`A Holistic Method for Conductor Ampacity and Sag Computation on

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‘A Holistic Method for Conductor Ampacity and Sag Computation on an
OHL Structure’
K. Kopsidas, S. M. Rowland, B. Boumecid,
IEEE Trans Power Delivery, Volume: 27, pp. 1047-1054 (2012)
A Holistic Method for Conductor Ampacity
and Sag Computation on an OHL Structure
Konstantinos Kopsidas, Member, IEEE, Simon M. Rowland, Senior Member, IEEE, and Boud Boumecid
Abstract—The rating current (ampacity) of a conductor erected
on a particular overhead line (OHL) structure installed at a specified location is influenced by the conductor, the OHL structure, as
well as weather and operational parameters. Many studies have
been carried out regarding calculating an aerial bare conductor’s
ampacity at a steady-state conductor temperature, but without
considering the OHL structure as part of the system. In this paper,
a holistic methodology for calculating the conductor’s ampacity
and sag at any temperature and power frequency, erected onto
a prespecified OHL structure is developed, considering together
the mechanical and electrical parameters of the overall system.
This methodology incorporates the conductor’s basic material
properties allowing the calculations to be applied to newly developed high-temperature low-sag composite conductors. In this way,
it becomes possible to identify, at the system level, the potential
benefits that may result from the improved performance of these
conductors as well as to indicate new sizes that may better fit a
prespecified system, optimizing its performance. The methodology
is also validated with a real system application, resulting in correct
predictions of the performance of a four-span double-line system.
Index Terms—Aging, ampacity, bare conductors, conductor
creep, high temperature, low sag, overhead line, reconductoring,
re-tensioning, sag, tension, thermal rating, upgrating, uprating.
HE NEED TO increase the power transfer capacity of
existing distribution and transmission lines has resulted
from the growth in demand for, and generation patterns of, electrical power. However, this is not always easy to achieve in a
deregulated environment where competition forces utilities to
operate existing lines at maximum rated capacities due to environmental and cost barriers. As a result, utilities may attempt
to extract the most out of the installed lines by operating them
closer to the thermal limit, and even temporarily exceeding it,
causing a loss of strength in the conductor and increased sags,
leading to higher risk operation and occasional blackouts [1].
Manuscript received June 15, 2009; revised April 18, 2011; accepted
November 09, 2011. Date of publication June 14, 2012; date of current
version June 20, 2012. This work was supported by the EPSRC Supergen V,
U.K. Energy Infrastructure (AMPerES) Grant in collaboration with the U.K.
electricity network operators working under Ofgem’s Innovation Funding
Incentive scheme. Paper no. TPWRD-00449-2009.
K. Kopsidas and S. M. Rowland are with the School of Electrical and Electronic Engineering, The University of Manchester, Manchester M13 9PL, U.K.
(e-mail: k.kopsidas@manchester.ac.uk; s.rowland@manchester.ac.uk).
B. Boumecid is with the National Grid Electricity Network Investment Department-Asset Policy, Warwick CV34 6DA, U.K. (e-mail:
Color versions of one or more of the figures in this paper are available online
at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TPWRD.2012.2187464
An economical method to increase the capacity of an existing
system can be achieved by enabling existing lines to operate
at higher temperatures. This may infringe ground-clearance
requirements, but can be accomplished by re-tensioning the
conductor, a technique mainly based on the increase of the conductor’s clearance to the ground by increasing its tension on the
power line. This, therefore, will increase the thermal limit of the
line since more thermal expansion can be allowed, permitting
more current to flow through it. Another way to achieve power
capacity increase is to replace conductors (reconductoring)
with larger all aluminum alloy conductors (AAAC) or aluminum conductor steel reinforced (ACSR) wire or conductors
with new materials which permit higher temperature operation
and developing lower sags. The latter are usually referred to
as high-temperature low-sag (HTLS) conductors. Either way,
the computations for the new maximum current capacity and
conductor sag have to be performed for the new conductor
which will become part of the existing OHL system.
This paper presents a holistic methodology that enables
evaluation of the maximum sag and ampacity of a conductor
considering its structure and the electromechanical properties
of its basic materials as well as OHL structure constraints.
This methodology provides the flexibility to evaluate the performance of nonstandard conductors on a prespecified OHL
system. The validation of this methodology with a real system
application is also presented.
Many studies have dealt with calculating an aerial bare conductor’s sag and ampacity at a steady-state conductor temperature. Most of this paper makes use of the widely accepted “ruling
span” method of sag-tension calculation for multiple suspension
spans. The method provides solutions to the parabolic and hyperbolic equations which define the relationship between span,
sag, and tension.
