Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures Zuheir Barsoum

Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures Zuheir Barsoum
KTH Engineering Sciences
Residual Stress Analysis and Fatigue
Assessment of Welded Steel Structures
Zuheir Barsoum
Doctoral Thesis
Stockholm, Sweden 2008
Division of Lightweight Structures
Department of Aeronautical and Vehicle Engineering
School of Engineering Sciences
Kungliga Tekniska Högskolan
Department of Aeronautical and Vehicle Engineering
Kungliga Tekniska Högskolan (KTH)
SE-100 44 Stockholm
Academic dissertation which with permission of Kungliga Tekniska Högskolan in Stockholm
is presented for public review and doctoral examination on April 25th 2008 at 10.00 in E1,
KTH, Lindstedtsvägen 3, Stockholm.
TRITA-AVE 2008:11
ISSN 1651-7660
© Zuheir Barsoum, 2008.
PREFACE
The work in this doctoral thesis has been carried out at the Division of Lightweight Structures
at the Department of Aeronautical and Vehicle Engineering at KTH between September 2003
and April 2008. The work done in this thesis is a part of the Nordic R&D project QFAB Quality and Cost of Fabricated Advanced Welded Structures which was brought to
conclusion in 2006. The research was continued in the Swedish R&D project LOST - Light
Optimized Welded Structures, which was started in 2007.
Some of the work has also been presented and discussed within Commission XIII (Fatigue of
Welded Structures) in the International Institute of Welding (IIW) during the Annual
Assembly in Prague 2005, Quebec 2006 and Dubrovnik 2007. The work in this thesis has
been funded by Volvo Construction Equipment (Volvo CE), SSAB, The Swedish Vehicle
Research Program (PFF), the Nordic Innovation Centre (NiCe), VINNOVA and the MERA
program.
First and foremost, I would like to express my sincere gratitude to my supervisor, Prof. Jack
Samuelsson (Volvo CE and KTH), for his continued support, encouragement and patience
throughout the process of this work. His guidance as a scientist, mentor and friend is greatly
appreciated. Mr. Mats Gustafsson at SSAB, with whom I have been working during the
research project is also acknowledged.
I would also like to thank the people within the Division of Lightweight Structures. Special
thanks are also due to Tekn. Lic. Anders Björkblad, Dr. Johan Martinsson, Prof. Dan Zenkert
and Mr. Bo Magnusson. Thank you Mr. Sohrab Kazemahvazi for valuable discussions about
research, mountain biking and for never getting tiered of my endless questions. Thank you
Tekn. Lic. Markus Kaufmann for always helping me with computer problems.
Prof. Lars-Erik Lindgren and Tekn. Lic. Andreas Lundbäck, at Luleå University of
Technology is acknowledged for valuable discussion on welding simulation and residual
stress measurements.
I am also grateful to the staff within Volvo Group for their co-operation during the span of
this work e.g. Mr. Bertil Jonsson, Tekn. Lic. Magnus Byggnevi, Dr. Fethi Abdulwahab, Mr.
Mirsattar Hejazifar, Mr. Nenad Mrden, Mr. Franjo Jakopovic and Mr. Kjell Eriksson.
Special and hearty thanks go to my twin brother Dr. Imad Barsoum at the Department of
Solid Mechanics (KTH) who I have had the pleasure to work with and write one of the papers
within this thesis. Thank you also for being a friend, a brother and a mentor from my very
first day in this world.
To Gina, Carla, Abbe and Amelin et al. just to mention a few of you and not being longwinded; thank you for a solid friendship, I know that it will last for a lifetime!
Last but not least, I am forever grateful to Nadira and Samir, my mother and father, Fadi, my
younger brother, for their unconditional love and support. God bless you!
Stockholm, February 2008
Zuheir Barsoum
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ABSTRACT
This doctoral thesis is concerned with fatigue life of welded structures. Several topics related
to fatigue of welded structures are treated such as; weld defects and their influence on fatigue
performance of welded structures, fatigue life prediction using LEFM (Linear Elastic Fracture
Mechanics), fatigue testing, welding simulation, residual stress prediction and measurement
and their influence on fatigue life.
The work that is reported in this doctoral thesis is part results of a Nordic R&D project QFAB
(Quality and Cost of Fabricated Advanced Welded Structures) and a Swedish R&D project
LOST (Light Optimized Welded Structures). One of the main objectives is to compare
different welding processes with respect to fatigue performance and gain understanding of the
weld defects, their appearance for different welding processes and their influence on fatigue
life. Another main objective is to study welding residual stresses and their effect on fatigue.
Fatigue design rules are in some cases conservative and further knowledge about the residual
stress field, especially on the weld root side, may improve the accuracy of life prediction
calculations. The aim is to develop simplified procedures for analysis of residual stresses,
their relaxation and influence on fatigue life.
Fatigue testing of Hybrid Nd: YAG laser/MAG and MAG welded (tandem arc solid wire, flux
cored wire, tandem flux cored wire) non-load carrying cruciform joints was carried out. Four
batches were produced, tested and the results were compared. The local weld geometry of the
cruciform welded joints was measured and analyzed. Residual stress measurement was
carried out close to the toe region using X-ray diffraction. Weld defects, in most cases cold
laps, in the cracked specimens were measured.
Further fatigue testing, weld defect assessment and residual stress and local weld geometry
measurements were carried out on joints welded with flux cored and metal cored arc wires.
Two-and three dimensional LEFM crack growth analysis was carried out in order to assess
the influence of weld defects, local weld geometry and residual stresses.
Residual stresses in multi-pass welded tube-to-plates were studied for two different tubular
joint configurations; a three-pass single-U weld groove for maximum weld penetration and a
two-pass fillet (no groove) weld for minimum weld penetration. Torsion fatigue tests were
performed in order to study crack propagation from the weld root. Mode III propagation from
the lower and upper weld toe on the same tubular joints was also studied. Some tubes were
stress relieved (PWHT) and some were fatigue tested with internal static pressure.
A three dimensional finite element welding simulation of the multi-pass welded tubular joint
was carried out. The calculated temperatures in the transient thermal analysis were compared
with measured temperatures. The FE predicted residual stresses in the as-welded conditions
were verified with hole drilling strain gage measurements. The residual stresses were used as
internal stresses in the finite element model for the torsion fatigue simulation in order to study
the cycle by cycle relaxation of the residual stresses in constant amplitude torsion loading.
A two dimensional finite element welding simulation procedure was developed in order to
predict welding residual stress. These were used together with a developed 2D LEFM
subroutine to predict the fatigue life, crack path and the effect of residual stresses on weld root
defects. The developed simulation subroutines were verified with results found in the
literature.
Residual stresses measurement and two-and three dimensional welding simulations were
carried out in fillet welded joints in order to study the three dimensional effects of the welding
process, boundary conditions and modelling technique on the formation of residual stresses.
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APPENDED PAPERS
Paper A
Barsoum Z. and Samuelsson J., Fatigue Assessment of Cruciform Joints Welded with
Different Methods, published in Steel Research International, Vol.77, No.12, 2006
This paper was also presented and discussed within Commission XIII (Fatigue of Welded
Structures) in the International Institute of Welding (IIW) during the annual assembly, Prague
2005.
Paper B
Barsoum Z. and Jonsson B., Fatigue Assessment and LEFM Analysis of Cruciform Joints
Fabricated with Different Welding Processes, accepted for publication in Welding in the
World, 2007.
This paper was also presented and discussed within Commission XIII (Fatigue of Welded
Structures) in the International Institute of Welding (IIW) during the annual assembly,
Dubrovnik, Croatia, 2007.
Paper C
Barsoum Z., Residual Stress Analysis and Fatigue of Multi-pass Welded Tubular
Structures, published in Engineering Failure Analysis, article in press, 2008.
This paper was also presented and discussed within Commission XIII (Fatigue of Welded
Structures) in the International Institute of Welding (IIW) during the annual assembly, Prague
2005.
Paper D
Barsoum Z. and Samuelsson J., Residual Stress Prediction and Relaxation in Welded
Tubular Joint, published in Welding in the World, Vol. 51, Issue 1/2, 2007.
This paper was also presented and discussed within Commission XIII (Fatigue of Welded
Structures) in the International Institute of Welding (IIW) during the annual assembly,
Quebec City, Canada, 2006.
Paper E
Barsoum Z. and Barsoum I., Residual stress effect on fatigue life of welded structures
using LEFM, submitted for publication, 2008.
Paper F
Barsoum Z. and Lundbäck A., FEM welding simulation of fillet welds – 3D effects on
residual stresses, to be submitted for publication, 2008.
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PAPERS NOT INCLUDED IN THIS THESIS
Barsoum Z. and Gustafsson M., Spectrum Fatigue of High Strength Steel Joints Welded
with Low Temperature Transformation Consumables, Presented at 2nd International
Conference on Fatigue Design, Senlis, France, 2007.
DIVISION OF WORK BETWEEN AUTHORS
Paper A
Z. Barsoum carried out fatigue testing, measurements, defect detection, evaluating results,
planned the work and wrote the paper. J. Samuelsson contributed to the paper with valuable
comments and discussion.
Paper B
Z. Barsoum carried out fatigue testing, measurements, defect detection, evaluating results,
FEM analysis, fracture mechanical analysis, planned the work and wrote the paper. B.
Jonsson carried out the 3D fracture mechanics analysis.
