[A. hobbacher] fatigue_design_of_welded_joints_and

The International Institute of Welding
Fatigue design of
welded joints and
components
Recommendations of IIW
Joint Working Group XIII-XV
XIII-1539-96IXV-845-96
A Hobbacher
ABINGTON PUBLISHING
Woodhead Publishing Ltd in association with The Welding Institute
Cambridge England

This document contains contributions from:
Prof Dr A Hobbacher, FH Wilhelmshaven, Germany
Prof P Haagensen, Inst of Technology, Trondheim, Norway
M Huther, Bureau Veritas, France
Prof Dr K Iida, Inst of Technology, Shibaura, Japan
Dr Y F Kudriavtsev, Paton Welding Institute, Kiev, Ukraine
Dr H P Lieurade, CETIM, Senlis, France
Dr S J Maddox, TWI, Cambridge, UK
Prof Dr Ch Miki, Inst of Technology, Tokyo, Japan
Prof Erkki Niemi, Lappeenranta Univ of Technology, Finland
A Ohta, NRIM, Tokyo, Japan
Oddvin 0rjasreter, SINTEF, Trondheim, Norway
Prof Dr H J Petershagen, Univ Hamburg, Germany
DR V van Delft, Delft Univ of Technology, The Netherlands
Published by Abington Publishing
Woodhead Publishing Limited, Abington Hall, Abington,
Cambridge CB21 6AH, England
www.woodheadpublishing.com
First published 1996, Abington Publishing
© 1996, The International Institute of Welding
Conditions ofsale
All rights reserved. No part of this publication may be reproduced or
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No responsibility is assumed by the Publisher for any injury and/or damage to
persons or property as a matter of products liability, negligence or otherwise, or
from any use or operation of any methods, products, instructions or ideas
contained in the material herein.
British Library Cataloguing in Publication Data
A catalogue record for this book is available from the British Library.
ISBN-13: 978-1-85573-315-2
ISBN-I0: 1-85573-315-3
Printed by Victoire Press Ltd, Cambridge, England

TABLE OF CONTENTS
1 GENERAL. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 6
1.1 INTRODUCTION .............................. 6
1.2 SCOPE AND LIMITATIONS ....................... 6
1.3 DEFINITIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 7
1.4 SYMBOLS .................................. 12
1.5 BASIC PRINCIPLES . . . . . . . . . . . . . . . . . . . . . . . . . . .. 13
1.6 NECESSITY FOR FATIGUE ASSESSMENT. . . . . . . . . . . .. 13
1.7 APPLICATION OF THE DOCUMENT ................ 14
2 FATIGUE ACTIONS (LOADING) ......................... 17
2.1 BASIC PRINCIPLES ................... . . . . . . . .. 17
2.1.1 Determination of Actions .................... 17
2.1.2 Stress Range . . . . . . . . . . . . . . . . . . . . . . . . . . .. 17
2.1.3 Types of Stress Raisers and Notch Effects ......... , 18
2.2 DETERMINATION OF STRESSES AND STRESS INTENSITY
FACTORS ................................ 19
2.2.1 Definition of Stress Components . . . . . . . . . . . . . . .. 19
2.2.2 Nominal Stress .......................... 20
2.2.2.1 General ......................... 20
2.2.2.2 Calculation of Nominal Stress . . . . . . . . . . .. 22
2.2.2.3 Measurement of Nominal Stress .......... 22
2.2.3 Geometric Stress (hot spot stress) .. . . . . . . . . . . . .. 23
2.2.3.1 General ......................... 23
2.2.3.2 Calculation of Geometric Stress .......... 24
2.2.3.3 Calculation of Geometric Stress by Parametric
Formulae . . . . . . . . . . . . . . . . . . . . . . . .. 26
2.2.3.4 Measurement of Geometric Stress ......... 26
2.2.4 Effective Notch Stress . . . . . . . . . . . . . . . . . . . . .. 28
2.2.4.1 General ......................... 28
2.2.4.2 Calculation of Effective Notch Stress ....... 29
2.2.4.3 Measurement of Effective Notch Stress . . . . .. 29
2.2.5 Stress Intensity Factors ..................... 30
2.2.5.1 General ......................... 30
2.2.5.2 Calculation of Stress Intensity Factors by Para-
metric Formulae .................... 30
2.2.5.3 Calculation of Stress Intensity Factors by Finite
Elements . . . . . . . . . . . . . . . . . . . . . . . .. 31
2.3 STRESS HISTORY . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 31
page 3

3 FATIGUE RESISTANCE . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 34
3.1 BASIC PRINCIPLES ..... . . . . . . . . . . . . . . . . . . . . . .. 34
3.2 FATIGUE RESISTANCE OF CLASSIFIED STRUCTURAL DE-
TAILS ................................... 34
3.2.1 Steel ................................ , 37
3.2.2 Aluminium . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 57
3.3 FATIGUE RESISTANCE AGAINST GEOMETRIC STRESS (HOT SPOT
STRESS) ................................. 73
3.3.1 Fatigue Resistance using Reference S-N Curve . . . . . .. 73
3.3.1.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 73
3.3.1.2 Aluminium . . . . . . . . . . . . . . . . . . . . . .. 73
3.3.2 Fatigue Resistance Using a Reference Detail ........ 73
3.4 FATIGUE RESISTANCE AGAINST EFFECTIVE NOTCH
STRESS ............ . . . . . . . . . . . . . . . . . . . . .. 75
3.4.1 Steel. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 75
3.4.2 Aluminium . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 75
3.5 FATIGUE STRENGTH MODIFICATIONS .............. 76
3.5.1 Stress Ratio ............................ 76
3.5.1.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 76
3.5.1.2 Aluminium ......... . . . . . . . . . . . . .. 76
3.5.2 Wall Thickness .......................... 77
3.5.2.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 77
3.5.2.2 Aluminium ................... . . .. 78
3.5.3 Improvement Techniques .................... 78
3.5.4 Effect of Elevated Temperatures. . . . . . . . . . . . . . .. 79
3.5.4.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 79
3.5.4.2 Aluminium ........ . . . . . . . . . . . . . .. 79
3.5.5 Effect of Corrosion. . . . . . . . . . . . . . . . . . . . . . .. 79
3.6 FATIGUE RESISTANCE AGAINST CRACK PROPAGATION.. 80
3.6.1 Steel. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 80
3.6.2 Aluminium . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 80
3.7 FATIGUE RESISTANCE DETERMINATION BY TESTING . .. 81
3.8 FATIGUE RESISTANCE OF JOINTS WITH WELD IMPERFEC-
TIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 83
3.8.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 83
3.8.1.1 Types of Imperfections. . . . . . . . . . . . . . .. 83
3.8.1.2 Effects and Assessment of Imperfections '" .. 83
3.8.2 Misalignment ........................... 85
3.8.3 Undercut . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 86
3.8.3.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 86
3.8.3.2 Aluminium ....................... 87
3.8.4 Porosity and Inclusions ..................... 87
3.8.4.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . .. 88
3.8.4.2 Aluminium .................. . . . .. 88
page 4

3.8.5 Cracklike Imperfections . . . . . . . . . . . . . . . . . . . .. 89
3.8.5.1 General Procedure. . . . . . . . . . . . . . . . . .. 89
3.8.5.2 Simplified Procedure .,. . . . . . . . . . . . . .. 90
4 FATIGUE ASSESSMENT .............................. 94
4.1 GENERAL PRINCIPLES ......................... 94
4.2 COMBINATION OF NORMAL AND SHEAR STRESS ...... 94
4.3 FATIGUE ASSESSMENT USING S-N CURVES .......... 95
4.3.1 Linear Damage Calculation by "Palmgren-Miner" Summa-
tion ................................ 95
4.3.2 Nonlinear Damage Calculation . . . . . . . . . . . . . . . .. 99
4.4 FATIGUE ASSESSMENT BY CRACK PROPAGATION CAL-
CULATION . . . . . . . . . . . . . . . . . . . . . . . . . . . ... . .. 100
4.5 FATIGUE ASSESSMENT BY SERVICE TESTING. . . . . . . .. 101
4.5.1 General. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 101
4.5.2 Safe Life Verification ...................... 101
4.5.3 Fail Safe Verification ...................... 103
4.5.4 Damage Tolerant Verification ................. 103
5 SAFETY CONSIDERATIONS . . . . . . . . . . . . . . . . . . . . . . . . . . .. 104
5.1 BASIC PRINCIPLES ...... . . . . . . . . . . . . . . . . . . . . .. 104
5.2 FATIGUE DESIGN STRATEGIES ................... 104
5.2.1 Infinite Life Design ....................... 104
5.2.2 Safe Life Design ......................... 105
5.2.3 Fail Safe Design ......................... 105
5.2.4 Damage Tolerant Design .................... 105
5.3 PARTIAL SAFETY FACTORS ..................... 105
5.4 QUALITY ASSURANCE ......................... 106
6 APPENDICES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 107
6.1 LOAD CYCLE COUNTING ....................... 107
6.1.1 Transition Matrix . . . . . . . . . . . . . . . . . . . . . . . .. 107
6.1.2 Rainflow or Reservoir Counting Method . . . . . . . . . ., 107
6.2 FRACTURE MECHANICS .... . . . . . . . . . . . . . . . . . . .. 108
6.2.1 Rapid Calculation of Stress Intensity Factors ........ 108
6.2.2 Dimensions of Cracks .. . . . . . . . . . . . . . . . . . . .. 109
6.2.3 Interaction of Cracks . . . . . . . . . . . . . . . . . . . . . .. 110
6.2.4 Formulae for Stress Intensity Factors . . . . . . . . . . . .. 110
6.3 FORMULAE FOR MISALIGNMENT ................. 115
6.4 STATISTICAL CONSIDERATIONS ON SAFETY ......... 119
6.4.1 Statistical Evaluation of Fatigue Test Data. . . . . . . . .. 119
6.4.2 Statistical Evaluation at Component Testing . . . . . . . .. 120
6.4.3 Statistical Considerations for Partial Safety Factors .... 122
7 REFERENCES ..................................... 123
page 5

1 GENERAL
The IIW, every other body or person involved in the preparation and publication of
this document hereby expressly disclaim any liability or responsibility for loss or
damage resulting from its use, for any violation of any mandatory regulation with
which the document may conflict, or for the infringement of any patent resulting from
the use of this document.
It is the user's responsibility to ensure that the recommendations given here are
suitable for his purposes.
1.1 INTRODUCTION
The aim of these recommendations is to provide a basis for the design and analysis of
welded components loaded by fluctuating forces, to avoid failure by fatigue. In
addition they may assist other bodies who are establishing fatigue design codes. It is
assumed that the user has a working knowledge of the basics of fatigue and fracture
mechanics.
The purpose of designing a structure against the limit state due to fatigue damage is to
ensure, with an adequate survival probability, that the performance is satisfactory
during the design life. The required survival probability is obtained by the use of
appropriate partial safety factors.
1.2 SCOPE AND LIMITATIONS
The recommendations present general methods for the assessment of fatigue damage
in welded components, which may affect the limit states of a structure, such as ul-
timate limit state and servicability limited state [1].
The recommendations give fatigue resistance data for welded components made of
wrought or extruded products of ferritic/pearlitic or bainitic structural steels up to
fy =700 MPa and of aluminium alloys commonly used for welded structures.
The recommendations are not applicable to low cycle fatigue, where aqnom> l.S·fy, for
corrosive conditions or for elevated temperature operation in the creep range.
page 6

1.3 DEFINITIONS
Characteristic value
Classified structural
detail
Concentrated load
effect
Constant amplitude
loading
Crack propagation
rate
Crack propagation
threshold
Cut off limit
Design value
Effective notch
stress
Equivalent stress
range
Loads, forces or stresses, which vary statistically, at a
specified fractile, here: 95 % at a confidence level of the
mean of 75% .
A structural detail containing a structural discontinuity
including a weld or welds, for which the nominal stress
approach is applicable, and which appear in the tables of
the recommendation. Also referred to as standard struc-
tural detail.
A local stress field in the vicinity of a point load or reac-
tion force, or membrane and shell bending stresses due
to loads causing distortion of a cross section not suffi-
ciently stiffened by a diaphragm.
A type of loading causing a regular stress fluctuation
with constant magnitudes of stress maxima and minima.
Amount of crack tip propagation during one stress cycle.
Limiting value of stress intensity factor range below
which crack propagation will not occur.
Fatigue strength under variable amplitude loading, below
which the stress cycles are considered to be non-dama-
ging.
Characteristic value factored by a partial safety factor.
Notch stress calculated for a notch with a certain effec-
tive notch radius.
Constant amplitude stress range which is equivalent in
terms of fatigue damage to the variable amplitude loading
under study, at the same number of cycles.
page 7

Fatigue
Fatigue action
Fatigue damage ratio
Fatigue life
Fatigue limit
Fatigue resistance
Fatigue strength
Fracture mechanics
Geometric stress
Hot spot
Hot spot stress
Local nominal stress
Local notch
Detoriation of a component caused by crack initiation
and/or by the growth of cracks.
Load effect causing fatigue.
Ratio of fatigue damage sustained to fatigue damage
required to cause failure, defined as the ratio of the
number of applied stress cycles and the corresponding
fatigue life at constant amplitude.
Number of stress cycles of a particular magnitude re-
quired to cause fatigue failure of the component.
Fatigue strength under constant amplitude loading corre-
sponding to infinite fatigue life or a number of cycles
large enough to be considered infinite by a design code.
Structural detail's resistance against fatigue actions in
terms of S-N curve or crack propagation properties.
Magnitude of stress range leading to a particular fatigue
life.
A branch of mechanics dealing with the behaviour and
strength of components containing cracks.
See 'hot spot stress'
A point in a structure where a fatigue crack may initiate
due to the combined effect of structural stress fluctuation
and the weld geometry or a similar notch.
The value of structural stress on the surface at a hot spot
(also known as geometric stress).
Nominal stress including macro-geometric effects, con-
centrated load effects and misalignments, disregarding
the stress raising effects of the welded joint itself. Also
referred to as modified nominal stress.
A notch such as the local geometry of the weld toe,
including the toe radius and the angle between the base
plate surface and weld reinforcement. The local notch
does not alter the structural stress but generates nonlinear
stress peaks.
page 8

