by Start Pilot, employing volatile
ether-based starting fuels which are pumped into
the engine air inlet manifolds to access the combustion chambers and promote combustion. Such
systems not only improve starting performance in
low ambient temperatures but they also increase
the engine's rate of load acceptance;
4. periodic pre-priming of lubricant. Oil fed from a
separate electric motor driven pump energized
from the primary power source is circulated at
periodic intervals through the oilways to flush the
cylinder liners and generally wet the engine moving parts in readiness for an automatic start.
Periodic priming also has the advantage of reducing the severe wear that can occur on cylinders
due to condensed combustion products when an
engine is at standstill.
High confidence levels can be expected from well
designed and maintained starting equipment. During the fuel crises of the 1970s, with their attendant
winter power shortages and interruptions, the British Post Office (which had some 600 MW of standby
power installations) reported an average of less than
one failure to operate in 600 incidents [7].
It is good practice to arrange for standby plant to
be 'exercised' at frequent intervals - typically once a
week, and preferably on load. Diesel fuel standing
unused for long periods is prone to deterioration.
Regular exercising also ensures that fuel in service
tanks is dispersed and 'freshened' by replacement.
Where manning is impracticable it is advisable to
incorporate automatic exerciser facilities within the
generator controls. This will allow the standby set to
be run to a convenient periodic programme which
mayor may not include load transfer. In any event
the controls should be so arranged as to ensure that
the standby unit immediately takes over the load
should power outage occur during any exercise
Short-break systems
11.3.2 Standby battery/inverter schemes
These schemes combine the energy-storage capabilities of batteries with inverters to give shortbreak standby supplies. They are particularly costeffective when applied to emergency lighting and
those signalling and telecommunication installations
where short interruptions of power supply are tolerable. Inverters are available from fractions of a
kilovolt-ampere in single-phase versions, through to
hundreds of kV A in modular form for 3-phase
The equipment consists of a storage battery,
rectifierlbattery charger, and solid-state inverter,
arranged in the form in Figure 11.4. The critical load
circuits are normally supplied direct from the mains,
hepce the term 'passive standby' often applied to
this type of inverter equipment. On mains failure,
switch Sl automatically closes to start-up the inverter. Switch S2then changes over and the inverter
feeds the load for as long as battery capacity allows.
Since the rectifier/charger is only required to charge
the battery and maintain it at float level and does
not supply any power direct to the inverter, it need
only have a small rating. Overall system efficiency is
high because the mains supply is normally used and
there are no losses due to conversion. Changeover
time is anything between 80 and 800 ms, depending
on the type and size of the battery and load transfer
switches used, and on the 'build-up' time of the
When the inverter is operating its output voltage
is proportional to the battery voltage and varies, to
some degree, with load. It therefore needs to be
regulated for all but constant loads, such as emergency lighting. The input operating range of the
inverter should also be matched to the total voltage
drop as the battery is discharged. See Figure 11.5 for
typical performance characteristics of small inverters [8].
Figure 11.5 Performance curves of typical inverters at
no-load and full-load
Battery charger controls are arranged to automatically provide a high rate constant-current 'boost
charge' after a period of battery discharge in order
to restore full capacity in the shortest possible time.
Once a predetermined terminal voltage is reached
the unit reverts to 'float charge'. It is important that
the batteries are well maintained; if not, there is a
danger that the inverter commutating circuits will
malfunction on start up. Some proprietary systems
incorporate self-checking circuits which switch the
system on periodically to feed a dummy load. If
performance "isoutside specification during such test
runs alarms are initiated. Should a high level of
reliability be required, an auxiliary inverter system
may be employed, and automatically connected on
failure of the primary system. In yet another arrangement two equally rated inverter units may be
used in parallel, each of which is capable of supporting the critical load. If one fails the other takes over.
In some standby systems certain inverter circuits
are used in their reverse direction to operate as the
battery charger. These are usually referred to as
'reversible inverter systems'. They came about
because it was realized that many inverter components (e.g. feedback diodes, output transformer and
filter choke) were being duplicated in the more
conventional systems [9].
The inverter is used as a battery charger in the
reversible system. When mains failure occurs circuits are reverted to their normal inversion mode by
means of contactors. The merits of this system are:
• overall cost is lower than for the type illustrated
in Figure 11.4;
• reliability is enhanced because of the commonality of components in the inverter and battery
charger modes;
• the equipment is more compact; and
• switching contactors are confined to a.c. circuits,
obviating the need for the more expensive d.c.
Where the critical load requires a 'clean' supply it is
possible to arrange for the inverter to supply power
Emergency and standby A.C. power supplies
continuously. Transfer to mains only occurs if the
inverter fails or for routine maintenance purposes.
Operating in this mode the rectifier/charger must
provide enough power to charge the battery within a
reasonable time after a period of mains loss. It must
also supply the critical load through the inverter. Its
rating therefore will be of the order of 1.5 times the
critical load demand.
Combinations of emergency generator and inverter systems are to be found on ships' installations, which must comply with the requirements of
the 1974 Convention for the Safety of Life at Sea
(SaLAS). This calls for a self-contained source of
power of sufficient capacity to afford a supply for all
services necessary for the safety of passengers and
crew in an emergency. Due regard must be paid to
those services that may have to operate simultaneously. The period for which such power must be
provided on passenger ships is 36 hours.
The power source may be either an emergency
generator with an independent fuel supply or a
battery, which must be capable of carrying the
emergency load without recharging or excessive
voltage drop. The British Authority stipulates that
the voltage drop (at the end of the 36 hour emergency period) must not exceed 12.5 % of nominal
system voltage. Also, the voltage should not be
more than + 10 % in the fully-charged condition.
Any accummulator chosen as the sole source of
emergency power is likely to be of significant physical size. It must also be maintained in a fullycharged state at all times. Little wonder, therefore,
that most operators install a generator as their
source of emergency power. SaLAS stipulates that
when a generator is used for this duty, an accumulator has also to be installed. The accumulator
must be of sufficient capacity to automatically and
continuously supply specified services (such as
emergency lighting and electrically operated watertight doors) for 30 minutes while the generator
is started. This complementary battery-powered
supply is normally derived from a solid-state inverter.
11.3.3 Multiple generator schemes
It may be necessary or desirable on larger installations or where loads are highly critical to provide
two (or more) standby units. Several configurations
are possible and some typical schemes are shown in
Figures 11.6 to 11.9.
generating sets
The first scheme, illustrated in Figure 11.6(a), is
similar to that shown in Figure 11.1 but with the
addition of a second selector transfer switch - ATS
2. If the utility supply fails (a so-called blackout
condition) or if it drops below acceptable levels (a
brownout condition), the system controls start both
standby units. The circuit breaker of the designated
lead unit closes as soon as its generator reaches
correct voltage and frequency and the load is connected through transfer switch ATS 2. Once the lead
unit accepts the load and system stability is reached,
the other unit is shut down. The slave unit is started
and put on line if the running set fails at any time
during operation. A master 'duty switch' gives selection of either set as the lead unit. When the utility
supply is restored the load is switched to it and the
duty set is stopped.
The scheme of Figure 11.6(b) shows another
'standby-to-standby' arrangement. Controls are arranged so that the operation of switch ATS 2 is
'slugged' to give it a longer operating time than ATS
1. On mains power failure, standby Unit 1 assumes
the load. If it does so stably within the time delay
period preset for ATS 2 the start circuits of Unit 2
are inhibited. Should Unit 1 fail at any time, standby
Unit 2 will start and transfer to the critical load
through ATS 2.
An alternative configuration is to have each of
two standby units supply its own separate portion of
the essential load through sectionalized bus bars
which are provided with interconnection facilities.
This enables a standby supply to be available when
one machine fails in service or is taken out of service
for maintenance purposes [10].
On installations where there are two or more
levels of load criticality, systems such as those shown
in Figure 11.7 may be employed. The mode of
operation is much the same as that for the system in
Figure 11.6(a). The lead set (or the first machine to
reach acceptable output) is connected to the emergency terminals of transfer switch ATS 1 through
A TS 3. The latter may be a multi-pole double throw
switch or a suitable arrangement of electrically and
mechanically interlocked contactors. A TS 3 simultaneously connects the emergency terminals of mains
transfer switch ATS 2 to the second generator.
When that generator reaches acceptable output
ATS 2 transfers the lower priority load to it. Should
the lead set fail at any time, the priority load is
transferred to the other set. On restoration of the
primary power source both loads are reconnected to
it and the two standby units are shut down.
Paralleled generating sets
Where the decision is made to install multiple
standby sets several alternative arrangements are
possible. Some of these are described below.
Figure 11.8 shows a typical two-unit scheme.
Again, either set may be selected as the lead unit,
which is to supply the higher-priority essential loads.
The second unit feeds the lower-priority loads and
Short-break systems
Figure 11.6
Basic schematics for dual standby unit
acts as standby to the lead set. If the normal power
source fails the master controls start both sets. The
lead machine is connected to the emergency bus bar
through its circuit breaker as and when its output
reaches an acceptable level. Transfer device ATS 1
then connects the first priority load to the bus bar.
After the second set has started it is automatically
synchronized with the lead set and paralleled to the
emergency bus bar . ATS 2 transfers the lowerpriority load to this busbar. Load sharing devices are
then operated to ensure that both sets assume equal
or predetermined proportions of the total essential
If there is sufficient combined capacity in the
paralleled sets other less critical loads may then be
manually switched to the emergency bus bar. In
schemes of this type the lead set should be able to
take-up the higher-priority load within 10 to 20 s.
Automatic synchronizing would add a further 10 s
delay before transfer switch ATS 2 can be operated.
Figure 11.9 shows a multiple standby generator
system designed to provide an emergency power
supply for three levels of critical load within commercial building complexes. In such installations,
the load priority classifications could well be as
Priority 1. Any loads related to safety of life and
property and where provision of an emergency
supply is legally mandatory. These may include
evacuation lighting, emergency health-care facilities, fire protection, and communication systems.
Priority 2. Those loads of a less critical nature,
but where emergency power is also a legal requirement. These may include emergency lighting in
working and machinery spaces, elements of the
Emergency and standby A.C. power supplies
Figure 11.8
Basic schematic for two parallel-running
standby units
vertical transportation system, and critical controls
for environmental conditioning services.
Priority 3. Those loads where power loss may
cause personnel discomfort or damage to a process
or to a building's plant and facilities. These may
include mechanical utility systems such as boilers,
compressors, water pumps, ventilation fans, and
space conditioning in key areas.
The multiple generator system works in much the
same way as the dual-generator system described
earlier. It will have additional controls to handle the
starting, synchronizing and paralleling of the greater
number of standby units, and the sequential switching of the priority loads.
11.3.4 Standby generators in peak load
The installation of parallel running generators for
purely standby purposes is very seldom a good
economic proposition. There must, perforce, be a
degree of in-built redundancy in such systems since a
similar service can be provided by a single standby
set. The latter, however, offers less reliability as far
as higher-priority essential loads are concerned.
Where two equally-rated sets are employed the
probability of supply failure to the more critical
elements of the load is halved. Having said that, it is
worth mentioning that, above a rated output of
about 800 kV A, the capital outlay for two generating sets may actually be less than that for a single
generator of the equivalent total capacity even when
taking automatic synchronizing and load sharing
controls into account.
One way of improving the economic viability of a
limited-use standby installation is to use it to reduce
peak demands on the public utility supply. A peaklopping (or peak-shaving) plant may be defined as
one that is designed to generate all or part of a
consumer's maximum electrical demand during national or regional peak demand periods. Its purpose
Short-break systems
is to reduce the Maximum Demand (M.D.) tariffs
levied by supply authorities on purchased electricity.
These demand charges are additional to unit
electricity charges and are related to the highest
value of units metered in any half-hour during the
utility's accounting period. As measured in the
United Kingdom, the M.D. (in kW or kVA) is twice
the largest number of kilowatt-hours (or kilovoltampere-hours) supplied in any 30 consecutive
minutes during the month for monthly and seasonal
tariffs, or during the year for annual tariffs. The
demand curve in Figure 11.10 illustrates the method
of measurement over a 30 minute period. If 200 kW
are drawn for the first ten minutes, 550kW for the
next ten minutes, and 450 kW for the last ten
minutes, this represents an average of 400 kW taken
from the supply over the half-hour. This is equivalent to 200 kWh. By definition, the M.D. is twice
this figure, i.e. 400 kW.
Maximum demand is perhaps the most unpredictable element of a consumer's electricity bill because
it is the cumulative effect of many coincidental
loads. This is evident from the graph of Figure 11.11
[11]. The fluctuation in demand over a typical
working day may be as shown in Figure 11.12. The
area under the graph is the energy consumed. While
the peaks may represent relatively small energy
requirements they can equate to very high Maximum Demand charges. One way of reducing these
charges is to eliminate or transfer this energy to
other times in the working day. Murray [11] cites
examples in manufacturing and processing establishments where significant reductions were achieved by
shedding loads that were not essential to the production or the treatment process, during peak periods.
This approach is known as 'demand control'. Where
there is difficulty in identifying sufficient 'sheddable'
load that will give any realistic savings, the alternative method of using idle standby generators to
Figure 11.12 Typical daily load curve
supply power for peak loads offers a way of obtaining financial return on the investment in standby
If the consumer's M.D. tariff is a monthly one it is
almost certain to vary seasonally - the highest
surcharges coinciding with those months in which
the public utility itself experiences its greatest demands. In the United Kingdom these are the winter
months and the M.D. tariffs reflect this. Typically,
the demand charge may increase by 350 % for the
months of November to March compared with the
April to October period. Therefore, if the consumer
can generate power in the November to March
period, attractive savings can be made.
Two modes of on-site power plant operation are
1. independent of the utility supply; or
2. in parallel with it.
If the first method is chosen, the installation's load
distribution network must make it possible to supply
isolated circuits independently.
The isolated load option
Figure 11.13 illustrates two typical schemes for
independent peak-power generation. Both offer a
method of effectively utilizing the otherwise idle
Emergency and standby A.C. power supplies
Figure 11.13
Typical independent peak-lopping schemes
standby generator of the schemes described in Figurel1.6
In the first arrangement (a), the utility supply
feeds the emergency loads. The standby generators
are paralleled and brought on-line on receipt of a
signal from a monitoring power zone relay or a
prediction-type maximum demand controller set to
detect any instantaneous peak demand that exceeds
a preset level. Should the primary power supply fail,
the peak loads are disconnected and the generators
power the emergency loads through the transfer
switch. The alternative configuration of diagram (b)
is very similar. In this case the peak load controlled
by circuit breaker CB 3 may either be left on or
dropped for the emergency.
The advantage of both schemes is that the emergency loads are supplied with power of short-break
duration - since the standby generators are already
running. Variations on these themes are possible.
These might include push button starting of a selected duty set at an agreed hour in each day of the
'peak' tariff season. This set may run alone until it is
loaded to, say, 80 % of its capacity. The second (or
slave) set is then paralleled with it, either manually
or automatically. As load demand decreases the
slave set may be off-loaded and stopped.
The parallel function option
When it comes to operating peak-lopping generators
in parallel with the utility supply, certain technical
requirements must be met and compliance with
them verified (by regional Electricity Supply Boards
(ESB) in the UK). The conditions governing the use
of private generating plant may vary for each UK
Area Board. They all include a requirement for
compliance with the Electricity Council's guidelines
G59 - Recommendations for the connection of private generating plant to the Electricity Boards' distribution system.
In essence, the following conditions apply [7, 10,
12, and 13].
1. the need to have the design and installation of the
plant vetted by the supply authority;
2. the need to have the plant operated by competent personnel who are always in attendance
and ready to take instructions from the supply
authority's control engineers;
3. parallel operation should not prejudice either the
utility supply or the consumer's private system,
nor should it introduce hazards to the utility's
Short-break systems
4. means should be provided for preventing power
being fed into the utility's network (unless 'exporting' has been specifically agreed with the
public supply authority). This infers the provision
of a reverse power relay operating into the
control circuit of the incoming supply circuit
5. compensation is required against low power
factor - to minimize the reactive energy drawn
from, or fed to, the supply authority's system.
This implies that the peak-lopping generator(s)
should operate at a power factor not greater than
that of the normal site load;
6. the neutral earthing shall conform to the supply
authority's regulations. These are:
• the Electricity Supply Regulations 1988 (as
amended); and
• the Energy Act which deals with private generating stations and requests by private generators and suppliers.
In the parallel-running mode, only the incoming
supply system is earthed at the neutral point. A
neutral earthing switch should therefore be provided and so arranged as to be open when the
peak-lopping units are in parallel with the utility
supply and closed when generating independently. The earthing and interconnection of generator and incoming supply neutrals should always be discussed and agreed with the supply
authority. It is occasionally possible to obtain
permission to use the supply authority's earth
system and avoid the need to provide a separate
independent earthing system. Consideration
should also be given to the question of whether 3or 4-pole transfer switching is required for the
standby arrangements. Much depends upon the
earthing system employed. See Sub-section
10.4.1 of Chapter 10.
Where parallel-running with the primary power
source is intended, it is not necessary to designate
particular loads as peak loads. Nor are the peakload transfer switching devices of Figure 11.13
necessary. Figure 11.14 shows a typical arrangement
for a utility-paralleled peak-lopping system. On loss
of incoming supply the peak-shaving operation is
suspended and the generators are transferred to the
emergency loads. The duration of supply loss is
governed by the changeover period of the transfer
switches - since the generators are already running
and on-line. In this configuration engine governor
and generator excitation controls are arranged to
ensure that the peak-lopping generators only assume that portion of the load which is in excess of a
preset peak demand limit. (See Chapter 8 and the
discussions related to Figure 8.17 in particular.) The
generators continue to do this for all peak loads up
to their joint capacity. Thereafter any increases in
demand are absorbed by the utility supply. This
arrangement offers the most economical method of
'peaking' operation.
General considerations
Multiple generator installations used in standby
and/or peak-lopping functions should include units
which have the individual capacity to supply peak
load and sufficient combined capacity to meet consumer's emergency requirements.
When appraising the economics of on-site peaklopping schemes certain cost factors should be considered. Any UK Area Board will expect the consumer to meet the costs associated with modifying or
installing equipment on its supply system to enable it
to accept the possible fault conditions imposed by
connection to the consumer's power system [12].
Many area board tariffs include a service charge
which reflects the consumer's potential needs and is
closely related to his maximum demand. It may be
considered to be a payment for the board's having to
maintain its capital equipment in readiness for the
consumer's demands. Where the consumer operates
his own generating plant, be it in a standby or in a
peak-lopping mode, this service charge appears in
special tariffs as a standby charge.
There is a subtle difference between the two
charges. The service charge is a specific payment
(per kV A) for the demand that the consumer agrees
he is likely to raise within the year. The standby
charge is a payment for the difference between what
the consumer said he was likely to require and the
demand which he actually makes in the accounting
period. It follows therefore that when negotiating a
special tariff with the board the consumer must
ensure that the hypothetical demand on which his
standby charge is based is a realistic one [14]. Any
sizeable charges that then arise would be entirely
due to unscheduled outages of his plant. It would be
prudent to allow for this contingency when preparing cost analyses for on-site generators.
The greatest economy would appear to be obtained if peak-lopping is one-third of the maximum
demand. If emergency power needs are approximately two-thirds the M.D. a two-set installation
(with one unit peak-lopping as in Figure 11.14)
seems to be the ideal. The area board then does not
have to reserve capacity as a contingency against
failure of the on-site generators. Generally speaking, if not less than 75 % of a consumer's total
annual electricity consumption is taken from the
area board, normal tariff rates should be obtainable
Of course the very highest reliability (close to
100 % security) is only achieved using the utility
supply with peak -lopping unites) operating in parallel
Emergency and standby A.C. power supplies
Figure 11.14 Typical peak-lopping system employing
parallel operation with a utility supply
lei with it and with standby capacity also available
from the utility. The only likelihood then of a
complete supply outage on site is if the peak lopping
unit(s) fail during a utility blackout.
Induction generators
It is appropriate to mention at this stage the robust
and economically attractive induction generator,
and to say that its use would not be applicable to the
schemes we have been discussing. It is effectively an
induction motor or asynchronous motor driven
above synchronous speed by its prime mover and
drawing its magnetizing current from other synchronous machines connected into the system, or from
capacitive loads within that system [16]. It is best
employed when run in parallel with a larger capacity
supply system. Accordingly, it finds favour on those
applications where on-site generators use the byproducts of a local processing plant (e.g. biogas
from effluent treatment works and creameries, and
refuse derived fuel (RDF) from waste landfill sites)
in co-generation schemes, supplying combined heat
and power locally and exporting any surplus electrical power to a utility. Where the generator relies
on the power system for its excitation a fault on the
system would result in the generator's output current quickly decaying. A well-planned protective
relay scheme is required to remove the generator
from the system busbars in the event of a fault.