Paper E
Z. Barsoum developed the FE subroutines for; welding simulations, LEFM automatic
simulation of fatigue crack propagation. Z. Barsoum also carried out the validation,
implementation of the developed subroutine on a welded structure, planned the work and
wrote the paper. I. Barsoum contributed with the remeshing algorithm.
Paper F
Z. Barsoum carried out the 2D welding simulations in Ansys and assisted in the 3D welding
simulations. Z. Barsoum planned the simulation work, measurements, fabrication of specimen
and wrote the paper. A. Lundbäck carried out the welding simulations in Marc and
contributed to the writing of the paper.
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vii
CONTENTS
ABSTRACT ............................................................................................................................... .i
APPENDED PAPERS ............................................................................................................... ii
PAPERS NOT INCLUDED IN THIS THESIS........................................................................ iii
DIVISION OF WORK BETWEEN AUTHORS...................................................................... iii
CONTENTS .............................................................................................................................. iv
INTRODUCTION...................................................................................................................... 1
Background ............................................................................................................................ 1
Research aim .......................................................................................................................... 3
Research approach.................................................................................................................. 3
FATIGUE LIFE ASSESSMENT............................................................................................... 4
Fatigue testing ........................................................................................................................ 5
Effective notch stress method ................................................................................................ 5
Linear elastic fracture mechanics (LEFM)............................................................................. 8
Mixed mode crack growth.................................................................................................... 10
WELD DEFECTS .................................................................................................................... 13
HEAT EFFECTS OF WELDING............................................................................................ 15
Thermal modelling and simulation ...................................................................................... 16
Welding residual stresses ..................................................................................................... 20
Finite element modelling of welding residual stresses......................................................... 20
Material modelling ............................................................................................................... 21
RESIDUAL STRESS MEASUREMENT TECHNIQUES ..................................................... 22
RESIDUAL STRESS EFFECT ON FATIGUE OF WELDS.................................................. 23
Effect of external loading..................................................................................................... 24
Effect on crack propagation ................................................................................................. 25
DISCUSSION .......................................................................................................................... 28
CONCLUSIONS...................................................................................................................... 29
SUGGESTION FOR FUTURE WORK .................................................................................. 30
SUMMARY OF APPENDED PAPERS.................................................................................. 30
REFERENCES......................................................................................................................... 31
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
INTRODUCTION
Background
Fatigue failure is still a dominating cause for breakdown of welded structures in construction
and mining equipment, trains, ships, agricultural machinery, bridges and off shore equipment,
hence leading to substantial costs. Structural details and components in these types of
structures are continuously subjected to variable amplitude loading during operation.
The demand of a more sustainable society require structures with lower weight, better
performance and in case of vehicles reduced fuel consumption. This will support the use of
efficient and more accurate fatigue design methods and the design methods must be connected
to quality requirements which can be understood and managed during production.
Eighty percent of the main structures and components in construction machinery are welded
steel structures fabricated from a variety of different steel grades. Figure 1 shows several
typical components in a Volvo Wheel Loader that are welded. Many of this structures are
complex regarding both geometries and loading conditions.
But welding without any improvement gives rise to local stress concentration, residual
stresses and different types of defects, these features combined with high applied cyclic and
complex service loading give rise to failure due to fatigue.
Figure 1. Examples of components that are welded in a Volvo Wheel Loader.
Stress concentrations at the weld toe and root are caused by the geometrical discontinuities
and, thus, fatigue cracks are easily initiated at these locations. Stress concentrations may also
result from weld defects, e.g., at weld toe from cold laps and undercuts, and at the weld root
from incomplete fusion and small effective throat thickness. These defects are, more or less,
sharp macro cracks and promote the use of the more accurate methods such as LEFM (Linear
Elastic Fracture Mechanics) together with crack propagation analysis and the effective notch
stress method to predict the fatigue life as in [1-5].
Residual stress that arises in welded joints by rapid heating and rapid cooling is another factor
in fatigue assessment of welded structures. It has been shown that tensile residual stresses in
welded structures can be as high as the yield strength of the material and have a detrimental
effect on the fatigue behaviour. Conversely, compressive residual stresses could have a
favourable effect on fatigue life [6]. However, spectrum loading may relax part of the residual
stress field which will affect the final fatigue life. The combination of welding residual
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
stresses and operating stresses to which engineering structures and components are subjected
can promote failure by fatigue.
The risk of failure can be reduced by various stress relieving processes, such as post weld heat
treatment (PWHT). In the case of root cracking, which often shows residual stresses in
compression, the PWHT may reduce the fatigue life [6]. However, the stress distribution for a
complex welded structure is usually not known, and conservative assumptions are made of the
residual stress distribution when fatigue life predictions are assessed [7].
The most widely used and most straightforward tool for structural stress analysis and LEFM
analysis is the finite element method. When it comes to finite element modelling and
simulation of the welding process in order to study the heat transfer, deformation and residual
stresses many phenomena have to be considered. Despite several simplifications, e.g.,
neglecting the micro structural changes during heating and cooling, lack of material data at
elevated temperatures and constant heat source modelling, finite element modelling of
welding is still a complex task. Three-dimensional models are used more frequently in
welding simulations due to the three-dimensional nature of the welding process although
these require a large amount of computational effort. Thus two-dimensional models are still
important in the early phase of the design and optimisation of welded structures.
The work that is reported in this thesis is part of the results of a Nordic and a Swedish welding
research project. The Nordic project, which involved the co-operation of 10 organisations
(KTH, Volvo, SSAB etc.) from Sweden, Finland, Denmark and Norway, was one of several
Nordic research and development projects that have been conducted during the last 20 years.
The main goal of these projects has been to improve the design, fatigue strength and
optimisation of welded structures. More recently reducing production costs has become an
important factor. Figure 2 illustrates the development of the Nordic R&D projects in the last
17 years.
Figure 2. Development of R&D in welded structures in Nordic projects.
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
The Nordic project QFAB - Quality and Cost of Fabricated Advanced Welded Structures was
completed in 2006. An international symposium was held in 2007 for the presentation of the
results from the QFAB-project [8]. The primary aim of this project was to reduce production
costs by fabricating components using high speed/deposition rate welding processes without
producing components which will be more sensitive to failure by fatigue. To achieve the
technical objectives, the programme of work has been divided into nine work packages (WP).
WP1 - High-speed Welding Process Development and Defect Detection – focuses on the
economic benefits of high productivity processes, innovation in the form of novel arc welding
techniques, consumables and increase in welding speed and deposition rate. For fatigue
sensitive components, it is equally important that an acceptable weld bead profile is produced
without forming toe defects especially cold laps. The welding processes investigated were be
based on the tandem arc process which uses two wires instead of the conventional single wire
process to achieve high welding speeds and high deposition rates. WP5 - Residual Stress
Prediction – gives attention to welding residual stresses and their effect on fatigue. Design
rules are, in most cases, conservative. Especially for fatigue assessment of the weld root better
knowledge about the residual stress field may improve the accuracy of fatigue life predictions.
In former Nordic R&D projects, many of the case studies were related to weld root problems
and the actual residual stress distribution was known in only a few cases. In this work package
the aim was to develop welding simulation procedures and also predict and measure residual
stress distributions. Effects of the welding residual stress will be incorporated in the fatigue
assessments of analyzed welds by local approaches i.e. LEFM. The research within this
doctoral thesis was continued in the Swedish R&D-project LOST - Light Optimized Welded
Structures.
Research aim
The research work in this doctoral thesis aims to increase the accuracy of fatigue design of
welded steel structures. More specifically, the research question(s) and issues addressed are:
ƒ
Establishing a link between weld quality and fatigue life.
ƒ
Improved engineering methods for prediction of welding residual stresses.
ƒ
Incorporation of residual stresses into fatigue design methods.
Research approach
The research can be divided in four different items:
1. Fatigue assessment – testing and fatigue life prediction.
Manufacture of cruciform joint test pieces for fatigue testing, in order to compare the
performance of different welding methods with respect to fatigue resistance. Manufacture of
tubular joint test pieces in order to study torsion fatigue and Mode III crack growth. Finally,
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
evaluation of their fatigue lives using LEFM, FEA and crack propagation analysis and
comparison with the experimental result.
2. Weld defect detection and characterisation.
After the fatigue testing, the weld defects at the fatigue starting points were studied (root and
toe defects). These were also characterised as cold laps, overlaps, weld spatter etc.
3. Welding simulations – Residual stress and deformation.
Finite element modelling and simulation of the welding process, 2-and 3 dimensional, was
carried out using sequentially coupled non-linear thermal-mechanical analysis with
temperature-dependent material properties and plasticity. Also, redistribution of the residual
stress due to service cyclic loading and crack growth from weld defects was studied.
4. Residual stress measurements.
Residual stress measurements were carried out using the X-ray diffraction method saw cutting
with stain gauges and the hole-drilling strain gauge method. These were made in order to
study the actual residual stress in the vicinity of the weld joint and to verify the finite element
calculations.
FATIGUE LIFE ASSESSMENT
There are several methods for fatigue life assessment which are frequently used in connection
with welded structures and components. They can be divided into global approaches; Nominal
and Structural Method, and local approaches; Effective Notch Stress Method and LEFM
(Linear Elastic Fracture Mechanics). These are outlined in [9] and a detailed procedure for
implementing them is described. Assessment and comparison of these methods can be found
in Martinsson [2] and Pettersson [3]. In this thesis the effective notch stress method is used to
evaluate the fatigue test results in Paper B. The LEFM method is used to evaluate the fatigue
test results in Paper C and to predict the fatigue life in Paper B, C and E. Figure 3 shows a
schematic illustration of work effort required for fatigue analysis of welded joints for the
different assessment methods.