~acro-geomeUic
discontinuity A global discontinuity, the effect of which is usually not
taken into account in the collection of standard structural
details, such as a large opening, a curved part in a
beam, a bend in a flange not supported by diaphragms or
stiffeners, discontinuities in pressure containing shells,
eccentricity in a lap joint (see fig. (2.2)-3).
~acro-geomeUic effect A stress raising effect due to macro-geometry in the
vicinity of the welded joint, but not due to the welded
joint itself.
~embrane stress Average normal stress across the thickness of a plate or
shell.
~ner sum Summation of individual fatigue damage ratios caused by
each stress cycle or stress range block above a certain
cut-off limit according to the Palmgren-~er rule.
~salignment Axial and angular misalignments caused either by detail
design or by poor fabrication or welding distortion.
~odified nominal stress See 'Local nominal stress'.
Nominal stress A stress in a component, resolved using general theories,
e.g. beam theory. See also local nominal stress.
Nonlinear stress peak The stress component of a notch stress which exceeds the
linearly distributed structural stress at a local notch.
Notch stress Total stress at the root of a notch taking into account the
stress concentration caused by the local notch, consisting
of the sum of structural stress and nonlinear stress peak.
Notch stress concentration
factor The ratio of notch stress to structural stress.
Paris' law An experimentally determined relation between crack
growth rate and stress intensity factor range.
Palmgren-Miner rule
Fatigue failure is expected when the Miner sum reaches
unity.
page 9

Rainflow counting
Range counting
Shell bending stress
S-N curve
Stress cycle
Stress history
Stress intensity
factor
Stress range
Stress range block
Stress range exceedances
Stress range occurrences
Stress ratio
A standardized procedure for stress range counting.
A procedure of determining various stress cycles and
their ranges from a stress history, preferably by rainflow
counting method.
Bending stress in a shell or plate-like part of a com-
ponent, linearly distributed across the thickness as as-
sumed in the theory of shells.
Graphical presentation of the dependence of fatigue life
N on fatigue strength S (~O"R or ~TJ, also known as
Wohler curve.
A part of a stress history containing a stress maximum
and a stress minimum, determined usually by a range
counting method.
A time based presentation of a fluctuating stress, defined
by sequential stress peaks and troughs (valleys), either
for the total life or for a certain sample.
Main parameter in fracture mechanics, the combined
effect of stress and crack size at the crack tip region.
The difference between stress maximum and stress mini-
mum in a stress cycle, the most important parameter
governing fatigue life.
A part of the total spectrum of stress ranges which is dis-
cretized in a certain number of blocks.
A tabular or graphical presentation of the cumulative
frequency of stress range exceedances, i.e the number of
ranges exceeding a particular magnitude of stress range
in a stress history. Here, frequency is the number of
occurrances. (Also referred to as "stress spectrum" or
"cumulative frequency diagram").
A tabular or graphical presentation of stress ranges, usu-
ally discretized in stress range blocks. See also "stress
range exceedances".
Ratio of minimum to maximum algebraic value of the
page 10

stress in a particular stress cycle.
Stress intensity factor ratio Ratio of minimum to maximum algebraic value of the
stress intensity factor of a particular load cycle.
Structural discontinuity
Structural stress
Structural stress
concentration factor
A geometric discontinuity due to the type of welded
joint, usually to be found in the tables of classified struc-
tural details. The effects of a structural discontinuity are
(i) concentration of the membrane stress and (ii) for-
mation of secondary shell bending stresses (see fig.
(2.2)-6).
A stress in a component, resolved taking into account the
effects of a structural discontinuity, and consisting of
membrane and shell bending stress components. Also
referred to as geometric stress.
The ratio of structural (hot spot) stress to modified (lo-
cal) nominal stress.
Variable amplitude loading A type ofloading causing irregular stress fluctuation with
stress ranges (and amplitudes) of variable magnitude.
page 11

1.4 SYMBOLS
K
Kmax
Kmm
Mk
Mk,m
M ..,b
R
Y
ar
e
fy
ku.
ks
~
m
t
4K
4KS,d
~
aO'
aO'S,d
aO'R,L
aT
'YM
r M
0'
stress intensity factor
stress intensity factor caused by O'max
stress intensity factor caused by O'min
magnification function for K due to nonlinear stress peak
magnification function for K, concerning membrane stresses
magnification function for K, concerning shell bending stresses
stress ratio
correction function for K, taking into account crack form, aspect ratio,
relative crack size etc.
correction function for K, concerning membrane stress
correction function for K, concerning shell bending stress
depth of a surface crack or semi length of a through crack
initial depth of a surface crack
crack size at failure
eccentricity, amount of offset misalignment
yield strength of the material
stress magnification factor due to misalignment
stress concentration factor due to structural discontinuity
stress concentration factor due to local notch
exponent of S-N curve or Paris power law
plate thickness, thickness parameter (crack center to nearest surface)
stress intensity factor range
design value of stress intensity factor range caused by actions
threshold stress intensity factor range
stress range
design value of stress range caused by actions
characteristic value of fatigue limit
shear stress range
partial safety factor for fatigue resistance in terms of stress
partial safety factor for fatigue resistance in terms of cycles
normal stress
shell bending stress
effective notch stress
(local) notch stress
stress maximum in stress history
membrane stress
Subscripts:
S fatigue actions
R fatigue resistance
O'min stress minimum in stress history
O'nlp nonlinear stress peak d design value
O'nom nominal stress k characteristic value
O'geo geometric stress, structural stress
T shear stress
page 12

1.5 BASIC PRINCIPLES
According to the ISO format for verification of structures [1], fatigue action and
fatigue resistance are clearly separated. Fatigue resistance is given in terms of tentative
data. The representation of tentative data has also been separated from the assessment
curves used for damage calculation, because different damage calculation methods may
require special modifications to the resistance S-N curve, which is usually based on
constant amplitude tests. Thus, the flexibility and possibility for continuous updating
of the document is maintained.
No recommendations are given for the fatigue load (action) side, nor for the partial
safety factor on fatigue actions 'YF.
The different approaches for the fatigue assessment of welded joints and components
considered are: nominal stress, geometric stress, effective notch stress, fracture mecha-
nics method and component testing.
1.6 NECESSITY FOR FATIGUE ASSESSMENT
Fatigue assessment is generally required for components subject to fluctuating loads.
In the following cases, detailed fatigue assessment is not required:
a) The highest nominal design stress range satisfies
steel /loS.d ~ 36 [MPa] I YM
aluminium: /loS.d ~ 14 [MPa] I YM
'YM should be taken from an applicable design code.
This paragraph is not applicable to tubular joints.
b) A Miner sum (4.3.1) equal to D=O.S using a FAT fatigue class accor-
ding to (3.2) of FAT 36 for steel or FAT 14 for aluminium corresponds
to a fatigue life greater than 5 million cycles.
c) For a detail for which a constant amplitude fatigue limit /lUR,L is speci-
fied and all design stress ranges are under the design resistance fatigue
limit
page 13

d) For a crack, at which all design stress intensity factors are under the
threshold level 4~ for crack propagation.
for steel
for aluminium
I1Ks,d ~ AKth I YM
4~=2.0 MPa"m
4~=O.7 MPa"m
1.7 APPLICATION OF THE DOCUMENT
Based on the initial information about the welded joint, the stress type to be used in
the fatigue assessment is defined and determined. The fatigue resistance data are then
selected according to the stress type of the fatigue action. The corresponding types of
fatigue action and resistance are:
Tab. {l}-l: Fatigue actions and resistances
IFatigue action IFatigue resistance I
Nominal stress Resistance given by tables of structural
details in terms of a set of S-N curves
Geometric stress (hot spot stress) Resistance against geometric stress in terms
of S-N curve
Effective notch stress Resistance against effective notch stress in
terms of a universal S-N curve
Stress intensity at crack tip Resistance against crack propagation in
terms of the material parameters of the
crack propagation law
The fatigue assessment procedure depends on the presentation of fatigue resistance
data. The chosen procedure has to be performed using adequate safety factors.
Tab. {1}-2: Assessment procedures
Presentation of fatigue resistance data Assessment procedure
S-N curves Linear cumulative damage, (in special
cases nonlinear damage calculation)
Material parameters of crack propaga- Crack propagation calculation
tion law
No data available Fatigue testing
page 14

Tab. {1}-3: General guidance for the application of the document
Item Initial Infor- Fatigue Fatigue
mation Action Resistance
(1) Does joint determine look up go to
correpond to a nominal fatigue (6)
tabulated yes - stress (2.2.2) then - resistance
structural class (FAT)
detail? in tables
(3.2)
if no +
(2) Is geometric determine look up re-
stress assess- geometric sistance S-N
ment ap- yes - stress (2.2.3) then - curve for go to
plicable? geometric (6)
stress (3.3)
if no +
(3) Is effective determine look up re-
notch stress effective sistance S-N
assessment yes - notch stress then- curve for go to
applicable? (2.2.4) effective (6)
notch stress
(3.4)
if no +
(4) Are cracks or determine look up
cracklike stress inten- resistance
imperfections yes - sity factor then - against go to
present? (2.2.5) crack pro- (7)
pagation
(3.6 and
3.8)
if no +
page 15

(5) Test entire go to (8)
component
(4.5)
test structural go to (1)
detail (3.7)
Modifications and Assessment Procedures
(6) Modify resis- is Miner calculate perform
tance S-N rule ad- design resis- summation
curve (3.5) equate tance S-N then - (4.3.1)
for all effects (4.3)1 curve (4.3.1) giving life
not yet co- yes - using 'YM (8) cycles,
vered assess if OK
if no - calc. dimen-
sionless crack
propagation
paramo from
resistance
S-N curve
(4.3.2) using
'YM (8)
then ~
(7) calc. design perform
crack propa- crack pro-
gation resis- then- pagation calc.
tance data (4.4) giving
using r M(8) life cycles
assess if OK
Safety Considerations
(8) define 'YM according to safety considerations (chapter 5)
page 16

2 FATIGUE ACTIONS (LOADING)
All types of fluctuating load acting on the component and the resulting stresses at
potential sites for fatigue have to be considered. Stresses or stress intensity factors then
have to be determined according to the fatigue assessment procedure applied.
The actions originate from live loads, dead weights, snow, wind, waves, pressure, ac-
celerations, dynamic response etc. Actions due to transient temperature changes should
be considered. Improper knowledge of fatigue actions is one of the major sources of
fatigue damage.
Tensile residual stresses due to welding decrease the fatigue resistance, however, the
influence of residual weld stresses is already included in the fatigue resistance data
given in chapter 3.
2.1.1 Determination of Actions
The actions in service have to be determined in terms of characteristic loads. Partial
safety factors on actions 'YF have to be applied as specified in the application code
giving the design values of the actions for fatigue assessment.
In this document, there is no guidance given for the establishing of design values for
actions (loads), nor for partial safety factors 'YF for actions (loads).
2.1.2 Stress Range
Fatigue assessment is usually based on stress range or stress intensity factor range.
Thus, the actions have to be given in these terms.
ll.K :;: Kmax - Kmm
page 17

The maximum and the minimum values of the stresses are to be calculated from a
superposition of all non permanent, i.e. fluctuating, actions:
a) fluctuations in the magnitudes of loads
b) movement of loads on the structure
c) changes in loading directions
d) structural vibrations due to loads and dynamic response
e) temperature transients
Fatigue analysis is based on the cumulative effect of all stress range occurrences
during the anticipated service life of the structure.
2.1.3 Types of Stress Raisers and Notch Effects
Different types of stress raisers and notch effects lead to the calculation of different
types of stress. The choice of stress depends on the fatigue assessment procedure used.
Tab. {2}-1: Stress raisers and notch effects
Type Stress raisers Stress determined Assessment procedure
A General stress analysis
using general theories
e.g. beam theory
B A + Macrogeometrical Range of nominal Nominal stress ap-
effects due to the design stress (also modi- proach
of the component (also tied or local no-
effects of concentrated minal stress)
loads and misalignments)
C A + B + Structural Range of struc- Geometric stress (hot
discontinuities due to the tural geometric spot stress) approach
structural detail of the stress (hot spot
welded joint stress)
D A + B + C + Notch Range of elastic a) Fracture mechanics
stress concentration due notch stress (total approach
to the weld bead (e.g. at stress) b) effective notch
toe or root) stress approach
a) actual notch stress
b) effective notch stress
page 18

2.2 DETERMINATION OF STRESSES AND STRESS
INTENSITY FACTORS
2.2.1 Definition of Stress Components
The stress distribution over the plate thickness is non-linear in the vicinity of notches.
----Notch stress = O'mem + O'ben + O'nlp
",--
o
Fig. (2.2)-1 Non-linear stress distribution separated to stress components
The stress components of the notch stress 0"1n are [2]:
O"mem membrane stress
O"hen shell bending stress
O"Dlp non-linear stress peak
If a refmed stress analysis method is used, which gives a non-linear stress distribution,
the stress components can be separated by the following method:
The membrane stress O"mem is equal to the average stress calculated through the
thickness of the plate. It is constant through the thickness.
The shell bending stress O"hen is linearly distributed through the thickness of the
plate. It is found by drawing a straight line through the point 0 where the
membrane stress intersects the mid-plane of the plate. The gradient of the shell
bending stress is chosen such that the remaining non-linearly distributed com-
ponent is in equilibrium.
The non-linear stress peak O"Dlp is the remaining component of the stress.
The stress components can be separated analytically for a given stress distribution u(x)
for x=O at surface to x=t at through thickness:
page 19

x=t
amem = 1.. Ja(x)·dx
t x=o
x=t
6 J ta = _. o(x)·(--x)·dx
ben 2 2
t ,,=0
a.J..(x) = a(x)-a -(l-x),o
'Nt' mem 2 ben
2.2.2 Nominal Stress
2.2.2.1 General
Nominal stress is the stress calculated in the sectional area under consideration, dis-
regarding the local stress raising effects of the welded joint, but including the stress
raising effects of the macrogeometric shape of the component in the vicinity of the
joint, such as e.g. large cutouts. Overall elastic behaviour is assumed.
The nominal stress may vary over the section under consideration. E.g. at a beam-like
component, the modified (also local) nominal stress and the variation over the section
can be calculated using simple beam theory. Here, the effect of a welded on attach-
ment is ignored.
~--------------------------------------~
Weld
'-......CJ nom
rM
~M
Fig. (2.2)-2 Nominal stress in a beam-like component
The effects of macrogeometric features of the component as well as stress fields in the
vicinity of concentrated loads must be included in the nominal stress. Consequently,
macrogeometric effects may cause a significant redistribution of the membrane stresses
across the section. Similar effects occur in the vicinity of concentrated loads or
reaction forces. Significant shell bending stress may also be generated, as in curling of
a flange, or distortion of a box section.
The secondary bending stress caused by axial or angular misalignment needs to be
page 20