The induction generator is not the best choice for
'black start' applications [17]. It needs selfexcitation by a variable capacitive device if it is to be
suitable for isolated operation. An automatic voltage regulator is used to switch-in capacitive VAr as
the load on the machine is increased. Deleterious
overvoltage and magnetizing current transients may
be experienced on capacitor-excited induction generators when load is suddenly removed from them.
Unfortunately, this is the modus operandi in standby
generator systems. The protection equipment in
induction-generator installations should always be
arranged to stop the prime mover [18]. A selfexcited induction generator may be operated in
No-break systems
parallel with other generators under certain conditions. On failure of that connection the machine will
continue at a voltage and frequency determined by
its control equipment [19].
11.4 No-break systems
These may be defined as those systems which give
continuous voltage and frequency within the limits
of the load elements connected to them. The limits
will vary with the nature of the connected equipment and their supply tolerance specifications.
Typical examples are illustrated in Table 11.1 [20].
No-break systems may be conveniently sub-divided
1. those that use rotary machines; and
2. those that employ purely static or solid-state
Group 1 systems characteristically employ kinetic
energy storage devices to help maintain frequency
until the primary emergency power source is
brought on line. The stored energy source in Group
2 systems (and sometimes the supplementary source
in Group 1 plants) is usually an accumulator. There
are also sub-groups of these two categories employing both rotary and static equipment in, so-called,
hybrid systems.
11.4.1 Kinetic energy systems
Systems in this category use the kinetic energy
stored in rotating masses. This energy is released to
a driven generator to maintain frequency during a
short period of primary power interruption until
input power is restored from a standby source.
Engine-flywheel units
The several variants of this type of plant are
described in the diagrams and texts that follow. The
original flywheel-clutch arrangements shown in Figures 11.15 and 11.16 have largely, if not entirely,
been superseded over the last 15 years by the more
refined and efficient configurations. Descriptions of
these earlier versions are included merely to delineate the origins of their modern counterparts.
Table 11.1 Typical supply parameters for technical loads
The earliest enginelflywheeUmotor/alternator
(EFMA) sets were of the type shown in Figure
11.15. The critical load is connected to the generator
and there is no direct electrical tie between the
mains supply and the load. When the primary power
supply is healthy (as shown in Stage 1), the generator and the heavy steel flywheel are driven by the
induction motor (on the higher-rated sets this was
usually of the wound-rotor type with a 2 % full-load
slip. Additional requirements were that it should be
capable of accelerating the flywheel and generator
rotor from standstill to full speed on no load, and
re-accelerating the plant on full-load from about
90 % speed to normal operating speed. The motor
was required to give this performance against a 10 %
reduction in supply voltage - an essential requisite
for least-developed countries where many of these
plants were installed.) The engine is stationary and
the magnetic clutch is de-energized. Transient fluctuations in the supply frequency and voltage are
'buffered' by the flywheel energy and the generator's a. V.r., respectively.
On any deviation from preset limits (Stage 2) the
incoming supply to the motor is interrupted and the
generator is powered by the energy stored in the
flywheel mass. The engine is simultaneously and
automatically started. This may be either electrically, through starter motors, or by direct
compressed-air admission to the cylinders. The
clutch is energized when the engine is at, or near,
synchronous speed. The engine then takes the generator load and also supplies the windage and friction losses of the motor and the flywheel. Some
versions of these plants employed the more drastic
method of immediately energizing the clutch on
supply failure. This cranked the engine at high
speed - in a similar manner to that described in the
short-break arrangement of Figure 11.2.
On positive restoration of the primary supply the
motor is reconnected to it. The magnetic clutch is
de-energized and the fuel supply to the engine is
cut-off. The plant then reverts to Stage 1 operation.
Advantages of the system
• When the unit is operating off the primary supply
the motor-generator 'cushions' the critical load
against voltage fluctuations caused by the switching of other large loads on the supply, and against
Emergency and standby A.C. power supplies
Figure 11.15 No-break diesel generator incorporating
separate induction motor and generator
all transient frequency deviations. In this sense,
the primary supply is effectively isolated from the
Moderate capital and maintenance costs.
The system can operate in an emergency for a
considerable time, limited only by its fuel storage
It is not necessary to synchronize supplies when
returning to primary power drive because the
load is always fed from the generator.
Quick recovery of flywheel energy after supply
Both electrical machines may be of the brushless
type. (Where very large inertia flywheels were
required on units of high power rating, the induction motor was usually of the slip-ring type.)
Disadvantages of the system
• The combined motor-generator efficiency is low
(approximately 85 %) so that there is wasteful
energy loss where loading is continuous for long
• The load's steady-state frequency is always less
than that of the utility supply; typically 2 % less.
• Since only a very small percentage of the kinetic
energy stored in the flywheel can be used without
exceeding acceptable frequency tolerances, extremely large flywheels are required for high
power ratings. This makes the total assemblies
rather bulky. Flywheels are, typically, rated to
sustain the load for 10s during changeover to
engine drive. The optimum economic design tolerance for frequency dip is 8 % to 10 %. This
would be intolerable to many modern technical
loads (see Table 11.1). (The larger flywheels were
flame cut from rolled steel plate to BS 970 Part 1
of EN 43A 50 carbon with an ultimate tensile
stress of 63 kgf/mm2 and a yield stress of 28 kgf/
mm2, and machined. They were then dynamically
balanced using weights fitted and screwed into
dovetail slots around the sidewall periphery of the
wheel. Typically, the 10 s wheel on a 500 kVA,
50 Hz, 1500 rpm unit (giving a frequency drop of
10 % during changeover to engine drive) weighed
8940 kg and had a diameter of 1750 mm and
thickness of 460 mm.)
• The large inertias not only make synchronizing a
lengthy procedure but also make stable parallel
operation impracticable [21].
The system illustrated in Figure 11.16 used a
combined-duty synchronous machine running as a
motor and driving the flywheel while the primary
power supply was healthy. On power failure the
clutch was energized and the engine was cranked
and started. The synchronous machine then ran as a
generator providing power to the critical load for the
duration of the emergency. The major advantages of
this system over the one described in Figure 11.15
were economic ones.
• The overall physical size and the cost of the unit
was reduced by the elimination of the separate
induction motor.
• Under normal utility supply conditions the motor
could be used as a synchronous condenser to
improve the installation's power factor.
No-break systems
The disadvantages were;
• The critical load was not electrically isolated from
the utility supply.
• When the switch in the primary power line was
opened (and during the transition from motor to
generator mode) there was some distortion of
supply to the critical load.
• The motor/generator had to be re-synchronized
with the restored primary supply.
One of the few diesel generating set manufacturers
who still make EFMA no-break units is the British
firm of Auto Diesels Power Plant. Its equivalent of
the older arrangement shown in Figure 11.15 uses a
fluid drive coupling and a one-way clutch combined
in a special unit (Figure 11.17). This permits torque
to be transmitted from the diesel engine to the
generator and flywheel but not in the opposite
direction [22]. The design of the clutch is such that it
can transmit a positive torque to the generator even
when the engine speed is below the generator/
flywheel speed. This means that from the moment
the engine attains self-sustaining speed the clutch is
helping to arrest the run-down of the flywheel. The
cushioning effect of the fluid converter-coupling also
ensures progressively smooth application of the load
to the prime mover. This avoids any risk of stalling
the engine.
Initial start-up of the unit is under manual control.
The engine is started and the clutch unit drives
Figure 11.17 Layouts of no-break EFMA sets using
combined fluid coupling and one-way clutch power
transmission drive (Courtesy: Auto Diesels Power Plant Ltd)
the generator/flywheeUmotor assembly up to speed
without any need for operator intervention. Once
the set has reached synchronous speed utility supply
and load are connected and the engine is stopped.
Equipment controls may then be set for the automatic mode.
The characteristics of the clutch unit enable
smaller flywheels (than those of earlier EFMA sets)
to be used. This results in lower costs and a reduction in the overall size of plant.
The manufacturer normally designs for a maximum frequency droop of 5 % (i.e 2.5 Hz at 50 Hz),
after 10 s of primary power supply failure. Where
well-maintained engines use pre-heating and starting aids much better performances (as low as 1 Hz)
are achieved in practice. The output characteristic of
an EFMA set of the type just described is shown in
Figure 11.18.
In a considerable refinement of the combined
synchronous machine system of Figure 11.16, Holec
N.V. uses the kinetic energy of a two-flywheel mode
induction coupling to replace the conventional
heavy flywheel. The composite diagrams of Figure
11.19 help explain the operation of the system.
The coupling comprises two concentric rotating
parts, the outer of which is fitted with a 3-phase a.c.,
and a d.c. winding. It is mechanically coupled to the
synchronous machine. The inner rotor is separated
from the engine by a free-wheel clutch. Both rotors
Emergency and standby A.C. power supplies
Typical output characteristics of an EFMA set
using a fluid drive coupling (Courtesy: Auto Diesels Power
Plant Ltd)
Figure 11.18
are capable of spinning independently of each other
and of the housing frame (see Figure 11.19(a)).
During normal operation (with the utility supply
healthy) the synchronous machine operates as a
motor and drives the outer rotor of the coupling at
1500 rpm on 50 Hz supplies. Excitation is applied
through shaft-mounted slip rings to the 4-pole,
3-phase winding of the outer armature. Both rotors
are then 'magnetically' linked. The inner rotor spins
at 1500 rpm relative to the outer rotor. In absolute
terms it is rotating at 3000 rpm.
On utility supply failure power flow is interrupted
to both the synchronous machine and the 3-phase
a.c. winding of the induction coupling. The outer
member of the coupling now drives the synchronous
machine which changes over from motor to generator operation. At the same time the d.c. winding of
the coupling is energized through a frequency measuring and control circuit and a start signal is given
to the diesel engine. Power torque is produced, by
eddy-current coupling, between the inner and outer
members of the coupling. The magnitude of this
torque is related to the difference in speed between
the two rotors and to the intensity of the magnetic
field (i.e. to the intensity of the current flowing in
the d.c. winding).
Thus in the first few milliseconds after normal
power supply failure the kinetic energy of the outer
rotor (alone) provides the power for the no-break
load. Thereafter the inner rotor dissipates its energy
(while it decelerates from 3000 to about I600rpm)
to the outer rotor which continues to drive the
generator. About 50 % of the initial KE of the inner
rotor is, therefore, delivered to the generator during
the changeover from synchronous motor to engine
drive. Since the outer rotor has also provided kinetic
energy the induction coupling may be described as a
'two-flywheel' device.
The diesel engine runs up to synchronous speed in
less than 3 seconds. It is able to do this because it is
free of any external inertia load. The free-wheel
clutch is energized when engine and coupling inner
rotor speeds coincide.
During emergency operation (Figure I1.I9(d))
the output frequency is kept constant irrespective of
engine speed variations. This is done by closed loop
control of the d.c. winding in the coupling.
On restoration of the utility supply the output of
the no-break set is automatically synchronized with
it. The synchronous machine then reverts to motor
duty. The free-wheel clutch disengages and the
inner rotor of the coupling accelerates to its (original) 3000rpm absolute speed. The diesel engine is
stopped some three minutes later.
During normal operation (Figure 11.19(b)), the
feed to the synchronous machine is taken from a
75 % tapping on the 'line inductor'. The reactance of
this portion of the inductor is three times the
subtransient reactance (Xn of the machine. The
of inductance
and idle-running
No-break systems
Emergency and standby A.C. power supplies
machine provides a stabilizing (or conditioning)
filter in the supply to the critical load. It acts as an
autotransformer maintaining the output voltage to
the load at 100 % and also attenuates transient
voltage disturbances and wave-form distortions. In
theory all peak and interference voltages should be
filtered out. In practice a reduction of some 95 % of
the higher harmonics in the utility supply is obtained. Residual values are within the tolerances
imposed by the most voltage-sensitive loads. During
emergency operation the UPS wave-form naturally
follows that of the synchronous generator.
The speed/time characteristics in Figure 11.20
summarize the dynamic conditions during the various phases of operation. Typical dynamic performances on failure of, or short-circuits on, the utility
supply, or during load surges up to ±50% are:
• maximum voltage excursions of -8 % to +6 %
for up to 50ms, recovering to nominal level
within 1 s; and
• frequency deviations of ± 1 %.
Power factor at the incoming utility supply terminals
is, typically, 0.98.
This type of UPS system cannot alter the frequency of the primary power supply. It should not be
installed in those areas where frequency problems
are experienced with utility supplies. A bypass
circuit using make-before-break switching is fitted as
Although an equalizing current flows within the
inductor circuit during changeover to utility supply
system output remains within specified tolerances.
There is also a phase displacement across the inductor. It has a maximum value of 30° at full load.
This diesel UPS system is available in three basic
1. Single systems using one UPS set with a bypass
circuit. Since load and utility supply voltages are
always in synchronism the unit can be by-passed
at any time. Load may thus be transferred direct
to the utility without the need to synchronize the
two supplies. Utility and UPS are said to be in a
permanent state of mutual hot standby-by. See
Figure 11.21(a) .
2. Parallel systems of two or more UPS units with a
common bypass circuit. The collective rating of
the units is sufficient to cover the total load
requirements (Figure 11.21(b».
3. Redundant systems of two or more UPS units
operating in parallel with a common bypass
circuit. The collective power rating exceeds the
load requirement by the rating of at least one set
(Figure 11.21(c».
Normal operation
Changeover to emergency operation
Emergency operation
Changeover to normal operation
Inner rotor of induction coupling
Synchronous machine and outer rotor of induction coupling
Diesel engine
Slip of induction coupling
Time between return of normal power supply and disengagement of free-wheel clutch
Return to condition 1; the diesel engine is shutdown
Figure 11.20 Speed/time characteristic of the diesel UPS
system shown in Figure 11.19 (Courtesy; Holec NV)
No-break systems
(c) Redundant system with bypass circuit
Four 400 kVA units are employed,
for a 1200 kV A UPS rating.
The power of one UPS unit is therefore
redundantly available.
Circuit breakers CB 1 and CB 2 should be
rated for 1600 kVA, so that at a later
stage in load growth, all 4 UPS units
may be operated in parallel as for (b)
Rgure 11.21 Examples of the application of diesel UPS
systems in various configurations (Courtesy: Holec NV)
The bypass circuits allow one or more sets to be
bridged for repair or routine maintenance purposes.
They also operate automatically in the event of unit
failure (in systems 1 and 2) to give a direct and
uninterrupted supply from a healthy utility. In the
case of the redundant mode system 3, security of the
critical load is maintained (albeit without reserve)
on failure of anyone UPS unit.
Full-load efficiencies lie between 86 % and 93 %,
depending upon the system configuration used and
on its power rating. Half-load efficiencies are between 83 % and 89 %.
The outline diagram of Figure 11.22 illustrates
unit equipment construction. Typically, the leading
dimensions of a 250kVA set are 5040mm long,
1250mm wide and 1510 mm high. The weight of the
unit is 7625 kg. Reading from left to right, the base
frame mounted assembly comprises the synchronous
compound machine with a tachogenerator, the induction coupling, the free-wheel clutch, and the
diesel engine complete with its starter battery.
The composite equipment cabinet consists of:
1. the diesel panel;
2. the induction coupling and generator control
3. the utility supply, load and by-pass switching gear
The line inductor/stabilizing filter equipment is
housed in the free-standing cubicle.
A typical diesel UPS installation is shown in
Figure 11.23. This illustrates a parallel-running
system consisting of two 300 kVA units (in the
foreground) supplying uninterruptible power for
lighting and signalling circuits for the Drechttunnel
near Rotterdam in Holland.
Emergency and standby A.C. power supplies
No-break systems
Rotary machines with backup energy sources
Systems in this group may be subdivided into those
that use d.c. drives and those that employ a.c.
drives. Considerable ingenuity has been demonstrated by designers of such UPS plants and many
configurations have been used. We shall confine our
attention to those most frequently applied.
D.C. drive systems. These systems use an accumulator as the emergency energy source. Batteries
may be the traditional vented lead acid and nickel
cadmium types or the increasingly popular sealed
lead acid type using gas re-combination (RE) technology. See Appendix B of this handbook for
descriptions of the various types, their electrical
characteristics, and their charging parameters.
The first system we shall examine (see Figure
11.24) uses the principles described in Section 11.4.1
but substitutes a battery bank and a coupled d.c.
machine for the diesel engine/flywheel/inductioncoupling drive of Figure 11.19. The battery is sized
to cover the expected duration of utility supply
outages. It is customary to design for periods of 15
to 30 minutes.
Under normal operating conditions the synchronous machine is connected to the utility supply - in
parallel with the critical load. It combines, as we
have seen, with the inductor to act as a stabilizing
filter in the network. The d.c. machine runs as a
generator float-charging the battery bank in readiness for emergency duty. The feeder breaker opens
on utility supply failure. The synchronous machine
then derives its power from the battery through the
d.c. machine, which now runs as a motor. On
restoration of the utility supply the d.c. machine
reverts to its generator mode and automatically
regulates the battery recharging current to levels
determined by the battery's discharged condition.
It is customary on battery-supported
systems to include a conventional standby diesel
generator as backup to the main supply source. This
enhances the UPS bridge-over capability following a
mains supply blackout. The standby generator is
started manually or automatically on prime-power
failure. It must have sufficient capacity to satisfy the
critical load demand and simultaneously provide the
power needed for recharging the battery.
The more conventional motor-generator scheme
is shown in the block schematic of Figure 11.25, and .
a typical three-unit installation of this type is illustrated in Figure 11.26. Such multi-set systems either
employ a common battery bank or have their individual unit-batteries paralleled.
The utility supply is converted to d.c. by matched
transformer/rectifier equipment. The d.c. is fed via a
starting and surge-limiting resistance (or a choke
coil) to the motor, which is coupled to a generator
suppling the critical load. The MG assembly incorporates a flywheel which is either supported on two
pedestal bearings or is mounted directly (if it is small
enough) on one of the machine shafts. The purpose
of the starting resistance, SR, is to control the motor
current during the run-up period. It is shorted-out in
graded steps by time-delayed contactors as the
motor approaches normal speed.
Thereafter, motor speed is controlled by a highaccuracy closed loop speed regulator, ASR. A voltage proportional to the speed of the motorgenerator set is obtained from a shaft-driven tachogenerator. This voltage is compared with a reference
level within regulator ASR and if any difference
exists the regulator adjusts the motor speed to hold
the generator frequency within specified and preset
The rectifier output float charges the battery
through a charge-limiting resistance (CR) when the
battery circuit breaker is closed. (Other proprietary
systems may use a separate fully-automatic battery
charger.) When the utility supply fails the a.c. input
contactor (UC) opens. The battery discharging
through diode D drives the motor. As battery
potential falls, ASR maintains the motor-generator
speed (and therefore the output frequency) within
specified limits. Typical performances, at 100 %
load changes are:
• ± 10 % VOltage variations for 150ms; and
• frequency deviations of ± 1 %.
Figure 11.24
Two 30 kVA battery-rotary UPS systems
installed in a Netherlands Post Office transmitter station
(Courtesy: Holec NV)
The length of time for which the battery is capable
of supporting the load is determined by its capacity.
Where major UPS installations require several
hours' stored-energy time it is often more economical to provide a standby diesel generator rather
than invest in a large capacity battery. Increasing the
Emergency and standby A.C. power supplies
No-break systems
battery size also involves increasing the size (and
cost) of the transformer/rectifier unit. Increasi.lg,
for example, the capacity of a 500 V nickel cadmium
battery on a 250 kVA system from a lO-minute to a
20-minute rating requires an additional battery bank
10m long, 0.7m deep and 104m high, and weighing
about 6000 kg.
Where high system reliability is sought it is usual
to arrange for several units to operate in a parallel
redundant mode.
A.C. drive systems. The simplest configuration
uses a motor-generator (FMG) set fitted with a
suitable flywheel to provide a continuous buffer
against 'spikes' or transients on the mains supply.
Following a primary power interruption the bridgeover time during which output frequency can be
maintained within prescribed limits is proportional
to the ratio of flywheel inertia to critical load. For a
2 % drop in frequency only 4 % of the initial kinetic
energy can be extracted [21].
The drive motor may be either a synchronous
machine or a low-slip wound rotor induction motor.