Figure 3. Schematic illustration of the relation between accuracy,
complexity and work effort required for fatigue analysis of welded joints.
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Fatigue testing
For welded joints the fatigue life under service loading is normally predicted on the basis of
laboratory tests data obtained using simpler loading. Such constant amplitude tests are run at
various applied stress amplitude levels, and the results are plotted as the stress range
(normally nominal stress) versus cycles to failure to give an S-N-curve. The fatigue life is
assumed to be consumed and the fatigue test is stopped when a large enough, approximately
half the plate thickness, visible crack is detected. For fatigue cracks propagating from the
weld toe (Papers A and B) the failure criterion is straight forward, but for fatigue failure from
the weld root (Papers C and E) it is more difficult to set a failure criterion since the crack is
not visible until it has propagated to the surface.
Papers A and B describe the fatigue tests on the non-load carrying cruciform welded joints,
which were tested with R=0 in tension, while Paper C presents those performed on the
welded tubular joints, in this case tested in reversed torsion (R=-1) and in some cases also
with internal static pressure. Figure 4 shows the fatigue test rigs.
a)
b)
Figure 4. Testing rigs used for the fatigue testing: a) Paper A and B; b) Paper C.
Effective notch stress method
IIW (International Institute of Welding) introduced the effective notch stress method in the
recommendations for fatigue design of welded structures and components in 1996.
Hobbacher [9] states that the effective notch stress is the stress at the root of a notch, e.g. at
the weld toe radius, obtained assuming linear-elastic material. To take into account variations
in the weld shape, the real weld contour is replaced by an effective notch root radius of 1 mm.
This fictitious notch radius has to be added to the actual notch radius, which is usually
assumed to be zero in a conservative way (worst case assumption). Therefore it is
recommended to assume generally Reffective=1 mm for design purposes, see figure 5. The
attractiveness of this fatigue assessment method for design purposes is high [10] and can it be
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
performed quickly. The effective notch stress method with a fictitious radius of 1 mm [9] is
valid for thicknesses above 5 mm. For thin structures a fictitious notch radius of 0.05 mm is
recommended.
Weld toe
Weld root
Figure 5. Fictitious rounding of weld toes and roots
utilized in the effective notch stress method [9].
For low strength steel welded joints under stress ratio R=0 and assuming a fictitious radius of
1 mm as the “worst case” assumption, Radaj [11] recommended a corresponding design
fatigue strength of 240 MPa (97.7% survival probability) at 2 million cycles on an S-N curve
of the form S3.N = constant.
Hobbacher [9] recommends a characteristic fatigue strength at 2 million cycles; FAT 225
(95% survival probability) for a fictitious radius of 1 mm with stress ratio R = 0.5 accounting
for residual stress effects. Olivier et al. [12-13] investigated the effect of stress ratio and
scatter on the notch stress approach. A problem arises for larger notch radii which have not
been verified by the aforementioned fatigue tests. For a ground welded joint having a small
stress concentration, the fatigue class is certainly far below FAT 225. Mild notches may occur
at weld toes with relatively large toe radius, small flank angle and/or for small plate
thicknesses. Fricke [14] recommends a reduction of the IIW effective notch fatigue class to
FAT 200 for notch radii larger than 1 mm.
In Fricke [15] extensive FE analysis was carried out using the effective notch stress method
outlined in the fatigue design recommendations by the IIW [9]. Several welded components
were analyzed for fatigue failure both from the weld toe and weld root with reasonable
accuracy. The analysis showed among other things significant differences between the
assumption of a keyhole and oval notch shape to represent the weld root. The problem with
modelling the notch at the root was also investigated by Pettersson [3]. Fricke [15] also gave
several recommendations for finite element modelling in the effective notch stress method,
including element size at the notch area and sub modelling.
In Paper B extensive fatigue testing of non-load carrying cruciform joints welded with
different methods were carried out. All the specimens failed by toe cracking. The test results
were evaluated according to the effective notch stress method outlined in the IIW
recommendations by Hobbacher [9]. Figure 6 shows the results from the evaluation together
with the IIW characteristic fatigue strength curve, FAT 225 MPa, for the effective notch stress
method. The fatigue test results are widely scattered, and the characteristic fatigue strength
they imply is below the FAT 225 curve.
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
1000
Effective Notch Stress Range [MPa]
P 50% = 272 MPa
Reffective = 1 mm (Kt=2.5)
Rmean = 1.4 mm
R (strd.deviation) = 1.2 mm
θmean = 46˚
θ(strd.deviation) = 6.8˚
Log Cmean (m=3) = 13.6
Log C (strd. deviation) = 0.23
Nr of strd. Deviations = 1.92
FATPf50% = 272 MPa
FATPf5% =194 MPa
R=0
failures
run outs
FAT 225
(m=3)
mean curve( m=3)
char. curve (m=3)
P 5% = 194 MPa
FAT 225 MPa
100
1,E+04
1,E+05
Cycles
1,E+06
1,E+07
Figure 6. Evaluation of fatigue test results using the IIW [9] effective notch stress method in Paper B.
The most important geometrical parameters of welds are the notch radius (weld toe and root),
the flank angle at the weld toe and the depth of penetration at the weld root. These depend on
welding process, weld filler and weld quality as has been illustrated in Papers A and B.
Jakubovski and Valteris [16] have shown that the statistical distribution of fatigue strength
can be attributed to the statistical distribution of the geometric parameters i.e.; the weld toe
radius and angle.
The weld toe radius and angle were measured for all the specimens in Paper B. Figure 7
shows the measured toe radii and angles together with the Kt = 2.5 (stress concentration
factor) line. This line defines a constant Kt for a combination of toe radius and angle. The
local weld geometry shows a large scatter even for the post weld treated (shot peened) welds.
Hence the large scatter in the local weld geometry will contribute to a large scatter in the
fatigue test results.
6
Toe radius R (m m )
5
θ
Aswelded
Kt=2.5
R
4
3
Mean value /
Strd. deviation
As welded
Radius
(mm)
1.4 / 1.2
Angle
(˚)
46 / 6.8
Shot peneed
0.9 / 0.8
37 / 6.8
2
1
0
20
30
40
50
60
70
Toe Angle θ
Figure 7. Local weld geometry; measured weld toe radii and angles in Paper B.
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
When employing the effective notch stress method for fatigue life assessment a smooth
idealized notch radius is assumed (minimum 1 mm) at the weld toe without any weld defects.
However, real welded joints do contain imperfections at the weld toe, such as undercut and, of
specific relevance in the present study, cold lap lack of fusion defects. These weld defects can
have a significant effect on the fatigue behaviour of the welded joint and reduce the fatigue
strength.
In Papers A and B weld defects in form of line and spatter induced cold laps were found in
the fatigue tested specimens. In most cases, fatigue cracks initiated at and grew from such
defects. Figure 8 shows the effect of weld toe defects (line cold lap) on the fatigue strength for
non-load carrying cruciform joints with a effective weld toe radius of 1 mm and a weld toe
angle of 45˚ (Kt = 2.5). Defect types and the scatter associated with the defects found in the
welded specimens in Papers A and B are discussed later.
Non-load carrying fillet weld
weld defect; line cold lap
FAT80 (nominal stress) - IIW
1
0,5
K t = 2.5
Toe radius = 1 mm
Toe angle = 45˚
FAT160 - IIW
Weld defect size (mm)
1,5
0
50
100
150
200
250
FAT Pf 5% (MPa)
Figure 8. Effect of weld a defect on the fatigue strength.
Linear Elastic Fracture Mechanics (LEFM)
LEFM is a method used to predict the behaviour of cracks in solids subjected to fatigue
loading. For welded structures LEFM is applied with the assumption of an existing of initial
crack(s). Three basic types of load case, or mode, are considered for crack propagation
analysis, see figure 9. The most common case is mode I, which is the crack opening mode.
The stress intensity factors (SIF) KI, KII and KIII in respective mode define the singular
stresses near the crack tip.
Figure 9. The three different modes in fracture mechanics.
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Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
In classical fatigue theory no distinction is made between the crack initiation and propagation
phases. The fracture mechanic approach to fatigue problems is concerned with the crack
propagation phase. The stress intensity factor is a suitable measure of the crack tip conditions.
The SIF can be calculated as a function of the load, crack size and the overall geometry of the
structure. For welds it is assumed that there exists a crack in the critical location. The sizes of
such pre-existing cracks vary from 0.05-1 mm [17-19] and can appear as e.g. cold laps,
overlaps, inclusions, gas blow, and undercut. For the weld root, the crack size varies from
zero to several mm from slag or lack of fusion [3]. In fillet welds without any joint
preparation there is a “design crack” with a size equal to the material thickness which in
combination with some lack of fusion may be even larger.
In Paper B weld toe failures are assessed using the measured defect sizes as initial cracks in
LEFM fatigue life prediction. In Papers C and E the root failures were analysed using LEFM
assuming the lack of fusion as the initial crack.