(a)
(J
nom
nom
(e) ...4;,-~..~..:;;:;;:.~..
,,1/
- P ~.-.-..
/
~-.-.-.-.-.-.-,~
Fig. (2.2)-3 Examples of macrogeometric effects
F
(c)
(f)
Fig. (2.2)-4 Modified Oocal) nominal stress near concentrated loads
(J
nom
F
considered if the misalignment exceeds the amount which is already covered by fatigue
resistance S-N curves for the structural detail. This is done by the application of an
additional stress raising factor km,efT (see 3.8.2). Intentional misalignment (e.g.allowa-
ble misalignment specified in the design stage) is considered when assessing the fatigue
actions (stress) by multiplying by the factor. If it is non-intentional, it is regarded as
a weld imperfection which affects the fatigue resistance and has to be considered by
dividing the fatigue resistance (stress) by the factor.
page 21

(a) (b)
c)
Fig. (2.2)-5 Axial and angular misalignement
2.2.2.2 Calculation of Nominal Stress
In simple components the nominal stress can be determined using elementary theories
of structural mechanics based on linear-elastic behaviour. In other cases, finite element
method (FEM) modelling may be used. This is primarily the case in:
a) complicated statically over-determined (hyperstatic) structures
b) structural components incorporating macrogeometric discontinuities, for
which no analytical solutions are available
Using FEM, meshing can be simple and coarse. Care must be taken to ensure that all
stress raising effects of the structural detail of the welded joint are excluded when
calculating the modified Oocal) nominal stress.
2.2.2.3 Measurement of Nominal Stress
The fatigue resistance S-N curves of classified structural details are based on nominal
stress, disregarding the stress concentrations due to the welded joint. Therefore the
measured nominal stress must exclude the stress or strain concentration due to the cor-
responding discontinuity in the structural component. Thus, strain gauges must be
placed outside of the stress concentration field of the welded joint.
In practice, it may be necessary firstly to evaluate the extension and the stress gradient
of the field of stress concentration (see 2.2.3.4) due to the welded joint. For further
measurements, simple strain gauge application outside this field is sufficient.
page 22

2.2.3 Geometric Stress (hot spot stress)
2.2.3.1 <ieneral
The structural geometric stress includes all stress raising effects of a structural detail
excluding all stress concentrations due to the weld profile itself. So, the non-linear
peak stress unlp caused by the local notch, e.g. by the weld toe, is excluded from the
geometric stress. The geometric stress is dependent on the global dimensional and
loading parameters of the component in the vicinity of the joint (type C in 2.1.3 table
{2}-l). They are determined on the surface at the point of the component which is to
be assessed.
a) b)
c) d)
Fig. (2.2)-6 Structural details and geometric stress
Structural geometric stresses ugeo are generally encountered in plate, shell and tubular
structures. They can be divided in two stress components, the membrane stress Umem
and the shell bending stress Uben. Because of the inclusion of stress raising effects of
structural discontinuities they are usually higher than the nominal stresses.
page 23

For fatigue assessment, the geometric stress has to be determined in the critical direc-
tion at the critical point of a welded joint (hot spot), where fatigue crack initiation is
expected. In general, maximum principle stress is used. Figure (2.2)-6 shows examples
of structural discontinuities and details together with the structural geometric stress
distribution.
computed total
stress
measuring points
geometric stress
1/ I
stress on surface
hot spot
Fig. (2.2)-7 Defmition of geometric stress
The geometric stress approach is recommended for welded joints where there is no
clearly defined nominal stress due to complicated geometric effects, and where the
structural discontinuity is not comparable to a classified structural detail.
It is important that the same stresses are determined for the fatigue action and the
fatigue resistance (see 3.3). The calculation or measurement procedures have to
correspond as closely as possible. When using measured geometric stress, no correc-
tion for misalignment is necessary. However, calculations have to refer to the actual
shape a joint including any possible misalignment.
The method is limited to the assessment of the weld toe.
2.2.3.2 Calculation of Geometric Stress
In general, analysis of structural discontinuities and details to obtain the geometric
stress is not possible using analytical methods. Parametric formulae are rarely availa-
ble. Thus, finite element (FEM) analysis is mostly applied. In this case the maximum
page 24

principal stress should be calculated.
Usually, geometric stress is calculated on the basis of an idealized, perfectly aligned
welded joint. Consequently, any possible misalignment has be taken into consideration
in the fatigue resistance data.
The FEM mesh must be fme enough near the critical point (hot spot) to enable the
stress and the·stress gradient to be determined at points comparable with the extrapo-
lation points used for strain gauge measurement (see 2.2.3.4).
For FEM analysis, sufficient expertise of the analyst is required. Guidance is given in
ref. [2]. In the following, only some rough recommendations are given:
a) Elements and number of integration points are to be selected on the basis of a
linear stress distribution over the plate thickness. 4-node thin shell or solid
elements or in cases of a steep stress gradient, 8-node thin shell elements or
20-node solid elements are recommended.
b) When thin shell elements are used, the structure is modelled at midface of the
plates or the tube walls. The stiffness of the weld intersection should be taken
into account, e.g. by modelling the welds with inclined shell elements [2]. This
is important when adjacent intersections are present, or when e.g. a longi-
tudinal gusset would cause a singularity behaviour.
c) The dimension of the first element adjacent to the weld toe, perpendicular to
the weld, must be such that valid results can be obtained at the extrapolation
points (see 2.2.3.4). The dimension of the element perpendicular to the inter-
section curve of the plates or tubes should be such, that the distance of the
center point or first integration point of the element to the weld toe is no more
than 0.4 t (distance a).
d) The FEM model should be able to represent the variation of stress along the
weld toe. For tubular joints, the dimension b of the element should be less than
1124 of the length of the intersection line.
e) The ratio between the biggest and smallest dimension of an element must not
exceed 3.
f) Transition of elements size should be gradual. For tubular structures, the maxi-
mum dimension of elements far from the weld toe is half the radius of the tube.
g) The stresses are calculated at the surface of the plate or shell.
h) When the weld size is not given in the drawings, it will be determined accor-
ding to the rules for welds for the type of the considered structure.
page 25

-1- -- B: race
r
a l1i 0.4 t
Fig. (2.2)-8 Maximum element size for plate and tubular
structures
2.2.3.3 Calculation of Geometric Stress by Parametric Formulae
For many joints between circular section
tubes parametric formulae have been
established for the stress concentration
factor ~ in terms of structural geometric
stress at the critical points (hot spots).
Hence the geometric stress ugeo becomes:
ft - k 00
VgllO S 110m
where Unom is the nominal axial. mem-
brane stress in the braces, calculated by
Possible crack initiation sites
elementary stress analysis. Fig. (2.2)-9 Example of tubular joint
2.2.3.4 Measurement of Geometric Stress
The geometric stress can be measured using two or three strain gauges attached at
particular distances from the weld toe. The closest gauge position must be chosen to
avoid any influence of the notch due to the weld itself (which leads to a non-linear
stress peak). The structural geometric stress at the weld toe is then obtained by
extrapolation.
Measurings have to be made at the critical points (hot spots) which can be found by:
a) measuring several different points
b) prior investigation by brittle lacquer
c) analysing the results of a prior FEM analysis
d) experience of existing components, which failed
e) photoelastic investigations
f) thermo-elastic investigations
page 26

Fig. (2.2)-10 Locations of strain gauges in plate structures
The placement of the strain gauges should lead to a reasonable extrapolation to the
critical point. The center point of the first gauge should be placed at a distance of 0.4 t
from the weld toe. The gauge length should not exceed 0.2 t. If this is not possible
due to a small plate thickness, the leading edge of the gauge should be placed at a
distance 0.3 t from the weld toe. The following extrapolation procedure and number
of gauges are recommended:
a) Two gauges and linear extrapolation in cases of mainly membrane stresses.
b) Three gauges and quadratic extrapolation in cases of shell bending stresses
caused e.g. by eccentric attachments in large diameter tubes or plane plates.
Figure (2.2-10) shows gauge positions for low bending stress due to low stiffness to
left, high bending stress due to high stiffness top right and bottom an example of a
thin walled structure.
For tubular joints, there exist recommendations which allow the use of linear extrapo-
lation using two strain gauges (see figure 2.2-11).
Usually the measurement of simple uniaxial stress is sufficient. For additional details
see ref. [2]
page 27

Brace
lit Real Stress
Distribution
•
Fig. (2.2)-11 Location of strain gauges at tubular structu-
res
2.2.4 Effective Notch Stress
2.2.4.1 General
Effective notch stress is the total stress at the root of a notch, obtained assuming
linear-elastic material behaviour. To take account of the statistical nature and scatter
of weld shape parameters, as well as of the non-linear material behaviour at the notch
root, the real weld contour is replaced by an effective one. For structural steels an
effective notch root radius of r =1 IDID has been verified to give consistent results.
For fatigue assessment, the effective notch stress is compared with a common fatigue
resistance curve.
The method is restricted to welded joints which are expected to fail from the weld toe
or weld root. Other causes of fatigue failure, e.g. from surface roughness or embedded
defects, are not covered. Also it is also not applicable where considerable stress com-
ponents parallel to the weld or parallel to the root gap exist.
The method is well suited to the comparison of alternative weld geometries. Unless
otherwise specified, flank angles of 300
for butt welds and 450
for fillet welds are
suggested.
In cases where a mean geometrical notch root radius can be defined, e.g. after certain
post weld improvement procedures, this geometrical radius plus 1 IDID may be used in
the effective notch stress analysis.
The method is limited to thicknesses t > = 5 IDID. For smaller wall thicknesses, the
method has not yet been verified.
page 28

2.2.4.2 Calculation of Effective Notch Stress
Effective notch stresses or stress concentration factors can be calculated by parametric
fonnulae, taken from diagrams or calculated from finite element or boundary element
models. The effective notch radius is introduced such that the tip of the radius touches
the root of the real notch, e.g. the end of an unwelded root gap.
Radius =1 mm
Fig. (2.2)-12 Effective notch stress concentration factors
Possible misalignment has to be considered in the calculations.
2.2.4.3 Measurement of Effective Notch Stress
Because the effective notch radius is an idealization, the effective notch stress cannot
be measured directly in the welded component. In contrast, the simple definition of the
effective notch can be used for photo-elastic stress measurements in resin models.
page 29

2.2.5 Stress Intensity Factors
2.2.5.1 General
Fracture mechanics assumes the existence of an initial crack 3j. It can be used to
predict the growth of the crack to a final size ar. Since for welds in structural metals,
crack initiation occupies only a small portion of the life, this method is suitable for as-
sessment of fatigue life, inspection intervals, crack-like weld imperfections and the
effect of variable amplitude loading.
The parameter which describes the fatigue action at a crack tip in terms of crack
propagation is the stress intensity factor (SIF) range .:1K.
Fracture mechanics calculations generally have to be based on total stress at the notch
root, e.g. at the weld toe. For a variety of welded structural details, correction func-
tions for the local notch effect and the nonlinear stress peak of the structural detail
have been established. Using these correction functions, fracture mechanics analysis
can be based on geometric stress or even on nominal stress. The correction function
formulae may be based on different stress types. The correction function and the stress
type have to correspond.
2.2.5.2 Calculation of Stress Intensity Factors by Parametric Formulae
First, the local nominal stress or the structural geometric stress at the location of the
crack has to be determined, assuming that no crack is present. The stress should be
separated into membrane and shell bending stresses. The stress intensity factor (SIF)
K results as a superposition of the effects of both stress components. The effect of the
remaining stress raising discontinuity or notch (non-linear peak stress) has to be
covered by additional factors M k•
where
K
O'mem
O'beo
Ymem
Ybeo
Mk,mem
stress intensity factor
membrane stress
shell bending stress
correction function for membrane stress intensity factor
correction function for shell bending stress intensity factor
correction for non-linear stress peak in terms of membrane ac-
tion
correction for non-linear stress peak in terms of shell bending
The correction functions Ymem and Yben can be found in the literature. The solutions in
page 30

ref. [14-16] are particularly recommended. For most cases, the formulae for stress
intensity factors given in appendix 6.2 are adequate. Mk-factors may be found in
references [19] and [20].
2.2.5.3 Calculation of Stress Intensity Factors by Finite Elements
Stress intensity factor determination methods are usually based on FEM analyses. They
may be directly calculated as described in the literature, or indirectly using the weight
function approach. For more details see appendix 6.2
2.3 STRESS mSTORY
2.3.1 General
The fatigue design data presented in chapter 3 were obtained from tests performed
under constant amplitude loading. However, loads and the resulting fatigue actions
(i.e. stresses) on real structures usually fluctuate in an irregular manner and give rise
to variable amplitude loading. The stress amplitude may vary in both magnitude and
periode from cycle to cycle. Because the physical phemomena implied in fatigue are
not chaotic but random, the fluctuations can be characterized by mathematical func-
tions and by a limited number of parameters.
The stress history is a record and/or a representation of the fluctuations of the fatigue
actions in the anticipated service time of the component. It is described in terms of
successive maxima and minima of the stress caused by the fatigue actions. It covers all
loading events and the corresponding induced dynamic response.
Fig. (2.3)-1 Stress time history illustration
In most cases, the stress-time history is stationary and ergodic, which allows the
definition of a mean range and its variance, a statistical histogram and distribution, an
energy spectrum and a maximum values probabilistic distribution from a representation
of a limited length. Therefore, the data needed to perform a fatigue analysis can be
page 31

determined from measurements conducted during a limited time.
A stress history may be given as
a) a record of successive maxima and minima of stress measured in a comparable
structure with comparable loading and service life, or a typical sequence of
load events.
b) a two dimensional transition matrix of the stress history derived from a).
c) a one- or two-dimensional stress range histogram (stress range occurrences)
obtained from a) by a specified counting method.
d) a one-dimensional stress range histogram (stress range exceedances, stress
range spectrum) specified by a design code.
The representations a) and b) may be used for component testing. c) and d) are most
useful for fatigue analysis by calculation.
2.3.2 Cycle Counting Methods
Cycle counting is the process of converting a variable amplitude stress sequence into
equivalent (in terms of fatigue damage) constant amplitude stress cycles. Different
methods of counting are in use, e.g. zero crossing counting, peak: counting, range pair
counting or rainflow counting. For welded components, the reservoir or rainflow
method is recommended for counting stress ranges [26 and 27].
2.3.3 Cumulative Frequency Diagram (Stress Spectrum)
The cumulative frequency diagram (stress spectrum) corresponds to the cumulative
probability of stress range expressed in terms of stress range level exceedances versus
the number of cycles. The curve is therefore continuous.
The spectrum may be discretized giving a table of discrete blocks. For damage cal-
culations 20 stress levels are recommended if more than 108 cycles are expected.
Below this number of cycles, 8 or 10 stress levels may be sufficient. All cycles in a
block should be assumed to be equal to the mean of the stress ranges in the block.
Besides the representation in probabilities, a presentation of the number of occurrences
or exceedances in a given number of cycles, e.g. 1 million, is used. An example
showing a Gaussian normal distribution is given below:
page 32