The set's steady-state output frequency is directly
proportional to input frequency. The decay in output frequency on loss of power is linear down to
85 % of speed [22]. A standby generator should be
incorporated if power interruptions exceeding all
but those of the shortest duration (up to 1.5 s) are to
be bridged (see Figure 11.27). The flywheel should
be sized for at least a 10 s power interruption, in
which case critical loads which cannot tolerate a
frequency drop of about 5 % should not be served
by this type of UPS system.
The system is commissioned using starting resistances designed for the long run-up time required
to accelerate the flywheel to full speed. If the FMG
unit is started on a standby generator supply, the
latter must be capable of accommodating the high-
impact starting load (see Chapter 5). The standby
generator should also have close-frequency governor controls if it is not to be affected by the
transients imposed by the peripheral equipment of
some critical loads. These may include essential
air-conditioning systems for computer installations,
off-line processing equipment, emergency lighting,
and could add up to a demand 5 to 6 times that of
the critical load itself.
Its simplicity, reliability, and use of wellestablished technology makes the FMG the most
popular equipment for 'buffering' large critical
loads. It should not be used on those loads which
cannot tolerate a frequency drop of 5 %.
IEEE Std 44~ 1980 [1] describes a system which
uses an eddy-current clutch between the flywheel
and the generator of the FMG set. An eddy-current
clutch works on the varying-slip principle and is
often used in drive systems. Varying the excitation
current to a winding (or to electromagnets) within
the clutch is equivalent to varying the supply voltage
to a conventional induction motor. The result is a
variation in torque for a given slip speed. In theory,
the speed range can go down to zero. It is always
limited, in practice, by the heat which can be
dissipated within the clutch [23].
On loss of primary power the generator receives
the stored energy of the flywheel. As it slows the slip
of the eddy-current clutch is automatically reduced
to hold the generator at synchronous speed. Equipment using this arrangement is capable of maintaining output frequency at ±0.5 %. Also, power is
sustained to the critical load for up to 15s after loss
of primary power. This gives any standby generator
a reasonable time in which to be automatically
started, run-up, and switched-in.
A serious disadvantage of the FMG system is its
low overall efficiency - of the order of 85 % because of the in-series mode of motor and generator operation. The situation worsens when the
FMG gets its power from a standby generator.
The overall configuration then becomes a threemachine, in-series system and the standby generator's efficiency has to be included. In these circumstances total system efficiency is, at best, some
11.4.2 Static systems
By far the most significant trend in the last 20 years
has been the swing away from rotary to static
technology in the design and application of UPS
systems. We have seen in Section 11.3.2 how the
stored energy in an accumulator may be used in
conjunction with a solid-state inverter to provide a
short-break standby power system (Figure 11.4).
Using the same basic elements, several UPS configurations are possible. Some of the more common
Emergency and standby A.C. power supplies
ones are described in this section. We shall first look
at the inverter 'block' itself.
Static inverter types
There are three basic designs to be considered.
Variants, offering combinations of their individual
features, are possible. A typical example of one such
'hybrid' will be described later.
The oldest of the inverter types is that which uses
a ferroresonant saturable transformer circuit. In
common with all inverter designs its first stage is a
d.c. filter whose function is to protect the circuit
from transient voltages in the d.c. source. The next
stage is a controlled bridge whose thyristors are fired
by an oscillator to give a square-wave output. The
final (and multi-purpose) element is a ferroresonant
transformer which shapes the wave-form and limits
output current. It employs an air-gapped, flux.saturated core. This combines with a secondary
circuit tuned to the inverter output frequency to give
voltage regulation (of the order ±5 %) and harmonic suppression. Output is restricted by heat losses
in this constant voltage transformer (c. v.t.), and
inverters of this type usually have a competitive
upper rating of about 50kV A. They may, however,
be used in multiples to give high output levels.
The design most frequently used on large threephase equipment is the stepped wave type. This
employs two inverter stacks per phase, operating in
antiphase to 'chop' the d.c. input into two rectangular wave-forms. These are fed into a summing
transformer to give a stepped (stairstep) wave-form.
In a three-phase arrangement the firing cycle of each
phase inverter is displaced 120° from the other two
and each is individually voltage-regulated. Since the
wave-form is so close to a sine wave, a relatively
small a.c. filter is required to remove harmonics and
give a near-sinusoidal output wave-form. The design
suffers from the disadvantage of a slower transient
response to input voltage and load step changes than
the other two types described. It does, however,
offer a very convenient way of combining thyristors
for high power outputs.
The most popular design in the medium power
range has tended to be that which uses pulse width
modulation (PWM) - sometimes referred to as a
fourth generation inverter. It works on the principle
of controlling the amplitude and repetition frequency of a continuous train of pulses generated from the
d.c. source (see also Section 5.4.3 of Chapter 5).
Again, because the output wave shape approximates
to a sine wave, a relatively small a.c. filter is
required. This design gives good transient voltage
response to input and load changes.
The introduction of power transistors has resulted
in very reliable and efficient low cost PWM inverters. Transistors require no commutation circuits, can be electronically protected, and, there-
fore, have good fault handling capabilities. The
higher switching frequencies made possible by transistors have resulted in even smaller filters, giving
improved speed of response to load changes.
PWM inverters using thyristors suffer from the
disadvantage that they require a fairly complicated
switching circuit and more commutations per load
cycle than those of the stepped wave design.
Because they have fewer components they are
potentially more reliable than stepped-wave designs.
In an effort to secure the best features of both
PWM and stepped designs in the middle range of
power ratings, one manufacturer has produced a
hybrid 3-phase design which is essentially a Step/
PWM inverter [24]. Using fewer power thyristors
than the conventional stepwave, and lower switching frequency than the PWM, it gives a higher
operational efficiency than would otherwise be obtained from the PWM inverter. The spectrum and
magnitude of the output wave-form require a lower
series inductance in the output filter than in the
stepped wave design. This enhances transient response and load compatibility.
Static inverter UPS configurations
Some of the more commonly used systems are
described below and are illustrated in single-line
block form.
The On-line system. During normal operation the
UPS system is continuously supplying power to the
critical load (see Figure 11.28). The rectifier/charger
has the following functions:
• to provide power to the inverter;
• to maintain the battery bank at full capacity; and
• to supply the UPS losses (the typical efficiency of
a PWM type inverter is 83 % )
So that, for example, a 12 kW rectifier would be
required for an inverter rated at 10kVA, 0.8 pJ.,
The battery and inverter filter a.c. line transients
provide regulated frequency and voltage to the load.
The battery is normally 'floating' on the d.c. input to
the inverter. When a.c. input fails the battery
supplies power to the inverter for the specified
period until it has discharged to a predetermined
minimum voltage.
The bypass switch may be manually or automatically operated. In most systems the inverter is
continuously operated in synchronism with the utility supply (in terms of both phase and frequency).
Uninterrupted bypass to the primary a.c. source is,
therefore, possible if the inverter fails. See Subsection 11.4.4 for descriptions of bypass switching.
The inverter is usually designed to switch to its
own crystal-controlled internal frequency standard
No-break systems
the battery-inverter circuit of (b) makes the
inverter less reliable; the inverter has also to
work harder in order to cope with the uncontrolled d.c. link.
3. The minimum supply voltage to the rectifier in
(b) must be greater that the maximum 'boost'
voltage, or oscillation between 'on mains' and 'on
battery' modes will occur.
4. Configuration (b) may be advantageous where an
existing charger is large enough to handle battery
recharge but not the inverter load [25].
Where extended primary power outages are the
norm the UPS system may be combined with a
standby diesel generator to provide power for indefinite periods. The configuration is used to maintain
an interrupted supply to less critical peripheral
loads, such as air conditioning and lighting in a
computer installation (see Figure 11.28(c». Diesel
engine start is usually delayed for primary power
outages of short duration. When on standby generator operation, the inverter's synchronizing circuit is
usually automatically disconnected. Stable frequency is, therefore, maintained at the critical load
- regardless of any large impacting loads on the
generator. Otherwise it is necessary to use a closeregulating governor on the prime mover and/or
introduce additional inertia into the standby generator's rotating masses.
The power required from the standby generator
may be expressed as [26]:
Figure 11.28 Continuous UPS system configurations
when the frequency of the utility supply deviates
from acceptable limits. It then automatically reverts
to the synchronized mode (within about 1 s) on
restoration of normal supply frequency.
Some systems may use a separate battery charger
and an unregulated rectifier to feed the inverter (see
Figure 11.28(b». The following points will have a
bearing on choice between system (a) or (b) of
Figure 11.28:
1. Arrangement (b) is less reliable because two
power switching circuits are required: the cost is
higher for this reason, and also because a larger
UPS cubicle is required.
2. While a rectifier may be more reliable than a
rectifier/charger, the blocking (isolating) diode in
is the generator output power (kW)
is the UPS output power (kW)
is the UPS system efficiency (usually
between 89 % and 93 %)
is the battery recharge power required
by the UPS
is the sum of the critical peripheral
loads which must be supplied by the
A is typically 15 % of the full load input kV A of a
single-module UPS. The true power requirement
(kW) is approximately 9 % of the UPS input power
- because of the low power factor. On a multimodule redundant UPS system there is always one
extra module to be considered. For a system with n
modules the maximum battery input power would
be (1.09n - I)A. It may be possible to limit the
battery charging during operation of the standby
supply in order to reduce the generator's capacity.
The UPS system manufacturer should be consulted
on the practicality of this approach.
If a 100kW UPS is the only load on a generator
Pa = (100/0.89)
= 122kW
Emergency and standby A.C. power supplies
Most standby generators use turbo-charged engines
which are capable of accepting only about 60 %
rated load in one step (see Section 2.5 of Chapter 2).
This means that the 122 kW requirement would
need a continuous rating of (122/0.6) kW = 203 kW.
Generator ratings of the order of 2 to 2.5 times UPS
rating are necessary with 6-pulse unfiltered
chargers. This requirement is reduced for 12-pulse
chargers and is further reduced if a diode charger is
used with no input transformer (B. Beech, 1989,
personal communication).
A feature available on some proprietary UPS
systems is a so-called input walk-in facility. This
ensures that the system draws no power on 'mains'
restoration other than transformer inrush current.
Then, over a period of some 15 s (the walk-in time),
the input power requirement is raised to a level
dictated by the sum of the power required to drive
the critical load and that preset power needed to
recharge the batteries.
Despite this soft start or ramp feature the full UPS
load is still taken-up by the generator in less than a
second. The reason will be clear from Figure 11.29
[26]. With no a.c. supply to the UPS the partially
discharged battery provides the inverter current.
The UPS rectifier (and therefore the generator) will
not supply current during the earlier part of the
ramp when its voltage is less than that of the battery.
The current demand shifts from battery to rectifier
when the rectifier voltage exceeds that of the battery - point A of the ramp characteristic. This shift
is, effectively, a step change in the generator load.
The ramp feature, therefore, does little to help with
the problem of turbo-charger lag (see Section 2.5 of
Chapter 2). Moody [26] reports that some interesting work is being undertaken by UPS manufacturers
to extend the ramp time over this critical period, the
objective being to reduce, or even eliminate, the
need to oversize generators which use turbo-charged
The reader will come across the term hot standby
operation. In this arrangement the inverter is
switched on but the critical load is fed from the
utility supply. On failure of the utility the load is
switched to the inverter. Because the inverter is
already in operation the voltage break to the load is
only about 50 to 100ms. The battery charger must
be capable of charging the battery, maintaining it in
its fully charged state, and supplying the no-load
current demand of the inverter.
Parallel system configurations. Several configurations are possible. They are usually arranged so as to
give some measure of redundancy in installations
• reliability is of paramount importance; or
• the primary power source is of such poor quality
that the critical load may not be exposed to it; or
• the utility supply frequency differs from that
required by the load.
Redundancy implies the presence of adequate reserves ready to be brought into immediate use.
In the two-paralleled-units arrangement each half
of the system is rated to the critical load demand.
Where multiple paths are used an additional unit
(over and above the number required to satisfy the
critical load demand) is installed to provide redundancy.
Figure 11.30 shows alternative approaches. The
first, (a), is essentially the same as the on-line
system of Figure 11.28, but it uses a second inverter
as an alternative source to the primary a.c. supply.
Transfer between inverters is performed by the
static switch, which gives only a small fraction of a
cycle of disturbance to the critical load. There is no
significant phase discontinuity (see Sub-section
Scheme (b) is for an application requiring large
blocks of critical power. Two or more 'on-line'
systems are operated in parallel. This calls for active
and reactive load-sharing equipment, and sensing
and protection logic to quickly detect individual
system problems. For example, a defective voltage
regulator reference which, unless it is isolated,
would drive the common bus voltage out of limits
due to the action of the load-balance circuits. Although not shown in Figure 11.30(b), static interrupters are usually included in each inverter line to
the common critical bus. They rapidly 'clear' a
malfunctioning inverter element and prevent serious
disruptions to the other systems or to the critical
load. The operation of the interrupters must be
rapid if the load is to be unaffected. It is not
sufficient to wait for load commutation. Forced
commutation switches are therefore necessary. This
has the disadvantage, however, that load is normally
carried by the switches and they have to operate
under fault current condition. An alternative ar-
No-break systems
figure 11.30 Typical parallel redundant systems
rangement uses filter circuits to limit the amount of
energy fed back into the failed element by the
'healthy' inverters. The power components of these
filters tend to be expensive [27].
Duplicate power and energy source elements are
shown in Figure 11.30(a). This is not usually necessary. The proven reliability of batteries will often
justify the installation of a common battery bank. In
the same way, two regulated d.c power supplies may
well suffice without unduly affecting overall system
reliability (see Figure 11.30(b».
11.4.3 Hybrid systems
We have already considered some hybrid systems in
Sub-section 11.4.1. The advantages of their good
load-carrying characteristics
capability are offset by:
and high overload
their relative inefficiency;
their physical size;
the complexity of their installation; and
the requirement for frequent maintenance particularly those systems using d.c. machines.
The first step taken towards improvement was to
replace the d.c. motor with a synchronous motor
driven by an inverter (see Figure 11.31). Several
benefits accrued. One was the elimination of a
commutator and brushgear. Another was faster
speed regulation and dynamic response - making it
possible to eliminate the flywheel or at least reduce
its size. A further spin-off was inverter control apd
Emergency and standby A.C. power supplies
Figure 11.31 Hybrid UPS system using an inverter and
synchronous motor-generator
commutation off the motor terminals. No complex
commutation circuits were required. The inverter
thyristors could be commutated by the reverse
e.m.f. in the motor and were, therefore, not subjected to high-voltage commutation spikes.
While this configuration achieved many of its
targeted criteria, it still did not give the efficiency
levels required. A radical change in the design of the
rotary machine was necessary in order to achieve
West German machine specialists Anton Piller
tackled the problem (in the early 1980s) with their
type DH 'hybrid' UPS system. This uses a power
converter (the Unibloc) which combines the windings
of both synchronous motor and generator in alternate slots of compact housing (see Figure 11.32). The
pony motor is used to commission the machine.
The 'Unibloc' operates as a rotary transformer.
When 3-phase a.c. power is applied to the motor
(primary) windings a rotating magnetic field is
created. This makes the d.c. -excited salient pole
rotor spin at synchronous speed. By controlling the
rotor excitation the secondary (generator) voltage
may be fixed - irrespective of primary (input) volt-
No-break systems
age variations. This means that relatively large
'mains' voltage disturbances are not passed on to the
generator output windings. In practice, even complete short-duration losses of a.c. primary supply do
not cause loss of a.c. output. High-voltage spikes
and frequency variations are also attenuated.
The other major difference between a conventional transformer and the Unibloc machine is that a
transformer will transmit harmonics from its primary to its secondary windings. The rotor of the
Unibloc is fitted with a set of damping windings.
Any harmonic voltage applied to the input winding
induces a rotating field in the usual manner.
Because the rotor is spinning at a speed related to
the fundamental frequency any harmonic of the
fundamental will produce a rotating field at a speed
higher than that of the rotor. It may not necessarily
be in the same direction of rotation but it causes an
alternating flux to be induced within the damper
winding. This has the effect of removing the harmonic or, at least, attenuating it. The Unibloc may
therefore be considered to act as a narrow band-pass
filter for the fundamental mains frequency.
More significantly, the composite machine
replaces the d.c. motor (2 bearings), flywheel (2
bearings) and the generator (2 bearings) of the
conventional FMG set with one 2-bearing machine.
Eliminating the flywheel and six bearings reduces
windage and friction losses and enhances the efficiency of the system. Also, because the Unibloc has
two completely separate sets of windings on its
stator, the frame size for a given rating is increased,
e.g. a 300kVA machine has, approximately, a
600 kV A frame size. This improves system efficiency. The larger its frame size the more efficient a
machine tends to be. Despite this, the Unibloc is
smaller and lighter overall than comparable dualmachine sets.
The basic single-line schematic of the UPS system
is shown in Figure 11.33. The Unibloc forms one of
two paths through the system. The second path is
through a simple diode rectifier, the emergency
battery, and a static inverter. As with the conventional FMG system, the inverter needs only to be of
a very simple design since it derives its commutation
from the Unibloc windings. Only six thyristors are
used. Up to 40 such devices may be necessary in
conventional forced-commutation units. This must
enhance system reliability.
In normal operation both circuit paths are used,
95 % of the power passing through the Unibloc. The
5 % flowing through the second path ensures that
the rectifier/inverter is always in an active state. This
increases system efficiency and allows the thyristor
inverter to idle in a constant proving mode. The
parallel path effectively introduces redundancy into
the single-module UPS system.
On primary power failure the battery/static inverter supplies the critical load. The system uses a
separate constant potential battery charger rated to
automatically recharge the battery to 95 % of full
charge within a period of ten times the discharge
Emergency and standby A.C. power supplies
period. As with the fully-static arrangements
described in Section 11.4.2, operating periods may
be indefinitely extended by installing a standby
generator to support the primary energy source.
When the unit runs off the standby generator
supply, the power path to the Unibloc is normally
through the rectifier and inverter. This is because
the standby generator's output cannot be assumed
to be continuously within frequency tolerance. A
signal from the standby generator ensures that this
path is selected by the DH system's controls.
Because the Unibloc's output voltage is controlled
by rotor excitation, reactive current may flow between the utility supply and the Unibloc, and vice
versa. The purpose of the series choke coil is to limit
these exchanges of reactive current. It is designed to
do this for input voltage variations of ± 15 %.
The Unibloc rotor stores a small amount of kinetic energy which helps maintain frequency stability.
Should a short-circuit occur on the input to the UPS,
power would be fed into the fault from the synchronous motor and this KE would be removed from the
system. To prevent this happening a thyristor switch
(arranged to open on reverse power) is inserted in
series with the Unibloc. It is always opened on
mains failure.
Parallel-running configurations are possible for
high power demand installations. The redundancy
feature built into each module may preclude the
necessity for parallel/redundant connections (i.e.
n + 1 modules) but this additional level of security
can be provided.
The UPS system (excluding battery bank and
battery charger) is packaged in a single, compact,
and acoustically treated housing. Noise level is of
the order of 70dB(A) at 1 metre. The enclosure
requires access on all four sides, which means it
would probably require more space than an equivalent static UPS system.
The power factor of the system varies with output
rating and is between 0.7 lagging and unity. It is not
feasible to use the synchronous motor for power
factor correction. The windings should not be overexcited, because they are used to regulate the Unibloc's output voltage.
11.4.4 Bypass switching
No UPS is truly uninterruptible unless it has some
form of automatic bypass device which transfers the
load to an alternate source of a.c. power on system
malfunction. Much depends on the input power
quality parameters required by the critical load fed
from the UPS system. There can be considerable
variance between computer and data processing
equipment manufacturers' requirements. The shorttime disturbances that their equipments will tolerate
are directly related to their in-built 'ride-through'
(or 'carryover') times. While frequency is an impor-
tant parameter, equipment limits (of the order of ± 1
to ± 1.5 %) are usually well within the performance
levels of modern UPS systems. Tolerances on voltage disturbances (and total outage) are of much
greater significance. Conditions are specified in
terms of amplitude and time duration. Typically,
they may be:
• transient dips of 25 to 30 % with recovery to
±5 % of nominal voltage within 150 ms;
• over-voltages up to 200 % for less than 200 f.Ls;
• a total maximum outage limit of anything between 2 and 20 ms.
Modern UPS systems of the type discussed in this
section are capable of meeting these transient requirements.