Figure 10 shows the crack growth per cycle as a function of the stress intensity factor range
(ΔK = Kmax - Kmin). The limit below which there is no crack growth is a material dependent
parameter called the threshold value for fatigue growth (∆Kth) which also depends also on the
local R-value (KImin/KImax). Region I is the threshold region, region II is the stable crack
growth region and region III is the unstable crack growth region. For large values of ∆KI (> 610 MPa√m) the crack growth rate versus SIF range is linear in a log-log-diagram and is
defined by the Paris law in equation 1. Only crack growth in region II is considered in this
thesis.
da
m
= C ⋅ (ΔK I , II , II )
dN
(1)
The finite element method is nowadays an established technique for LEFM evaluation of
cracks. The quarter point elements (QPE) were developed by Barsoum [20] and Henshell et
al. [21]. The QPE is used for the finite element modelling of the crack tip in order to capture
the singularity. This method is widely used and is implemented in several commercial FE
softwares, i.e. ANSYS [22]. There are two methods using the QPE technique to calculate the
crack tip displacements for evaluation of crack tip parameters (e.g. SIF) displacement
correlation technique (DCT) which uses all four nodes on the crack faces and the quarter point
displacement (QPD) which uses the two nodes closest to the crack tip [23].
Several 2D finite element software’s specifically developed for fracture mechanics fatigue
crack growth (FCG) analyses are available e.g. Franc2D [24-25]. An alternative to FE based
fracture mechanics codes is the weight function solutions for LEFM implemented in software
such as Afgrow [26], used for damage tolerance calculations and have the ability to include
the effect of residual stress on FCG. However, this weight function technique is outdated and
considers only a few elementary case studies. The initial residual stresses are redistributed
during fatigue crack growth and this effect is incorporated in the developed subroutine for
LEFM fatigue life assessment in Paper E.
Today there are a few finite element/boundary codes available that can carry out 3D fatigue
crack growth analysis e.g. Beasy [27] and Franc3D [28]. In Paper B a 3D FCG analysis is
carried out on spatter-induced cold laps at the weld toe using the Franc3D software.
Within the Nordic fatigue R&D projects several simulation tools have been developed for
automatic 3D fatigue crack growth analysis. Martinsson [2] developed a 3D FCG routine in
ANSYS in order to carry out fatigue assessments of weld defects at the weld toe or root.
Björkblad [29] has recently developed a state-of-the-art simulation tool (Castana) for 3D FCG
analysis of cracks emanating from various surface and internal casting defects.
9
Fast fracture
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
I
III
m
1
Threshold
log da/dN
II
ΔKth
log ΔK
ΔKc
Figure 10. Principal features of the crack growth law.
Mixed mode crack growth
Mixed mode crack growth may occur in structures. Several studies within the Nordic projects
have reported mixed mode failure in their applications [2-4]. Mixed mode means that there
exists an interaction between the different fracture modes. Pure mode III crack growth rates
are typically one or two orders of magnitude smaller than mode I crack growth rate. For the
case with only torsion load the crack growth rate in mode III is dependent on the applied
torque range. Higher ranges give higher crack growth rates. Figure 11 shows crack growth
rates for mode I and mode III [30]. Mode I growth is dominant when the stress intensity factor
is low while mode III growth dominates in the high stress intensity factor region. This
indicates that mode III could be neglected under mixed mode conditions in the initial phase of
crack propagation, Byggnevi [4]. The mode I/III crack growth data are much more
complicated and very dependent on the ratio of modes I and III. Tschegg [31-32] investigated
the influence of a superimposed static mode I on mode III fatigue crack growth. The test
showed that the crack growth rate was not influenced by small axial loads (a resulting KI
value of 0 – 3 MPa√m). For higher values of KI, 4 – 9 MPa√m, the mode III crack growth rate
was increased. The increased mode I led to reduced sliding crack closure influence (friction,
abrasion and mutual support between the fracture surfaces). Mode II and III fracture surfaces
are irregularly shaped, and friction forces play a major role in determining the crack growth
rate. One of the consequences of these friction forces is that the crack growth rate is not
uniquely correlated to the applied stress intensity factor but is crack-length dependent, [33].
In Paper C the welded tube-to-plates were also exposed to cyclic mode III torsion, with an
internal pressure static mode I loading to reduce the friction and enable crack growth from the
weld root. It was shown that mode III crack growth is sensitive to the magnitude of the static
mode I loading. In Tjernberg et al. [34] and Olsson et al. [35] crack growth data for Paris’ law
based on fatigue loaded welded specimens failing from the root side was proposed. The slope
m was set to 5 and the C value between 0.7E-14 and 1.7E-14 (units in MPa and m).
10
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
da/dN [m/cycle]
KI
e
KIII
ΔK [MPa√m]
Figure 11. Crack growth rates for mode I and III, [4].
Several 2D linear-elastic Mode I/II interaction criteria are available for numerical
computation of incremental crack growth e.g. the maximum circumferential stress criterion
(σθmax). The σθmax theory states that the crack will move in the direction of maximum
circumferential stress around the crack tip when a critical value of the stress is reached [36].
Other Mode I/II interaction theories are the maximum potential energy release rate (Gθmax)
and the minimum strain energy density (Uθmin) [37-38]. Bittencourt et al. [25] showed that if
the crack orientation is allowed to change in automatic crack growth simulations, the three
interaction theories provide basically the same results.
In Paper B 2D and 3D LEFM FCG analysis of cold laps (weld defects parallel to the parent
material) was carried out. For the line cold lap (2D) the crack experiences mode II and will
rapidly deflect and the crack will continue to grow in mode I. For crack growth from a
spatter-induced cold lap (3D) in Paper B the crack is experiences mode I, II and III at the
initial stage. However, immediately after the first crack growth increment mode I crack
growth is dominating. Figure 12 shows the stress intensity factors (I, II and III) for the FCG
analysis of a spatter-induced cold lap.
60
90˚
K_I (0˚, 180˚)
ΔKI,II,III (MPa√m)
50
K_II (0˚, 180˚)
K_III (0˚, 180˚)
40
K_I (90˚)
K_II (90˚)
30
180˚
0˚
crack front
K_III (90˚)
20
10
0
0
2
4
6
8
crack growth step
10
Figure 12. Mixed mode SIF for spatter induced cold lap in Paper B.
11
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Several mixed mode crack growth models have been developed for fatigue life prediction for
cracks under mixed mode loading conditions. All the models are based on fatigue life
assessment by the Paris law replacing ΔKI with an equivalent stress intensity factor (ΔKeq)
which includes mode I, II and III i.e. the crack tip displacement model in equation 2 proposed
by Tanaka [39];
ΔK eq
4
⎡ 4
⎤
8 ⋅ ΔK III
4
= ⎢ΔK I + 8 ⋅ ΔK II +
⎥
(1 − ν ) ⎦
⎣
0.25
(2)
A mixed mode FCG analysis was carried out on a seam weld root in a Volvo Wheel Loader
lifting frame, see figure 13. Paris law was used in order to calculate the life by using ΔKeq for
mixed mode crack growth. The crack growth constants, C and m, recommended by the IIW
[9] were used. ΔKeq was calculated using different mixed mode crack growth models found in
the literature [39-42]. Figure 13b shows the resulting fatigue lives calculated using the
different ΔKeq formulae. The different models show large differences in calculated fatigue
life, e.g. a four times longer life for the formula according to Richards [42] compared to that
from Tanaka [39]. However, it is a challenge to develop a unique ΔKeq formula that covers all
possible mixed mode crack growth.
25000
Crack tip displacement [39]
Strain energy density [40-41]
Pook
20000
Irwin
Analyzed weld
Life in hours (h)
Strain energy relase rate [40-41]
Richard [42]
15000
10000
5000
0
a)
b)
Figure 13. Mixed mode crack growth: a) analyzed Volvo lifting frame; b) predicted fatigue life using different
mixed mode crack growth models, ΔKeq.
12
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
WELD DEFECTS
It is known that the local weld geometry, toe angle, toe radius, undercuts and cracks strongly
influence the fatigue strength. The local geometry affects the local stress concentration and
together with defects of different type fatigue cracks may form during cycle loading and lead
to large scatter in fatigue life. This has been observed in Papers A and B. Cold laps, sharp toe
radii and large design root cracks are typical fatigue starting points. Toe radius is the
geometrical parameter that most significantly affects the stress concentration and, hence,
fatigue life for fatigue failure from the toe. Figure 14 shows the four most common types of
weld defect. In almost all specimens within the batches analyzed in Paper A and B weld
defects in the form of cold laps were found.
a)
b)
c)
d)
Figure 14. Different fatigue sensitive weld defects: a) cold lap at weld toe (~0.2 mm); b) crack at weld toe,
undercut, (~0.2 mm); c) root defects and initial crack; d) interbead crack.
High quality welds with smooth weld geometries and small defects have higher fatigue
strengths. Figure 15a shows a simulation of crack propagation from a 0.3 mm cold lap defect
to the half-plate thickness in a non-load carrying cruciform welded joint analyzed in Paper B.
Due to the horizontal shape of the cold lap mixed mode crack growth will occur and result in
a higher ΔKII than the line crack during the first crack growth increment. ΔKII decreases when
the crack has propagated to approximately 0.3 mm and the cold lap continues to propagate as
a straight-fronted edge crack, and after 1 mm of crack growth 70 % of the life is consumed.
Figure 15b shows the stress intensity factor (SIF) for different cold lap defect sizes at different
growth increments. The SIF reaches a fairly constant level for all cold lap sizes after ~ 0.3
mm of growth and the cold lap will grow in the shape of a line crack in the subsequent growth
increments. This illustrates the effect of the weld defect, size and type, on the fatigue life.