Tab. {2.3}-1: Stress range occurrance table (stress histogram or frequency)
# of block Relative Occurrence
stress range (frequency)
1 1.000 2
2 0.950 16
3 0.850 280
4 0.725 2720
5 0.575 20000
6 0.425 92000
7 0.275 280000
8 0.125 605000
Gaussian normal distribution
relative stress range
~
-~ ...
0,8
1-0...
.........
0,6
.......
to-.
0,4
~
0,2
o
10 100 1000 10000 100000 100000
cyclces
Fig. (2.3)-2 Cumulative frequency diagram (stress spectrum)
page 33

3 FATIGUE RESISTANCE
Fatigue resistance is usually derived from constant or variable amplitude tests. The
fatigue resistance data given here are based on published results from constant ampli-
tude tests. Guidance on the direct use of test data is given in section 3.7 and 4.5.
The fatigue resistance data must be expressed in terms of the same stress as that
controlled or determined during the generation of those data.
In conventional endurance testing, there are different definitions of failure. In general,
small specimens are tested to complete rupture, while in large components the obser-
vation of a through wall crack is taken as a failure criterion. The fatigue resistance
data are based on the number of cycles N to failure. The data are represented in S-N
curves
C
N=-
~GIII
or
In fracture mechanics crack propagation testing, the crack growth rate data are derived
from crack propagation monitoring.
All fatigue resistance data are given as characteristic values, which are assumed to
have a survival probability of at least 95 %, calculated from a mean value of a two-
sided 75 % confidence level, unless otherwise stated (see 3.7).
3.2 FATIGUE RESISTANCE OF CLASSIFIED STRUC-
TURAL DETAILS
The fatigue assessment of classified structural details and welded joints is based on the
nominal stress range. The (nominal) stress range should be within the limits of the
elastic properties of the material. The range of the design values of the stress range
shall not exceed 1.5 fy for nominal normal stresses or 1.5 f/,./3 for nominal shear
stresses.
In most cases structural details are assessed on the basis of the maximum principal
stress range in the section where potential fatigue cracking is considered. However,
guidance is also given for the assessment of shear loaded details, based on the maxi-
mum shear stress range. Separate S-N curves are provided for consideration of normal
or shear stress ranges, as illustrated in figures (3.2)-1 and (3.2)-2 respectively.
page 34

Care must be taken to ensure that the stress used for the fatigue assessment is the same
as that given in the tables of the classified structural details. Macrogeometric stress
concentrations not covered by the structural detail of the joint itself, e.g. large cutouts
in the vicinity of the joint, have to be accounted for by the use of a detailed stress
analysis, e.g. finite element analysis, or appropriate stress concentration factors (see
2.2.2).
The fatigue curves are based on representative experimental investigations and thus
include the effects of:
structural stress concentrations due to the detail shown
local stress concentrations due to the weld geometry
weld imperfections consistent with normal fabrication standards
stress direction
welding residual stresses
metallurgical conditions
welding process (fusion welding, unless otherwise stated)
inspection procedure (NDT), if specified
postweld treatment, if specified
Furthermore, within the limits imposed by static strength considerations, the fatigue
curves of welded joints are independent of the tensile strength of the material.
Each fatigue strength curve is identified by the characteristic fatigue strength of the
detail at 2 million cycles. This value is the fatigue class (FAT).
The slope of the fatigue strength curves for details assessed on the basis of normal
stresses (fig. (3.2)-1) is m=3.00. The constant amplitude fatigue limit is S· 106 cycles.
The slope of the fatigue strength curves for detailes assessed on the basis of shear
stresses (fig. (3.2)-2) is m=S.OO, but in this case the fatigue limit corresponds to an
endurance of 108 cycles.
The descriptions of the structural details only partially include information about the
weld size, shape and quality. The data refer to a standard quality as given in codes and
standard welding procedures. For higher or lower qualities, modifications may be
necessary as given in 3.5 and 3.8 . All butt welds shall be full penetration welds
without lack of fusion, unless otherwise stated.
All S-N curves of details are limited by the material S-N curve, which may vary due
to different strengths of the materials.
Disregarding major weld defects, fatigue cracks originate from the weld toe, and then
propagate through the base material, or from the weld root, and then propagate
through the weld throat. For potential toe cracks, the nominal stress in the base
page 35

log 11G ., limit by material S-N curve
.................
slope m - 3.00
186 286
FAT
Class Constant amplitude
fatigue limit
5e6 187 log N
Fig. (3.2)-1: Fatigue resistance S-N curves for m=3.00, normal stress (steel)
log A't
slope m-5
FAT
Class
286
fatigue limit
1e8 N cycles
Fig. (3.2)-2 Fatigue resistance S-N curves for shear stress (steel)
material has to be calculated and compared with the fatigue resistance given in the
tables. For potential root cracks, the nominal stress in the weld throat has to be
page 36

calculated. Ifboth failure modes are possible, e.g. at cruciform joints with fillet welds,
both potential failure modes have to be assessed.
3.2.1 Steel
The fatigue resistance values given below refer to welded joints in the as welded
condition unless otherwise stated. The effects of welding residual stress and axial
misalignment up to e/t=O.I (see 3.8.2) are also included.
NDT indicates that the weld must be inspected using appropriate methods to ensure
that it does not contain any significant imperfections. Arrows indicate the loading
direction.
Tab. {3.2}-I: Fatigue resistance values for structural details in steel assessed on the
basis of normal stresses.
No. Structural Detail Description FAT
(Structural steel)
I100 IUnwelded parts of a component I
111
~
Rolled and extruded products 160
~
1) Plates and flats
2) Rolled sections
-=::0 ~
3) Seamless hollow sections
m = 5
For high strength steels a higher
FAT class may be used if veri-
fied by test.
No fatigue resistance of a detail
to be higher at any number of
cycles!
121 Machine gas cut or sheared ma- 140
~ terial with no drag lines, cor-
/~
ners removed, no cracks by
inspection, no visible imper-
I /' ~IIUIl.l fections
,
m=3
page 37

No. StructuralI>etail Description FAT
(Structural steel)
122 Machine thermally cut edges, 125
7* corners removed, no cracks by
/~
inspection
I /' mllll1. m = 3
,
123 Manually thermally cut edges, 100
7* free from cracks and severe
/~
notches
I /' mU1l1. m = 3
,
124 Manually thermally cut edges, 80
7* uncontrolled, no notch deeper
/~
than.5 mm
m = 3
I /' IWIIIII.l
,
1 200 IButt welds, transverse loaded
I
211 Transverse loaded butt weld (X- 125
groove or V-groove) ground
flush to plate, 100% NDT
.-~~-+
212 Transverse butt weld made in 100
shop in flat position,
.-~--
toe angle < 300 , NDT
page 38

(Structural steel)
213 Transverse butt weld not satis- 80
fying conditions of 212, NDT
--~~~--.
214 Transverse butt weld, welded
on ceramic backing, root crack 80
-~~
215 Transverse butt weld on per- 71
manent backing bar
.-~~--
216
./'
Transverse butt welds welded
from one side without backing
bar, full penetration
root controlled by NDT 71
/ noNDT 45
page 39

(Structural steel)
217 Transverse partial penetration 45
butt weld, analysis based on
.-~-.
stress in weld throat sectional
area, weld overfill not to be
taken into account.
The detail is not recommended
for fatigue loaded members.
It is recommended to verify by
fracture mechanics (3.8.5.2)!
221 Slope Transverse butt weld ground
u:= 1-
flush, NDT, with transition in
thickness and width
slope 1:5 125
Slope
slope 1:3 100
slope 1:2 80
--- )-+-1 :z
For misalignement see 3.8.2
222 Transverse butt weld made in
--f Ie t- shop, welded in flat position,
weld profile controlled, NDT,
with transition in thickness and
+- width:
r::=-+ slope 1:5 100
slope 1:3 90
slope 1:2 80
For misalignment see 3.8.2
223 Slope Transverse butt weld, NDT,
-c:e=J- with transition on thickness and
width
slope 1:5 80
Slope slope 1:3 71
-~ 1- slope 1:2 63
page 40

(Structural steel)
224 Transverse butt weld, different 71
thicknesses without transition,
~~-
centres aligned.
In cases, where weld profile is
equivalent to a moderate slope
transition, see no. 222
225 Three plate connection, root 71
crack
-~T-
226 Transverse butt weld flange 112
~ ~
splice in built-up section welded
r
prior to the assembly, ground
/~,~b--=:1l (r~b)
flush, with radius transition,
NDT
231 /' Transverse butt weld splice in 80
rolled section or bar besides
flats, ground flush, NDT
~
232 Transverse butt weld splice in
I~I- ---lO
circular hollow section, welded
from one side, full penetration,
I " I root inspected by NDT 71
noNDT 45
page 41

(Structural steel)
233 Tubular joint with permanent 71
backing
~[O::1}
234 Transverse butt weld splice in
1-~-L--lD
rectangular hollow section,
welded from one side, full
penetration,
I
" I root inspected by NDT 56
noNDT 45
241 Transverse butt weld ground 125
edges flush, weld ends and radius
ground ground, 100% NDT at crossing
- - flanges, radius transition.
shop at flat position, weld pro-
file controlled, NDT, at cros-
- - sing flanges, radius transition
page 42

(Structural steel)
243 Transverse butt weld ground 80
r--- ground flush, NDT, at crossing flanges
.,/ ~/ with welded triangular tran-
--I J-- sition plates, weld ends ground.
~ V
Crack starting at butt weld.
----
244 Transverse butt weld, NDT, at 71
~
ground crossing flanges, with welded
L
~-
triangular transition plates, weld
---I ends ground.
""'- Crack starting at butt weld.
-
245 Transverse butt weld at crossing 50r--
flanges.
Crack starting at butt weld.
-I 1-
-----
300 Longitudinal load-carrying welds
311 Automatic longitudinal seam 125
welds without stop/start posi-
~ ~
tions in hollow sections
with stop/start positions 90
page 43

No. Structural Detail
(Structural steel)
312
313
321
322
323
Description FAT
Longitudinal butt weld, both 125
sides ground flush parallel to
load direction, 100% NDT
Longitudinal butt weld, without 125
stop/start positions, NDT
Continuous automatic lon- 125
gitudinal fully penetrated K-butt
weld without stop/start positions
(based on stress range in flange)
NDT
gitudinal double sided fillet
Continuous manual longitudinal 90
fillet or butt weld (based on
stress range in flange)
page 44

(Structural steel)
324
325
Description FAT
Intermittent longitudinal fillet
weld (based on normal stress in
flange u and shear stress in web
T at weld ends).
T/U = 0 80
0.0 - 0.2 71
0.2 - 0.3 63
0.3 - 0.4 56
0.4 - 0.5 50
0.5 - 0.6 45
0.6 - 0.7 40
> 0.7 36
Longitudinal butt weld, fillet
weld or intermittent weld with
cope holes (based on normal
stress in flange u and shear
stress in web T at weld ends),
cope holes not higher than 40%
of web.
T/U = 0 71
0.0 - 0.2 63
0.2 - 0.3 56
0.3 - 0.4 50
0.4 - 0.5 45
0.5 - 0.6 40
> 0.6 36
page 45

No.
331
332
Structural Detail
(Structural steel)
-- -
Description
Joint at stiffened knuckle of a
flange to be assessed according
to no. 411 - 414, depending on
type of joint.
Stress in stiffener plate:
A
a = a' f ·2·sin«
f LAst
Af = area of flange
ASt = area of stiffener
Stress in weld throat:
Aw = area of weld
throat
Unstiffened curved flange to
web joint, to be assessed accor-
ding to no. 411 - 414, depen-
ding on type of joint.
Stress in web plate:
F
a =-L
r·t
F
C1 = f
W T"La
Ffaxial force in flange
t thickness of web plate
a weld throat
page 46
FAT