Considerable work on the topic took place during the 1950s
and 1960s, mainly focusing on graphical and analytical methods
for sag-tension calculations [2]–[4] and methods for the estimation of the current-carrying capacity of ACSR conductors [5],
[6]. A method for sag-tension calculation based on stress-strain
and temperature elongation data obtained on ACSR conductors
was presented in [7] and had been further extended and developed to a computer program (STESS) for use in transmission line and operation [8]. A similar stress-strain approach is
also followed by [9] in an attempt to deal with the limitations
of existing methods and software concerning gap-type conductors. Other recent modifications to the methods include a hy-
0885-8977/$31.00 © 2012 IEEE
brid numerical method to calculate the sag of composite conductors [10] and an “aluminum stress method” which allows vibration constraints to be based on specified values of static tensile aluminum stress [11]. There are also commercially available
sag-tension programs which give good results for most practical
applications [12], [13].
These methods were initially developed to evaluate the performance of conventional (AAAC and ACSR) conductors by
also incorporating experimental measurements of existing conductors. Recent research in this area has been extended to include novel (HTLS) conductors in these evaluations [14]–[17].
A comparison of the performance of ACSS with ACSR and aluminum conductors aluminum reinforced (ACAR) is presented
in [15] assuming the system is under specific extreme heavy
loading conditions. With this specific comparison, the ACSS
conductor appears to perform better because of the different
maximum tension at the heavy load condition [15]. In [16],
ACSS, ZTACIR, and CTACSR conductors are also compared
with ACSRs due to their bimetallic similarities.
Most of the sag-tension methods and packages, even though
producing acceptable predictions, are limited to estimating
conductors’ performance based on data provided by conductors’ manufacturers. Hence, they are not flexible to investigate
how nonstandard conductors will perform. Furthermore, limited work has been done to examine conductor performance
incorporating the properties of the structure (wood pole or
lattice towers) to evaluate the real benefit of different HTLS
conductors when compared with ACSRs and AAACs.
In order to progress efficiently to these investigations, a more
holistic approach should incorporate, along with the weather
loading and conductor properties, the system limitations to
allow nonconventional conductors (in size and materials) to be
evaluated on prespecified systems. These considerations are
taken into account in the methodology introduced in this paper.
This allows the overall performance evaluation of lighter conductor technologies with different sag performance, instead of
limiting the comparisons only to conductors’ sag performance.
Furthermore, reconductoring scenarios can be investigated
considering new technologies and conductors of composite
materials and nonstandard sizes.
A. Outline of the Computations
The current rating of a conductor erected on a particular
OHL structure at a specified location is controlled by the
weather, the conductor, the OHL structure, and operational
conditions. This current rating specifies the maximum power
transfer capability of the OHL system. The overall performance
of the system is affected by properties that can be divided
into three distinct groups: 1) mechanical; 2) electrical; and 3)
aging. Consequently, the computations are performed in three
levels: 1) mechanical; 2) electrical; and 3) aging. Every level is
influenced by the weather, conductor structure, and operational
conditions. These three distinct computational levels are linked
together as shown in Fig. 1 in order to compute the conductor
maximum current capacity and maximum sag. The input data
are divided into four groups: 1) the OHL support structure
Fig. 1. Flowchart of the computations.
data; 2) the weather data; 3) the conductor data; and 4) the
operational data.
OHL data describe the structure at the time of installation
of the wires. These include the design installation tensile stress
of the electrically unloaded conductor, the ambient temperature
during the installation, the span length, the type of insulation
sets, and the difference between the insulation set attachment
levels as well as their heights from the ground. The last three
are determined by the support structure (i.e., the lattice tower or
wood pole) of the OHL system.
The weather data include the ice thickness, ice density, and
the wind speed at a given ambient temperature which define
the designed maximum weather loading. These data are usually described by weather maps [18]–[20] or can be derived
from historical data. These determine the absolute maximum
working tension (AMWT) that the conductors may experience
and, therefore, the maximum sag value developed at low temperatures. This group also includes weather data that are used
for the current-temperature calculations as described in [21].
The conductor data involve a conductor’s electrical and mechanical properties as well as its physical design. The most basic
properties are the density, modulus of elasticity, coefficient of
thermal expansion, tensile strength, conductivity, stranding pattern, grease pattern and density, and type of strands (trapezoidal
or cylindrical), as well as their diameter. These are usually provided by conductor manufacturers and are included in relevant
standards [22]–[29] for the most standard conductor types.