13
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
10
failure
8
0
7
0.1
6
0.2
0.3
5
0.4
4
0.5
3
1
Growth increment (mm)
Δ K MPa√m
9
2
1
Line cold lap
0
Spatter induced
cold lap
0
0,2
0,4
0,6
0,8
1
1,2
Cold lap (mm)
a)
b)
Figure 15. Cold lap weld defect: a) crack growth simulation; b) stress intensity factor as function of
defect size and growth increment.
During the welding process a common phenomenon causing cold laps is spatter. A small drop
of melted material falls ahead of the welding point and sticks to the surface of the plate.
However, the drop is too small to be fully fused to the cold plate. Somewhat later, when the
weld runs over, the small drop is partly melted and a small crack is left in the toe area of the
weld. The form of the cold lap is thus a semi elliptical defect (depth to surface length ratio a/c
= 1) rather than the usual assumption of a line cold lap (a/c = 0). This latter type is often
found when the welding speed is too high or if the plate is dirty or rusty. Similar behaviour
can also be expected if there is a dirty spot at the material surface which may cause lack of
fusion when welding, and hence a defect.
In Paper B the predicted fatigue lives for a line cold lap defect (2D LEFM) and a spatter
induced cold lap defect (3D LEFM) are compared. The spatter-induced cold lap had 2.7 times
longer life compared to the line cold lap.
6
Crack growth (mm)
5
3D LEFM analysis (A/C=1)
2D LEFM analysis (A/C=0)
4
Non load carrying cruciform joint:
3
Cold lap = 0.2 mm
Thickness = 12 mm
Throat thickness = 5 mm
Toe angle = 70˚
Toe radius = 0 mm
Nominal stress = 100 MPa
2
1
0
0,E+00
5,E+05
1,E+06
2,E+06
2,E+06
3,E+06
3,E+06
LIFE (cycles)
Figure 16. Comparison of fatigue life for a line cold lap (a/c=0) and a spatter- induced cold lap (a/c = 1).
14
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Lopez and Korsgren [19] studied the presence of weld toe defects by plastic replicas and
microscopy inspection. They concluded that the welding induced defects consist mainly of
cold laps (80 %). Basic coated electrodes and basic flux cored wire gave deeper cold laps than
rutile flux cored wire and solid electrodes. The solid wire showed the most cold laps (> 1.5
mm). In Paper B a comparison between solid wires, flux cored, hybrid laser/MAG and
tandem arc MAG-processes was made regarding e.g. defect characterisation and fatigue
strength. The flux cored wire showed almost no defects (>> 0.03 mm), low stress
concentration and higher fatigue strength (~30%) compared with the solid wire welded
cruciform joints. This was also in good agreement with Lopez and Korsgren [19] and
Samuelsson [17].
The weld defects vary in size (approximately 0.03-1.5 mm) and in many cases are very
difficult to detect and characterise. In Lopez and Korsgren [19] the mean value of cold laps
varied from 0.047 – 0.27 mm depending on weld method. Previous methods for detecting
defects in welds has been cutting, polishing etching the specimen and examination in a
microscope. Also plastic replicas of the weld can be examined. This is done in Paper A. In
Holst et al. [43] a new method was developed to make the detection of defects in fillet weld
toes easier. The method is totally destructive where the test specimen is put in a hydraulic
press with a successively increased force till it brakes at the toe and then the defects are
studied in a stereoscope.
Within the Nordic project QFAB a new method, Farajian and Järvstråt [44], was developed
for detection and characterisation of weld defects in weld bead on plate. In this method weld
toes defects could be opened and investigated longitudinally. The specimens are cooled down
in liquid nitrogen, hit by a pendulum of an impact test device and then studied in a
microscope. The method is under development for detecting weld defects in fillet welds.
HEAT EFFECTS OF WELDING
The heat input supplied by the welding arc produces complex thermal cycles in the weldment
and these, in turn, cause changes in the microstructure of the heat-affected zone. The high
heat concentration is necessary because metallic materials diffuse the heat and cause transient,
inhomogeneous thermal stress and metal movement. The transient temperature field causes
thermal expansion, stress and strain that usually plastically deforms in the weld
neighbourhood and results in residual stress and strain that remains when the weldment cools
and the structure is distorted from its original shape. The most significant early studies on
analysis of the pattern of heat conduction in solids and analytical solutions for a moving heat
source were performed by Rosenthal [45]. Finite element simulations in order to predict
temperature fields, residual stresses and deformation due to welding in 2- and 3 dimension is
used more frequently nowadays [46-48]. Accurate and reliable residual stress predictions are
essential for structural integrity and fatigue assessment of components containing residual
stresses. However, finite element simulation of residual stresses due to welding involves in
general many phenomena e.g. non-linear temperature dependent material behaviour, 3D
nature of the weld pool and the welding processes and microstructural phase transformation.
Despite the simplification by excluding various effects, welding simulations are still CPU
time-demanding and complex. Hence, simplified welding simulation procedures are required
in order to reduce the complexity and thus maintain the accuracy of the residual stress
predictions. Figure 16 shows the simulation scheme and coupling fields in welding analysis
used in this thesis.
15
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Heat source
Heat Flow:
Mechanics:
Temperature field, HAZ,
and Fusion Zone
Residual stress field and
Distortion
Thermal
Material
Properties
Mechanical
Material
Properties
Figure 16. Simulation scheme in welding analysis.
Thermal modelling and simulation
The principal parameter of the heat source for the temperature field is the heat input into the
welding spot (heat quantity) Q [J], and consequently the heat flow q [J/s], i.e. in arc welding
the total output is the product of current I [A] and voltage U [V] in the case of direct current.
Consequently the heat flow is expressed as
Q = η hUI .
(3)
Heat losses e.g. radiation in welding is taken into account by the heat efficiency ηh which
depends on the welding process. In computing heat flow in actual weldments, it is often
necessary to take into consideration the effect of the size and shape of the heat source. This
can be done by treating the heat source as being distributed over an area [J/mm2] or a volume
[J/mm3], see Radaj [49]. Fourier’s law of heat conduction describes the heat flow propagation.
The states that the heat flow density q [J/mm2] is proportional to the negative temperature
gradient ∂T/∂t [°C /mm] by the equation
q = −λ
∂T
,
∂n
(4)
where λ [J/(mm°C)] denotes the thermal conductivity and T [°C] the temperature.
In terms of FE modelling of the welding heat source the heat generation Q can be expressed
as
Q = Qsurface + Qvolume + Q filler ,
(5)
where Qsurface is the surface heat flux, Qvolume is the volume heat flux and Qfiller is the heat
generated by the addition of the weld filler material at a prescribed initial temperature.
One objective with heat transfer analysis in welding is to determine the temperature fields in
an object resulting from conditions imposed on its boundaries, the quantity sought being the
temperature distribution. Figure 17 shows a schematic model of the temperature distribution
when a surface weld bead is being deposited at a speed v. The coloured bands represent
isothermal areas.
16
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Welding direction
y
z
x
Weld path
Figure 17. Schematic of the welding thermal model.
The fundamental equation of heat conduction for a homogeneous and isotropic continuum
with temperature-independent material characteristic values in a solid is [49]
λ ⎛ ∂ 2T ∂ 2T ∂ 2T ⎞ 1 ∂Qv
∂T
⎜
⎟+
=
+
+
.
∂t cρ ⎜⎝ ∂x 2 ∂y 2 ∂z 2 ⎟⎠ cρ ∂t
(6)
The parameter c [J/g°C] is the mass-specific heat capacity and the parameter ρ [g/mm3] is the
density. The parameter Qv [J/mm3] is the rate of temperature change due to heat generated per
volume, and λ/cρ is the material-and temperature-dependent coefficient of thermal diffusivity
κ [mm2/s]. The temperature field in equation (3) could either be stationary (i.e. steady-state),
the temperature field is then time-constant at all points, i.e. ∂T/∂t = 0, or non-stationary (i.e.
transient), then the temperature field is determined on the material side by the thermal
diffusivity. The characteristic of heat flow during arc welding is that the heat source moves at
constant speed on the surface of the work piece, and that the size of the heat source is small
compared to the size of the work piece.
In Paper C a 2D axisymmetric welding analysis was carried out on a multi-pass welded
tubular joint. The heat input model used was with constant distribution of surface and volume
flux. The predictions showed reasonably good agreement with micro samples of the fusion
zone and the heat affected zone (HAZ), see figure 18b.
In Paper E development of a 2D FE subroutine was carried out for welding simulation
procedure in order to predict the residual stresses due to the welding process. The heat input
consisted of a constant volume flux and prescribed initial activation temperature of 1500 ˚C
for the weld filler. Also guidelines and recommendations for modelling of a 2D simplified
welding heat source are outlined in Paper E. Figure 18a shows the temperature predicted
using the heat source in Paper E compared to measurements. Figure 18c shows the fusion
zone and HAZ predictions compared with micro samples for a verification example found in
the literature [50].
17
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
800
Temperature (˚C)
FEM 15 mm
FEM 21 mm
600
FEM 27 mm
Measurement 15 mm
Measurement 21 mm
400
Measurement 27 mm
200
0
0
200
400
600
800
1000
Seconds
a)
b)
c)
Figure 18. Thermal modelling and simulation: a) temperature as function of time in Paper E; b)
fusion zone and HAZ in Paper C; c) and Paper E.