(Structural steel)
1 400 ICruciform joints and/or T-joints I
411 Cruciform joint or T-joint, K- 80
t e butt welds, full penetration, no
~
lamellar tearing, misalignment
e<0.15·t, weld toes ground,
toe crack
tl ~ e+
butt welds, full penetration, no
lamellar tearing, misalignmentt"::
-v////.....:~ ~vj'// /1- e<0.15·t, toe crack
+ ~
~
413 Cruciform joint or T-joint, fillet 63
tl ~ el
welds or partial penetration K-
~ butt welds, no lamellar tearing,
v/ / / / / :"-. f"/A// / A- misalignment e< 0.15·t,
+ ~~"" toe crack
~
414
(~ ~
Cruciform joint or T-joint, fillet 45
welds or partial penetration K-
L. ..1111~~ butt welds including toe ground
0)0<~t0"'/L joints,
~~~ weld root crack.
~ Analysis based on stress in weld
throat.
page 47

(Structural steel)
421
422
t:~-U:-:-I 0d]q
423
I~~::-J 0
r=it=J
424
t~U::ml [JJ
~
425
t-~[~ __I[]
~
Description FAT
Splice of rolled section with 45
intermediate plate, fillet welds,
weld root crack.
Analysis base on stress in weld
throat.
Splice of circular hollow section
with intermediate plate, single-
. sided butt weld, toe crack
wall thickness > 8 mm
wall thickness < 8 mm
Splice of circular hollow section
with intermediate plate, fillet
weld, root crack. Analysis
based on stress in weld throat.
Splice of rectangular hollow
section, single-sided butt weld,
toe crack
Splice of rectangular hollow
section with intermediate plate,
fillet welds, root crack
page 48
56
50
45
40
50
45
40
36

(Structural steel)
431
I500 INon-load-carrying attachments
511
512
513
Description
Weld connecting web and flan-
ge, loaded by a concentrated
force in web plane perpendicu-
lar to weld. Force distributed
on width b =2·h + 50 IDIn.
Assessment according to no.
411 - 414. A local bending due
to eccentric load should be
considered.
Transverse non-load-carrying
attachment, not thicker than
main plate
FAT
K-butt weld, toe ground 100
Two-sided fillets, toe 100
ground
Fillet weld(s), as welded 80
Thicker than main plate 71
Transverse stiffener welded on
girder web or flange, not thik-
ker than main plate. For weld
ends on web principle stress to
be used
Two-sided fillets, toe
ground 100
Fillet weld(s): as welded 80
thicker than main plate 71
Non-Ioadcarrying stud
as welded 80
page 49

No.
514
515
521
522
523
Structural Detail
(Structural steel)
I
i I I
I _/1t-rfull pene~tion
.. weld
_. - -
f~~lat ~d ~
/ ~.
i(ff/::~ 1
HI Hr
_ _ HI
-
• :-. t
f
t~9- (t)
.L~'- __ _ ,
r"......... u
--.-t
1.
Description FAT
Trapezoidal stiffener to deck 71
plate, full penetration butt weld,
calculated on basis of stiffener
thickness, out of plane bending
plate, fillet or partial penetra-
tion weld, calculated on basis of
stiffener thickness and weld
throat, whichever is smaller
Longitudinal fillet welded gus-
set at length I
I < 50 mm
I < 150 mm
I < 300 mm
1 > 300 mm
gusset near edge: see 525 "flat
side gusset"
set with radius transition, end
of fillet weld reinforced and
ground, c < 2 t, max 25 mm
r > 150 mm
set with smooth transition
(sniped end or radius) welded
on beam flange or plate.
c < 2 t, max 25 mm
r > 0.5 h
r < 0.5 h or cp < 20°
page 50
80
71
63
50
90
71
63

(Structural steel)
524 Longitudinal flat side gusset
r r welded on plate edge or beam
t~ • flange edge, with smooth tran-
- -- - sition (sniped end or radius).<J
..... I t =-T c < 2~, max. 25 mm(t.].)
.II:
....i. r > 0.5 h 50
r < 0.5 h or q; < 20° 45
For ~ < 0.7 t1, FAT rises 12%
welded on plate or beam flange
~-~
edge, gusset length I:
i..a..~~
I < 150 mm 50
I < 300 mm 45
1 > 300 mm 40
~
welded on edge of plate or
w beam flange, radius transition
~~
ground.
r> 150 or r/w > 113 90
-- ~ ~
116 < r/w < 113 71
r/w < 116 50
531 Circular or rectangular hollow 71
section, fillet welded to another
:rrc,=fJ ~
section. Section width parallel
- -- -- - to stress direction < 100 mm,,..,
I I I I
I I I I
else like longitudinal attachmentI I I I I I
I I I I ,
I ~ ],00 IIII1l
1600 1 Lap joints I
611 Transverse loaded lap joint with
fillet welds
Fatigue of parent metal 63
......VZ~~7I--+ Fatigue of weld throat 45
Stress ratio must be 0 <R < 1 !
page 51

(Structural steel)
612
F
+ C:~:
0"=-
A
r
613
~~~P1~~
t===:=::;1 --L.
700 Reinforcements
711
t. tD~
712
Description
Longitudinally loaded lap joint
with side fillet welds
Fatigue of parent metal
Fatigue of weld (calc. on
max. weld length of 40
times the throat of the
weld
Lap joint gusset, fillet welded,
non-load-carrying, with smooth
transition (sniped end with
cp<20° or radius), welded to
loaded element c<2·t,
max 25 mm
FAT
50
50
to flat bar 63
to bulb section 56
to angle section 50
End of long doubling plate on 1-
beam, welded ends (based on
stress range in flange at weld
toe)
to < 0.8 t 56
0.8 t < to < 1.5 t 50
to > 1.5 t 45
End of long doubling plate on
beam, reinforced welded ends
ground (based on stress range in
flange at weld toe)
to < 0.8 t 71
0.8 t < to < 1.5 t 63
to > 1.5 t 56
page 52

(Structural steel)
721
~
~~~
~EJ
731 ground
~ Dum
I800 IFlanges, branches and nozzles
811
812
I
821
Description FAT
End of reinforcement plate on
rectangular hollow section.
wall thickness:
t < 25 mm 50
Reinforcements welded on with
fillet welds, toe ground 80
Toe as welded 71
Analysis based on modified
nominal stress
Stiff block flange, full penetra- 71
tion weld
Stiff block flange, partial pene-
tration or fillet weld
toe crack in plate 63
root crack in weld throat 45
Flat flange with almost full
penetration butt welds, modified
nominal stress in pipe, toe
crack
71
page 53

(Structural steel)
822
~... I
~~ !
1,
831
f
t'-
I
" I
~~ i
832
841
..I
~
~
~
~~
~-
Description
Flat flange with fillet welds,
modified nominal stress in pipe,
toe crack.
Tubular branch or pipe penetra-
ting a plate, K-butt welds.
If diameter> 50 mm, stress
concentration of cutout has to
be considered
ting a plate, fillet welds.
be considered
Nozzle welded on plate, root
pass removed by drilling.
If diameter > 50 mm, stress
be considered
page 54
FAT
63
80
71
71

(Structural steel)
842
!
I
i
I
I
1900 1Tubular joints
911
912
t
~~,
913
I
t t
I
I~ i ~~~ ~
, ,
921
~
-:~V
~
¥
Description
Nozzle welded on pipe, root
pass as welded.
be considered
FAT
63
Circular hollow section butt 63
joint to massive bar, as welded
Circular hollow section welded 63
to component with single side
butt weld, backing provided.
Root crack.
to component single sided butt
weld or double fillet welds.
Root crack.
Circular hollow section with
welded on disk
Fillet weld, toe ground 90
Fillet welds, as welded 71
page 55

(Structural steel)
931
f-]
Tube-plate joint, tubes flatten- 71
l- ed, butt weld (X-groove)
Tube diameter < 200 mm
~4-> and
1-. plate thickness < 20 mm
932 Tube-plate joint, tube slitted
1-) So- and welded to plate
tube diameter < 200 mm and 63
:::CO
plate thickness < 20 mm
I
tube diameter > 200 mm or 45
plate thickness > 20 mm
Tab. {3.2}-2: Fatigue resistance values for structural details in steel assessed on the
basis of shear stresses.
Structural detail FAT log C for m=5 stress range at fati-
class gue limit [N/mm2]
Parent metal, full 100 16.301 46
penetration butt welds
Fillet welds, partial 80 15.816 36
page 56

3.2.2 Aluminium
The fatigue resistance values given below refer to welded joints in the as-welded
condition unless otherwise stated. Effects of welding residual stress and axial misalign-
ment up to e/t=O.l (see 3.8.2) are also included.
NDT indicates that the weld must be inspected using appropriate methods to ensure
that it does not contain any significant imperfections.
Arrows indicate the loading direction. All slopes are m=3.00 if not stated otherwise.
The grid of the S-N curves is given in fig. (3.2)-3 for normal stress and in fig. (3.2.)-
4 for shear stress.
log t.a II limit by material 8-N curve
...........
'.
slope m - 3.00
186 286
FAT
Class
'.........
Constant amplitude
fatigue limit
•..•..•..•.....1••••••••••••••••••••••
I
5e6 187 log N
Fig. (3.2)-3 Fatigue resistance curves for aluminium (normal stress)
page 57

log 11 't
slope m-5
FAT
Class
286
fatigue limit
1e8 N cycles
Fig. (3.2)-4 Fatigue resistance curves for aluminium (shear stress)
Tab. {3.2}-3: Fatigue resistance values for structural details in aluminium alloys
assessed on the basis of normal stress.
(Structural aluminium alloys)
I100 IUnwelded parts of a component I
111
~~
Rolled and extruded products or
components with edges machi-
ned, m=5
-::0 ~ AA 5000/6000 alloys 71
AA 7000 alloys 80
No fatigue resistance of a detail
to be higher at any number of
cycles!
page 58

122 Machine thermally cut edges, 40
comers removed, no cracks by
~
inspection
m = 3.0
200 Butt welds, transverse loaded
211 Transverse loaded butt weld (X- 50
groove or V-groove) ground
flush to plate, 100% NDT
....~~-.
shop in flat position,
.-~--
toe angle < 300 , NDT
213 Transverse butt weld, 32
toe angle < 500
.-~--
215 Transverse butt weld, 25
toe angle > 500
, or
~~~--
transverse butt weld on per-
manent backing bar
page 59

216 Transverse butt welds welded
./' from one side without backing
bar, full penetration
root controlled by NDT 28
/ noNDT 18
221 Slope Transverse butt weld ground
-cC 1-
flush, NDT, with transition in
thickness and width
slope 1:5 40
Slope
slope 1:3 32
slope 1:2 25
--- I-+-1 X
For misalignement see 3.8.2
222 Transverse butt weld made in
Ie j- shop, welded in flat position,
-+ weld profJ.le controlled, NDT,
with transition in thickness and
+-
width:
:----+ slope 1:5 32
slope 1:3 28
slope 1:2 25
223 Slope Transverse butt weld, NDT,
-a:J- with transition on thickness and
width
slope 1:5 25
Slope slope 1:3 22
--- 1- slope 1:2 20
+-1 :z
page 60

224 Transverse butt weld, different 22
thicknesses without transition,
.-~-
centres aligned.
In cases, where weld profile is
equivalent to a moderate slope
transition, see no. 222
225 Three plate connection, root 22
crack
-~T'+-
226 Transverse butt weld flange 45
r~. ~
splice in built-up section welded
~~,
prior to the assembly, ground
~b-::::::ll (ra})}
flush, with radius transition,
NDT
1300 1Longitudinalload-carrying welds I
311 Automatic longitudinal seam 50
welds without stop/start posi-
~ ~
tions in hollow sections
312 Longitudinal butt weld, both 50
~
sides ground flush parallel to
load direction, 100% NDT
,..
page 61

313
321
322
323
Description FAT
Longitudinal butt weld, without 45
stop/start positions, NDT
gitudinal fully penetrated K-butt
NDT
gitudinal double sided fillet
Continuous manual longitudinal 36
fillet or butt weld (based on
stress range in flange)
page 62

324
325
Description FAT
Intermittent longitudinal fillet
weld (based on normal stress in
flange (f and shear stress in web
T at weld ends).
Tier = 0 32
0.0 - 0.2 28
0.2 - 0.3 25
0.3 - 0.4 22
0.4 - 0.5 20
0.5 - 0.6 18
0.6 - 0.7 16
> 0.7 14
Longitudinal butt weld, fillet
weld or intermittent weld with
cope holes (based on normal
stress in flange (J and shear
stress in web T at weld ends),
cope holes not higher than 40%
of web.
Tier = 0 28
0.0 - 0.2 25
0.2 - 0.3 22
0.3 - 0.4 20
0.4 - 0.5 18
0.5 - 0.6 16
> 0.6 14
page 63

332
~./ r F
~I f
) , ~"'''!',T771'J7( ..
cr "'+((t) })
-
Description
Joint at stiffened knuckle of a
flange to be assessed according
to no. 411 - 414, depending on
type of joint.
Stress in stiffener plate:
A
(J = (J • f'2 .sinIX
f LAst
Ar = area of flange
ASt = area of stiffener
Aw = area of weld
throat
FAT
Unstiffened curved flange to ---
web joint, to be assessed accor-
ding to no. 411 - 414, depen-
ding on type of joint.
Stress in web plate:
F
(J = ~
r·t
F
(J = f
W r"La
Ffaxial force in flange
t thickness of web plate
a weld throat
page 64

1 400 ICruciform joints and/or T-joints I
t! ~ e!
lamellar tearing, misalignment
t:
e<O.15·t, weld toes ground,-v/////. l'...: ~VK///
+ ~ toe crack
~
t~ ~ e+
lamellar tearing, misalignment~
-v////.;:: ~ ~VK///j- e<O.15·t, toe crack
• ~t::::
413 Cruciform joint or T-joint, fillet 22
t~ ~ 1
welds, or partial penetrating K-
~ butt weld,
V/ / / / /'~r//V//A- misalignment e<O.15·t,
, "II~'" toe crack
;::;
414
~~ ~
Cruciform joint or T-joint, fillet 16
welds or partial penetrating K-
L .....~h.. butt welds (including toe ground
050<~~ffij- welds),
"..~,.. weld root crack.
~ Analysis based on stress in weld
throat.
page 65

1500 INon-load-carrying attachments
511
512
I)
513
514
Description
Transverse non-load-carrying
attachment, not thicker than
main plate
FAT
Two-sided fillets, toe 36
ground
Fillet weld(s), as welded 28
Thicker than main plate 25
Transverse stiffener welded on
girder web or flange, not thik-
ker than main plate. For weld
ends on web principle stress to
be used
Two-sided fillets, toe
ground 36
Fillet weld(s): as welded 28
thicker than main plate 25
Non-Ioadcarrying stud
as welded 28
plate, full penetration butt weld,
calculated on basis of stiffener
thickness, out of plane bending
page 66

515
521
-
522
-- t
"f
523
~ (t)
I~.L~'''' __ _ ,
... 0
-- t
"f
524
t~L~---- -+
..- I t ;"""T
(t1)
.r:
-±..
Description FAT
plate, fillet or partial penetra-
tion weld, calculated on basis of
stiffener thickness and weld
throat, whichever is smaller
set at length I
I < 50 mm
I < 150 mm
I < 300 mm
I > 300 mm
gusset near edge: see 525 "flat
side gusset"
set with radius transition, end
of fillet weld reinforced and
ground, c < 2 t, max 25 mm
r > 150 mm
set with smooth transition
(sniped end or radius) welded
on beam flange or plate.
c < 2 t, max 25 mm
r > 0.5 h
r < 0.5 h or cp < 20°
Longitudinal flat side gusset
welded on plate edge or beam
flange edge, with smooth tran-
sition (sniped end or radius).
c < 2t2, max. 25 mm
r > 0.5 h
r < 0.5 h or cp < 20°
For t2 < 0.7 t}, FAT rises 12%
page 67
28
25
20
18
32
25
20
18
16

welded on plate or beam flange
~
edge, gusset length I:
1 < 150 mm 18
1 < 300 mm 16
I > 300 mm 14
-.:::.. welded on edge of plate or
~ beam flange, radius transition
~"1IC
ground. 36
r> 150 or r/w > 113 28-- ~ ~
r 116 < r/w < 113 22
r/w < 116
600 Lap joints
611 Transverse loaded lap joint with
fillet welds
Fatigue of parent metal 22
+-v~;?~!-+
Fatigue of weld throat 16
Stress ratio must be 0 <R <1 !
612 Longitudinally loaded lap joint
F with side fillet welds(]'::-
i [:~:
A Fatigue of parent metal 18
r Fatigue of weld (calc. on
max. weld length of 40 18
times the throat of the
weld
page 68