The last group, operational data, includes the maximum operating temperature of the conductor, system frequency, methodology used for mitigation of the conductor creep (if any), predicted duration of the maximum conductor operating temperature, designed emergency loading, and duration of operation at
the elevated temperature.
Some of the variables that describe the overall structure have
to be predefined in order to initiate the computations. These variables are the ones that define the weather loading, the maximum
operating loading, and the OHL structure.
The methodology summarized in Fig. 1 and detailed in subsequent sections can then be used to calculate and compare
the electrical and mechanical performance of different conductors for the same OHL structure and identify the most suitable for the particular structure. Changes in the OHL structure
cause changes in the performance of the same conductor, as
do weather changes (e.g., maximum ambient temperature, maximum wind speed, ice loading, etc.).
Fig. 1 also shows that the computations at the aging level are
performed at the end in order to calculate the final conductor
sag. This computation level can also be used to evaluate the increase of initial conductor tension that is needed to balance the
plastic strain of the conductor. Furthermore, if pretensioning of
the conductor is considered in the initial input data (affecting the
OHL and conductor data groups), this part may be omitted, for
simplification, and the final system condition would be the conductor sag calculated at the mechanical level with the same conductor ampacity. However, conductor pre-tensioning may not be
permitted by the strength of the support structure.
B. Mechanical Computations
The mechanical part of the computations is performed to calculate the conductor sag and tension at the designed maximum
operating temperature of the conductor
. In order to
achieve this, the maximum conductor tension (MCT) of the
specified OHL system is required.
The MCT (i.e., the maximum working tension of the conductor on a particular OHL system) is controlled by the following limitations that are generally applied in any structure:
• the maximum permitted tension allowed by the weakest
system component which may be the insulator pin, the insulator, or the tower/pole structure; this tension limit is defined in the calculations as the structure maximum working
tension (SMWT);
• the self-damping vibration limit tension of the conductor
effectively defines the everyday tension (EDT) at a specified conductor temperature; this limit is usually employed
to reduce the aeolian vibrations to an acceptable level
[20] and determines the conductor’s vibration limited
maximum working tension (VLMWT);
• the absolute conductor working tension (ACWT) for the
specified weather loading, for example, 50% of the conductor rated breaking strength (RBS) at 5.6 C with combined wind and ice [20], [29]; the weather (ice and wind)
loading and the minimum temperature of the conductor influence the MCT since they affect the elastic (weight) and
thermal (temperature) elongation of the cable.
The procedure to determine the MCT of the power-line
design is illustrated in Fig. 2. Initially, the weather-loading
conditions of a power line (for example, load cases 1 to 6 in
[29] and [30]) are identified by its location and the help of
weather maps [20], [29] or historical data. The weather-loading
conditions also define the corresponding safety factors (SF)
[29]–[31]. These safety factors are applied to the conductor
Fig. 2. Flowchart of the procedure for MCT calculation of an OHL system.
and the structure (insulator/insulator pin), and the smallest
maximum allowable tension of these is set as the absolute
maximum working tension (AMWT) of the conductor applied
on the specified system (Fig. 2).
The Newton–Raphson iteration method is employed on the
change of state (1), derived from the catenary curve. In (1), the
elastic and thermal elongations of the conductor are included
with the plastic elongation computed separately since the latter
is affected by the operating conditions [32]
Thermal elongation
Elastic elongation
where subscripts
respectively, and
define the final and initial conditions,
horizontal conductor tensile force;
resulting conductor weight per-unit length;
span length;
conductor temperature;
conductor’s coefficient of thermal expansion;
conductors modulus of elasticity;
conductor cross-sectional area.
The everyday tension (i.e., the design stress in the unloaded
conductor that is applied to minimize the aeolian vibrations
of the conductor at a designed temperature) and the conductor
weight are set as initial conditions while the final conditions are
the conductor resulting weight (i.e., the vectorian sum of the
conductor, wind, and ice weight) and the VLMWT at the maximum-designed weather loading.
The output VLMWT of this iteration is then compared with
the AMWT and the smallest one defines the MCT of the OHL
system at the specified weather loading.
Once the MCT is known, then a second Newton–Raphson iteration of (1) takes place having as initial conditions the MCT
of the OHL system and the conductor resulting weight (i.e., the
sum of the conductor, wind, and ice weight) at the maximum
designed weather loading temperature, in order to identify the
conductor tension at any operating temperature
(Fig. 1). Once the conductor tension at the operating temperature is calculated, the sag is computed by using the catenary
curve. The output is then linked with the other two computational parts.