The user is concerned with what happens when
the UPS system goes into the bypass mode. Most
systems work on the 'reverse transfer' principle. In
normal operation the critical load is fed from the
UPS. It is only transferred to the utility supply on
UPS malfunction or (in the case of static systems) on
overload, caused perhaps, by a fault in a load branch
circuit (see Section 11.4.5)
Bypass devices may vary from:
• make-before-break
electromechanical switches
on rotary systems, which have inherent 'bridgeover' energy (magnetic or kinetic) during a
transfer; to
• static switches, which are an essential requirement for a truly uninterruptible power supply on
inverter-based systems.
Some equipment uses both mechanical and static
switching devices. For example Emerson Electric's
Accupower UPS module may use a static switch
operating in parallel with a bypass circuit breaker.
The signal to initiate a transfer is simultaneously fed
to circuit breaker and static switch. The static device
(consisting of back-to-back thyristors that conduct
load power when gated) closes within 5 f.LS.The
'mechanical' breaker then takes over from it.
When the load is transferred to a 'stiff' (large
capacity) utility supply that source will show no
transient drop in voltage. Depending upon the inverter type, a significant transient drop could be
experienced at the critical load's terminals when
load is switched to the UPS system. With PWM and
stepped wave designs the transfer may cause a
complete interruption for between 2 and 10 ms (see
Figure 11.34). This is not the case with ferroresonant type inverters because of the ride-through
provided by their stored magnetic energy.
A manually operated bypass circuit should be
provided external to the UPS module so that the
entire UPS can be bypassed for maintenance purposes. A word of warning here. Many manufacturers provide internal circuits which they claim give
No-break systems
Figure 11.34
Reconstructed wave-form of UPS output
voltage when transferring the load to a utility supply
through a static switch
this facility. On closer inspection this often proves
not to be strictly true, and input power is present at
the input and output terminals of the UPS module.
If this is so, the only safe approach is to use external
isolators and a circuit breaker (or a contactor) - as
shown in the single-line configuration of Figure
11.35. The bypass-for-maintenance routine would
then be:
1. The internal bypass circuit must be activated.
2. The external bypass breaker/contactor must be
closed, and only after this has been done may
isolators 1 and 2 be opened to give safe access to
all parts of the UPS module.
Switching the module back into service with zero
break in supply to the load entails closing isolators 1
and 2 and then activating the internal bypass circuit
before opening the external bypass-switching
11.4.5 Load protection co-ordination
One of the problem areas associated with UPS
systems is that of co-ordination of their protective
devices with those of the loads they serve. If a
branch load fault develops, sufficient current must
be 'generated' to trip the branch load's protective
device and avoid the loss of other loads. Unlike the
a.c. generator in a rotary UPS system the static
inverter does not have significant overload capacity.
We have established (in Section 4.3 of Chapter 4)
that the short-circuit current/time characteristic of
rotating machines is affected by the machine's subtransient reactance (X"d), its transient reactance
(X'd), and by its synchronous reactance (X'd) - in
that chronological order. The machines in rotary
UPS systems are likely to have an instantaneous
asymmetrical short-circuit current of 10 to 20p.u.
This decays to about 2 to 2.5p.u. within O.ls and
may rise again to a level of 3 p.u. within a further
0.1 s if a current-forcing circuit is employed. The
final level can be sustained for a further 10s without
thermally overstressing the machine.
A static inverter cannot generate these magnitudes of current. The best overload characteristics
are those obtained from ferroresonant type designs,
which can supply up to 150 % full load current
continuously. The most one can expect from PWM
and stepped-wave designs is an overload capacity of
150 % for 10s. Overcurrent capability is also limited
and is usually of a similar order, i.e. 150 % for a few
seconds. Above its overload capability, the UPS
system's output voltage is reduced progressively and
output current is automatically limited to a preset
figure (see Figure 11.36). If the currents of impacting loads (such as motor starting inrush currents or
fault currents) exceed the preset level, the static
bypass switch is used to transfer the demand back to
the 'stiff' primary power source. It does this without
interrupting the supply to the critical load.
Figure 11.37 shows typical prospective current/
time characteristics for a generator and a nonferroresonant inverter of similar rating. These characteristics are superimposed on a family of curves
for fuses to BS 88 Part 2 (lEC 269-2), and
Emergency and standby A.C. power supplies
Figure 11.37 Typical prospective current/time
characteristics of rotary and solid-state UPS configurations
compared with H.R.C. fuse characteristics
demonstrate the need for introducing the back-up
energy of the primary supply source whenever a
solid-state UPS system is used. This will ensure the
timely clearance of any protective devices in branch
load circuits.
11.4.6 System reliability
The availability (A) of any system is expressed in
terms of its mean time between failure (m.t.b.f.) and
the mean time to repair it (its m.t.t.r.):
(m.t.b.f. x 100%) -:- (m.t.b.f.
+ m.t.t.r.)
where the periods are all in the same units of
time - usually hours.
M.t.b.f. is a measure of the average time that a
particular item of equipment will remain in service,
while m.t.t.r. is related to the time it takes to repair
and return it to service. Factors that affect the
timescale of the repair operation are [28]:
1. the methods used to confirm that a defect exists;
2. the speed with which a defective component or
assembly is identified;
3. the rectification or replacement of the defective
devices; and
4. verification of correct function after repair.
The mean time to repair may vary from tens of
minutes to 24 hours, depending upon the complexity
and severity of failure. Where designs have been
based upon:
good component accessibility;
a sound modular assembly approach;
in-built diagnostic circuitry; and
where an adequate set of spares is stocked on site,
repair times as low as 30 to 40 minutes can be
achieved. This presupposes that adequately skilled
personnel are immediately available for repair
System m.t.b.f.s of 50000 to 200000 hours are
claimed by manufacturers of static and rotary UPS
systems. One obvious method of improving system
reliability is to introduce some measure of redundancy. Others are:
1. to reduce the component count; or
2. to provide bypass or isolation of failed devices or
It is important that the prospective user establishes
the source of the data upon which vendors' availability predictions are based. Where the equipment
offered has a proven pedigree and a verifiable track
record supported by documented field reliability
data, there should be no problem. Where a relatively new design is being considered reliability
claims should be substantiated by calculations based
on well-proven data obtained from sources such as
the United Kingdom Atomic Energy Authority, the
American Mil Standards Handbooks (e.g. MILHDBK-217), or from ANSI/IEEE 500.
11.4.7 Co-ordination of earthing systems
All technical loads are susceptible to electrical
'noise' input. This is usually caused by inductive
pick-up in power lines and by stray currents in earth
Some computer manufacturers, fully aware of the
electronic needs of their own equipment but not
necessarily familiar with power systems, recommend
computer clean earthing schemes that separate the
computer earthing system from that of the power
system. Not only is this unsafe practice but it also
contravenes most national codes (such as the American NEC, the British lEE, Wiring Regulations, and
No-break systems
the Electricity Supply Regulations) and is quite
It is important to avoid the creation of earth
loops, which are the commonest cause of inductive
interference. This is best done by employing insulated and radial earthing circuits to the individual
loads from a single point in the consumer's installation. This point may be the earth-bar at the utility
intake switchboard, the independent UPS source, or
at the computer system's own power distribution
unit (PDU). All three points should be bonded
together by an insulated conductor. It is also necessary to ensure that the metal enclosures of the
computer equipment are insulated from extraneous
metalwork within the building. This means that
insulated gland plates and insulated terminal boxes
should be used at the input to each individual load.
The load equipment enclosures should sit on insulated blocks or on an insulated floor. Where the
latter is chosen care should be exercised in selection
and construction in order to avoid problems with
'static' [29).
Where earthing electrodes are employed in the
consumer's installation (for example, in an authorized use of a protective multiple earthing system
(PME)) the use of separate electrodes for the computer load should be avoided. They may do more
harm than good. If they are used it is essential that
effective bonding to the main power system earthing
electrode is maintained. Where a high-frequency
earth grid is required it is usually located under the
raised floor of the computer centre and is formed by
the floor supports and stringers. The individual
cabinets of the computer suite should be connected
to this grid by the shortest possible insulated earth
conductors. The grid itself should be securely
bonded to the consumer's installation earth bar by a
single short connection [29).
11.4.8 Specifying and selecting UPS systems
The following comments are largely based upon the
sound advice offered by Holscher [25).
One is faced with a wide choice of equipment as a
result of the proliferation of UPS manufacturers in
recent years. Some are comparative newcomers into
the marketplace and cannot be expected to have the
field experience of the relatively smaller core of
experienced manufacturers. This does not imply any
inferiority since no one manufacturer has a patent
on acceptable technical performance. But as far as
the purchaser is concerned the best recommendation should be a substantial list of satisfied users.
Every manufacturer of static UPS systems seems
to have selected the inverter design type it will build
as standard, and naturally seeks to promote that
particular type. This, unfortunately, often leads to
confusing and misleading statements concerning
performance and compatibility with computer and
technical loads. The situation is often aggravated by
the tendency of some computer and data processing
equipment manufacturers to overstate their requirements.
Again, further complication may be introduced
when buyers' engineers, who are not necessarily
experienced inverter designers, write specifications
which combine the desirable features of a number of
different manufacturers' products. The result is a
build description that no one manufacturer can
meet, without exception. Most reputable suppliers
can meet each other's performance specifications
without much difficulty. The buyer's engineer
should keep his specifications relatively simple, with
emphasis on the operational requirements of the
technical load itself. Size and ambient conditions
should be given but no attempt should be made to
design inverter circuitry or construction. Rather,
vendors should be asked to describe their systems
and state their design philosophy and to give an
indication of the safety factors used in selecting
critical components.
In selecting suppliers it will be found that price
competition is fierce. The prospective user should
request a list of installations over the previous three
years of equipment similar to the type offered.
Market competitiveness unfortunately seems to
cause considerable 'turnover' in the industry and,
where any doubt exists, the financial soundness of
vendors should be established by a credit rating
11.4.9 Recent trends in static UPS systems
The size and cost of standard modules have reduced
considerably over the years. This continues, but at a
slower rate perhaps than in the last ten years. The
major advances will be in control and circuit
diagnostic techniques to reduce modes of system
failure, and will lead to simpler rather than more
complex circuits. Market pressures in computer
hardware have meant reducing prices for equivalent
performance. This is being reflected into the UPS
The systems described in this chapter have inferred the use of thyristors in inverter stacks and this
continues to be the case for high-power units. This
was also true of mid-range systems until about five
years ago when power transistor inverter technology
began to make inroads into this sector of the
market. The main barrier to their adoption had been
one of high component cost. This is no longer the
case and inverter systems using PWM power transistors controlled by CMOS digital logic cards are
now available in ratings up to 200kVA. In addition
to being easier to control, transistors require less
logic and offer better reliability than thyristors
(see Sub-section 11.4.2). Another spin-off is that
Emergency and standby A.C. power supplies
to be about 20 % higher in first cost for the larger
output ratings. This difference is more than offset by
installation costs on those plants which use diesel
While rotary no-break systems have been available
for many years, they were threatened to be overtaken by the rapid upsurge of static equipment in the
place. Rotary equipment
were quick to respond by applying the lessons
learned on static systems. If failures occurred on
rotary systems they were usually in the electromechanical components of control circuits and not
in the rotating machines themselves. The first stage
in the rotary 'comeback' was the adoption of solidstate control technology to effect improvements
system reliability. Significantly, with the advent of
such as that described under Figures
11.19 and Figure 11.33, the swing to static systems
has been arrested.
Some of the factors likely to
affect choice of either system are discussed below.
Maintenance. Apart from replacement of cooling
fans every five years or so, little maintenance should
be required on static systems. Routine maintenance
is minimal and is mainly concerned with ensuring
that cleanliness of interiors is preserved and cooling
circuits are operating effectively. On rotary systems
it is necessary to ensure that bearing lubrication is
at regular six-monthly intervals. This
may be done with the equipment in operation. A full
service necessitating shut down should be undertaken at least every two years. On such occasions,
system outage would be of the order of four hours
and repairs bias advantage in
favour of the rotary systems. Even the best training
courses offered by static system vendors are not
to give enough information
to repair
printed circuit boards or to provide component-level
of circuits. The user relies heavily on
the service support provided by the manufacturer
his accredited agent. The technology associated with
rotary systems is within the capacity of the competent electromechanical
technician. The fault assessment time on rotary sets is usually short. The
majority of faults occur in the control logic. User's
personnel often find difficulty in diagnosing static
system faults. Repairs may be quickly made but
system failure is likely to recur if the first diagnosis is
designs are considerably smaller
physical terms than their thyristor counterparts.
11.4.10 System choice considerations
Efficiency. There is now little to differentiate between the two systems, in this regard. The best of
both are on a par, with values between 80 % and
92 %, depending upon rating.
Acoustic noise levels.
Rotary plants using engines
as their primary power source will be noisy. Noise
levels can be as high as 102 dBA at 1 m (see Chapter
14). Those employing motor drives are likely to
have levels of 90 dBA at the same distance. Inverterbased systems may be expected to be of the order of
80 to 85 dBA. The nature of the noise (i.e. its
is quite different
in each case. That
emitted from rotating machines is almost entirely
due to 'windage' whereas that from static systems is
due to magnetostriction
in the iron cores of transformers and inductors - the so-called transformer
'hum'. Most observers find the low frequency of the
latter type of noise to be more irritating.
It is' not as easy to attenuate the noise in static
units as it is in rotary plant enclosures.
This is
because they require relatively large quantities of
cooling-air flow (approximately
30 %
more than rotary units of the same output rating)
along multi-directional
Performance and response times. Most static sets
use thyristors in their rectifier sections. Harmonics
may be introduced into the supply network, under
certain operating conditions (see Chapter 5). Rotary
sets tend to use diodes. Wave-form
distortion is
lower. While response
times and the regulation
of static
overcurrent capability is poorer than that of rotary
plant (see Sub-section 11.4.5).
11.5 References and bibliography
The easiest equipment to install is the
static types and the packaged rotary units. Systems
i.e. engines as the power source are
bulkier and heavier. They require larger foundation
masses and the use of vibration mounts, or sound
insulation pads, to'prevent
transmission of secondary noise to other parts of a building.
Cost comparisons. In a highly
it is injudicious to be precise in this
regard. Generally speaking, static equipment tends
11.5.1 References
1. IEEE recommended
practice for Emergency
and Standby Power Systems for Industrial and
IEEE Std 446-1980
- the 'Orange Book' Mar. 1980 published by
the IEEE
Inc. and distributed
by WileyInterscience.
2. Jones, D. and King, A.c. Standby power generation. A paper presented to the lEE's Second
Referencesand bibliography
Conference on Installation Engineering, Conference Publication No. 251 pp. 10-14 (lEE), June
3. Mahon, L.LJ. Diesel engine plant. Chapter 8
of Electrical Engineer's Reference Book, 14th
edn, Editors M.A Laughton and M.G. Say,
Butterworth, (1985)
4. Hodge, D. Starting gear and starting aids.
Chapter 14 of Diesel Engine Reference Book,
Edited by L.R.C. Lilly, Butterworth, (1984)
5. Tout-Kowsky, P. Factors in designing clutches.
Engineering, March, pp. 208-209 (1975)
6. Kent, A.J. Clutches and brakes Survey No. 42.
Engineers' Digest, 34(4) 59-65 (1973)
7. Shearer, A.B. Selection and use of standby and
peak lopping generating equipment. A paper
presented to the Diesel Engineers and Users
Association (now the Institution of Diesel and
Gas Turbine Engineers) February (1974)
8. Barness, B. Fail-on inverter power supplies.
Engineering July, pp. 574-577 (1977)
9. Smollinger, G.J. and Raddi, W.J. Reverse
energy transfer through an a.c. line synchronised pulse width modulated sinewave inverter.
A paper presented to the Third International
Telecommunications Energy Conference organised by the lEE in May 1981. Conference
Publication No.196 pp. 126-131 (1981)
10. Emergency Electrical Services, Hospital Technical Memorandum No.11, issued by the
Department of Health and Social Security and
obtainable through HMSO and booksellers,
revised (1974)
11. Murray, P. Saving cash through smoothing your
peaks and shedding your load. Electrical Times,
September 4, pp. 8-9 (1981)
12. West, G .S. Private generation - the interface
with the public supply. One of a series of articles
on Private Generation published in Electrical
Times, June 29, pp. 29-36 (1972)
13. Lake, T. Standby diesel generating sets - the
benefits of peak lopping operation and factors
affecting the choice of suitable plant. Electrical
Review, November 24, pp. 709-711 (1972)
14. Windett, AS. Reducing the Cost of Electricity
Supply, Gower Press Ltd. (1973)
15. Shearer, AB. Why private generation? Electronics and Power, February 8, pp. 37--40 (1973)
16. Eastham, J .F. Principles and characteristics of
induction generators. Electrical Review, November 11, pp. 809-813 (1960)
17. Murthy,
S.S., Jha, C.S., Singh, B.P.,
Vidyashankar, C.R., Nagami, C. and Tandon,
AK. Some studies on design and voltage regulation of capacitor self-excited induction generators. A paper presented at an lEE conference
on Electrical Machines - design and applications. Conference Publication No. 254 pp.
180-184 (1986)
18. Smith, I.R. and Sriharan, S. Transients in induction machines with terminal capacitors. Proceedings lEE, 115(4), 519-527 (1968)
19. Recommendations for the connection of private
generating plant to the Electricity Boards' distribution systems. Engineering Recommendation
G 59, The Electricity Council, June (1985)
20. Jones, D. Changing emphasis of power systems.
A paper presented to the Inpower '86 Conference on: Power systems now and for the next
decade, organized jointly by the Institution of
Diesel and Gas Turbine Engineers and Fuel and
Metallurgical Journals Ltd., pp. AI-AI4 September (1986)
21. King, A.C. Uninterruptible
power supply
equipments. A paper presented to the First
International Conference on Industrial Power
Engineering organized by the lEE. Conference
Publication No.272 pp. 38--42, December (1986)
22. Theunissen, S.G. No-break gen-sets for airports. Diesel and Gas Turbine Progress Worldwide, September (1975)
23. Laithwaite, E.R. Electrical variable-speed
drives, Survey No. 17. Engineers' Digest, October (1964)
24. Griffiths, D.C. A new hybrid UPS inverter. A
paper presented to the Third International Telecommunications Energy Conference organized
by the lEE in May 1981. Conference Publication No.196 pp. 156-160 (1981)
25. Holscher, John N. Static UPS systems: configurations and inverter specifications. Computer
Design, July (1974)
26. Moody, R.L. and Rounce, W.J. UPS and diesel
generating set compatibility. A paper presented
at a technical seminar jointly presented by Liebert Ltd., Emerson Electric Ltd., Dale Electric
of Great Britain Ltd., and Newage International
Ltd., March (1990)
27. James, M. Development and experience of UPS
systems during the 1970's. A paper presented to
the Third International Telecommunications
Energy Conference organized by the lEE in
May 1981. Conference Publication No.l96 pp.
201-204 (1981)
28. Shepherd, J., Morton, A.H. and Spence, L.F.
Higher Electrical Engineering, 2nd edn, Pitman
Publishing, reprinted (1971)
29. Westinghouse consulting application guide,
Catalog 55-000, 7th edn (1984-85)
11.5.2 Bibliography
While many of the references listed above contain
very useful cross-references to related papers on
topics covered, the reader's attention is particularly
directed to the lEE Conference Publications mentioned in [2] and [9]. In addition to the paper [2],
Conference Publication No. 251 contains the
Emergency and standby A.C. power supplies
following contributions which may be of help to the
researching reader:
Dilliway, J.G. Power requirements and systems, pp.
Garnham, E.J.B. Electrical Installations for large
computer systems, pp. 5-9.
May, G.J. and Turner, M. Sealed lead acid batteries
for standby power, pp. 15-19.
Conference Publication No. 196 [9, 24, 27] contains
several papers of interest under conference sessions
covering un interruptible power supplies, a.c. standby systems; reliability; and d.c. power plant and
lEE Conference Publication 287: Installation
Engineering - designing and maintaining successful
systems (lEE) June 1988 contains a number of
contributions on standby and un interruptible power
systems. Two papers of particular significance in our
context are:
Vernon, P. Uninterrupted power supplies today,
Hung, W.W., McDowell, G., Day, J. and Nutall,
R.E. An alternative approach to the design of
standby power systems, p. 28.
IEEE Standard 446--1980 [1] is a particularly useful
guide on this subject. Attention is directed to Section 3: General need guidelines; and Section 4;
Systems and hardware.