613
9-
+
"';'~m;J~~~
~==;;;:J-1..
I700 IReinforcements
711
t, to~
1; (""til"")"
712
721
Description
Lap joint gusset, fillet welded,
non-load-carrying, with smooth
transition (sniped end with
~<20° or radius), welded to
loaded element c < 2·t,
max 25 mm
FAT
to flat bar 22
to bulb section 20
to angle section 18
End of long doubling plate on 1-
beam, welded ends (based on
stress range in flange at weld
toe)
to < 0.8 t 20
0.8 t < to < 1.5 t 18
to > 1.5 t 16
End of long doubling plate on
beam, reinforced welded ends
ground (based on stress range
in flange at weld toe)
to ::;; 0.8 t
0.8 t < to < 1.5 t
to > 1.5 t
End of reinforcement plate on
rectangular hollow section.
wall thickness:
t < 25 mm
28
25
22
20
page 69

731 ground Reinforcements welded on with
~ ~ fillet welds, toe ground 32
Toe as welded 25
..- @ ---+- Analysis based on modified
nominal stress
1800 IFlanges, branches and nozzles I
811 Stiff block flange, full penetra- 25
tion weld
i W~
812 Stiff block flange, partial pene-
tration or fillet weld
I
a~~.. toe crack in plate 22
I root crack in weld throat 16
I
821 FIat flange with almost full 25
~
penetration butt welds, modified
nominal stress in pipe, toe
~~ !
~
crack
I
,
822 FIat flange with fillet welds, 22
~II..
modified nominal stress in pipe,
toe crack.
~~
,
~I,
page 70

831
r;;: I
~
~
I
t'
" ~~-S~
i~
832
v Tv
v Iv
-k"'-."'-."'~ ~ ~"'-.""-"'IIIIIr..:y
841
- ~~
I
~
~
I ~
-~)
:s
~-
842
I
~I ~
i ~..
I
."""""""'-I
1900 1Tubular joints
911
Description
ting a plate, K-butt welds.
be considered
ting a plate, fillet welds.
be considered
Nozzle welded on plate, root
pass removed by drilling.
be considered
Nozzle welded on pipe, root
pass as welded.
be considered
Circular hollow section butt
joint to massive bar, as welded
page 71
FAT
28
25
25
22
22

912
913
921
t
I
t
t
~ ~ ~
~
~
to component with single side
butt weld, backing provided.
Root crack.
Circular hollow section welded
to component single sided butt
weld or double fillet welds.
Root crack.
Circular hollow section with
welded on disk
18
Fillet weld, toe ground 32
Fillet welds, as welded 25
Tab. {3.2}-2: Fatigue resistance values structural details in aluminium alloys assessed
on the basis of shear stress.
Structural detail FAT log C for m=5 stress range at fati-
class gue limit [N/mm2]
Parent metal, full 36 14.083 16.5
Fillet welds, partial 28 13.537 12.8
page 72

3.3 FATIGUE RESISTANCE AGAINST GEOMETRIC
STRESS (HOT SPOT STRESS)
3.3.1 Fatigue Resistance using Reference S-N Curve
The S-N curves for fatigue resistance against structural geometric stress (2.2.3) are
given in the table {3.3}-1 for steel, where the definition of the FAT class is given in
chapter 3.2. The resistance values refer to the as-welded condition unless stated
otherwise. The effects of welding residual stress are included.
The design value of the geometric stress range shall not exceed 2·ry •
3.3.1.1 Steel
Tab. {3.3}-1: Fatigue resistance against geometric stress
INo. IDescription IFAT I
1 Flat butt welds, full penetration, with a possible
misalignment according to notch cases 211-213 (Tab.
3.2-1)
2 ftllet welds at toe,
toe ground 112
toe as welded 100
m=3
3 Cruciform joints with a possible misalignment, not
yet accounted for in determination of geometric
stress, to be assessed according to notch cases 411-
413 in (Tab. 3.2-1)
3.3.1.2 Aluminium
At present, no commonly accepted data for the resistance of aluminium alloys against
geometric stress are available. Therefore, the reference detail method outlined in 3.3.2
is recommended.
3.3.2 Fatigue Resistance Using a Reference Detail
The tables of the fatigue resistance of structural details given in 3.2, or fatigue data
from other sources which refer to a comparable detail may, be used. The reference
page 73

detail should be chosen as similar as possible to the detail to be assessed. Thus the
procedure will be:
a) Select a reference detail with known fatigue resistance, which is as similar as
possible to the detail being assessed with respect to geometric and loading
parameters.
b) Identify the type of stress in which the fatigue resistance is expressed. This is
usually nominal stress (as in tables in chapter 3.2).
c) Establish a FEM model of the reference detail and the detail to be assessed
with the same type of meshing and elements following the recommendations
given in 2.2.3.
d) Load the reference detail and the detail to be assessed with the stress identified
in b).
e) Determine the geometric stress ugeo,rd' of the reference detail and the geometric
stress ugeo,assess of the detail to be assessed.
t) The fatigue resistance for 2 million cyles of the detail to be assessed FATassess
is then calculated from fatigue class of the reference detail FATref by:
(J
FAT = geo,Te/. FAT
aste&S' (J ref
geo,assus
page 74

3.4 FATIGUE RESISTANCE AGAINST EFFECTIVE
NOTCH STRESS
3.4.1 Steel
The effective notch stress fatigue resistance against fatigue actions, as determined in
2.2.4 for steel, is given in table {3.4}-1. The defInition of the FAT class is given in
chapter 3.2. The fatigue resistance value refers to the as-welded condition. The effect
of welding residual stresses is included. Possible misalignment is not included.
Tab. {3.4}-1: Effective notch fatigue resistance for steel
No. Quality of weld notch Description FAT
I Effective notch radius Notch as-welded, normal 225
equalling 1 mm replacing welding quality
weld toe and weld root
notch m=3
3.4.2 Aluminium
At present, no commonly accepted data can be given.
page 75

3.5 FATIGUE STRENGTH MODIFICATIONS
3.5.1 Stress Ratio
3.5.1.1 Steel
For stress ratios R <0.5 a fatigue enhancement factor f(R) may be considered by
multiplying the fatigue class of classified details by f(R). The fatigue enhancement
factor depends on the level and direction of residual stresses. The following cases are
to be distinguished:
I: Base material and wrought products with negligible residual stresses « 0.2·fy),
stress relieved welded components, in which the effects of constraints or
secondary stresses have been considered in analysis.
f(R) = 1.6
f(R) = -0.4 • R + 1.2
f(R) = I
for R < -1
for -1 S R < 0.5
for R > 0.5
II: Small scale thin-walled simple structural elements containing short welds. Parts
or components containing thermally cut edges.
f(R) = 1.3
f(R) = -0.4 • R + 0.9
f(R) = 1
for R < -1
for -1 < R < -0.25
for R > -0.25
ill: Complex two- or three-dimensional components, components with global
residual stresses, thickwalled components.
f(R) = 1 no enhancement
The ranking in categories I, II or III should be done and documented by the design
office. If no reliable information on residual stress is available, f(R) =1.
It has to be noted in this respect that stress relief in welded joints is unlikely to be
fully effective, and long range residual stresses may be introduced during assembly of
prefabricated welded components. For such reasons, it is recommended that values of
f(R) >1 should only be adopted for welded components in very special circumstances.
3.5.1.2 Aluminium
The same regulations as for steel are recommended.
page 76

Factor f(R)
1.61'C""'"----.:....;-----,-----,.------,---,------,
1.5i-----""k:----+----+----t---t-----i
1,41-----+---~---+-----i---+------l
1.3-1c----t----+----"";c---t----+----i
1.21-----"......"..---+---_+--~o;:----+_-____l
1.1 i-----t---""Ioo;::----+----t------"''Ioc-----i
-0.75 -0.5 -0.25
Stress ratio R
o 0.25 0.5
-I: low resld. stress ~ II: medium res. str. .....111: high resid. str
Fig. (3.5)-1 Enhancement factor f(R)
3.5.2 Wall Thickness
3.5.2.1 Steel
The influence of plate thickness on fatigue strength should be taken into account in
cases where cracks start from the weld toe on plates thicker than 25 mm. The reduced
strength is taken in consideration by multiplying the fatigue class of the structural
detail by the thickness reduction factor f(t). The thickness correction exponent n is
dependent on the effective thickness terrand the joint category (see table {3.5}-1) [21].
Tab. {3.5}-1: Thickness correction exponents
Joint category Condition n
Cruciform joints, transverse T-joints, as-welded 0.3
plates with transverse attachments
Cruciform joints, transverse T-joints, toe ground 0.2
plates with transverse attachments
Transverse butt welds as-welded 0.2
Butt welds ground flush, base material, any 0.1
longitudinal welds or attachements
The plate thickness correction factor is not required in the case of assessment based on
effective notch stress procedure or fracture mechanics.
page 77

j(t)
= (;;r where t>2Smm
~tOJ
~JIIf LIt ~ 2 then tef! = O.S·L ~
toe""""","L
else tef! = t
Fig. (3.5)-2 Toe distance
3.5.2.2 Aluminium
The same regulations as for steel are recommended.
3.5.3 Improvement Techniques
Post weld improvement techniques may raise the fatigue resistance. These techniques
improve the weld profile, the residual stress conditions or the environmental conditions
of the welded joint. The improvements methods are:
a) Methods of improvement of weld profile:
Machining or grinding of weld seam flush to surface
Machining or grinding of the weld transition at the toe
Remelting of the weld toe by TIG-, plasma or laser dressing
b) Methods for improvement of residual stress conditions:
Peening (hammer-, needle-, shot- or brush-peening)
Coining
Overstressing
Stress relieving thermal treatment
c) Methods for improvement of environmental conditions:
Painting
Resin coating
The effects of all improvement techniques are sensitive to the method of application
and the applied loading, being most effective in the low stress / high cycle regime.
They may also depend on the material, structural detail and dimensions of the welded
joint. Consequently, fatigue tests for the verification of the procedure in the endurance
range of interest are recommended (chapters 3.7 and 4.5).
page 78

3.5.4 Effect of Elevated Temperatures
3.5.4.1 Steel
For higher temperatures, the fatigue resistance data may be modified with a reduction
factor given in fig. (3.5)-3. The fatigue reduction factor is a conservative approach and
might be raised according to test evidence or application codes.
Reduction factor
1 ..............
~...........
~
r"-.
","~
~"-
"
0,9
0,8
0,7
0,6
0,5
0,4
100 150 200 250 300 350 400 450 500 550 600
Temperature T [deg Celsius]
Fig. (3.5)-3 Fatigue strength reduction factor for steel at elevated temperatures
3.5.4.2 Aluminium
The fatigue data given here refer to operation temperatures lower than 70°C. This
value is a conservative approach. It may be raised according to test evidence or an ap-
plicable code.
3.5.5 Effect of Corrosion
The fatigue resistance data given here refer to non-corrosive environments. Normal
protection against atmospheric corrosion is assumed.
A corrosive environment or unprotected exposure to atmospheric conditions may re-
duce the fatigue class. The fatigue limit may also be reduced considerably. The effect
depends on the spectrum of fatigue actions and on the time of exposure.
No specific recommendations are given for corrosion fatigue assessment.
page 79

3.6 FATIGUE RESISTANCE AGAINST CRACK PRO-
PAGATION
The resistance of a material against cyclic crack propagation is characterized by the
material parameters of the "Paris" power law of crack propagation
da
- = C ·I!..Km if I!..K < I!..Kth then
dN 0
da
- = 0
dN
where the material parameters are
Co constant of the power law
m exponent of the power law
AI{ range of cyclic stress intensity factor
~ threshold value of stress intensity
R ratio Kmin/K.ou, taking all stresses including residual stresses into ac-
count (see 3.5.1)
In the absence of specified or measured material parameters, the values given below
are recommended. They are characteristic values.
3.6.1 Steel
Co = 9.5 .10-12
Co = 3.0 '10-13
m = 3
(units in MPav'm and m) or
(units in N*mm-3n and mm)
I!..~ = 6.0 - 4.56'R but not lower than 2
~ = 190 - 144'R but not lower than 62
3.6.2 Aluminium
Co = 2.6 .10-10
Co = 8.1 _10-12
m = 3
(units in MPav'm and m) or
(units in N*mm-312 and mm)
I!..~ = 2.0 - 1.5 .R but not lower than 0.7
I!..Kth = 63 - 48 .R but not lower than 21
page 80
(units in MPaVm) or
(units in N*mm-3n)
(units in MPaVrn) or
(units in N*mm-3n)

3.7 FATIGUE RESISTANCE DETERMINATION BY
TESTING
Fatigue tests may be used to establish a fatigue resistance curve for a component or a
structural detail, or the resistance of a material against (non critical) cyclic crack
propagation.
It is recommended that test results are obtained at constant stress ratios R. The S-N
data should be presented in a graph showing log(endurance in cycles) as the abscissa
and log(range of fatigue actions) as the ordinate. For crack propagation data, the
log(stress intensity factor range) should be the abscissa and the log(crack propagation
rate per cycle) the ordinate.
Experimental fatigue data are scattered, the extent of scatter tends to be greatest in the
low stress/low crack propagation regime (e.g. see fig. (3.7)-1). For statistical evalua-
tion, a Gaussian log-normal distribution should be assumed. The number of failed test
specimens must be equal or greater than
5. For other conditions, special statistical
considerations are required. log .t.o
scatter
scatterband
,
"
mean curve
''', scatter
/ ""~.-""'::"';--"'-....
characterlstlc '"
curve
,
log N
Many methods of statistical evaluation
are available. However, the most com-
mon approach for analysing fatigue data
is to fit S-N or crack propagation curves
by regression analysis, taking log(N) or
log(da/dN) as the dependent variable.
Then, characteristic values are establi-
shed by adopting curves lying approxi-
mately two standard deviations (2 Stdv at Fig. (3.7)-1 Scatterband in SN curve
a greater number of specimens) of the
dependent variable from the mean. In the case of S-N data, this would be below the
mean, while the curve above the mean would be appropriate in case of crack propaga-
tion data.
Thus, more precisely, test results should analysed to produce characteristic values
(subscript k). These are values at a 95 % survival probability in reference to a two-
sided 75 % confidence level of the mean. They are calculated by the following proce-
dure:
a) Calculate 10glO of all data: Stress range .&U and number of cycles N, or
stress intensity factor range .&K and crack propagation rate daldN.
b) Calculate exponents m and constant logC (or logCo resp.) of the for-
mulae:
page 81

for S-N curve logN =logC -m 'logAa
da
for crack propag. log- = logCo - m·logAK
dN
by linear regression taking stress or stress intensity factor range as the
independent variable. If the number of data n <15, or if the data are not
sufficiently evenly distributed to determine m correctly, a fixed value of
m should be taken, as derived from other tests under comparable condi-
tions, e.g. m=3 for welded joints.
c) Calculate mean xm and standard deviation Stdv of logC (or logCo resp.)
using m obtained in b).
d) If Xj are the logs of tentative data, the formulae for the calculation of
the characteristic value Xk will be:
Stdv =
S-N data: xk =x", -k'Stdv
Crack propagation rate: xk = x", +k'Stdv
The values of k are given in table {3.7}-1.
Tab. {3.7}-1: Values of k for the calculation of characteristic values
n 5 10 15 20 25 30 40 50 100
k 3.5 2.7 2.4 2.3 2.2 2.15 2.05 2.0 1.9
For more details and information, see appendix 6.4.1 and ref. [35].
In case of S-N data, proper account should be taken of the fact that residual stresses
are usually low in small-scale specimens. The results should be corrected to allow for
the greater effects of residual stresses in real components and structures. This may be
achieved either by testing at high R-ratios, e.g. R=O.5, or by testing at R=O and
lowering the fatigue strength at 2 million cycles by 20% .
page 82