The MCT computation (described in Fig. 2) is performed
based on the conductor design and material properties as illustrated in Fig. 3. This condition is set as a reference point and
then using (1), the final tension and sag can be evaluated at different temperatures with an iteration process using small (1 C)
temperature steps. The aluminum and steel tensions are computed from the conductor tension incorporating the conductor
creep. The iteration continues up to the knee point, which defines the temperature at which the aluminum contributes zero
force to the conductor. The tension and sag at this point are
then set as initial conditions to perform calculations for temperatures above the knee point. A new iteration takes place to
evaluate the final tension and sag conditions above this point.
This is performed using a modified version of (1), where the
that affect the thermal and
conductor properties
elastic elongations are replaced with the values that correspond
, since the core deto the conductor’s core
fines the elastic and thermal elongations of the equation above
the knee-point temperatures.
C. Electrical Computations
The electrical part of the computation is used to calculate the
of any conductor at any temperature deac resistance
fined in the previous part of the computations instead of using
the linear interpolation and the tabulated data of [33] for the
standard conductor sizes described in [21]. This increases the
flexibility in the type and size of conductors that can be examined, since this method is not limited to the conductors included
in [33]. It also improves accuracy.
The detailed computations within this step of the process are
illustrated in the flowchart of Fig. 4. The basic electrical and
physical properties of the materials used for the conductor, as
well as the conductor itself are used to calculate the dc resistance
at 20 C based on the conductivity, thermal coefficients,
diameter and number of strands, and spiraling factors specified by ASTM for cylindrical and trapezoidal strands [22]–[29].
For wires of distinct strength (core) and conductive (aluminum)
is calculated separately and every part is
members, the
then corrected to the operational temperature by using the apof the
propriate temperature coefficients [22]. The overall
conductor is then computed by considering core and outer conductive member resistances in parallel configuration. When the
Fig. 3. Flowchart of mechanical computations for the conductor tension and
sag at final system conditions.
conductor’s core is of the same material and shape of strands,
then the
is calculated by omitting the core’s calculations.
, the skin factor is calculated
To compute conductor
based on the physical structure of the conductor. The skin
factor calculation is based on the work of Dwight [34]–[36]
and Lewis and Tuttle [37] which is further simplified with the
use of [38] in the code presented in this paper. The skin factor
calculation is performed on cylindrical and tubular conductor
shapes, in order to evaluate the effect of non-conductive core
materials on conductors.
calculations is then
A second correction factor for the
applied for cases when steel material is used for the core design
of the conductor, in order to address the magnetization effect
. This effect is negligible
of the core on the conductor’s
on ACSR conductors with even numbers of aluminum layers
[21], [33], [37], [39] and so, it is not considered in these cases.
For the ACSR conductors with an odd number of aluminum
layers, however, the magnetization factor is used to accurately
(Fig. 4). In the case of single-layer ACSR,
calculate the
the correction method described in [21] is used, while for the
three-layer conductors, the approach of [39] is employed.
Since the magnetization factor is influenced by current flow
through the conductor, an iteration with the use of [21] takes
place to correct the calculation as illustrated in Fig. 4. This
makes the resistance of steel core conductor technologies dependent on conductor temperature and current flow, unlike other
conductor technologies whose resistance is dependent on the
conductor temperature.
Fig. 4. Synopsis of the ac resistance calculation methodology used.
The method described here is employed for calculating
of any aluminum alloy conductors (AAC), ACAR,
AAAC, and ACSR or any other bimetallic and “bimaterial”
conductor technology, type, and size. It can, furthermore, be
used for “tubular stranded” conductors or more practically
for non-conductive composite core stranded conductors. The
computation between
difference in the final result of the
of the
different types of conductors is determined by the
overall cross-section area of the conductor, the skin factor, and
the magnetization factor differences.
D. Aging Computations
Within this part of the computation, the nonelastic elongation
(i.e., conductor creep) of the conductors under the permanent
tensile load is calculated for the operating conditions and conductor temperature defined during the mechanical computation
Aging computations are divided into two clusters depending
on the type of conductor. The first one regards the well-established conductor types AAC, AAAC, ACAR, and ACSR and the
second is in reference to the recently developed aluminum conductor composite reinforced (ACCR) and aluminum conductor
composite core (ACCC) composite conductors [14], [40]. In the
first cluster of computations, the predictor equations of Table I
are used under the prespecified designed operating conditions of
the system [41]–[43]. These conditions are defined as follows.