The following literature, reports and papers also
cover the various topics of this chapter in some
1. Everson, H.K. Emergency power - standby or
uninterrupted. Section 7 of Switchgear and Control Handbook, Edited by Robert W. Smeaton,
McGraw-Hill, (1977)
2. Jones, D.L. and Tetchner, G. Packaged units
produce total energy in creamery. Modern
Power Systems, November (1984)
3. Goodchild, G. Battery-based standby power
systems. Electrical Equipment,
4. Tanzer, R. Choosing static inverters for nobreak power supplies. Electrical Review,
201(11), 37-38, (1977)
5. Petrie, J.M. Static inverter power supply for
Electrical Review,
201(18), 29-31, (1977)
6. Marriner, G.R. No-break power evolves to
higher efficiencies with greater reliability. Electrical Review, 207(3), 30-31, (1980)
7. Taylor, M. Static u.p.s.s. - a guide for buyers
and users. Electrical Times, May 1, pp. 8-9
8. Taylor, B.G. Computer centre power engineering for the tropics. Power Engineering
Journal, 1(1), 37-42, (1987)
9. West, Ian Un interruptible power sources for
computer installations. Electronics and Power,
July/Aug. pp. 571-574 (1983)
to. Higgins, J.C. Problems in the specification and
assessment of electronic-equipment reliability.
Proceedings lEE, 113(9), 1413-1419 (1966)
11. Hinsley, AJ.A. Our earth. Electrical Design,
March, pp. 37-39 (1987)
12. Barnett, P. Changing trends in the design of
uninterruptible power supplies. Electronics and
Power, July, pp. 457-460 (1987)
13. Taylor, M. Choosing the right UPS. Electronics
and Power, July, pp. 461-463 (1987)
14. Harris, W.P. UPS systems - static or rotary?
Electronics and Power, July, pp. 465-467 (1987)
15. Bailey, AC. Transportation - Ships. Chapter
31 of Electrical Engineer's Reference Book, 14th
edn, Editors M.A Laughton and M.G. Say,
Butterworth, (1985)
16. Hung, W.W. and McDowell, G.W.A Hybrid
UPS for standby power systems. Power Engineering Journal, 4(6), 281-291 (1990)
Fuels and
lubricating oils
Oil refining
12.1 Introduction
This chapter describes how the various liquid and
gaseous fuels used in reciprocating internal combustion (RIC) engines are derived from petroleum
sources, and explains some of the more common
terms associated with crude oils and their refining
The quality of fuels is determined by a number of
parameters and by the proportions of impurities
within them. The methods used to measure these
physical properties will be outlined and the relevance of certain key properties, to engine performance will be examined.
In the same way, the significant properties of
lubricating oils will be defined and consideration
given to potential problems in service.
Those specifications which are most widely accepted as criteria for engine fuels and lubricants will
be reviewed
12.2 Petroleum
Petroleum is a remnant of organic matter, formed
mostly from creatures rather than plants. The main
chemical elements of its compounds are hydrogen
and carbon (hydrocarbons). It may be:
1. solid (e.g. kerogen, the oil-producing material in
shale) ;
2. liquid (as in crude oil); or
3. gaseous (natural gas).
The organic materials which are the source of petroleum accumulate in beds consisting of silt and mud.
A porous rock such as sandstone, then forms over
these source beds. Heat from the pressure of materials above the beds and from the earth's interior
enhances petroleum formation.
Liquid or gaseous petroleum forced out of the
source bed may escape from the sandstone unless
some peculiar geographical disturbance forms a trap
preventing its migration toward the earth's surface.
The petroleum explorer's task is to locate these
traps and drill into them. Priest reports [1] that in
1980 one in four drillings produced a commercial
petroleum well and one in fifty identified a significant new reserve.
The recovery of liquid oil from shale and tar sands
(sandstone containing natural tar in its porous structures) is hampered by unfavourable economics.
Shale oil recovery also faces technological, environmental and social barriers that may indefinitely
delay its becoming a significant proven energy
reserve [1]. There has been limited commercial
exploitation of Canadian tar sands.
Since the temperature-pressure-time
process is so variable, the composition of liquid
petroleum (crude oil) differs according to its source.
The quality from a particular source (a well or
producing area) may be quite constant over a period
of years. The major sources are now the Middle
East, Africa, Russia, Indonesia, North and South
America, and the North Sea.
Prior to the mid-1970s oil companies were able to
select crudes for their refineries to suit the finished
products required. Now, however, there is considerably less flexibility in the choices available. This
seriously affects the ability to make the fullest range
of products compatible with the demands of the
fuels and chemicals markets.
There is considerable variation in the yield (or
product breakdown) from the different crudes.
These variations are reflected in the finished products. The need to blend crudes to get a satisfactory
end product is exemplified by the sulphur content of
residual fuels, which is very expensive to control by
refinery processing. It is necessary to blend the
residue of a high sulphur crude (such as that from
Kuwait) with one from a low sulphur crude (such as
Nigerian) to give the low value required by engine
performance requirements and environmental considerations - such as the need to reduce acid rain
caused by sulphur oxides [2].
12.3 Oil refining
While it is sometimes possible to use crude oil in its
original state as a diesel fuel, most crude oils are
refined. (Contrary to popular belief most crudes are
not very viscous. They are fairly 'thin' liquids which
flow easily [3].) The prime purpose of refining crude
oil is to split it into a number of component parts
(fractions or cuts) of controlled properties according
to boiling range. The resulting fractions may then be
used directly or re-formed to make a range of
commercial products. See Figure 12.1.
The first step in the process is distillation (or,
more accurately, fractional distillation). Crude oil is
heated to partly vaporize it. The resultant mixture of
liquid and vapour is then fed into a fractionation
column (or bubble tower). The vapours rise and are
progressively condensed as the temperature falls.
The condensates are drawn off as side-cuts at various
points in the tower by means of liquid collecting
plates and draw-off pipes. The heavier fractions
condense in the lower part of the column. The
lighter fractions rise towards the top of the tower
before they condense and liquefy. Some do not
condense and remain as vapours (gas).
This is a very simplistic description of the fractionation process. The reader's attention is directed to
the text-books in [3] and [4] for a more detailed
treatment of the subject.
Reading from the top of the column the gas
fraction is progressively followed by gasoline, kerosine and naphtha, and a middle distillate called
Fuels and lubricating oils
sarily confined to the lighter fractions. There are
examples appropriate to heavier fuel oils which are
derived either directly or by the use of subsidiary
products of conversion - such as blending [5].
Those refining processes which are designed to
purify the fraction or product are known as treatment processes. Their function is to remove or
reduce the impurities inherited from the parent
crude oil. For example, the sweetening process is
concerned with the removal or neutralization of
harmful sulphur compounds from the lighter distillates.
The same process, using very high pressure and
temperature conditions, may be employed to reduce
the sulphur content of residual fuel oils. The equipment is expensive and fuel and catalyst usage is large
[3]. Increasing use of these fuels in land-based
medium speed c.i. engines coincides with public
demand for reductions in sulphur-originating pollution of the environment. Most modern refineries are
now equipped with desulphurization plants of this
Separation processes such as vacuum distillation,
crystallization and solvent extraction, are used to
'fashion' cuts from crude oil for special purposes and
to remove unwanted materials from products. A
variety of solvent extractions is used by lubricating
oil manufacturers (often in three or four separate
processes) to remove, for example:
Figure 12.1 Diagrammatic layout of crude oil refinery
processes for fuel and lubricating oils
gas (or diesel) oil. (The description derives from the
fact that it was used to enrich town gas [4].) That
part of the crude oil which does not vaporize but
remains in the column after the distillates have been
drawn off is called the residue. It is black and very
viscous. The residuals are not free-flowing at normal
temperatures. They have to be heated for storage
and handling and for use in RIC engines.
The fractionation process is followed by a series of
conversion processes whose purpose is to improve
the quality and balance the quantity of finished
refined products [3]. They are principally aimed at
the improvement of the lighter fractions and are
named according to the function which they perform. The so-called cracking processes essentially
break large molecules into smaller ones, and the
reforming processes alter molecular structure [5].
The action of the cracking process may be controlled
by reaction conditions such as elevated temperatures or pressures, or by the use of chemical reagents (catalysts). Diesel fuels produced as a result
of thermal cracking are more difficult to ignite than
those produced by distillation, and fuels produced
by catalytic cracking generally have good ignition
qualities [6]. Conversion processes are not neces-
1. aromatics, since these make the oil too thin at
higher temperatures; and
2. asphalt, for the preparation of the more viscous
grades of oil [3].
The rebuilding of distillate fractions into lighter
products uses a catalytic reformer (cat. reformer)
process. The by-products of the process are large
quantities of light hydrocarbon gases such as methane, propane and butane. The gases from the
catalyst cracking processes also contain a high proportion of the olefinic gases ethylene, propylene,
and various forms of butane [3].
In the context of fuels for RIC engines we may
summarize as follows [5]:
1. The crude oil is separated into temperaturerelated fractions - leading to the fuels required.
2. The lighter fraction distillate fuels are improved
by various conversion processes and their impurities removed, reduced or neutralized.
3. The heavier residual fuels may be further separated and blended with selected distillates to give
a range of fuels (principally identified by their
viscosities); some property control is possible.
12.4 Liquid fuel specifications
The liquid fuels used in c.i. engines are categorized
as either distillate grades or residual or blended
Propertiesof fuel oils
fuels. These are further subdivided into classes
based upon viscosity. Two of the most widely recognized specification authorities are the British
Standards Institution (BSI) and the American Society for Testing of Materials (ASTM).
The British standards applicable to fuel oils for
engines are:
BS 2869
- Fuel oils for non-marine use;
BS MA 100 - Specification for petroleum fuels for
marine oil engines and boilers
The latest (1988) revision of BS 2869 incorporates all
amendments since 1983 and is issued in two parts:
Part 1 - Specification for automotive diesel fuel
(class A 1). This anticipates the preparation of a
European Standard (EN) specification for light distillate automotive diesel fuel.
Part 2 - Specification for fuel oil for agricultural
and industrial engines and burners. It discriminates
between a general purpose distillate grade class A 2
fuel (heavier than class A 1) which is specifically
marketed for use with oil engines and those distillate
and residual or blended grades (seven classes - Cl
to H) which are traditionally used as burner fuels. It
recognizes that there is widespread use of the latter
fuels on larger c.i. engines and cautions that the user
should satisfy himself that any fuels so applied are
suitable for his intended application.
Diesel fuels which comply with BS 2869 Classes A
1 and A 2 and ASTM D-975 Grades I-D and 2-D,
and fuel oils to ASTM D-396 Grades 1 and 2 are
distillate oils widely used on high-speed and
medium-speed engines. ASTM D-975 also has a
Grade 4-D which, although it is considered to be a
'distillate' oil, is intolerable to some high speed
engines and is best suited to medium- and low-speed
BS 2869 Classes E to H, and ASTM D-396 Grades
5 and 6 fuel oils are of the residual or blended type.
Preheating before combustion may be required for
the BS Class E and ASTM Grade 5 (light) fuels.
Preheating will be required for burning, storage and
handling of Class F, G, H and Grades 5 (heavy) and
6 fuels.
Almost all medium-speed engines will operate on
the BS Class E and F and the ASTM Grade 4 and 5
(light) residual fuels. The crosshead type.2-stroke
engines ' (and even some of the slower running
medium-speed 4-stroke engines) can operate
successfully on Class G and Grade 5 (heavy), and, to
a lesser extent, on Class H and Grade 6 heavy
residual fuels. Having said this the considerable
price advantage offered by using fuels composed
partly, or entirely, of residual components (versus
the heavy distillates) has to be weighed against a
high investment in fuel storage, handling and treatment facilities and, possibly, greater plant maintenance costs. These considerations have led to
many ope~ators considering heavy distillate fuels of
intermediate quality and cost [7].
12.5 Properties of fuel oils
The quality requirements of fuels are specified
under those properties or characteristics which are
appropriate to the efficient utilization of the fuels.
These include:
ignition quality,;
carbon residue;
flash point;
water and sediment content;
ash content;
sulphur content;
copper corrosion;
pour point, cloud point, and cold filter plugging
• density;
• smoke point;
• distillation range and end point.
Other characteristics affecting engine operation are:
• the specific energy (or the calorific value) of the
fuel; and
• asphaltene content (where heavy residual oils are
12.5.1 Viscosity
The viscosity of a fluid is its resistance to shear,
flow, or change of shape. It is a measure of the
'flowability' or 'pumpability' of the fluid. It is
expressed by the number of seconds taken by a
certain volume of oil to flow through standard
orifices in viscosimeters or viscometers. The measurement must always be related to a stated temperature - typically, 38°C (lOO°F) and 50°C (122°F).
The Redwood I method is used in the UK for
distillate fuels, the Saybolt Universal in the USA,
and the Engler in Europe. Redwood II and Saybolt
Furol are used for residual fuels. Values for Redwood and Saybolt are expressed in seconds (hence:
SRI for Seconds Redwood I; and SSU for Seconds
Saybolt Universal).
The Engler figure is a 'specific viscosity', i.e. it is a
relative measurement. Degrees Engler is the ratio of
the time of flow of a similar volume of water. For
example, an oil of viscosity 6 degrees Engler has a
run-through time six times that of the same volume
of water, at 20°C (70°F). All these units of viscosity
are really 'flow times' . They have gradually been
superseded by kinematic viscosity.
Kinematic viscosity is the ratio of the coefficient of
viscosity to the density of a fluid. It is calculated
from the time taken for a sample of fuel to flow
though a glass capillary of known length and diameter, under constant pressure at a controlled temperature [7]. A typical instrument is the suspended
level capillary viscometer. The test method is
defined in BS 2000: Part 71 and IP 71.
Fuelsand lubricating oils
Kinematic viscosity (v) is expressed in stokes or
m2/s. The centistoke (cSt) is used in practice, and
equates to 1 mm2/s. The kinematic viscosity scale is
linear - unlike the others, which are not linear at
low viscosities. This means that a BS 2869 Class G
fuel with a viscosity of 40 cSt is twice as viscous as a
Class F fuel which has a viscosity of 20 cSt. The
kinematic viscosity reference temperature for Class
A 1 to D fuels in BS 2869: 1988 is 40°C, and 100°C
for the residual fuels Classes E to H.
The more viscous a fluid the more it 'drags'. The
viscosity of a fuel decreases as it is heated. A
viscosity/temperature relationship chart is given in
Figure 2 Appendix C of BS 2869: Part 2: 1988.
Viscosity conversion tables are widely available but
they are only valid for the conversion of viscosity
values into different units at the same temperature.
12.5.2 Ignition quality
The ignition quality of a fuel affects the time between injection and the start of combustion. A fuel
of high ignition quality is one which will ignite at a
low temperature. Cetane number is used to measure
the ignition characteristics of a fuel.
The most. widely accepted test method is that
defined in ASTM D-613: IP 41/81 (BS 5580 and
ISO 5165). The technique involves the comparison
of a fuel sample with a reference fuel in a special
single-cylinder test engine (known as a Cooperative
Fuel Research (CFR) engine) which has an adjustable compression ratio. The reference fuel is a range
of blends in which cetane (C16H34) is given a value of
100 and alpha-methylnaphthalene (CllHIO) is O.
The ignition delay is measured in the CFR engine
using the fuel to be tested. A similar test is then
conducted with differing blends of the reference
hydrocarbons until a mixture is found which shows
the same time-delay as the test fuel. If the test fuel
matches with a blend, say, of 45 % cetane and 55 %
alpha-methylnaphthalene, its cetane number is said
to be 45.
In practice, because the reference hydrocarbons
are expensive and difficult to obtain it is customary
to use secondary (or sub-standard) reference fuels
calibrated in terms of cetane number [7].
It is not always convenient or possible to conduct
CFR engine tests. The alternative parameter of
calculated cetane index (CCI) is one which provides
a good indication of ignition quality. Its use in the
routine quality control of fuels is permitted by
BS 2869 since it closely approximates to the determined cetane number for those fuels that do not
contain additives to improve ignition quality. It is
unreliable for gas oils, which may contain large
amounts of iso-paraffins or thermally cracked components [7].
The 1989 issue of BS 2869: Part 1 specifies limiting
values for both cetane number and cetane index.
BS 2000: Parts 218 and 364 (identical with Institute
of Petroleum specifications IP 218 and 364, respectively) give equations for CCI and nomographs for
ease of reference. Part 218 should be used for a fuel
with a cetane index of 55 and above and Part 364 for
a fuel with a cetane index below 55.
12.5.3 Carbon residue
This is a measure of the tendency of a fuel to form
'coke' when heated. The Conradson and Ramsbottom tests are used to establish the percentage of
carbon residue in a fuel after evaporation of a
sample under controlled conditions and within a
closed space. Both tests are commonly associated
with the lighter fuels and give much the same
results. The Ramsbottom method, described in
BS4451 and IP14/65, is specified in BS2869. A
quantity of the specimen oil is heated until the
volatile matter in it has evaporated. That which
remains is the carbon residue. It is weighed and the
result is expressed as a percentage of the quantity of
fuel tested. Since the values in distillate fuels are
usually low (less than 0.05 %) the test is often made
on the 10 % residue from the distillation test (see the
sub- section which follows). Allen [7] points out that
carbon and asphaltene contents of fuel oils generally
move in parallel. The asphaltene content is about
one-half that of the carbon value.
12.5.4 Volatility
The volatility of a fluid is the measure of its readiness to pass into vapour. The lighter fuel oils are
tested for this characteristic by a distillation recovery
test (BS2oo0: Part 123; IP 123). A 100mi sample of
the fuel is distilled at a prescribed rate in a standard
laboratory apparatus. Observations are made of
temperature versus volume of distillate, and of the
amount of residue. Those fuels which boil above
350°C have a tendency to decompose (or crack)
before their heavy 'ends' are boiled off. It is,
therefore, usual to record the percentage volume
distilled to a stated temperature (e.g. 200°C and
Another method (specified in ASTM D-975) records the temperature at which 90 % of the fuel
distils. The limits for a Grade 2-D distillate are, for
example, 282°C (min) and 338°C (max). The lower
the temperature the higher the volatility of the fuel.
12.5.5 Flash point
The flash point is defined as the lowest temperature
at which the fuel gives off sufficient flammable
vapour to produce a momentary flash when a small
flame is applied. Alternatively, it is that temperature at which an 'air-fuel vapour' mixture can be
ignited by a flame or spark.
Propertiesof fuel oils
Test methods give a purely empirical determination and the results vary according to the apparatus
used [7]. The tests are conducted in a vessel whose
cover is removable. In the closed flash point test the
specimen fuel is heated in the covered vessel. A
shutter in the cover is removed at frequent intervals
and a test flame is passed over the oil. Two closed
test methods may be specified. The Pensky-Martens
(BS2ooo: Part 34 and IP 34) is the one usually
specified in statutory and insurance regulations. The
non-statutory Abel test is designated in BS 2000:
Part 170 and IP 170 for Class C 1 and C 2 burner
The open flash point is determined using an open
vessel. The figure obtained for a given fuel is of the
order of lOOChigher than its closed figure.
The flash point only marks the presence of a
certain proportion of ignitable vapours in a specified
set of conditions. It does not indicate the general
volatility of the fuel [7]. The distillation characteristics are the yardstick for that (see the previous
sub-section). Nor does the flash point signify how a
fuel will ignite in an engine. That depends upon the
ignition quality of the fuel (its cetane number). In
fact, gasoline, which has a very low flash point, is a
very poor diesel fuel because of its poor ignition
quality [6].
There is also a fire point temperature - that at
which the vapour given off by the fuel bums steadily. It is usually some 30°C above the open flash
point. The legal limit set in the UK is 55°C.
12.5.6 Water and sediment content
These two parameters are indicators of the purity of
the fuel. Tests use 'distillation and extraction techniques (BS4385 and BS4382). The fuel specimen
may be centrifuged to separate both water and
sediment. Content levels may be specified separately (as in BS 2869) or a combined-content limit
may be defined (as in ASTM D-975 and ASTM
D-396). Water content and combined water-andsediment content are expressed in percentage by
volume of the sample fuel, whereas sediment content is usually given in percentage by mass.
12.5.7 Ash content
The ash content is determined by burning a known
quantity of the fuel; and weighing the inorganic
material which remains (BS 4450). Distillate fuels
leaving the refinery usually contain much less than
0.01 % (by mass) of ash-forming materials. A higher
ash content may be taken as an indication of contamination in handling or storage [7].
Incombustible ash levels are higher in residual
fuels - anything up to 0.25 %. Constituents may
include impurities such as highly abrasive sand and
rust, sodium, vanadium, aluminium, and nickel.