3.8 FATIGUE RESISTANCE OF JOINTS WITH WELD
IMPERFECTIONS
3.8.1 General
3.8.1.1 Types of Imperfections
The types of imperfections covered in this document are listed below. Other imperfec-
tions, not yet covered, may be assessed by assuming similar imperfections with com-
parable notch effect.
Imperfect shape
All types of misalignment including centre-line mismatch (linear misalignment)
and angular misalignment (angular distortions, roofing, peaking).
Undercut
Volumetric discontinuities
Gas pores and cavities of any shape.
Solid inclusions, such as isolated slag, slag lines, flux, oxides and metallic
inclusions.
Planar discontinuities
All types of cracks or cracklike imperfections, such as lack of fusion or lack of
penetration (Note that for certain structural details intentional lack of penetra-
tion is already covered, e.g. at partial penetration butt welds or cruciform
joints with fillet welds)
If a volumetric discontinuity is surface breaking or near the surface, or if there
is any doubt about the type of an embedded discontinuity, it shall be assessed
like a planar discontinuity.
3.8.1.2 Effects and Assessment of Imperfections
At geometrical imperfections, three effects affecting fatigue resistance can be dis-
tiguished, as summarized in table {3.8}-1.
page 83

Increase of general stress level
This is the effect of all types of misalignment due to secondary bending. The
additional effective stress concentration factor can be calculated by appropriate
formulae. The fatigue resistance of the structural detail under consideration is
to be lowered by division by this factor.
Local notch effect
Here, interaction with other notches present in the welded joint is decisive.
Two cases are to be distinguished:
Additive notch effect
If the location of the notch due to the the weld imperfection coincides with a
structural discontinuity associated with the geometry of the weld shape (e.g.
weld toe), then the fatigue resistance of the welded joint is decreased by the
additive notch effect. This may be the case at weld shape imperfections.
Competitive notch effect
If the location of the notch due to the weld imperfection does not coincide with
a structural geometry associated with the shape geometry of the weld, the
notches are in competition. Both notches are assessed separately. The notch
giving the lowest fatigue resistance is governing.
Cracklike imperfections
Planar discontinuities, such as cracks or cracklike imperfections, which require
only a short period for crack initiation, are assessed using fracture mechanics
on the basis that their fatigue lives consist entirely of crack propagation.
After inspection and detection of a weld imperfection, the fIrst step of the assessment
procedure is to determine the type and the effect of the imperfection as given here.
If a weld imperfection cannot be clearly associated to a type or an effect of imperfec-
tions listed here, it is recommended that it is assumed to be cracklike.
page 84

Tab. {3.8}-1: Categorization and assessment procedure for weld imperfections
Effect of imperfection Type of imperfection Assessment
Rise of general stress Misalignment Formulae for effective
level stress concentration
Local additive Weld shape imperfec- Tables given
notch tions, undercut
effect
competitive Porosity and inclusions Tables given
not near the surface
Cracklike imperfection Cracks, lack of fusion Fracture mechanics
and penetration, all types
of imperfections other
than given here
3.8.2 Misalignment
Misalignment in axially loaded joints leads to an increase of stress in the welded joint
due to the occurrence of secondary shell bending stresses. The resulting stress is
calculated by stress analysis or by using the formulae for the stress magnification
factor km given in appendix 6.3.
Secondary shell bending stresses do not occur in continuous welds longitudinally
loaded or in joints loaded in pure bending, and so misalignment will not reduce the
fatigue resistance. However, misalignment in components, e.g. beams, subject to
overall bending may cause secondary bending stresses in parts of the component,
where the through thickness stress gradient is small, e.g. in a flange of a beam, where
the stress is effectively axial. Such cases should be assessed.
Some allowance for misalignment is already included in the tables of classified struc-
tural details (3.2). In particular, the data for transverse butt welds are directly ap-
plicable for misalignment which results in an increase of stress up to 30%, while for
the cruciform joints the increase can be up to 45% . In these cases the effective stress
magnification factor km,eff should be calculated as given below.
Tab. {3.8}-2: Effective stress magnification
IType of welded joint I1<;. already
covered I~~ Ibutt welds, transverse 1.30 ~/1.3 at least 1.0
cruciform joints 1.45 ~/1.45 at least 1.0
page 85

For the simultaneous occurrence of linear and angular misalignment, both stress mag-
nification factors should be applied simultaneously using the formula:
(22)
As misalignment reduces the fatigue resistance, the fatigue resistance of the classified
structural detail (3.2) has to be divided by the effective stress magnification factor.
3.8.3 Undercut
The basis for the assessment of undercut is the ratio ult, i.e. depth of undercut to plate
thickness. Though undercut is an additive notch, it is already considered to a limited
extent in the tables of fatigue resistance of classified structural details (3.2).
Undercut does not reduce fatigue resistance of welds which are only longitudinally
loaded.
3.8.3.1 Steel
Tab. {3.8}-3: Acceptance levels for weld toe undercut in steel
Fatigue class Allowable undercut ult
butt welds fillet welds
100 0.025 not applicable
80 0.075 0.05
71 0.10 0.075
63 0.10 0.10
56 and lower 0.10 0.10
Notes: a) undercut deeper than 1 mm shall be
assessed as a crack-like imperfection.
b) the table is only applicable for plate
thicknesses from 10 to 20 mm
page 86

3.8.3.2 Aluminium
Tab. {3.8}-4: Acceptance levels for weld toe undercut in aluminium
Fatigue class Allowable undercut ult
butt welds fillet welds
40 0.075 0.05
36 0.10 0.075
32 0.10 0.10
28 and lower 0.10 0.10
Notes: a) undercut deeper than 1 mm shall be
assessed as a crack-like imperfection.
b) the table is only applicable for plate
thicknesses from 10 to 20 mm
3.8.4 Porosity and Inclusions
Embedded volumetric discontinuities, such as porosity and inclusions, are considered
as competitive weld imperfections which can provide alternative sites for fatigue crack
initiation than those covered by the fatigue resistance tables of classified structural
details (3.2).
Before assessing the imperfections with respect to fatigue, it should be verified that the
conditions apply for competitive notches, i.e. that the anticipated sites of crack initia-
tion in the fatigue resistance tables do not coincide with the porosity and inclusions to
be assessed and no interaction is expected.
It is important to ensure that there is no interaction between multiple weld imperfec-
tions, be it from the same or different type. Combined porosity or inclusions shall be
treated as a single large one. The defect interaction criteria given in (3.8.5) for the
assessment of cracks also apply for adjacent inclusions. Worm holes shall be assessed
as slag inclusions.
If there is any doubt about the coalescence of porosity or inclusions in the wall
thickness direction or about the distance from the surface, the imperfections shall be
assessed as cracks. It has to be verified by NDT that the porosity or inclusions are
embedded and volumetric. If there is any doubt, they are to be treated as cracks.
The parameter for assessing porosity is the maximum percentage of projected area of
porosity in the radiograph; for inclusions, it is the maximum length. Directly adjacent
inclusions are regarded as a single one.
page 87

3.8.4.1 Steel
Tab. {3.8}-5: Acceptance levels for porosity and inclusions in welds in steel
Fatigue class Max. length of an inclusion in Limits of
mm porosity in %
as-welded stress relieved +
of area * **
100 1.5 7.5 3
90 2.5 19 3
80 4 58 3
71 10 no limit 5
63 35 no limit 5
56 and lower no limit no limit 5
* Area of radiograph used is length of weld affected by po-
rosity multiplied by width of weld
** Maximum pore diameter or width of an inclusion less than
114 plate thickness or 6 mm
+ Stress relieved by post weld heat treatment
3.8.4.2 Aluminium
Tab. {3.8}-6: Acceptance levels for porosity and inclusions in welds in aluminium
Fatigue class Max. length of an Limits of porosity
inclusion in mm ** in % of area * **
as-welded
40 and higher 1.5 0+)
36 2.5 3
32 4 3
28 10 5
25 35 5
15 and lower no limit 5
* Area of radiograph used is length of weld affec-
ted by porosity multiplied by width of weld
** Maximum pore diameter or width of an inclusion
less than 114 plate thickness or 6 mm
+) Single pores up to 1.5 mm allowed
Tungsten inclusions have no effect on fatigue behaviour and therefore do not need to
be assessed.
page 88

3.8.5 Cracklike Imperfections
3.8.5.1 General Procedure
Planar discontinuities, cracks or cracklike defects are identified by non-destructive
testing and inspection. NDT indications are idealized as elliptical cracks for which the
stress intensity factor is calculated according to 2.2.5.
2e 2e CLADDING
1
-' a t-r&
Il }
Ei!
LlIMIH1tR
~ IHDICATION
2& 2& ,
J
t , .
~---.
1
I
2a ,~.
G; 2&
l' t
-2e
~
Fig. (3.8)-1 Transformation ofNDT indications to an elliptic or semi-elliptic cracks
For embedded cracks, the shape is idealized by a circumscribing ellipse, which is
measured by its two half-axes a and c. The crack parameter a (crack depth) is the
half-axis of the ellipse in the direction of the crack growth to be assessed. The remai-
ning perpendicular half-axis is the half length of the crack c. The wall thickness
parameter t is the distance from the center of the ellipse to the nearest surface. If the
ratio alt > 0.75, the defect is to be recategorized as a surface defect.
Surface cracks are described in
terms of a circumscribing half-
ellipse. The wall thickness
parameter is wall thickness t. If
the ratio of alt> 0.75, the
defect is regarded as being fully
penetrating and is to be recate-
gorized as a centre crack or an
edge crack, whichever is ap-
plicable.
c c
b
b =distance to nearest edge t =distance to nearest surface
Fig. (3.8)-2 Crack dimensions for assessment
For details of dimensions of cracks and recategorization see appendix 6.2.
page 89

3.8.5.2 Simplified Procedure
The simplified procedure is based on the integration of the crack propagation law (4.4)
from an initial defect size ~ to defect size of 0.75 times wall thickness using the
material resistance against crack propagation as given in 3.6.1 for steel.
In the tables the stress ranges at 2*1()6 cycles corresponding to the definition of the
fatigue classes (FAT) of classified structural details (3.2) are shown. The tables have
been calculated using the correction functions and the weld joint local geometry
correction given in 6.2.4. (see tab. {6.2}-1 and tab. {6.2}-3).
In assessing a defect by the simplified procedure, the stress range AOi for the initial
crack size parameter ~ and the stress range AO'e for the critical crack size parameter
ae are taken. The stress range Au or the FAT class belonging to a crack propagation
from 3j to ae at 2.106 cycles is then calculated by:
~ 3 3Il.a = Il.a· -Il.ar c
For aluminium, the tables may be used by dividing the resistance stress ranges at 2·
106 cycles (FAT classes) for steel by 3.
L .1
L=toe distance
Fig. (3.8)-3 Toe distance I for simplified
procedure
Tables {3.8}-7: Stress ranges at 2.106 cycles (FAT classes in N/mm2) of welds contai-
ning cracks for the simplified procedure (following 3 pages)
page 90