• EDT and the ambient temperature at which this tension
occurs as well as its predicted duration.
• Maximum operating conductor temperature, conductor
tension at that temperature, and its predicted duration.
• Conductor’s MCT at the designed ambient temperature at
which this tension occurs and its predicted duration.
The computations include the prediction of elevated temperature creep (Table I) which occurs above 75 C for aluminum
conductors and 100 C for the steel-reinforced conductors [42].
The second cluster of the aging computation is very similar to
the first one but instead of using the well-established predictor
equations (Table I) adapted from [41], the stress-strain curves
produced by testing the conductors are used in the absence of
standard data. It should also be noted that the aging computations use an iteration process, since the stress is reduced with
time as the creep strain gradually increases.
The methodology described in this paper is used to predict
the sag and identify the ampacity of a 400-kV L2-type lattice
tower OHL system after 29 years from the initial installation
of the conductor. The system was strung with a twin-bundle
Rubus AL5 aluminum alloy conductor. The conductor’s electrical and mechanical properties are computed using the described methodology.
Table II shows the conditions relevant to the conductor’s installation along with the surveyed and predicted sag values after
29 years of the OHL’s operation. As can be observed, the sag
prediction is very good and the calculated creep for the 29 years
. The differences in span lengths beis 597.2 microstrains
tween the left and right circuits in the first and fourth spans
are due to the use of angle towers at the tension points. This
difference in tower type results in nonparallelism of the contiguous tower cross-arms and, therefore, in different conductor
span lengths for the right and left circuits.
Fig. 5. Surveyed profile of the critical span complemented with calculated catenary curves at surveyed and maximum electrical loading temperatures.
3 Difference between the measured sag and the calculated predicted values.
The ampacity calculations at the surveyed operating conductor temperatures resulted in similar values as those given by
the survey data for both circuits as illustrated in Table III. The
sag at 75 C (maximum electrical loading) is also calculated
(Table II) and the catenary curve of the most critical span of this
OHL system is fitted on the conductor (surveyed) sag profile.
This is illustrated in Fig. 5 and as can be seen, the conductor
preserves the 7.6 m of phase-to-ground required clearance [30],
[44]. This, therefore, allows for higher operating conductor
temperatures without the requirement of re-tensioning the conductor. Consequently, it enables increasing the line’s thermal
rating, but with an increase in conductor creep that will be
caused by the elevated operating temperature according to [41].
In order to compute the new maximum operating temperature
(and, therefore, the system’s ampacity), the calculations should
be performed again from the beginning and the estimated
amount of hours that the line would operate above 75 C should
also be specified. This step is important since the elevated
conductor temperature could affect the conductor aging and,
thus, the ground clearance after a long period of the conductor
in service.
The methodology discussed here links the conductor properties with the properties defining the OHL structure in a way
Fig. 6. Flowchart for choosing the appropriate technique for uprating the
thermal limit of an aerial power system.
that enables investigating the benefits afforded by uprating techniques, such as re-tensioning and reconductoring, as illustrated
in the flowchart in Fig. 6. It can also be used to evaluate the improvement in performance that different conductor sizes and/or
types may result in, and compares this with different conductors
and system operating conditions on a prespecified OHL system
(Fig. 6). Hence, it enables the overall performance evaluation
of lighter conductor technologies on a specified system, rather
than comparing only their sag performances.
The methodology can then be used to compare the electrical
and mechanical performance of different conductors for the
same OHL structure and identify the one that allows more electrical power to be transferred through the specified structure
and evaluate the losses. Changes in the OHL structure cause
changes in the system and, thus, the performance of the same
conductor would be different. Weather changes (e.g., ambient
temperature, maximum wind speed, etc.) also influence the
performance of the overall system.
This methodology can be used to identify the effect of increase in operating temperature as well as the impact of vibration dampers on the conductor or the effect of increases in ambient temperature (due to global warming) [45]. Furthermore, it
can be used to identify the effect of changes on the OHL design
(i.e., increasing the height of the conductor attaching points or
their maximum tensile strength). It can also compute the additional creep effect caused by the initial overtension applied to
negate the conductor’s plastic deformation.
The electrical computation section addresses all of the
bimetallic conductors and with small modifications in the
spiraling factors, skin effect, and magnetization factor, it can
also be used for composite core conductors with or without
conductive cores (i.e., the conductors in [40] and [14]). Consequently, it increases the flexibility of the calculations at any
temperature for any conductor technology and design (even
for those for which the properties are known but still not fully
developed), and allows the direct comparison of the performance of different conductor sizes and technologies on the
same OHL structure. For example, modeling the performance
of the different nonexisting ACCC/TW core to aluminum
ratio conductors can be performed to find the optimal fit for a
particular OHL system at a given location.