In Paper D a 3D welding simulation was carried out on the multi-pass welded tubular joint
analyzed in Paper C. The heat source consisted of a constant distribution of surface flux and
volume flux. To simulate the 3D moving heat source it was necessary to model the heat
source during each time increment. In this work the moving heat source is simplified by
assuming that the welding arc stayed at an element with constant qs and qv and then moved to
the next element at the end of the load step as the welding was finished. Figure 19a shows the
temperature predictions in Paper D using the simplified constant distributed 3D moving heat
source in figure 19b on the 2 pass butt welded verification case [50], see figure 19c.
In Paper F 3D and 2D welding simulations were carried out on T-fillet welds. To model the
transient heat input and addition of filler material, the built-in welding functionality of
MSC.Marc [51] was used. The user defines a weld path, weld speed and the applied effect.
The shape of the heat input is the double ellipsoid proposed by Goldak [52] and the user
should supply the width, height, forward and rear length of the heat source. Figure 20 shows
the results from the thermal analysis in Paper F when using the Goldak heat source.
18
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Temperature Measurement 12mm
800
Temperature Measurement 15mm
T em epra ture [°C]
700
Temperature Measurement 21mm
Temperature Measurement 27mm
600
Butt Weld - 2 Pass 12.7mm
500
Butt Weld - 2 Pass 14.4mm
400
Butt Weld - 2 Pass 26.5mm
Butt Weld - 2 Pass 21.3mm
300
200
100
0
0
200
400
Time [s]
600
800
1 000
a)
b)
c)
Figure 19. 3D welding simulations in Paper D: a) temperature predictions; b) time simplification of moving
heat source in 3D welding simulation; c) 2 pass butt weld.
a)
b)
Figure 20. 3D welding simulation of T-fillet welds in Paper F: a) weld penetration profile of the fused zone; b)
3D welding simulation with Goldak moving heat source.
19
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Welding residual stresses
The residual stresses caused by inhomogeneous temperatures are termed thermal stresses.
These elastic thermal stresses disappear after removing the inhomogeneous temperatures
which cause them. Where major differences in temperature exist, the thermal stresses give rise
to plastic deformation and, after removal of the temperature differences and complete cooling,
residual stresses remain. During the welding process the weld area is heated up significantly
relative to the surrounding area and fused locally. The material expands as a result of being
heated. The heat expansion is restrained by the surrounding colder area, which gives rise to
elastic thermal stresses. The thermal stresses partly exceed the yield limit, which is lowered at
elevated temperatures. Consequently the weld area is plastically hot-compressed and, after
cooling down, it thus displays tensile residual stresses, and the surrounding area compressive
residual stresses.
The fundamental equations in the thermo-elastic-plastic analysis of welds are:
1. A change in temperature, ΔT, causes a volumetric strain αΔT.
2. At each point, the strain increment can be defined as
dε ijTot = dε ijElastic + dε ijPlastic + dε ijThermal
(7)
−1
dε ijElastic = Dijkl
dσ kl
(8)
dε ijPlastic = λ
∂F
∂σ ij
(9)
dε ijThermal = αΔTδ ij
(10)
3. The material obeys isotropic/kinematic hardening hypothesis.
4. The material properties e.g. the Young’s modulus and the yield strength are assumed
to be temperature dependent.
Finite element modelling of welding residual stresses
Several authors have summarized information on the application of the finite element method
to the simulation of residual stresses during and after welding, e.g. Lindgren [46-48].
Lindgren [46-47] discussed the approach for accounting for thermo mechanical couplings and
corresponding computational strategies and the difficulties of the material modelling during
the welding process. In Lindgren [48], the increasing accuracy of welding simulations
depending on the modelling process and the opportunity to use larger models is outlined.
Radaj [53] discussed the intelligent solution and simplification using the finite element
method to simulate the complex welding process.
Small strain approximations are often used, although welding often leads to visible
deformation. Most analyses are performed in two steps. The thermal analysis is followed by
the mechanical analysis. In the case of completely decoupled calculation of the temperature
20
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
field and the residual stresses, the transient temperature field is first stored for all time steps
and is then used as an input data file for the residual stress calculations. Identical meshes and
time steps are mandatory when this method is used. The thermal and mechanical analysis can
be performed in a so-called staggered approach or simultaneously. The thermal dilatation,
which is the sum of the thermal expansion and the volume changes, is the driving force and
therefore each time step is usually started by solving for the temperatures, Tn+1, for time tn+1
and thereby using the geometry xn. The geometry is updated in the subsequent mechanical
analysis for this time step. Thus the updating of the geometry in the thermal analysis lags one
time step behind the mechanical analysis. Some analysis only takes the thermal expansion
into account, and there exist different approaches for including the effect of the
microstructure, that is the dependence of the material properties on the history of the thermo
mechanical process.
A coarse element mesh may be adequate. In the case of a more detailed analysis the chosen
mesh fineness may not have much influence on the accuracy of the calculation result. The
residual stress calculation requires a mesh refinement which is extended some distance from
the moving heat source. Such an adaptive mesh design oriented towards the result error may
considerably increase the efficiency of calculation. This advantage is even more important
when solving three dimensional problems.
In multi-pass welding different techniques exist in order to simplify the simulation. Analyzing
multi-pass welds as a series of single-pass welds is rigorous. Lumping successive passes
together is one way to reduce the computing time. Other simplification of the thermal and
mechanical process is to either merge some weld passes into larger welds or accounting for
some of the weld passes.
To be able to add the filler material, especially in multi-pass welding the technique of element
“birth and death” is suitable as described in the FE-software ANSYS [22]. All the elements
are defined in the model and born in the later stage of the analysis. Another approach by
Lindgren [46] is called quiet element where the whole structure is included in the
computational model. The elements, corresponding to non laid welds, should be given
material properties so that they do not affect the rest of the model. The elements are given
normal material properties at the start of the weld pass, and all the strains and stresses that
have been accumulated are removed.
In Papers C, D E and F extensive welding simulations were carried out on fillet welds, butt
welds and pipe welds focusing on accurate prediction of the residual stress after welding.
Material modelling
The thermal material characteristic values, expressed in equation (6) are in reality
temperature-dependent and not constant, the extent of the change varying according to the
kind of material. When the values of thermal properties are treated as variables that change
with the temperature, equation (6) becomes non-linear and the mathematical analysis becomes
extremely complex.
However, the greatest limitation in welding simulation is the problem of obtaining accurate
temperature-dependent material properties, especially at the higher temperature regime, T >
700 ˚C. These are at best reasonable approximations at present. The phase change during
heating and cooling of typical low-alloy steels also has an influence on accurate modelling of
the temperature field and the residual stress formation. It is evident that the material model
needs to represent the real material behaviour with sufficient accuracy.
21
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Material properties are divided in two groups; the temperature-dependent thermo-physical
properties (thermal conductivity, specific heat and density) used for the thermal analysis and
the temperature-dependent mechanical properties (thermal expansion, Young’s modulus,
Poisson’s ratio and yield stress) used for the coupled thermo-mechanical analysis. The metal
adjacent to a weld is exposed to rapid thermal cycles and often undergoes changes in
microstructure when it is subjected to elevated temperatures. These phase changes in turn
influence the thermal properties, making them time- as well as temperature-dependent. The
microstructure and hardness of the heat-affected zone depend upon the cooling rate, and this
rate is influenced by various factors including plate thickness, welding conditions, preheat, the
length of the weld, joint geometry, etc.
For typical carbon steels the yield stress is considerably reduced for temperatures above 800
˚C and naturally vanishes at the melting temperature. The high temperature range, T ≥
0.5*Tmelting, is of little significance for the formation of residual stresses because of the
disappearing yield limit in this range [49]. Therefore a cut-off temperature of Tcut-off =
0.5*Tmelting is used in the mechanical material model, i.e. if the temperature from the thermal
analysis is higher than Tcut-off then material properties are evaluated at the Tcut-off temperature.
The microstructure changes during heating (from ferrite and pearlite to austenite) and cooling
(from austenite to ferrite, pearlite, bainite and/or martensite) depend on temperature and
cooling rates. The evolution of microstructure can take the form e.g. of grain growth and
precipitation and is introduced by e.g. annealing processes and will have an influence on the
material properties.
The microstructure evolution during welding can be an important factor for the final
formation of the residual stresses. For example the austenite-martensite transformation will
create compressive residual stresses since the volume change to martensite is greater. The
closer the transformation temperature to room temperature, the more significant it is.
Nowadays it is possible to include the microstructural evolution in the FE welding analysis in
several FE codes, i.e. ANSYS [22] or ABAQUS [54]. Unfortunately, dealing with the lack of
high temperature material data relevant to microstructural dependency and the resulting
material properties in the heat affected zone is one of the most challenging tasks in
computational weld mechanics. However, this is out of the scope of this thesis.
RESIDUAL STRESS MEASUREMENT TECHNIQUES
Many techniques have been used for measuring residual stresses in metals. These can be
categorized in terms of the degree of damage to the metal they introduce: destructive, semidestructive or non-destructive methods. In the destructive and semi-destructive methods the
residual stresses are measured, by mean of stress-relaxation, by measuring the elastic-strain
release that takes place when a sample is sectioned, drilled or milled using electrical or
mechanical strain gauges. Resistance strain gauges, detachable extensometers and photo
elastic surface layers are mainly used for measurements in practice. Although they are
destructive or semi-destructive, the stress relaxation techniques provide reliable data and are
the most widely used and frequently referred to [53, 55-57]. Hoe-drilling is semi-destructive
in that the hole can be ground out or peened to prevent it acting as a source for fatigue crack
initiation. In crystalline structures, elastic strain can be determined non-destructively by
measuring the lattice parameter with X-ray diffraction. The most severe limitation is that it
can be applied non-destructively only on the surface, the surface strain can be determined
22
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
within a small area, to a depth of approximately 0.0025 mm, and slightly more by electro
polishing the surface, layer by layer [58].