12.5.8 Sulphur content
Legislation in most countries restricts the sulphur
content of fuels. For example, the EEC Directive
75/716/EEC sets the limits of type A and type B gas
oils (equivalent to BS 2869 Classes AI, A 2, and D
fuels) at 0.30 % and 0.50 % mass, respectively. The
content in residual fuels is higher. The limit set by
BS 2869: Part 2, is 3.5 % for the so-called burner
fuels (Classes E to H).
The sulphur in a burned fuel is converted into
sulphur oxides. When combined with condensate in
exhaust gases these oxides form corrosive acids
which have a detrimental effect on cylinder liners
and piston rings. Problems are largely overcome in
those engines which can operate on residual and
blended fuels by using highly-alkaline cylinder oils,
and by maintaining high jacket water temperatures.
This is discussed in Sub-section 12.7.2.
Various test methods are employed for determining sulphur content. These include energy dispersive
X-ray fluorescence (IP 336), the lamp method
(BS2000: Part 107) for distillates, and the bomb
calometric method (BS 2000: Part 61) for residual
fuels. EEC Directive 75/716/EEC stipulates the use
of the Wickbold combustion method (EN 41;
BS 5379).
12.5.9 Copper corrosion
Fuels must be tested for the presence of free sulphur, which might corrode copper pipes or attack
certain barrel and tank coatings [4]. BS 2869 specifies the use of a three-hour copper strip tarnish test
(described in BS 2000: Part 154 and IP 154).
12.5.10 Pour point, cloud point and cold filter
plugging point
These three characteristics are a measure of an oil's
low temperature properties.
The pour point of an oil is defined as the lowest
temperature at which an oil remains a fluid, i.e. the
temperature at which the oil will just flow when the
containing (test) vessel is tilted for 5 seconds. It
relates to the lowest temperature at which a fuel can
be pumped without difficulty. Knowledge of a fuel's
pour point is of particular value in storage and
transport. When used in conjunction with viscosity
values it helps determine the heating needs of an oil
store. The test method is defined in BS 2000: Part 15
and IP 15. The alternative terms setting or solidifying points define the temperature at which the oil is
no longer fluid, i.e. the temperature at which it
solidifies or congeals. This is some 3°C below the
pour point.
The cloud point is the temperature at which the oil
becomes cloudy or turbid due to the separation of
traces of paraffin wax from the fuel. It is the point at
Fuels and lubricating oils
which crystallization of the paraffin wax begins, and
marks the temperature at which fuel line and filter
blocking might be experienced. The test method is
defined in BS2000: Part 219 and IP 219.
The cold filter plugging point (CFPP) is now
recognized as being the most appropriate method of
assessing the low temperature flow properties of
oils. The dynamic test method using a 340 mesh
screen (BS 6188; EN 116 and IP 309) has shown
better correlation with practice than the cloud point
test [4]. It is now adopted as the standard test
method in BS 2869 and in many European standards. The CFPP varies with the type of crude from
which the gas oil has been extracted. It is usually
between 1 and 6°C below the cloud point. Pour
point and CFPP may both be influenced by
'depressant' additives, whereas cloud point is not.
It should be noted that these low temperature
requirements only apply to distillate fuels. Engines
operating on residual fuels are equipped to handle
high-viscosity fluids and although it is not normal
practice, it is possible to pump fuel oils below their
pour point - which may be as high as 38°C for the
heavier grades [7].
12.5.11 Density
The density of a fuel is its mass per unit volume
(expressed in g/ml). It should not be confused with
relative density (replacing the older term specific
gravity), which is the ratio of the density of the fuel
at a specified temperature (often 15°C) to the density of water at the temperature of its maximum
density (4°C). Relative density (r.d.) is a pure
number - numerically equal to the density in g/ml.
A substance with an r.d. of less than 1 will float on
It is necessary to know the density of a fuel in
order to convert its bulk volume into mass for
operational purposes, e.g. storage and heating
needs. Also, fuel consumption rates are universally
quoted on a weight basis (e.g. g/kW/h or g/hp/h).
Density is usually measured by means of a hydrometer (BS 4714 and IP 160) or by a pyknometer
(BS4699 and IP 189), which is a glass vessel graduated to hold a definite volume of liquid at a given
temperature. By weighing it full of liquid at different
temperatures the variations in density may be found
An oil of high density is termed a heavy oil and
one of light density, a light oil. Since the density of
an oil depends upon the densities of its individual
components it is possible to have fuels of identical
densities but widely varying constituent component
densities. Fuels of the same density may, therefore,
differ widely in terms of viscosity and ignition
In spite of this (because it has been shown to have
a relationship with those two properties), the dens-
ity of a fuel is used in empirical formulae to obtain
sufficiently accurate values of the cetane index and
calorific values of a fuel. See Sub-sections 12.5.2 and
12.5.12 Smoke point
This parameter is only mentioned because it is
included in Table 2 of BS 2869: Part 2 as one of the
specified properties of the kerosene-based burner
fuel classes C 1 and C 2. It is a measure of the carbon
present in a flame (its luminosity) and is of particular
significance for gas turbines, flueless heating appliances, and vaporizing and atomizing burners for
domestic use. It is necessary to control the combustion quality of the fuels used in this equipment in
order to prevent excessive radiant heating of critical
The smoke point is defined as the maximum
height (mm) at which a fuel can be burnt in a
wick-fed lamp without smoking. The test method is
given in BS 2000; Part 57 and IP 57.
12.5.13 Specific energy
The calorific value, or specific energy, of a fuel is a
measure of the quantity of heat produced by a unit
mass of the fuel, on complete combustion. It is
expressed in kilojoules (or megajoules) per
kilogram - kJ/kg. Fuel consumption is related to the
calorific value of the fuel - for a given engine and
The fuel supplier can determine the gross calorific
value at constant volume (Qgv) and the net calorific
value at constant pressure (Qnp), using a bomb
calorimeter and the test methods of BS 2000: Part 12
and IP 12. The gross calorific value includes the
latent heat energy of the water vapour formed by
combustion of the hydrogen. The water vapour does
not contribute to the power output from an engine
since it passes out of the engine exhaust. The plant
user, therefore, needs to know the net calorific
value. The difference between gross and net values
is usually about 2500--2900 kJ/kg, depending upon
the hydrogen percentage of the fuel [9].
Because tests can be costly, it is usual to assume
that the calorific value of a given fuel is the same as
that previously found by testing a fuel of the same
The density of a fuel is closely related to its
heating or calorific value. Fuel is mainly composed
of carbon and hydrogen. Carbon has a much lower
heating value than hydrogen but a greater atomic
weight. The reason why a high density (heavy) oil is
heavier than a light oil is that it contains more
carbon and less hydrogen. The heavy oil must
therefore have less heating value on a weight basis,
i.e. high density equates with low heating value per
unit mass [6].
The effect of fuel properties on performance
The relationship between density and calorific
value has been studied by a number of authorities
and empirical expressions have been developed as a
result. For example, the formulae given in BS 2869:
Part 2 (Appendix C, Section C 5) enable sufficiently
accurate calculations to be made for Qgv and Qnp
using the density of the fuel and applying corrections
for any sulphur, water and ash present. The British
Standard also offers charts (derived from these
equations) which enable rapid estimates to be made
of the values. The equations and figures are identical with those published in ISO 8217: 1987.
12.6 The effect of fuel properties on
The effects on engine performance of the various
fuel properties are summarized in the sub-sections
which follow.
12.6.1 Viscosity
The lubricating property of a fuel oil reduces as its
viscosity reduces. The range of kinematic viscosity
in commercial distillate fuels is 1.5 to 5.5 cSt at 40°C
(BS2869). Injection pump wear may be accelerated
if the kinematic viscosity falls below 2 cSt at that
temperature. The fuel itself provides the lubrication
for the plungers in fuel pumps. Distillate fuels with
viscosities between 1 and 2 cSt are 'dry' and lack
sufficient lubricity to adequately lubricate the cams,
tappets and bearings of the injection pumps. This
could result in scuffing and ultimate failure of these
parts. Fuels of this type are widely used in Canada
and are referred to as 'arctic' fuels.
Special precautions also need to be taken when
high-speed engines are required to use the 'dry' type
aviation fuels obtained from kerosene-naphtha fractions. These may include Avtur 40/50, Avtag lP-4
and Avcat lP-5. It becomes necessary to regularly
drain the fuel injection pump and fill it with lubricating oil, or to employ a separate lubricating oil tank
and pump for automatic lubrication of the injection
Heavy residual fuels must be heated (to reduce
their viscosity) before they are admitted into the fuel
injection systems of medium-speed and slow-speed
engines. In practice, a composite system of heating,
treating and filtering is used in order to ensure that
the fuel arrives at the engine pumps and injectors in
the correct condition for combustion. The fuel's
temperature is progressively increased to about 95°C
before it is transferred to the daily service tank. The
general aim is to maintain a fuel viscosity between
10 and 20cSt during treatment. Thereafter, the
correct injection pressure, injection timing, and
running temperatures must be sustained.
It is normal practice to employ gravity feeds from
day tanks to engines when operating on distillate
and intermediate fuels. This is not feasible with
residual fuels and special steps have to be taken to
avoid the oil viscosity being too high at the fuel
injection pumps because of temperature drop due to
low velocity in the delivery pipes. One engine
manufacturer [10] uses a bus-rail system around
which a small pump circulates oil at a rate 4 times
greater than the engine's full-load consumption. A
bus-rail heater placed between this pump and the
engine compensates for any heat losses from the
lagged delivery pipes between daily tank and engine
and from the bus rail.
Incidentally, when heavy residual fuels are used
the engines must be started and stopped on light
distillate oil,
12.6.2 Density
We established in Sub-section 12.5.11 that the density of a fuel is related to its calorific value. This
means that density has a direct bearing on an
engine's fuel consumption and on its power output.
Also, the 'burnability' (or the ignition quality) of a
fuel is influenced by its cetane number, which also
has a direct relationship with density.
Heavier residual fuels have lower heating values
on a weight basis but increasing values on a volumetric basis. Since fuels are bought on a volumetric
basis there is an incentive to select those of high
density. Any economic advantages so gained may be
offset by the operational problems that could arise.
These include engine knocking, after-burning, uneven burning, variation in ignition delay, and a
steeper ignition pressure gradient (see Figure 1.4 of
Chapter 1). As Wilbur and Wight point out [9]
' ... these factors contribute to increased fatigue of
engine components, excessive thermal loading, increased exhaust emissions, and critical piston ring
and liner wear. The long-term effects on an engine
are la significant increase in fuel consumption and
component damage. The greatest fouling and deposit build-up will occur when the engine is operated at
reduced or very low loads '.
Another important consideration with high density residual fuels is that with values approaching 1.0
glml (they can be as high as 0.995 for some lowgrade fuels) there will be separation problems during pre-treatment. This is because above a density of
1.0 it is not possible to separate water from the fuel
oil by normal centrifugal action [11]. More sophisticated pre-treatment plant is required.
The performance figures for high-speed and
medium-speed engines operating on distillate fuels
are usually based on fuels complying with BS 2869
and ASTM D-975, i.e. those with densities around
0.835 to 0.865. Since fuel pump settings are 'calibrated' on volume of fuel it becomes necessary to adjust
Fuels and lubricating oils
the engine maker's declared performance figures if
the density of the fuel to be used varies from the
calibrated value. This is of particular significance
when intermediate grade fuels, produced by mixing
the lighter residual fuel oils with distillates, are to be
12.6.3 Low temperature characteristics
The cold filter plugging point (CFPP) for distillate
fuels should always be below the minimum starting
temperature at which the engine is to operate. This
ensures satisfactory fuel flow through the filters and
prevents wax crystals collecting on them and blocking them. The nearer the operating temperature is
to a fuel's CFPP the shorter the life of the filters. In
those areas of the world where high day-time temperatures are followed by low night temperatures, the
CFPP should be below the night-time temperatures.
One way of using fuels with a high CFPP is to
arrange for an auxiliary supply of low CFPP fuel to
start and run the engine until it is thoroughly
warmed, after which, the high CFPP fuel may be
used. Operation reverts to the low CFPP fuel just
before the engine is stopped. The engine is run on
this fuel for a sufficient time to ensure that the
filters, fuel injection pumps, and nozzles are saturated with the fuel. Similar precautions are necessary with the distillates used for starting those
engines which operate on the heavier residual fuels.
Problems are more likely to occur with mobile
and transportable equipment operating in low ambient conditions. Shipboard generators, those installed in power stations, and standby generators
housed in heated plant rooms should experience no
12.7 Harmful constituents in fuel oils
We shall now discuss the effects of those constituents most harmful to engine operation. Problems
may be magnified where combinations of several
harmful constituents occur.
12.7.1 Carbon residue
Prolonged operation on a fuel with a high carbon
content is likely to result in the formation of hard
combustion residues. These will be particularly evident at injector nozzles and in the piston crown, and
will also lead to excessive top ring groove wear and
sticking of rings. It is important that the fuel injection characteristics of engines using residual fuels
(which can have carbon residues between 5 and
15 %) are satisfactorily maintained.
The main hindrance to clean injection will be the
build-up of carbonaceous deposits on nozzles (trum-
peting) due to nozzle temperature. This causes intermittent sticking of the needle, resulting in a wet
nozzle and carbon formation on the tip. The trumpet formation has the effect of reducing combustion
efficiency. More fuel is required for a given load on
the engine and cylinder temperature rises. The
conditions under which pistons, rings, and exhaust
valves must operate then become progressively
more arduous. In practice, the carbon builds up on
the nozzles to a point where the trumpets detach
themselves from the nozzles, whereupon, injection
conditions return to near normal, and the cycle
recommences. A periodic rise and fall of cylinder
exhaust gas temperatures at steady load is evidence
of this condition occurring. In extreme cases the
temperatures may reach such high values that the
needle seat may begin to lose its hardness and the
needle rapidly hammers its way into its seat until the
nozzle is ruined [10]. Various methods may be used
to control nozzle temperature. These include cooling by water or by a separate distillate fuel circuit.
Engine wear rate may be greatly increased where
high carbon content coincides with high sulphur
12.7.2 Sulphur content
Residual fuels may contain up to 5 % sulphur by
weight. Part of this sulphur (in the form of sulphur
trioxide, S03) combines with water vapour, which is
another combustion product, to form sulphuric acid
(H2S04), If the temperature of the acid vapour falls
below its dew point when in contact with the relatively cool surfaces of the the cylinder walls the
vapour condenses on the cylinder walls. (Note: the
dew point is that temperature at which the vapour
saturates and begins to condense - anywhere between 110°C and 1800C). This causes serious corrosion on liners, piston-ring grooves, and piston rings.
Careful contr61 of cylinder liner temperatures is
therefore required to ensure that they are maintained above the acid dew point. Corrosive wear
may be minimized by using a highly-alkaline cylinder lubricating oil.
Problems are not confined to engines running on
residual fuels. They may also occur on higher speed
engines using distillate fuels with a combined sulphur and carbon content exceeding 2 %. Again,
wear rate is minimized by maintaining high jacket
water temperatures and using heavy duty lubricating
12.7.3 Ash content
The ash-forming constituents of most concern when
operating on heavy fuels are vanadium, sodium,
aluminium and silicon.
Gaseous fuels
High vanadium content may lead to:
• hot corrosion on parts such as exhaust valve seats,
valve plates, and the turbine blades of turbochargers; and
• the formation of hard, brittle deposits on these
Partial breaking away of deposited layers may cause
blow-by at the exhaust valves. The turbine efficiency
is reduced because its passageways are narrowed,
and its blade contours are altered by the deposits
These problems are increased when appreciable
quantities of sodium are also present in the fuel
- because the melting point of the vanadium-sodium
eutectic falls to about 550°C. Methods of preventing, or delaying, deterioration of valve seats include
the use of exhaust valve rotation and water-cooled
valve cages. The fuel may also be water-washed to
remove the sodium compounds. This is not a feasible solution in marine applications.
Problems are not confined to residual fuels - in
which vanadium and sodium contents may be as
much as 400 parts per million (ppm) and 150ppm,
respectively. They are also likely to occur on high
speed engines using the lighter distillates. Contents
of 50-60ppm may cause trouble. Limiting values of
lOppm are usually recommended if frequent exhaust valve and turbocharger rotor renewals are to
be avoided.
The aluminium-silicate compounds used by refineries in catalytic cracking processes may not be
fully removed from end products. The compounds
are hard and can cause wear on injection pump
plungers, injector nozzles, and piston rings.
12.8 Gaseous fuels
The gaseous fuels used in spark-ignition, dual-fuel
and alternative fuel engines may be classified under
two broad groupings:
• natural gas (NG); and
• by-product gases.
They differ in their properties and in their commercial availability.
12.8.1 Natural gas
This is defined in BS 1179 as one 'consisting mainly
of methane, occurring naturally in underground
accumulations'. The terms associated and nonassociated are applied to reserves of natural gas.
Those in which the gas is in contact with crude oil
in what is predominantly an oilfield are known as
associated reserves. The gas may be in a dissolved
state within the crude oil (solution gas), or it may be
in contact with gas-saturated crude (gas cap). Gas
production is very dependent upon the development
programme for the crude oil and on the market
potential for that gas which is surplus to oilfield
requirements (e.g. not re-injected or re-cycled)
Those reserves from which it is only economical to
produce gas, or where gas production is not significantly related to oil production, are known as nonassociated reserves. The gas may be retained in its
reservoir until suitable marketing outlets have been
identified. Although it is usually found in deeply
buried bed formations the gas is fairly easy to collect
and has the advantage of being available at high
pressure [12, 13].
The natural gas industry's activities may be
broadly classified as:
storage; and
This includes treatment and processing. It is not
within the scope of this chapter to describe these
operations in any detail. Very briefly, the field gas
which is brought up to the well-head passes through
a treatment station. This may be at the well-head
site or at a location central to several well-heads. In
the case of off-shore fields the station will be a shore
The operations at the treatment station may include the separation of any brine, condensed water,
and liquid hydrocarbons present in so-called wet
gases. They may also include desulphurization; carbon dioxide and nitrogen removal, and dust and
oil-fog removal. (Note: the term dry gas is used to
describe that gas which contains no (or very little)
liquid hydrocarbons.)
Where natural gas needs to be converted to its
liquified form for transportation it must be purified.
The process consists of the removal of impurities
such as water, carbon dioxide and hydrogen sulphide, and any heavy hydrocarbons such as propane
and butane. Purification may take place either before or during liquefaction. In the latter case the
impurities are frozen-out during the process.
The critical temperature of any gas is defined as
that temperature above which it cannot be liquified
by pressure alone. Since methane, the major constituent of NG, has a critical temperature well below
ambient, it needs to be cooled to about -160°C
before it can be transformed into liquid at atmospheric pressure.
The major natural gas production areas are the
USSR and North America - accounting for about
55 % of the world's output. The rest comes mostly
Fuelsand lubricating oils
from the Middle East, Africa and Western Europe
(including the North Sea). A typical analysis [13] for
North Sea gas is:
% volume
Other hydrocarbons (e.g. pentanes)
Carbon dioxide
Long-distance transmission of NG is achieved in two
1. by passage as gas through pipelines; or
2. by transportation as a liquid in ships, rail cars and
road tankers.
An example of how both methods may be integrated
into the one system is provided by the British Gas
Council's transmission system for natural gas from
the North Sea [13].
The welI-head gas pressure is usualIy reduced to
about 70 bar (7 MN/m2) before transmission along
pipelines. It may be necessary to employ intermediate gas compression stations, spaced along the
length of the main transmission lines. They serve to
restore the gas pressure in the system to near its
initial inlet level.
Liquefied natural gas (LNG) must be transported
and stored at temperatures below its boiling point of
-160°C. The insulation and refrigeration problems
this initialIy posed have now been overcome but the
cost of transportation and storage is high when
compared with that for other petroleum fuels.
Peebles and Clayton [12] point out that this cost
differential is further exacerbated by the fact that
the density of LNG is about half that of crude oil,
and its calorific value (per unit weight) is not much
more. In the 1970s the capital cost of a large crude
oil carrier was about one-sixth of that of an LNG
tanker of the same 'heat-carrying' capacity. There is
no reason to believe that this cost ratio has changed
since then. Despite this, the movement of natural
gas in liquified form has opened up the possiblities
of developing gas reserves in locations distanced
from large energy-consuming markets.