This methodology has already been used to show how different conductor technologies behave in standard U.K. OHL
structures [45]–[49], highlighting the potential of reconductoring with novel HTLS conductors.
The current rating and sagging performance of a conductor
erected on an OHL structure are influenced by the weather,
the conductor, and its installation method, the OHL structure,
and operational conditions. The methodology presented in this
paper links all of these parameters together in order to evaluate
the actual performance of a conductor on a prespecified structure. It enables performance comparison of different conductor
types and sizes on the same structure. This is important since,
in some cases, the structure limits the conductor’s performance.
The accuracy of this methodology in predicting conductor sag
and current flow is also evaluated. The validation process with a
four-span double-line system resulted in correct predictions for
the 29-years operation time.
The main contribution of this methodology is the flexibility
on calculating the conductor properties rather than obtaining
them from a database as well as incorporating OHL structure
constraints (weight and tension). It can be used to evaluate potential uprating by reconductoring, and increased temperature
of operation by retensioning on a given line. Furthermore, it
can help with the choice of the most suitable conductor for
a given system and, thus, identify the best option in terms of
conductor size and technology type, suggesting new conductor
[1] V. T. Morgan, “The loss of tensile strength of hard-drawn conductors
by annealing in service,” IEEE Trans. Power App. Syst., vol. PAS-98,
no. 3, pp. 700–709, May 1979.
[2] P. F. Winkelman, “Sag-tension computations and field measurements
of bonneville power administration,” Trans. Amer. Inst. Elect. Eng. Part
III, Power App. Syst., vol. 78, pp. 1532–1547, 1960.
[3] T. P. Harley, “A direct method for sag-tension calculations,” Trans.
Amer. Inst. Elect. Eng. Power App. Syst., vol. 72, pp. 603–608, 1953.
[4] M. Landau, “Incremental method for sag-tension calculations,” Trans.
Amer. Inst. Elect. Eng., vol. 70, pp. 1564–1571, 1951.
[5] H. E. House and P. D. Tuttle, “Current-carrying capacity of ACSR,”
Trans. Amer. Inst. Elect. Eng. Part III Power App. Syst., vol. 77, pp.
1169–1173, 1959.
[6] J. H. Waghorne and V. E. Ogorodnikov, “Current carrying capacity
of ACSR conductors,” Trans. Amer. Inst. Elect. Eng., vol. 70, pp.
1159–1162, 1951.
[7] O. Nigol and J. S. Barrett, “Characteristics of ACSR conductors at high
temperatures and stresses,” IEEE Trans. Power App. Syst., vol. PAS100, no. 2, pp. 485–493, Feb. 1981.
[8] J. S. Barrett, S. Dutta, and O. Nigol, “A new computer model of ACSR
conductors,” IEEE Trans. Power App. Syst., vol. PAS-102, no. 3, pp.
614–621, Mar. 1983.
[9] I. Albizu, A. J. Mazon, and I. Zamora, “Flexible strain-tension calculation method for gap-type overhead conductors,” IEEE Trans. Power
Del., vol. 24, no. 3, pp. 1529–1537, Jul. 2009.
[10] A. Alawar, E. J. Bosze, and S. R. Nutt, “A hybrid numerical method
to calculate the sag of composite conductors,” Elect. Power Syst. Res.,
vol. 76, pp. 389–394, 2006.
[11] J. S. Barrett and Y. Motlis, “Allowable tension levels for overhead-line
conductors,” Proc. Inst. Elect. Eng., Gen., Transm. Distrib., vol. 148,
pp. 54–59, 2001.
[12] Southwire. [Online]. Available: http://www.sag10.com/support/SupportDocumentation.htm
[13] A. K. Deb, Powerline Ampacity System: Theory, Modeling, and Applications. Boca Raton, FL: CRC, 2000.
[14] A. Alawar, E. J. Bosze, and S. R. Nutt, “A composite core conductor
for low sag at high temperatures,” IEEE Trans. Power Del., vol. 20, no.
3, pp. 2193–2199, Jul. 2005.
[15] H. W. Adams, “Steel supported aluminum conductors (SSAC) for overhead transmission lines,” IEEE Trans. Power App. Syst., vol. PAS-93,
no. 5, pp. 1700–1705, Sep. 1974.