In Papers A and B the X-ray diffraction technique was used in order to measure the residual
stresses in the vicinity of the weld toe (> 1 mm) where the fatigue failure occurred. In Paper
F the X-ray diffraction technique was used in order to validate the residual stress predictions.
In Paper D the hole drilling method was used to measure the residual stresses and validate the
FE predictions of the residual stresses.
RESIDUAL STRESS EFFECT ON FATIGUE OF WELDS
Together with stress concentrations and weld defects residual stress fields are one of the
determinant parameters controlling the fatigue strength of welded joints. The effects of
residual stresses may be either beneficial or detrimental, depending on the magnitude, sign
and distribution of the stress with respect to the load-induced stresses. Often the welding
residual stresses are detrimental tensile and of material yield strength magnitude. Speaking in
terms of fracture mechanics; the tensile residual stresses will reduce the fatigue life of the
structure by increasing the growth rate of the fatigue crack. Tensile stresses keep the crack
open and prevent crack closure while compressive residual stresses increase crack closure
thus decreasing the fatigue crack growth rate. When applying external loads to a welded
structure containing residual stresses, the residual stresses will be redistributed depending on
the level of the external stress.
Maddox [59] discussed the effects on the form of the standard S-N curve in the presence of
residual stress fields in welds. The slope of the S-N curve is the negative reciprocal of the
crack growth curve and the variation in R-value has the predominant effect of altering the
intercept but not the slope of the initial portion of the S-N curve. Figure 21 shows the effect of
compressive and tensile residual stresses on the S-N curve.
log S
Shift of sloped line
due to increase in R
Standard curve (R = 0)
Increasing R value
lowers fatigue limit
With Residual Stress
Slope = (-1/m)
log N
a)
log S
log S
log N
log N
b)
c)
Figure 21. S-N curves (schematic): a) general form and effect of changing R; b) tensile residual
stresses with remote tensile loading; c) compressive residual stresses with remote tensile loading.
23
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Bogren and Martinez [60] reported spectrum fatigue testing and residual stress measurements
on non-load carrying fillet welded test specimens, similar to those used by Maddox [59]. It
was found that the residual stresses in the as-welded condition were close to the yield stress of
the material. These residual stresses were found to relax very rapidly; within 8 % of the total
specimen life 50 % or more of the initial stresses were relaxed.
Martinez et al. [61] studied the relaxation of the residual stresses in as-welded and TIG-dressed
welded specimens. Several residual stress measurement techniques were used namely neutron
diffraction, X-ray diffraction and ultrasonic. The tensile stresses close to the weld toe were
decreased by 40 % as a result of the TIG-dressing operation. It was also shown by the relaxation
study that the application of a static load also relaxed the residual stresses.
Effect of external loading
Recently the author [62] has been carrying out fatigue studies on high strength steel joints
welded with low temperature transformation filler materials, LTT. This weld filler material
reduces the tensile residual stress in the weld (in some cases introduces compressive residual
stresses) due to the volume expansion in the martensite transformation at lower temperature.
The test specimen studied was a plate with out-of-plane fillet welded gussets and it was
welded with conventional and LTT weld fillers.
The fatigue testing was carried out in both constant amplitude (CA) and variable amplitude
(VA) loading. The residual stresses in the vicinity of the weld toes (> 1 mm) were measured
at different stages of the fatigue life in order to study the relaxation/redistribution of the
residual stress in LTT and conventional welded specimens in CA and VA, respectively.
Figure 22 shows the specimen studied. Figure 23 - 24 show the fatigue test results and
residual stress relaxation measurements. The LTT joints showed ~ 40 % increase in mean
fatigue strength compared to the conventional joints under CA loading and ~ 12 % increase
under VA loading, respectively. The improvement of the fatigue strength is less significant in
variable amplitude testing mainly due to the severe relaxation of the compressive residual
stresses in comparison to CA.
Figure 22. Out-of-plane gusset fillet welded specimens studied by the author [62].
24
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
400
Residual Stress (MPa)
σr [MPa]
1000
100
R-ratio = 0
Constant Amplitude
OK Tubrod 15.55 (LTT)
OK Autrod 12.51 (Conventional)
Mean curve (LTT)
Mean curve (Conventional)
IIW - FAT 71 [MPa]
10
1,E+04
300
200
100
0
-100
1,E+00
1,E+05
1,E+06
Cycles to failure
OK Autrod 12.51 - Conventional
OK Tubrod 15.55 - LTT
1,E+07
a)
1,E+02
1,E+04
1,E+06
Total numbers of cycles
b)
Figure 23. Effect of residual stress on fatigue: a) constant amplitude loading (CA); b) residual stress relaxation
due to CA loading.
1000
250
Residual Stress (MPa)
σeq [MPa]
200
100
Variable R-ratio
Variable Amplitude SP2
OK Tubrod 15.55 (LTT)
OK Autrod 12.51 (Conventional)
run out
Mean curve (LTT)
Mean curve (Conventional)
IIW - FAT 71 [MPa]
10
1,E+05
1,E+06
Cycles to failure
OK Autrod 12.51 - Conventional
OK Tubrod 15.55 - LTT
150
100
50
0
-50
-100
-150
1,E+00
1,E+07
a)
1,E+02
1,E+04
Total numbers of cycles
1,E+06
b)
Figure 24. Effect of residual stress on fatigue: a) variable amplitude spectrum loading (VA); b) residual stress
relaxation due to VA loading.
Effect on crack propagation
A quantitative assessment of the influence of the residual stresses on the LEFM fatigue life
prediction can be made by the principle of superposition based on the effective stress intensity
factor (Keff);
K eff = K applied + K residual
(5)
where Kapplied is the stress intensity factor due to the external applied cyclic loading and
Kresidual is the stress intensity factor due the residual stress. The stress intensity factor, Kresidual,
is usually obtained using weight function solutions [26] by applying the residual stress field
25
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
on the un-cracked component. However, these weight function techniques only gives
qualitative results, are outdate and are valid only for elementary cases. In order to achieve a
more accurate determination of Kresidual the residual stress distribution has to be known at each
stage of the crack growth [64].
Linear elastic fracture mechanics (LEFM) is frequently employed for fatigue assessment of
cracks and flaws in critical welds. The approaches available for the incorporation of residual
stresses into the LEFM fatigue life assessment are the mean stress method and the crack
closure method. The mean stress method changes the crack growth rate by an alteration of the
mean stress at the crack tip by an effective stress ratio Reff. The crack closure approach
assumes an open crack if Keff > 0 and a closed crack if Keff < 0.
In the SINTAP defect assessment procedure, [65], the superposition of stress-intensity factors
as shown above is employed and stress redistribution is accounted for using a correction
factor for Kres.
Michaleris et al. [66] presented a finite element methodology for the incorporation of residual
stress effects into fracture assessment. Following the welding simulation, interpolation is used
to transfer the computed residual stresses onto fine meshes for evaluation of fracture
mechanical parameters.
Additionally, the effect of the welding residual stress field on uncertainties in fatigue life
predictions was modelled by Josefson et al. [67].
Figure 25 shows an example of a cold expanded hole analyzed in Paper E where the effect of
the residual stress on the FCG was analyzed by FEM and LEFM. The procedure was then
applied to a welded steel structure.
A welded ship engine frame box at MAN B&W was analyzed in Paper E regarding residual
stresses and fatigue assessment from the weld root using LEFM. The welding simulation
procedure, residual stress mapping algorithm and LEFM subroutine developed in Paper E
were used for the analysis. Figure 26a shows the welded engine part analyzed while figure
26b shows the fatigue life predictions in comparison with fatigue testing.
26
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
0,5
8 mm
3 mm
σ residual stress / σ yield stress (-)
0.3 mm
0
-0,5
No crack
a=0.3 mm
a=3 mm
a= 8 mm
-1
Radial Residual Stress-σy
-1,5
0
1
r / R (-)
2
3
b)
a)
1,4
Analytical solution
2D FE (LEFM) - External load
1,1
Keff / σ yield stress(π R)2 (-)
2D FE (LEFM) - External + Residual stress
0,8
2D FE (LEFM) - Residual stress
2D FE (non cold worked), Pavier et al
0,5
2D FE (cold worked), Pavier et al
3D FE (mid thickness), Pavier et al
0,2
-0,1
-0,4
-0,7
0
0,5
1
r / R (-)
1,5
2
2,5
c)
Figure 25. Incorporation of residual stress effect on crack propagation in Paper E: a) FCG and LEFM analysis
of cold expanded hole in plate; b) redistribution of residual stress due to crack growth; c) effective stress
intensity factor, Keff.
100
Rig force (kN)
run outs
PWHT experiment
PWHT prediction
As Welded experiment
As Welded Prediction
WELDED ENGINE FRAME BOX
10
1,E+04
a)
1,E+05
Nf (Cycles)
1,E+06
1,E+07
b)
Figure 26. Analysis in Paper E: a) a two stroke diesel engine and welded frame box part analysed; b) fatigue
life predictions compared with fatigue testing.