Large quantities of natural gas may be stored in
aquifers which are underground ieservoirs similar in
structural form to NG fields but with no gas in the
porous layer [13]. Long-term storage under high
pressure in underground steel reservoirs, and by
solution in propane, are other possibilities. Variations in weekly and seasonal demands are met in this
Pipeline storage is useful in meeting peak demands of shorter duration. The technique is known
as line-packing. Gas is stored in the pipelines at a
higher pressure than is theoreticalIy needed so that
the various elements in the piping network act as
high-pressure gas holders. Very accurate and centralized control of the system is then required to cope
with variations in demand from different parts of the
total system. The British Gas Council operates one
such computerized system which controls the transmission of gas from the off-shore gas fields, through
shore terminals, to the bulk distribution grids of the
Area Boards [13].
Six hundred cubic metres of NG reduce to 1 m3 of
LNG. Storage of NG in its liquid form thus requires
much less space at the receiving terminal than it
does in its gaseous form. The LNG must be regasified and the calorific value of the resultant gas
adjusted to that of the gas distributed to consumers.
Gas is metered at the transmission line take-off
points and its pressure is adjusted to about 5 bar
(0.5 NMlm2) for the medium pressure distribution
feeders to major industrial consumers. In areas of
heavy demand there may be further low-pressure
feed-in points.
It may be necessary to compress the gas at the
inlet to a dual-fuel engine to compensate for the
supply pressure fluctuations that would otherwise
occur when the engine operates at varying load
levels. TypicalIy, the gas may be compressed to
2.75 bar (0.275 MN/m2).
The simplified schematic diagram of Figure 12.2
[12] shows the activity phases for both NG and LNG
12.8.2 By-product gases
There are four natural sources of methane gas
(CH4). The first is natural gas. The second is marsh
gas, associated with marshy areas in which the
methane is produced by the decomposition of plants
by bacteria in the absence of air [14]. The third
source is that produced by the digestion of animal,
human and industrial waste materials (bio-mass).
The recovery process makes use of digesters. The
most widely used large-scale digesters are those
found in sewage works. The fourth source, and one
which is being increasingly exploited, is bio-gas from
landfill sites which contain domestic and industrial
wastes. Here, the gas is piped from large waste tips.
Gaseous fuels
It is usually dried, but it may also be scrubbed,
before use. The scrubber is designed to remove
impurities (particularly those of a corrosive nature)
and to increase the calorific value of the gas. A
large-scale method of scrubbing the gas is to bubble
it through water at high pressure.
Aerobic digestion takes place in the presence of
free oxygen while the anaerobic process occurs in
the absence of free oxygen. Both systems rely on
bacteria to break down organic matter. Sewage
works may use a combination of them in a two-stage
process in which the secondary digesters are open
The quantity and composition of the bio-gas produced depends upon the organic feed materialsthe major constituents of which are carbohydrates,
fat, and protein. Gas composition may be in the
range 60 to 70 % methane and 30 to 40 % carbon
dioxide. Other gases, such as carbon monoxide,
hydrogen, nitrogen, oxygen and hydrogen sulphide,
may also be present. Their proportions depend on
the digested feedstock.
The methane content of landfill gas may be as low
as 30 %. This is reflected in the comparison of its
lower calorific value with those typical of other
gaseous fuels:
Lower calorific value (MJ/m3)
Well head gas
Piped natural gas
Sewage gas
Landfill gas
One of the problems with landfill gas is that it may
contain trace elements based on fluorines and chlorides, which can be very corrosive on engine parts
Large sewage works are usually self-sufficient in
terms of electrical energy. They use dual-fuel
engines which may run on diesel alone or on a
mixture of diesel and bio-gas. It is standard practice
to use the waste heat energy derived from the
generator to heat the plant's digesters.
Small scale digesters may be installed on farms.
The gas may be used to provide heat for intensive
animal rearing or to power refrigerators and generate electricity for milking machinery on dairy
Other types of by-product gas are those obtained
from refinery processes. These consist mainly of C1
Fuels and lubricating oils
to C4 hydrocarbons with variable amounts of hydrogen, nitrogen and, possibly, hydrogen sulphide.
Typical examples are commercial butane (C4HIO)
and propane (C3HS)' which are capable of being
kept in the liquid state by bottling under pressure.
'Sweeting plants' are used at the refinery to remove
the hydrogen sulphide (HzS) from them. They have
high calorific values - typically 94 MJ/m3 for propane and 122MJ/m3 for butane.
Bottled gas (NG is also available in this form) is
more appropriate to domestic and commercial consumers located in remote areas where it is used in
burners for space heating and for air conditioning. It
is unlikely to be used for other than very small
12.8.3 Gas analysis
Engine makers request complete analyses for the
gases on which their engines are to operate. They
are particularly concerned with the presence of any
inert gases which would have the effect of reducing
the engine's power output on site. Also, hydrogen
sulphide contents in excess of 500ppm may be
problematic. The corrosive effects of HzS are similar
to those resulting from excessive sulphur content in
liquid fuels. Jones [15] reports that much of the
matching of lubricating oils to his company's (Dorman Diesels) range of gas engines was carried out in
well-head gas applications, and it has been found
that this oil technology has also been applicable to
engines operating on 'sour' bio-gas produced in both
digesters and landfill sites. (Note: a sour gas is one
containing acid gases, such as hydrogen sulphide
and carbon dioxide.) Correct timing of the lubricating oil changes is critical when engines operate on
bio-gases which contain acidic products as the oil
loses its reserve alkalinity quicker.
The methods for the analysis of fuel gases are
described in BS 3156 (see referenced standards in
Section 12.11 of this chapter).
Some fuels, particularly those of the bio-gas type,
are wet gases. This means that they contain moisture
which needs to be removed before they are used.
This is particularly necessary when the gas has been
cooled below its dew point. Knock-out pots (vessels
which act as receivers for condensate and remove
water droplets) may be employed in the gas stream.
An alternative approach (used on larger installations) is to chill the gas and then allow its temperature to rise again to reduce the moisture content
Gas supplies should not be assumed to be clean.
Rust which may have accumulated in pipework can
be scoured off with the passage of gas. It is, therefore, desirable to have efficient filtration close to
installed engines.
12.8.4 Gas properties
The significant properties of fuel gases are:
• Calorific value.
• Relative density.
• Wobbe number (related to both of the above
properties) .
The calorific value (cv) of a gaseous fuel is a
measure of the heat liberated by the complete
combustion under specified conditions of unit volume of the gas. It is conveniently expressed in
The gross calorific value relates to the condition in
which it is assumed that the latent heat of the water
vapour produced by the combustion of the fuel is
released. The net calorific value, on the other hand,
assumes that the water produced by combustion
remains as a vapour. The latter may also be referred
to as the lower calorific value (see Sub-section
12.8.2). Methods for the determination of calorific
value are specified in BS 3804.
The descriptions rich gas and lean gas are applied
to gases of relatively high calorific value (those
containing mainly low boiling point hydrocarbons)
and of relatively low cv, respectively. The term lean
gas should not be confused with 'lean burn', which
refers to the running of a gas engine on a high
air-to-fuel ratio. See Section 1.8 of Chapter 1.
The relative density of a gas is the ratio of the mass
of unit volume of dry gas to that of unit volume of
dry air under the same conditions of temperature
and pressure (BS 1179). It is therefore a pure
number. Typical figures for natural gas may vary
from 0.59 to 0.64.
The reader may meet the term Wobbe number in
a gas supplier's specification. It represents a measure of the heat released when a gas is burned at
constant gas supply pressure and is given by the
Associated terms are gas modulus, which is given by
the formula:
and gas group, which is the classification of a gas
according to its Wobbe number.
A supplier's specification may also include the
octane number (or the motor octane number) of the
gas. It is a measure of the limit of detonation, and
corresponds to the percentage of iso-octane in a
mixture (of iso-octane and heptane) which is equal
to the product gas in knock characteristics. Knock-
ing describes the detonation in a gas engine cylinder
due to over-compression of the air-gas mixture
before sparking. The octane number is analogous to
the cetane number of a distillate fuel oil. Both relate
to the ignition quality of a fuel.
Jones [15] advises that a 'mean octane number'
calculated from those of the various constituents in a
gas is not always a good guide to a fuel's ignition
quality. The heavy fractions in the gas can give rise
to detonation where the mean number might suggest
that no problem exists. This particularly applies to
turbocharged engines where compression pressures
are high. Also, because most engines of this type are
charge cooled, the temperature of the air-gas mixture entering the cylinder is critical. If heavy fractions are present in the gas the inlet manifold air
temperature must be kept low.
12.9 Lubricating oil
The many tasks that a lubricating oil has to perform
within an engine were discussed in Chapter 1 (Subsection 1.7.1). When selecting oils for service requirements one needs to take account of the conditions
that exist between engines of different speeds and
piston connections. For instance combined cylinder
and bearing lubrication is used in high-speed and
medium speed trunk piston engines, whereas separate lubrication of cylinders and bearings is customarily applied to low-speed, crosshead type
engines. Cylinder lubrication conditions can be particularly severe in those engines which use the
combined system. The period between engine overhauls will largely depend on the rate of cylinder
wear and the condition of the cylinders themselves
Modern oils fur engines are of the mineral type,
manufactured from crude oil (see Figure 12.1). Most
are now refined using solvent extraction processes,
in which undesirable compounds are selectively dissolved and removed [17]. Further processes follow,
such as refrigeration prior to dewaxing and clay
filtering, to give a rationalized range of oils from the
one refinery. These oils are then blended and compounded with additives to give the many grades and
special properties needed for engine duty.
12.9.1 Properties of lubricating oils
The chief properties of lubricating oils measurable
by physical and chemical laboratory tests are listed
below. The first nine are common to those of liquid
fuel specifications (see Section 12.5) and the tests
used to determine them are similar. We shall consider the significance of the properties in relation to
engine lubrication.
relative density
carbon residue
ash content
sulphur content
water content
pour point
cloud point
flash point
demulsification number
insolubles content
diesel fuel diluent
oxidation stability
This is perhaps the most important of the oil properties since it determines the effectiveness of the oil
film separating moving and rubbing surfaces. It
controls the load-carrying ability and affects the
friction and wear of bearings, governs the oil's
sealing effect and the rate at which the oil is
consumed [6].
While every attempt is made to standardize on the
use of kinematic viscosity (cSt) in oil specifications,
the arbitrary (flow time) units of Saybolt, Redwood
and Engler still persist within the industry. Also, the
Society of Automotive Engineers (SAE) viscosity
classifications continue to be widely applied.
Devised initially for the automotive industry, the
SAE system uses numbers which represent a range
of viscosity at given below-zero temperatures for the
lower viscosity oils and at 100°C (212°F) for those of
higher viscosities.
The oil grades are classified in the order of their
viscosity (see Table 12.1). The viscosity of the six
thinner grades (carrying the suffix W for 'winter') is
the dynamic viscosity, at sub-zero temperaturesexpressed in centipoise (cP). That for the four
Table 12.1 SAE Grades for engine oils
Max. viscosity
(cP) at °C
3250 at
3500 at
3500 at
35po at
4500 at
6000 at
- 30
- 20
Kinematic viscosity at ]()(f'C
Fuelsand lubricating oils
thicker grades is the kinematic viscosity (cSt) at
100°C. (Note: the dynamic viscosity is the preferred viscosity for bearing technology. It is equal to
the kinematic viscosity multiplied by the density of
the oil at the same temperature. Measurement is by
specialized viscometers.)
Any oil which is described by one of the above
numbers (e.g. SAE 20W or SAE 30) is said to be a
single-grade (or monograde) oil. One which meets
the requirements of more than one grade is termed a
multigrade oil. For example, a lOW/30 oil will have a
maximum viscosity of 3500 cP at -20°C and a
kinematic viscosity within the range 9.3 to 12.5 cSt
at 100°C.
All engine manufacturers publish lists of the proprietary and certificated monograde and multigrade
lubricating oils approved for use with their engines.
They have satisfied themselves that the multigrades
selected have been shown to have performance
levels equivalent to the oil manufacturer's certificated monograde oils.
The reader may also come across references to
ISO, ASTM and BSI grades. These are eighteen
rationalized grades classified by viscosity in the
range 2cSt to 1500cSt at 40°C (BS4231; ISO 3448).
They are not generally applied to engine oils, which
are conventionally classified by one of the designations discussed earlier (namely kinematic viscosity,
Saybolt, Redwood and Engler units, or SAE grade).
The viscosity of any oil changes with its temperature - decreasing with temperature rise and vice
versa. A useful chart is included in BS4231. The
rate of change is indicated by the viscosity index
(VI). It is calculated from equations involving kinematic viscosities determined at 40°C and 100°C, and
from tables giving the viscosities (at 40°C) of O-VI
and 100-VI reference oils that have the same viscosity at 100°C as that of the oil under test. The normal
range of VI is from 0 to 100. The higher the index
the less the change in viscosity with temperature.
Oils are broadly classified by index as follows:
The density of a lubricating oil has no bearing on its
quality. It is merely useful for costing purposes and
for computing weights and volumes. In conjunction
with viscosity it can give an indication of the source
of the oil or its method of refining [17].
Carbon residue
We have discussed (in Sub-section 12.7.1) the significance of high carbon content in fuel oils. The
carbon residue in lubricating oil is thought to contribute to the build-up of carbon deposits in combustion chambers. This is debatable, but if it does, the
contribution is very small indeed. Some of the
best-performing lubricants have high carbon residues [18]. The amount of residue varies with different crude oils, refining methods and viscosities. For
example, naphthenic-base oils show less carbon
residue than paraffinic oils [19].
Carbon residue tests are no guide as to the nature
of the carbon deposits that will form in an engine.
The paraffin-based oils produce a hard gritty deposit
in piston ring grooves. This may cause the rings to
stick fast. Oils with a naphthene base produce soft
fluffy carbon which may blowout with the exhaust
gases as sooty chunks [6].
Ash content
Unused oils have a very low ash content - unless
they have been contaminated in handling and storage. If ash is present in any quantity, it may contribute to deposits in combustion chambers and on
exhaust valves; but not near as much as ash from the
fuel [17]. Used oils, on the other hand, will contain
metal particles derived from engine wear. Attempts
are made to assess the rate of this wear from the
ferrous content of the ash [20].
Sulphur content
• High viscosity (HVI) above 80.
• Medium viscosity (MVI) between 30 and 80.
• Low viscosity (LVI) below 30.
During normal running an engine's oil temperature
is high (of the order of 98°C at the cylinders). Where
the engine is required to start in cold conditions the
VI gives an indication of the change in viscosity to
be expected in the oil between the two temperature
extremes in which it is required to operate.
It is possible to increase the viscosity index of an
oil by using so-called VI improvers. Multigrade SAE
oils are treated in this way (see Sub-section 12.9.2).
One important advantage of using a VI -improved oil
is that when an engine is shut down, the oil drains
very slowly from the hot bearings, pistons and
cylinder walls; a better lubricant film is then present
on engine restart [18].
Sulphur content has little significance in relation to
lubrication. The presence of some additives may
increase the sulphur content [17].
Water content
New oils rarely contain water. If it is present in used
oils it comes from combustion products, coolant
leakage, or atmospheric condensation.
Pour point
Low pour point oils are necessary for engines required to start in cold ambients. The pour point
must then be below the lowest temperature at which
the engine is required to start. Additives may be
used to depress the pour point.
Cloud point
This is being replaced by the cold filter filling point
(CFFP). Both are indicative of the temperature at
which wax begins to crystallize from the oil. They
have no significance during normal running.
oil changes with the charge-oil's immediate (tested)
condition in order to determine the best action to be
taken, i.e. either replace the oil charge or continue
for the present.
Demulsification number
Flash point
Most lubricating oils have sufficiently high flash
points (typically 1S00C closed) to eliminate fire
hazard during handling and storage. Flash point has
no significance in relation to the oil's performance in
an engine.
The measure of an oil's acidity or of its alkalinity is
given by its neutralization number. The neutralization value of an acid oil (its acid number) is expressed in milligrams of potassium hydroxide
(KOH) necessary to neutralize 1 gram of oil, i.e. it is
expressed in mg KOH/g. The term total acid number
(TAN) is an indication of the total acidic constituents of the oil, i.e. the combination of organic and
inorganic acids present. The strong acid number
(SAN) is a measure of the oil's inorganic acidity
only. The difference between the two numbers
(TAN and SAN) represents the organic acidity and
is a good indicator of the acidic 'weakness' of an oil
and of the effects of some additives. Highly-refined
straight mineral oils contain very small amounts of
acidic matter and may have a neutralization number
of 0.10 or less [19].
The total base number (TBN) is a measure of the
alkalinity of an oil. It is determined by titration of
the oil with hydrochloric acid or perchloric acid
(HCI04). The result is expressed as that amount of
acid necessary to neutralize 1 gram of the oil.
Unused oils containing dispersant additives are
usually alkaline. Reduction in alkalinity when in use
is a measure of the depletion of such additives in an
oil [17].
The neutralization number of a used oil is higher
than when it is new. This is because the oil oxidizes
in service and organic acids are produced. The
neutralization number does not distinguish between
corrosive and non-corrosive acids. It is not, therefore, a measure of the corrosive tendency of an oil
Some engine manufacturers may advise that oil is
drained and renewed when its neutralization
number exceeds a stated level. Gunther [19] suggests that the more appropriate yardstick is the rate
at which the oil reaches that level, and that' ... a
decided rise in neutralization number, accompanied
by an increase in viscosity and other pertinent
changes, is an indication that a serious condition has
been reached, or is being approached '. One needs
to correlate the history of the engine's operational
The demulsification number of an oil is the time (in
seconds) required for an oil which has been emulsified with water to separate under defined test
conditions. The emulsification may be achieved by
stirring or by steam blowing.
Modern mineral oils possess good properties of
water separability. They will lose this property when
contaminants are allowed to enter the oil and remain there indefinitely. If water is present in a
service oil' carbon dioxide from the air will combine
with it to form a weak acid (carbonic acid). This
results in a sludge which could add to the emulsification properties of the oil. Any tendency to emulsify
is increased by the presence of combustion products
in the lubricating oil. Since emulsification permits
oil to retain water particles it is undesirable for the
following reasons [19]:
1. it leads to lubricant deterioration;
2. it could cause mechanical failures (because oil is
displaced by water in frictional areas);
3. it may cause rust and corrosion; and
4. it could cause the lubricant to 'foam'.
lnsolubles content
Tests (typically BS 2000: Part 143 and IP 143) are
applied to used oils to determine the amount of
contamination present. In the first part of the test
precipitation by normal heptane is used to give a
composition of inorganic, carnbonaceous, and asphaltic matter. The asphaltenes are extracted from
this composite using a solvent and they are weighed
after the solvent has evaporated. The remaining
residue then consists of inorganic and carbonaceous
matter. The inorganic element contents are determined (as 'sulphated ash') by using the test method
of BS 2000: Part 163 and IP 163, in which the residue
is heated to free it from its carbon matter. The
sulphated ash may be further analysed by chemical
or spectographic methods [17].
The nature of lubricant insolubles is fairly complex. Van HeIden and Hengeveld [21] report that
extensive research has shown that the two major
constituents of oil contaminants in large residualfuelled engines are combustion soot and calcium
sulphate. They quote a typical analysis for extracted
separator sludge as showing 20 % soot, 6S % inorganic calcium salts and 10 % hydrocarbons, with the
remainder comprising degraded additives and other
inorganics. These inorganics might include, for
example, metallic wear debris, corrosion products,
and inorganic contaminants from external sources.
Fuelsand lubricating oils
They conclude [21] that' ... lubricant contamination
is becoming more severe with modern highly-rated
engines, widespread use of residual fuels and trends
to reduce oil consumption'; and, since ' ... engine
cleanliness is related to oil contamination and the
fouling rate seems to be proportional to insolubles
concentration, ... this can provide some guidance to
engine builders to set limits to maximum tolerable
contamination levels'.
Diesel fuel diluent
It is possible to get dilution of lubricating oil by fuel
leakage into the crankcase but it can also occur by
loss of fuel from multi-hole injection nozzles. The
effect is progressive. The best safeguard against
undue dilution is to maintain the fuel injection
equipment in good order [17]. One immediate remedy is to change the oil or to drain off part of the
charge in use and top-up with fresh oil. It may be
desirable to use a heavier grade oil for the top-up
Any appreciable contamination of the lubricating
oil by fuel causes a marked reduction in the viscosity
of the oil and hence its delivery pressure [16].
Determination of the degree of dilution by a comparison of the viscosities of used and unused oil is,
however, unreliable. Far more dependable methods
are, to compare the boiling point distributions in the
two oils by gas chromatography [17], or to compare
their flash points.