[16] I. Zamora, A. J. Mazon, P. Eguia, R. Criado, C. Alonso, J. Iglesias, and
J. R. Saenz, “High-temperature conductors: A solution in the uprating
of overhead transmission lines,” in Proc. IEEE Power Tech, Porto, Portugal, 2001, pp. 1–6.
[17] A. G. Exposito, J. R. Santos, and P. Cruz Romero, “Planning and operational issues arising from the widespread use of HTLS conductors,”
IEEE Trans. Power Syst., vol. 22, no. 4, pp. 1446–1455, Nov. 2007.
[18] Overhead electrical lines exceeding AC 1 kV up to and including AC
45 kV—Part 1: General requirements—Common specifications, British
Standard BS EN 50423-1, 2005.
[19] Overhead electrical lines exceeding AC 45 kV—Part 1: General
requirements—Common specifications, British Standard BS EN
50341-1, 2001.
[20] ENATS 43-40: Single circuit overhead lines on wood poles for use at
high voltage up to and including 33 kV, ENATS 43-40, 2004, no. 2,
Energy Netw. Assoc. Tech. Spec..
[21] IEEE Standard for Calculating the Current-Temperature of Bare Overhead Conductors, IEEE Std. 738-2006 (Revision of IEEE Standard
738-1993), 2007, 1–59.
[22] “ASTM B 193-02: Test method for resistivity of electrical conductor
materials,” Annual Book of ASTM Standards, vol. 02.03, pp. 68–72,
[23] “ASTM B 230/B 230M-99: Specification for aluminum 1350-H19
wire for electrical purposes,” in Annual Book of ASTM Standards. Philadelphia, PA: ASTM, 2005, vol. 02.03, pp. 91–94.
[24] “ASTM B 231/B 231M-04: Specification for concentric-lay-stranded
aluminum 1350 conductors,” in Annual Book of ASTM Standards.
Philadelphia, PA: ASTM, 2005, vol. 02.03, pp. 95–105.
[25] “ASTM B232/B 232M-01: Specification for concentric-lay-stranded
aluminum conductors, coated-steel reinforced (ACSR),” in Annual
Book of ASTM Standards. Philadelphia, PA: ASTM, 2005, pp.
[26] “ASTM B233-97(2003): Specification for aluminum 1350 drawing
stock for electrical purposes,” in Annual Book of ASTM Standards. Philadelphia, PA: ASTM, 2005, vol. 02.03, pp. 122–125.
[27] “ASTM B 498/B 498M-98(2002): Specification for zinx-coated
(galvanized) steel core wire for aluminum conductors, steel reinforced
(ACSR),” in Annual Book of ASTM Standards. Philadelphia, PA:
ASTM, 2005, vol. 02.03, pp. 218–221.
[28] “ASTM B 779-03: Standard specification for shaped wire compact concentric-law-stranded aluminum conductors, steel reinforced (ACSR/
TW),” in Annual Book of ASTM Standards. Philadelphia, PA: ASTM,
2005, vol. 02.03, pp. 309–314.
[29] Conductors for overhead lines—Round wire concentric lay stranded
conductors, British Standard BS EN 50182, 2001.
[30] Overhead electrical lines exceeding AC 45 kV—Part 3: Set of national
normative aspects, British Standard BS EN 50341-3, 2001.
[31] National Electrical Safety Code 2007 Edition, IEEE Standard C2-2007,
Aug. 2006.
[32] K. Kopsidas, “Modelling thermal rating of arbitrary overhead line
systems,” Ph.D. dissertation, Elect. Electron. Eng. Dept., Univ. Manchester, Manchester, U.K., 2009.
[33] , L. Kirkpatrick, Ed., Aluminum Electrical Conductor Handbook.
Washington,, DC: Aluminum Assoc., 1989.
[34] H. B. Dwight, “Skin effect in tubular and flat conductors,” AIEE Trans.,
vol. 37, pp. 1379–1403, 1918.
[35] H. B. Dwight, “Skin effect and proximity effect in tubular conductors,”
AIEE Trans., vol. 41, pp. 189–198, 1922.
[36] H. B. Dwight, “A precise method of calculation of skin effect in isolated
tubes,” AIEE Trans., vol. 42, pp. 827–831, 1923.
[37] W. A. Lewis and P. D. Tuttle, “The resistance and reactance of
aluminum conductors, steel reinforced,” AIEE Trans., vol. 77, pp.
1189–1215, 1959.
[38] F. W. J. Olver, “9. Bessel functions of integer order,” in Handbook
of Mathematical Functions, M. Abramowitz and I. A. Stegun, Eds.