27
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
DISCUSSION
Fatigue failure is still the dominating cause for breakdown of many welded structures in
construction and mining equipment, trains, ships, agricultural machinery, bridges and off
shore equipment. The fatigue design of these structures is governed by relatively conservative
fatigue design codes with no or small differentiation for different welding quality. The global
challenge of saving resources will promote the development of more light-weight and
effective structures for many types of machinery and installations.
Fatigue design of welded joints has been developed during the last 35 years. For many years
the only applicable method for the fatigue assessment of welded structures was the nominal
stress method or structural testing of welded parts or components. The major type of welding
process was manual arc welding. Nowadays more sophisticated fatigue assessment methods
have been developed e.g. notch stress method and LEFM and the welding processes are faster
and more productive e.g. robotized hybrid laser and arc welding methods. The trend moves
towards optimized welded structures regarding cost and weight and this requires high speed
welding method producing high quality welds and reliable analysis tools for fatigue life
assessment.
A draft of a weld quality rule based on local approach fatigue assessment methods, principally
the effective notch stress and LEFM methods, was developed within the Nordic R&D project
QFAB and is nowadays introduced as a weld class system corporate standard within Volvo
and in the international community e.g. IIW (International Institute of Welding). Within the
framework of this research work many aspects of fatigue design and manufacturing of welded
joints has been studied and introduced into the industrial environment.
It is clearly obvious that residual stress prediction and welding simulation in complex
structures is elusive and employs many researchers for compilation of residual stress
distribution for welded joints. Two of the important issues regarding welding simulation,
besides the complexity and high degree of nonlinearity of the problem, are the quantification
of net heat input supplied by the welding arc and the modelling of material behaviour at
higher temperatures. Both are important in order to have an accurate model in successful
welding simulations. As a general rule it can be stated that all production components and
structures contain residual stresses that influence fatigue strength.
Lack of penetration, lack of fusion or “design root cracks” in load-carrying fillet welded
joints, cold laps and undercuts will all act as initial cracks from which fatigue cracks can
propagate. If the applied stress intensity variation is high enough and the residual stress level
is in tension, then the fatigue life will be shorter than if the residual stress is in compression.
The computational effort needed to predict the residual stress field in complex welded structures
and their relaxation during spectrum fatigue loading will be the major future challenge in the
design of advanced fabricated structures.
Another aspect of the design of welded structures against fatigue is the introduction of post weld
improvement techniques which have evolved the international industrial and scientific
community. Significant progress has been made with weld improvement methods as
developed for fatigue design rules for improved welded joints within IIW Commission XIII
(Fatigue design of welded structures) [68]. The weld improvement techniques such as shot
peening, hammer peening [69] and Ultrasonic Impact Treatment (UIT) [70-71] focus on
improving the weld toe region by enlarging the toe radius, reducing the weld defects and
introducing compressive residual stresses. Improvement methods will also support the use of
higher strength materials, an alternative which traditionally has been limited by older design
rules. Following effective improvement of the weld toe failure tends to move to other locations
such as interbead toes and the weld root. The assessment of the weld root will be more important
in connection with weld improvement.
28
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
Finally, the challenge for the international scientific and industrial community, in order to be able
to design and manufacture lightweight, effective welded structures may be listed as
ƒ
Apply local design methods,
ƒ
Assess residual stress fields including relaxation,
ƒ
Apply weld improvement methods in highly stressed areas,
ƒ
Specify and monitor the needed weld quality needed,
ƒ
Improve weld inspection tools for general purpose.
CONCLUSIONS
ƒ
The majority of weld defects found in high speed single run MAG welding are cold
laps in connection with spatter. The existence of cold lap in connection with good
weld geometry reduces the weld quality. In the case of small toe radius the problem
with cold laps is overturned by high stress concentration factor.
ƒ
It is indicated that spatter-induced cold laps could be expected to have 2-3 times
longer fatigue life than line cold laps. Special effort is needed to avoid or eliminate
these kinds of flaws e.g. by an alternative welding process.
ƒ
Good fatigue strength and relatively small scatter was observed in the local weld
geometry and the fatigue life for specimens welded with manual FCAW. In some
specimens welded with FCAW almost no weld defects were found in form of cold
laps.
ƒ
The lack of weld root penetration has a major influence on the fatigue life of filletwelded joints. However, it has been shown that the weld root in e.g. the welded
tubular joints is under favourable compressive residual stresses and this will enhance
the fatigue life. The residual stress at the vicinity of the weld root is at present difficult
to measure hence welding simulations are used to quantify these.
ƒ
The developed welding simulation procedures with 2D models, simplified material
models and heat source tend to predict residual stresses with sufficient accuracy and
with 30-100 times shorter computational time compared with 3D welding analysis.
The developed welding simulation procedure is a valid tool for residual stress
prediction for further accurate fatigue life predictions.
ƒ
The fatigue life prediction for the diesel engine welded frame box using the developed
LEFM subroutine and the residual stress mapping procedure confirms the fatigue life
enhancement observed in the fatigue testing of the as-welded specimens compared
with the stress relived. This illustrates the significant effect of residual stress on the
fatigue life and hence should be included in the fatigue life assessment using LEFM.
29
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
SUGGESTION FOR FUTURE WORK
ƒ
Tools for better fatigue life prediction and equivalent mixed mode crack growth
models are required.
ƒ
Better and more effective detection and characterisation methods of weld defects (cold
laps) in connection with fatigue testing.
ƒ
General models for residual stress relaxation in variable amplitude spectrum loading.
SUMMARY OF APPENDED PAPERS
In Paper A fatigue testing and weld quality of Hybrid Nd: YAG laser/MAG and MAG
welded (tandem arc solid wire, flux cored wire, tandem flux cored wire) non-load carrying
cruciform joints is examined. Four batches were produced, tested and the results were
compared. The local weld geometry of the cruciform welded joints was measured and
analyzed. Residual stress measurement was carried out close to the toe region using the X-ray
diffraction. Weld defects (cold laps) in the cracked specimens was measured. The MAG flux
cored wire welded batch showed best fatigue resistance (FAT 136 MPa), low residual stresses
(-40 to -180 MPa) and almost no weld defects (cold laps << 0.03 mm).
In Paper B fatigue testing and defect assessment were carried out on specimens welded with
robotic and manual welding using flux cored (FCAW) and metal cored (MCAW) filler
materials in order to study the effect of the welding method on the fatigue strength, weld
quality and weld defects. The few largest defects were removed by the shot peening process,
although small defects survived. This led to a smaller scatter in fatigue live for the shot
peened specimens. The fatigue life predictions using 2D LEFM FE-model for simulating a
continuous cold lap defect along the weld toe showed a qualitative agreement with the fatigue
test results. A 3D LEFM analysis of crack growth from a spatter -induced cold lap defect was
also carried out. This showed similar trends in crack growth compared to the 2D analysis of a
continuous cold lap, although the spatter-induced cold lap defect (semi-elliptical) had a longer
fatigue life (x2.7), and hence is less dangerous from a fatigue point of view.
In Paper C the residual stresses in multi-pass welded tube-to-plates were studied. Two
different tubular joint configurations were studied; a three-pass single-U weld groove for
maximum weld penetration and a two-pass fillet (no groove) welded tube-to-plates for
minimum weld penetration. The three pass welded tubular joint showed nearly double the
residual compressive stress level near the weld root (-230 MPa) as compared to the two-pass
weld (-120 MPa). Torsion fatigue tests were performed in order to study crack propagation
from the weld root, lower and upper weld toe in Mode III. Some tubes were stress relieved
(PWHT) and some was fatigue tested with internal static pressure. LEFM was employed in
order to assess the effect of the residual stress.
In Paper D a three dimensional finite element welding simulation of the 3 pass welded
tubular joint was carried out. The calculated temperatures in the transient thermal analysis
were compared with temperature measurements. The FE predicted residual stresses in the as30
Z. Barsoum
Residual Stress Analysis and Fatigue Assessment of Welded Steel Structures
welded conditions were verified with hole drilling strain gage measurements. The residual
stresses was used as internal stresses in the finite element model for the torsion fatigue
simulation in order to study the relaxation of the residual stresses in constant amplitude
loading.
In Paper E a welding simulation procedure is developed using the FE-software ANSYS in
order to predict residual stresses. A subroutine for LEFM analysis was developed in 2D in
order to predict the crack path of propagating fatigue cracks. A subroutine was developed in
order to incorporate the predicted residual stresses and their relaxation during crack
propagation by stress mapping between meshes without and with cracks, respectively. The
objective was to investigate fatigue test results from special designed test bars from a welded
ship engine frame box at MAN B&W where all test failed from the non-penetrated weld root.
The LEFM fatigue life predictions shows good agreement with the fatigue test result when the
residual stresses are taken into account in the crack growth analysis.
In Paper F two- and three dimensional finite element welding simulations have been carried
out on a T-fillet weld. The objective is to study the residual stresses and the 3D effect of the
welding process. Moreover, welding simulations using solid models and contact models in the
unfused weld roots were carried out in order to investigate the possible effect with respect to
the residual stresses. Residual stress measurements were carried out using X-ray diffraction
technique. The 2D residual stress predictions shows good agreement with measurements, and
is suitable for residual stress predictions for incorporation in further fatigue crack growth
analysis from weld defects emanating from the weld toe and root.
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