An oil which has air entrapped within it will foam
(or froth) when it is churned. The air may enter the
oil in many ways. Gunther [19] cites several, including:
• through the intake of an unsubmerged pump;
• situations where oil is discharged into a sump at a
point above the oil surface; and
• splash lubrication systems, where oil breaks up
into small droplets.
In dry sump systems (see Sub-section 1.7.1 of
Chapter 1) the scavenge pump must have a larger
capacity than the circulating pump in order to
ensure that the sump remains 'dry'. At times, therefore, the scavenge pump is likely to be drawing in
some of the atmosphere of the crankcase together
with oil. When system pressure is reduced the air
entrapped in the oil expands and causes foaming
[17]. New oils have a tendency to foam more rapidly
than used oils. Also, the greater an oil's viscosity the
greater the likelihood of foaming.
The rate at which bubbles on the surface of an oil
break-up depends upon the surface tension of the
oil. Some additives (such as dispersants) and contaminants increase surface tension and may intensify
frothing. All oils will foam. Anti-foam agents (addit-
ives or suppressants) may be included in an oil to
prevent formation of a stable foam. Small concentrations of silicones (typically polymeric organic
siloxanes) are used for the purpose. The concentrations need to be small (less than 0.001 %) because
one does not want to reduce the surface tension too
much and so diminish the oil's lubricating property.
The 'foam-breaking' agents allow the small bubbles
that form on the oil's surface to combine and form
larger ones which then break up more readily [9].
Oils are aerated under standard conditions
in sample tests. The volume of foam produced
immediately after cessation of air-blowing is then
compared with that which remains after the sample
has settled for ten minutes. The test is repeated on
a further sample at elevated temperature [17].
Oxidation stability
The oxidation of any substance implies the combination of oxygen with that substance or the removal of
hydrogen from it. This process takes place in lubricating oil and changes occur within the oil due to the
presence of entrained air (or water). It is perhaps
the most important of the internal contamination
processes which affect the service life of an oil.
Oxygen in the air combines with the hydrocarbons
of an oil to form aldehydes, which then convert into
organic acids. The process is accelerated by an
increase in temperature, by aeration, and by the
presence of catalytic metals. These conditions are
inherently prevalent in RIC engines. For example:
• high combustion temperatures;
• aeration caused by the splashing effect of crank
• water due to condensation from combustion; and
• the presence of catalytic copper and ferrous metals. The oxidation effect is more pronounced
when the surfaces of such metals are fresh, and
are being worn [19, 20].
The organic acids resulting from the oxidation produce various harmful effects. Those acids that are
more volatile (i.e. those with low boiling points) will
attack bearing and other surfaces causing pittage
corrosion. Those with high boiling points react with
the oil to produce sludges, gums and carbon residues. These products may be soft or hard, depending upon engine temperature and the relative proportions of the organic acids.
The sludges are deposited in sumps, in oilways, on
filters and strainers, and in heat exchangers. The
gums form in the combustion chambers and adhere
to piston rings and grooves, valves and valve
chambers. In their more extreme form, the 'sludges'
appear as hard varnishes or lacquers (thin layers of
converted varnish) which cause valve and piston ring
sticking and even valve burning. The effects of
Lubricating oil
carbon residues have been discussed earlier and in
Sub-section 12.7.1.
There are many oxidation tests that may be
conducted in laboratories. Each has been devised to
simulate different service conditions. Results, therefore, cannot be directly compared. (Lansdown [18]
suggests that the method of IP 176, Oil oxidation
and bearing corrosion test (Petter Engine Test) is
most suitable for engine oils.) More reliance is
placed on actual engine tests.
12.9.2 Additives
Straight mineral oils are no longer able to meet the
operating conditions imposed on modern engines.
Consequently, lubricating oils must now have a
range of chemical compounds added to them to
improve their quality and performance beyond that
which can be achieved by refining alone. A very
brief description follows of some of the main additives.
Corrosion and rust inhibitors
Corrosion inhibitors act to form a protective film
that adheres to (or is adsorbed on) metal surfaces,
or they may act to break the chain reaction of
oxidation and acid formation. See also below.
Typical compounds are organic amines and soaps
such as calcium, sodium or barium sulphonates.
The purpose of rust inhibitors is to provide a thin
film of oil on engine surfaces during operation and
shutdown periods. The essential feature of the compounds used is their high polar attraction to metal
surfaces, giving a plating action. They include metal
sulphonates, amines, fatty acids and metallic soaps
of fatty acids [19].
Metal salts, sulphur and phosphorous compounds, and amines are some of the ingredients of
Rust inhibitors and those additives designed to
reduce engine wear, also have a passivating effect on
catalytic metal surfaces. Conversely, oxidation inhibitors also act as corrosion inhibitors.
Detergents and dispersants
The term detergent implies a cleaning agent. It
relates to products used in solution for washing or
cleaning by action other than by simple dissolution.
Detergents are metal soaps soluble in oil. They do
not correct the cause of oxy-product development
and only act to prevent its build-up on clean surfaces
exposed to lubricating oil. Detergents do not remove existing dirt deposits from dirty surfaces. They
may be more accurately described as detergentdispersants.
The function of a dispersant is to keep contaminant particles isolated from each other. In other
words, it acts to keep the fine particles in suspension
within the oil mass. This may lead to difficulties with
choking of fine-particle filters. The normal filters
used in engines are unlikely to give trouble [17].
Contaminant particles in suspension are preferable to their being deposited on engine parts. Dispersant oils do look dirty in use. The suspended dirt
is fairly harmless even though the lubricating qualities of the oil may not be enhanced by its presence.
Dispersant oils can be safely used until their recommended change period is reached. The periodicity of
oil changes is materially increased by the use of
detergent-dispersants and anti-oxidants.
Organo-metallic salts, such as salicylate ester
salts, are the usual additives for high operating
temperatures, while polymer compounds are generally used for low temperature conditions.
Oxidation inhibitors
Oxidation inhibitors (or anti-oxidants) are used to
prolong the useful life of an oil. They act to prevent,
retard or modify the oxidation process, and to
inhibit corrosion of bearing and metal surfaces.
They may do this in different ways. Those that
compete with the oil for oxygen (i.e. act as 'decoys'
for any environmental oxygen) are known as preferential anti-oxidants. They are expendable. When
they are 'spent' the oxidation stability of the parent
oil reverts to its base characteristic. The subsequent
service life of the charge then depends upon the oil's
basic qualities.
The so-called metal deactivators (also called anticatalysts and passifiers) serve to neutralize the catalytic acti<m of any metal particles suspended in the
oil and form a barrier, or protective coating, over
those engine surfaces which are in contact with the
lubricating oil.
Viscosity-index improvers
These are high molecular weight polymers which act
to reduce the rate of change of viscosity with temperature. They raise the viscosity of an oil by a
greater proportion at higher temperatures than at
lower temperatures. In so doing, they ensure good
oil circulation at low temperatures and maximum
viscosity at high temperatures. VI improvers are
used in multigrade oils meeting SAE requirements.
Pour-point depressants
Paraffinic base oils are usually de-waxed as part of
the refining process. The process is costly and if the
de-waxing is too severe there is a danger of removing desirable constituents. In practice, a combination of de-waxing and adding selected depressants is
Fuelsand lubricating oils
used. The depressant materials are complex
polymers of high molecular weight.
When a paraffinic lubricating oil is cooled below
its pour point it coagulates because the wax separates out to form a congealed matrix of needleshaped crystals. This matrix acts as a sponge to hold
the oil. Some pour-point depressants function by
surrounding the wax crystals with a protective coating, thus preventing them linking up to form the
matrix. The viscosity of the free oil is not affected.
Some viscosity-index improvers also function well
as pour-point depressants. Their mode of action is
somewhat obscure. They appear to modify or guide
the growth of the wax crystals in a three-dimensional
mode around a nucleus, so preventing the interlocking matrix from being formed [17].
Foam depressants
Anti-foaming agents function to reduce the surface
tension of an oil, thus releasing any air entrained in
it. They are usually high molecular weight silicone
polymers (such as the polymethyl siloxanes) in concentrations of less than 0.001 %.
12.9.3 Oil specifications
We have discussed the SAE system of classification
in a previous Sub-section (12.9.1). We shall now
consider the specifications that are commonly
quoted in lubrication practice.
The American bodies API, ASTM and SAE have
co-operated in producing a classification of lubricating oils by type of use in both gasoline and diesel
engines. Those applicable to the latter are:
CA - for light duty service with high-quality fuels;
CB - for moderate duty service by naturally aspirated engines operating on lower quality fuels
with higher sulphur content;
CC - for moderate to heavy duty service by
pressure-charged engines; and
CE - for severe duty service typical of turbocharged engines in high speed, and low speed,
high output duty.
Lubricant manufacturers market their products
under brand names (e.g. Shell Rotella TX and
Silkolene Hardwick). They issue data sheets which
describe the typical properties of the product. These
give prospective users a yardstick by which a particular brand may be assessed for intended service.
It is customary for engine manufacturers to cooperate with oil refiners in undertaking thorough
engine test programmes on various brands of oil in
several kinds of service. They then issue reference
lists of the specified oils to be used consistent with
the duty and rating of an application and the fuel
which it is intended to use.
If a plant operator should need to consider an
unlisted oil he must compare the properties of the
brand he intends using with those of one, or more,
of the approved brands on the engine manufacturer's list. Lubricant data sheets should be used for
such comparisons. Better still, the operator should
ensure that the oil has an official certificate of
conformance for the appropriate specification
designated by the engine manufacturer. Any change
effected should always be followed by a careful
monitoring of the new oil's performance in the
engine's systems.
Among the widely accepted specifications are the
American military specification MIL-L-2104, and
its British equivalent (the DEF-2101 series) issued
by the UK Ministry of Supply. Modifications to both
specifications are denoted by a letter suffix: B, C, D,
The American companies Caterpillar and General
Motors were the first to devise extensive test programmes for the evaluation and approval of lubricants for use in their engines. Much of the subsequent military standardization was based on the
earlier work of these two companies. Often-quoted
specifications are those derived by the Caterpillar
Tractor Company - its Series 1, 2 and 3 oils,
designated Cat 1, etc. Later designations generally
have the higher performance levels. It is outside the
scope of this chapter to give full details of the many
standard engine tests, and their associated specifications. Those readers who wish to study the topic in
some depth should make [17] or [20] their starting
12.10 Corrosion inhibition in transit
and storage
It is prudent to arrange for engines to be given
temporary protection against atmospheric corrosion
during transport or storage. Practical measures include:
1. the application of impervious coatings to insulate
internal surfaces from their environment; and
2. the use of chemical inhibitors to reduce corrodibility of such surfaces. In particular, to reinforce
oxide films so that they can persist and suppress
further reaction even in the presence of moisture.
Combinations of both are possible and are effectively used on engines for storage periods of 12
months or more [22].
Typical 'temporary' protectives are the Shell Ensis Engine Oils. They are available in various viscosity grades and are applied by using the correct grade
in place of normal lubricating oil for a short period
before engine lay-up. They contain additives which
neutralize the acids formed during fuel combustion.
Reference standards
Their compatibility with service lubricants means
that no flushing procedures are necessary and they
may even be retained as the service lubricant in
some types of engine until the first scheduled oil
Fuel systems may be treated by using a special
'calibration fluid' as a fuel. The fluid is left in the
fuel lines during lay-up.
The cooling system too should be treated with a
mixture of water and an inhibitor (such as Shell
Dromus Oil B).
The engine to be treated is run until it is thoroughly warm. It is then stopped, drained of lubricant and coolant, and refilled with the recommended protectives. The fuel filter housings and any
integral service tank are drained of fuel and the fuel
system is primed with the calibration fluid. The
engine is run for a short period and allowed to cool
after it is shut-down. Additional protective lubricant
is then sprayed into inlet and exhaust manifolds care being taken to seal all open flanged joints with
well-seated, and gasketed, blind flanges. Unpainted
external surfaces should be swabbed with a protective (such as Shell Ensis Fluid 260) which has dewatering properties and will give a thick film coating
(about 0.01 mm) of medium-hard consistency. A
drying time of about four hours is recommended in
temperate climates.
12.11 Referenced standards
Listed below are those UK standards, issued by the
British Standards Institution and the Institute of
Petroleum, to which reference has been made in this
chapter. Corresponding International (ISO), European (EN) or Institute of Petroleum (IP) standards
are shown in brackets. The extent of their agreement with the British Standard varies. Some are
completely identical and have dual-numbering.
Others have substantial technical equivalence, with
differences perhaps in wording and presentation.
BS 526 - Definitions of the calorific value of fuels
BS 1179 - Glossary of terms used in the gas industry
BS 2000 - Methods of test for petroleum and its
Part 12 Heat of combustion of liquid hydrocarbon fuels
Part 15 Pour point of petroleum oils
Part 34
Flash point by Pensky-Martens
closed tester
Part 57 Smoke point of kerosine
Part 61 Sulphur in petroleum products
(bomb method)
Part 71
Kinematic viscosity of transparent
and opaque liquids and calculation of
dynamic viscosity
Part 107 Sulphur in petroleum products (lamp
Part 123 Distillation of petroleum products
Part 143 Asphaltenes in petroleum products
(precipitation with normal heptane)
Part 154 Detection of copper corrosion from
petroleum products by the copper
strip tarnish test
Part 163 Sulphated ash from lubricating oils
and additives
Part 170 Flash point by the Abel apparatus
(non-statutory method)
Part 218 Calculation of cetane index of diesel
fuels (range 55 and above)
Part 219 Cloud point of petroleum oils
Part 230 Pumpability test for industrial fuel
Part 364 Calculated cetane index of diesel fuel
oils (range below 55)
(Note: these BS 2000 Parts have correspondinglynumbered equivalents in standards issued by the
Institute of Petroleum; for example, Part 163 is
identical with IP 163.)
BS 2869 - Fuels for non-marine use
Part 1 Specification for automotive diesel fuel
(class A 1)
Part 2 Specification for fuel oil for agricultural
and industrial engines and burners
(classes A 2, C 1, C 2, D, E, F, G and
BS 3156 - Analysis of fuel gases (this is a multi-part
standard which deals with:
• determination of: the main combustible and inert
the minor constituents, and
the combustion characteristics
of common fuel gases;
• the separation of constituents by chromatographic analysis; and
• determination of the constituents and characteristics of gases at pressures above 1.7 bar.)
BS 3195 - Methods for sampling petroleum products
Part 1 (ISO 3170 and ISO 1995) Light hydrocarbons: manual sampling
BS 3804 - Methods for the determination of the calorific value of fuel gases
Part 1 Non-recording methods
BS 4231 (ISO 3448) - Classification for viscosity
grades of industrial liquid
BS 4382 (ISO 3735 & IP 53) - Method for determination of sediment in
crude petroleum and
fuel oils (extraction
Fuelsand lubricating oils
BS 4385 (ISO 3733 & IP 74) - Method for determination of water in
petroleum products
and bituminous materials (distillation
BS 4450 (EN 7 & IP 4) - Determination of ash from
petroleum products
BS 4451 (ISO 4262 & IP 14) - Method for determination of carbon residue of petroleum
products (Ramsbottom method)
BS 4699 (IP 189 & IP 190) - Methods for determination of density or relative density of petroleum
and petroleum products
(pyknometer methods)
BS4714 (ISO 3675 & IP 160) - Method for laboratory determination
of density or relative density of crude
petroleum and liquid
(hydrometer method)
BS5379 (EN 41) - Determination of the sulphur
content of petroleum products
by the Wickbold combustion
- Dieselfuels-determination of ignition quality Cetane method
BS 6188 (EN 116 & IP 309) - Method for determination of cold filter
plugging point of diesel and domestic fuels
BS MA 100 - Specifications for petroleum fuels for
marine oil engines and boilers
A related specification is that of ISO
8217 - Petroleum products - Fuels (class
F) - Specifications of marine fuels)
IP 336/86 - Sulphur in petroleum products by energy
dispersive X-ray fluorescence (nondispersive X-ray fluorescence)
12.12 References and bibliography
12.12.1 References
1. Priest, J. Energy - Principles, Problems, Alternatives 3rd edn, Addison-Wesley, (1984)
2. Carpenter, P.R. Fuels of the future. A paper
presented to the Diesel Engineers and Users
Association (now the Institution of Diesel and
Gas Turbine Engineers) Publication No. 365,
March (1975)
3. Roberts, W.G. The quest for oil. Revised edition, Methuen, March (1975)
4. Kirtley, J.G. Fuel and combustion. Chapter 4 of
Diesel Engine Reference Book, edited by
L.C.R. Lilly, Butterworth (1984)
5. Horsley, M.E. Oil fuels. Chapter 3 of Industrial
Fuels, edited by Peter C. Bell, The Macmillan
Press Limited (1971)
6. Kates, E.J. and Luck, W.E. Diesel and High
Compression Gas Engines, 3rd edn, American
Technical Society (1974)
7. Allen, R.L. Diesel and turbine fuels. Chapter 5
of Industrial Fuels, edited by Peter C. Bell, The
Macmillan Press Limited (1971)
8. Uvarov, E.B., Chapman, D.R. and Isaacs, I.
The Penguin Dictionary of Science, 5th edn,
Penguin, (1982)
9. Wilbur, C.T. and Wight, D.A. Pounder's
Marine Diesel Engines, 6th edn, Butterworth
10. Greenhalgh, R. Turbocharged-diesel generating plant burning residual fuels. In Hawker
Siddeley Technical Review Industrial Edition,
1(3) 16-23 (1964)
11. Syassey, O. Adapting medium-speed engines to
future fuels qualities, Diesel and Gas Turbine
Worldwide (Diesel and Gas Turbine 'Publications) XIII(9) 90-94 (1981)
12. Peebles, M.W.H. and Clayton, R.S. Natural
gas. Chapter 8 of Industrial Fuels, edited by
Peter C. Bell, The Macmillan Press Limited
13. Scott Wilson, D. Natural gas. Chapter 4 of
Materials and Technology Vol. IV edited by
L.W. Codd et al. Longman Group, (1972)
14. Meynell, P.J. Methane - Planning a Digester,
2nd revised edn, Prism Press, (1982)
15. Jones, D.L. Gas engines for generating sets and
CHP systems. A paper read at the INPOWER
88 Conference jointly organized by the Institution of Diesel and Gas Turbine Engineers and
Fuel and Metallurgical Journals Ltd., September (1988)
16. The Lubrication of Oil Engines, Shell International Petroleum Company,
17. Pohl, W. Lubrication and lubricating oils.
Chapter 16 of Diesel Engine Reference Book,
edited by L.C.R. Lilly, Butterworth (1984)
18. Lansdown, A.R. Lubrication - A Practical
Guide to Lubricant Selection Pergamon Press,
19. Gunther, R.C. Lubrication, Bailey Brothers
and Swinfen Ltd (1972)
20. Pugh, B. Practical Lubrication, Butterworth,
21. Van Heiden, A.K. and Hengeveld, J. Crankcase lubricants - an integrated view of engine
performance and oil treatment. A paper read at
the INPOWER 88 Conference jointly organized
Referencesand bibliography
by the Institution of Diesel and Gas Turbine
Engineers and Fuel and Metallurgical Journals
Ltd., September (1988)
22. Shell Temporary Protectives for Corrosion Prevention, Shell International Petroleum Company Limited (1972)
12.12.2 Bibliography
The text-books listed in the references have been
used in the preparation of this chapter. Those which
are recommended as starting points for research by
the reader are:
1. For fuel oils: [4, 5 and 7].
2. For gaseous fuels:
[12 and 14]. The latter is
particularly recommended for its explanation of
the digestion process and the engineering requirements of digesters in the production of bio-gas.
Not included under the references is a work by
M. Medici entitled The Natural Gas Industry,
Butterworth, (1974), which is targeted at a wide
readership but presents specialized knowledge on
many aspects of the industry.
3. For lubricants:
[17, 18, 19 and 20]. Another
useful text-book, on the wider subject of tribology, is that edited by J. Halling entitled Principles of Tribology, The Macmillan Press Ltd,
issued in paperback in 1978. The contribution by
R.B. Howarth on Lubricant properties and testing (Chapter 9) is of particular interest in this
Other useful sources of information are:
1. The house journals and technical literature of the
major engine manufacturers and oil companies
[10, 16 and 22].
2. International trade journals and magazines specializing in prime-movers and engine-powered
equipment such as Diesel Engineering, Diesel and
Gas Turbine Worldwide, High Speed Diesel
Report and European Power News.
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