Washington, D. C.: U.S. Dept. Commerce, 1964, pp. 378–385.
[39] B. S. Howington and L. S. Rathbun, “AC resistance of ACSR—Magnetic and temperature effects,” IEEE Trans. Power App. Syst., vol. PAS104, no. 6, pp. 1578–1584, Jun. 1985.
[40] 3M. Aluminum Conductor Composite Reinforced Technical Notebook (477 kcmil Family) Conductor & Accessory Testing. 2006.
[Online]. Available: http://www.energy.ca.gov/2004_policy_update/documents/2004-06-14-workshop/public_comments/2004-0628_3M_PART2.PDF
[41] Overhead Conductor Manual, R. Thrash, A. Murrah, M. Lancaster, and
K. Nuckles, Eds. Carrollton, GA: Southwire, 2007.
[42] IEEE Guide for Determining the Effects of High-Temperature Operation on Conductors, Connectors, and Accessories, IEEE Standard
1283-2004, 2005, 1–28.
[43] “CIGRE SC22-WG05: Permanent elongation of conductors. Predictor
equation and evaluation methods,” Electra N 75, pp. 63–98, 1981.
[44] “ENATS 43-8: Overhead line clearances,” Energy Netw. Assoc. Tech.
Spec., no. 3, 2004.
[45] K. Kopsidas and S. M. Rowland, “A performance analysis of reconductoring an overhead line structure,” IEEE Trans. Power Del., vol. 24, no.
4, pp. 2248–2256, Oct. 2009.
[46] K. Kopsidas, S. M. Rowland, M. N. R. Baharom, and I. Cotton, “Power
transfer capacity improvements of existing overhead line systems,” in
Proc. IEEE Int. Symp. Elect. Insul., 2010, pp. 1–5.
[47] K. Kopsidas and S. M. Rowland, “Evaluating opportunities for increasing power capacity of existing overhead line systems,” Inst. Eng.
Technol. Gen., Transm. Distrib., vol. 5, pp. 1–10, 2011.
[48] K. Kopsidas and S. M. Rowland, “Investigating the potential of re-conductoring a lattice tower overhead line structure,” in Proc. IEEE Power
Eng. Soc. Transm. Distrib. Conf. Expo., 2010, pp. 1–8.
[49] K. Kopsidas, I. Cotton, and S. M. Rowland, “Towards a holistic
perspective of the existing overhead line power network to facilitate
its flexible expansion,” in Proc. IEEE/Power Energy Soc. Power Syst.
Conf. Expo., 2011, pp. 1–7.
Konstantinos Kopsidas (M’06) was born in Lefkas,
Greece. He received the B.Eng. degree in electrical
and electronic engineering from The University of
Manchester Institute of Science and Technology
(UMIST), Manchester, U.K., in 2004, and the
M.Sc. (Hons.) degree in electrical power system
engineering, and the Ph.D. degree in electrical power
engineering from The University of Manchester,
Manchester, U.K., in 2005 and 2009, respectively.
He was a Research Assistant/Associate with the
SuperGen V (AMPerES) and is now a Lecturer with
the School of Electrical and Electronic Engineering, The University of Manchester, also collaborating with Arago Technology Ltd., Manchester, U.K.
Dr. Kopsidas received the Scottish Power “Power Learning Award” and the
UMIST course prize twice.
Simon M. Rowland (SM’07) was born in London,
U.K. He received the B.Sc. degree in physics from
the University of East Anglia and the Ph.D. degree
from London University, London, U.K.
He has worked for many years on dielectrics
and their applications. He has also been Operations
and Technical Director within multinational manufacturing companies. He joined The School of
Electrical and Electronic Engineering, The University of Manchester, Manchester, U.K., as a Senior
Lecturer in 2003. He was appointed Professor of
Electrical Materials in 2009.
Dr. Rowland received the IEE Duddell Premium Award in 1994 and became a
Fellow of the Institute of Electrical Engineers in 2000. Currently, he is President
of the IEEE Dielectric and Electrical Insulation Society.
Boud Boumecid received the M.Sc. degree in
civil engineering from the University of Salford,
Manchester, U.K., in 1994.
He previously worked with consultants and
contractors, where he was involved in the civil and
structural design of power stations, substations, and
overhead lines. He is a Design Engineer (Overhead
Lines) with the Electricity Network Investment
Department—Asset Policy, National Grid, U.K.
Mr. Boumecid is an active member of CIGRE
Study Committee SC B2 (Working Group 23 on
Foundations) and of the BSI-PEL/011 Committee.
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