ISurface cracks at fillet weld toes I
a i long surface crack near plate edge, fillet welds (lit = 2.5)
25.0 a a 0 0 0 0 0 0 0 0 0 0 4 6 10 19
20.0 0 0 0 0 0 0 0 0 0 0 a 6 8 10 13 22
16.0 0 0 0 0 0 0 0 0 0 0 7 9 12 14 17 24
12.0 0 0 0 0 0 0 a a 0 8 12 14 17 19 22 28
10.0 a a a a a a a 4 7 11 15 18 20 22 24 29
8.0 a a a a a a 6 9 12 15 19 22 24 25 28 31
6.0 a a a a a 9 12 15 17 21 24 27 28 29 31 34
5.0 0 0 0 0 8 13 16 19 21 25 27 29 31 32 33 35
4.0 0 0 0 7 13 17 21 23 25 28 31 32 34 34 35 36
3.0 0 0 10 14 20 24 27 29 31 33 35 36 37 37 38 38
2.0 8 15 20 23 29 32 34 36 37 39 40 40 41 41 41 39
1.0 26 32 36 38 42 43 44 45 45 46 46 46 45 45 44 42
0.5 43 47 49 50 51 51 51 51 51 50 50 49 48 48 47 43
0.2 59 60 60 60 59 58 57 56 56 54 53 52 51 50 49 44
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
a i long surface crack (a:c = 1: 10) , fillet welds (lit = 2.5)
25.0 a 0 0 0 0 0 0 0 0 a 0 0 6 10 15 26
20.0 0 0 0 0 0 a a 0 0 0 a 8 12 15 20 30
16.0 0 0 0 0 0 0 0 0 a 0 10 15 18 21 25 33
12.0 a a 0 0 0 0 0 0 0 12 18 22 25 27 30 38
10.0 0 a 0 0 0 0 0 7 12 17 23 26 29 31 34 40
8.0 0 0 0 0 0 0 10 15 18 23 28 31 33 35 38 44
6.0 0 0 0 0 a 14 19 23 26 31 34 37 39 40 42 48
5.0 a a a a 13 20 24 28 31 35 38 40 42 43 45 50
4.0 0 0 0 11 20 26 31 34 36 39 42 44 46 47 49 53
3.0 0 a 16 21 29 34 38 40 42 45 48 49 51 52 54 56
2.0 13 23 30 34 40 44 47 49 50 53 55 57 58 58 59 61
1.0 38 45 50 53 57 59 61 63 64 66 67 67 68 68 68 66
0.5 59 64 67 69 72 74 75 75 76 76 76 76 76 75 74 71
0.2 83 86 87 88 89 89 89 88 88 87 86 84 83 82 81 75
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
ai short surface crack (a:c = 1:2), fillet welds (lit = 2.5)
25.0 0 0 a 0 a 0 0 0 a 0 0 0 13 20 28 41
20.0 0 0 a 0 a a 0 0 0 0 0 18 25 29 35 46
16.0 a a 0 0 0 a 0 0 a 0 21 28 32 36 41 49
12.0 0 0 a 0 0 a 0 0 0 25 33 38 41 44 48 54
10.0 0 0 0 0 0 0 0 15 24 32 39 43 46 48 51 56
8.0 0 0 0 0 0 0 21 29 34 40 46 49 52 53 55 59
6.0 0 0 0 0 0 28 35 40 44 49 53 56 58 59 60 62
5.0 0 0 0 0 27 36 42 47 50 54 58 60 61 62 63 64
4.0 0 0 a 23 38 45 50 54 56 60 62 64 65 65 66 66
3.0 a 0 31 40 50 55 59 62 64 66 68 68 69 69 69 68
2.0 26 42 51 57 63 67 69 71 72 73 74 74 74 74 74 70
1.0 63 71 75 78 81 82 83 83 83 83 83 82 81 80 79 74
0.5 87 91 92 93 94 93 93 92 91 90 88 87 86 85 83 76
0.2 109 109 108 107 105 103 101 100 99 96 94 92 90 88 86 78
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
page 91

ISurface cracks at butt weld toes I
a j long surface crack near plate edge, butt welds (lit = 1)
25.0 0 0 0 0 0 0 0 0 0 0 0 0 4 6 10 19
20.0 0 0 0 0 0 0 0 0 0 0 0 6 8 10 13 23
16.0 0 0 0 0 0 0 0 0 0 0 7 9 12 14 17 26
12.0 0 0 a a a 0 a a a 8 12 14 17 19 23 30
10.0 0 a a a a 0 a 4 7 11 15 18 21 23 26 32
8.0 0 0 a a a a 6 9 12 15 19 22 25 27 29 34
6.0 0 0 a a a 9 12 15 17 22 25 28 30 31 33 37
5.0 0 a a a 8 13 16 19 21 25 29 31 33 34 36 39
4.0 0 a a 7 13 17 21 24 26 30 33 35 36 37 39 41
3.0 0 a 10 14 20 24 28 30 32 35 38 39 40 41 42 43
2.0 8 15 20 24 30 34 36 38 40 42 43 44 45 45 46 45
1.0 26 33 38 41 45 47 48 49 50 51 51 51 51 51 51 48
0.5 46 50 53 54 56 57 58 58 58 57 57 56 55 55 54 50
0.2 65 67 67 67 67 66 65 65 64 62 61 60 59 58 56 51
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
a j long surface crack (a:c = 1:10), butt welds (1ft = 1)
25.0 0 a a 0 a a a a a 0 0 0 6 10 15 27
20.0 0 a a a a a 0 a 0 0 a 8 12 15 20 31
16.0 0 a a a a a a a a a 10 15 18 21 25 35
12.0 0 0 a a 0 0 a a a 12 18 22 25 27 31 40
10.0 0 a a 0 a 0 a 7 12 17 23 26 29 32 35 43
8.0 0 a a 0 a 0 10 15 18 23 28 32 34 36 39 46
6.0 0 a a 0 a 14 19 23 26 31 35 38 40 42 44 51
5.0 0 a a a 13 20 24 28 31 35 39 42 44 45 48 53
4.0 0 a a 11 20 26 31 34 37 41 44 46 48 50 52 57
3.0 0 a 16 21 29 35 39 42 44 47 50 52 54 55 57 61
2.0 13 23 30 34 41 46 49 51 53 56 58 60 61 62 64 66
1.0 39 46 51 55 59 63 65 67 68 70 71 72 73 73 74 73
0.5 61 67 70 73 76 78 80 81 82 82 83 83 82 82 82 78
0.2 88 91 93 95 96 96 96 96 96 95 94 93 92 91 90 84
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
aj short surface crack (a:c = 1:2), fillet welds (lit = 1)
25.0 0 a a a a a a a a a 0 a 13 20 28 42
20.0 0 0 0 0 0 0 0 a a a 0 18 25 29 35 47
16.0 0 0 a a a 0 a 0 0 a 21 28 32 36 41 51
12.0 0 0 a a 0 0 a 0 0 25 33 38 41 44 48 56
10.0 0 0 a 0 0 0 a 15 24 32 39 43 47 49 53 59
8.0 0 0 a a 0 0 21 29 34 40 46 50 53 55 57 62
6.0 0 0 a a a 28 35 40 44 50 54 57 60 61 63 66
5.0 0 0 0 a 27 36 42 47 50 55 59 62 63 65 66 68
4.0 0 0 0 23 38 45 51 54 57 61 65 67 68 69 70 71
3.0 0 0 31 40 50 56 60 63 65 69 71 72 73 74 74 74
2.0 26 42 51 57 64 69 72 74 75 77 79 79 80 80 80 78
1.0 64 72 77 81 85 87 88 89 89 90 90 89 89 88 87 82
0.5 91 95 97 99 100 101 101 100 100 99 98 96 95 94 92 85
0.2 116 117 117 117 115 114 112 111 110 107 105 103 101 99 97 88
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
page 92

Embedded cracks
ai embedded long crack near plate edge
25.0 0 0 0 0 0 0 0 0 0 0 0 0 7 12 18 29
20.0 0 0 0 0 0 0 0 0 0 0 0 11 15 19 23 33
16.0 0 0 0 0 0 0 0 0 0 0 13 18 21 24 28 37
12.0 0 0 0 0 0 0 0 0 0 15 21 25 28 30 34 41
10.0 0 0 0 0 0 0 0 9 14 21 26 30 32 34 37 44
8.0 0 0 0 0 0 0 12 18 22 27 31 35 37 39 41 47
6.0 0 0 0 0 0 17 23 27 30 34 38 41 42 44 46 52
5.0 0 0 0 0 16 23 28 32 34 38 42 44 46 47 49 54
4.0 0 0 0 14 24 30 34 37 40 43 46 48 50 51 53 58
3.0 0 0 19 26 33 38 42 44 46 49 52 54 55 56 58 62
2.0 16 27 34 39 45 49 51 53 55 58 60 61 62 63 65 68
1.0 43 50 55 58 62 65 67 68 69 71 73 74 75 76 77 79
0.5 65 69 73 75 78 80 81 82 83 85 86 87 88 88 89 91
0.2 90 93 95 97 99 100 101 102 103 104 105 105 106 106 107 108
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
a, embedded long crack (a:c = 1:10) apart from plate edge
25.0 a 0 0 0 0 0 0 0 0 a 0 0 7 12 18 32
20.0 0 0 a 0 0 0 0 0 0 a 0 12 17 20 26 37
16.0 0 a 0 a 0 0 0 0 0 a 15 21 25 28 32 43
12.0 0 a a a 0 0 0 0 0 19 26 30 33 36 40 48
10.0 0 0 a 0 0 0 0 11 18 25 31 35 38 40 44 52
8.0 0 0 0 0 0 0 16 22 27 33 38 41 44 46 49 57
6.0 a 0 0 0 0 22 28 33 36 41 45 48 50 52 55 62
5.0 a 0 0 0 21 29 35 39 42 46 50 52 54 56 59 66
4.0 0 0 0 18 30 37 42 45 48 52 55 57 59 61 64 71
3.0 0 0 25 32 41 46 50 53 55 59 62 64 66 68 70 77
2.0 20 34 42 47 54 58 61 63 65 69 72 74 76 77 80 85
1.0 53 60 65 69 73 77 80 82 84 87 89 91 93 94 96 100
0.5 77 83 86 90 94 97 100 102 103 105 107 109 110 111 112 115
0.2 108 113 117 119 123 125 127 128 129 131 132 133 134 134 135 138
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
ai embedded short crack (a:c = 1:2) apart from plate edge
25.0 0 0 0 0 0 0 0 a 0 0 0 0 14 23 32 49
20.0 0 0 0 0 0 0 0 0 0 0 a 21 29 34 40 54
16.0 a a 0 0 0 0 0 0 a 0 24 33 38 42 47 59
12.0 a a a a 0 0 0 0 0 29 38 44 48 51 56 66
10.0 a a a 0 0 0 0 17 27 38 45 50 54 57 61 70
8.0 a a a a 0 0 24 33 39 47 53 57 61 63 67 74
6.0 a a a a 0 32 41 47 51 58 63 66 69 71 74 80
5.0 a a 0 0 31 42 49 54 58 64 68 71 74 75 78 84
4.0 a 0 0 27 44 53 59 63 66 71 75 77 79 81 83 89
3.0 a a 36 46 58 65 69 73 76 79 83 85 87 88 90 95
2.0 30 49 59 66 74 79 83 86 88 91 94 95 97 98 100 104
1.0 74 83 89 93 99 102 105 106 108 110 112 113 114 115 116 119
0.5 105 111 115 117 121 1:3 125 127 128 129 131 132 132 133 134 136
0.2 140 144 146 148 150 152 153 154 155 156 157 158 158 159 160 161
t = 3 4 5 6 8 10 12 14 16 20 25 30 35 40 50 100
page 93

4 FATIGUE ASSESSMENT
4.1 GENERAL PRINCIPLES
In fatigue assessment, the fatigue actions and the fatigue resistance are related by
means of an appropriate assessment procedure. It must be ensured that all three ele-
ments (actions, resistance and assessment procedure) correspond. Three procedures
may be distinguished:
a) Procedures based on S-N curves, such as
nominal stress approach
geometric stress approach
effective notch stress approach
b) Procedures based on crack propagation considerations
c) Direct experimental approach by fatigue testing of components or entire
structures
4.2 COMBINATION OF NORMAL AND SHEAR
STRESS
If normal and shear stress occur simultaneously, their combined effect shall be con-
sidered. Three cases may be distinguished:
a) If the equivalent nominal shear stress range is less than 15% of the
equivalent normal stress range or if the damage sum due to shear stress
range is lower than 10% of that due to normal stress range, the effect
of shear stress may be neclected.
b) If the normal and shear stress vary simultaneously in phase, or if the
plane of maximum principal stress is not changed significantly, the
maximum principal stress range may be used.
c) If normal and shear stress vary independently out of phase, in damage
calculation the damage sums shall be calculated separately and finally
added. A Miner damage sum of l:Di =O.5 or the usage of 112 of the cal-
culated life cycles is recommended.
Fracture mechanics crack propagation calculations should be based on maximum
principal stress range.
page 94

4.3 FATIGUE ASSESSMENT USING S-N CURVES
Fatigue verification is carried out using
and
the design spectrum of fatigue actions in terms of stress ranges .dUi,S,d' in which
the stresses of the characteristic spectrum .dUi,S,k have been multiplied by the
partial safety factor 'YF for fatigue actions.
the design resistance S-N curve based on design resistance stresses .dUR,d, in
which the characteristic resistance stress ranges .dUR,k have been divided by the
partial safety factor 'YM for fatigue resistance.
The design resistance S-N curve may be modified further according to the needs of the
damage calculation procedure.
For constant amplitude loading, the characteristic stress range .dUR,k at the required
number of stress cycles is firstly determined. Secondly the fatigue criterion is checked:
IlCJR,t
IlCJSd = .dCJn ·YF s: .dCJR,d = - -
" YM
At variable amplitude loading, cumulative damage calculation procedure is applied.
Usually a modified "Palmgren-Miner"-rule, as described in 4.3.1, is appropriate. For
load spectra which are sensitive to the position of the fatigue limit or cut-off limit, or
in which the spectrum changes during the service time, additional assessment using the
nonlinear damage calculation method described in 4.3.2 is recommended.
In fields of application, where no test data nor service experience exist and the shape
of the stress spectrum is not close to constant amplitude, it is recommended that only
half of the calculated life should be assumed. This is equivalent to a Miner sum (4.3.1)
of Dd=O.S.
4.3.1 Linear Damage Calculation by "Palmgren-Miner" Summation
If the maximum design stress range .dUmax,S,d of the load spectrum is lower than the
design fatigue limit .dUL,R,d of the design fatigue resistance S-N curve, or if it is lower
than the design cut-off limit .dUcut, R,d in cases where no fatigue limit is given, the life
of the welded joint can be assumed to be inimite and no further damage calculation is
necessary.
If the constant amplitude fatigue limit of the resistance S-N curve corresponds to an
endurance less than 108 cycles, the fatigue resistance curve has to be modified accor-
page 95

ding to fig. (4)-1. Then the slope ml of the S-N curve from the constant amplitude
fatigue limit (5· 106 cycles) up to lOs cycles is assumed to be ml = 2· m1 - 1 [30] 1.
log stress range
2E6 Nc 1E8
igue limit
cut off
1.0E+04 1.0E+05 1.0E+06 1.0E+07 1.0E+08 1.0E+09
log N
Fig. (4)-1 Modification of fatigue resistance Wohler S-N curve for Palmgren-
Miner summation
1Although it is accepted that the stresses below the constant amplitude fatigue
limit must be included in cumulative damage calculation relating to welded joints,
there are currently different opinions how this should be achieved. The method
presented here (fig. (4)-1) appears in a number of codes, including Eurocode 3.
However, recent research indicates that it can be unconservative. Here, this question
is partially solved by recommending a Miner sum of ED=O.S if the spectrum is not
close to constant amplitude. Other suggestions recommend that the S-N curve should
be extrapolated further down before the slope change is introduced. For critical cases
or areas of doubt, the user should consult relevant published literature.
page 96

For fatigue verification it has to be shown that:
i n.
"E.Dd = L -' ~ 0.5...1
1 Ni
where EDd damage by summation (note restrictions in 4.2 and 4.3)
i index for block number in load spectrum of required design life
nj number of cycles of design load stress range 40i,s,d in load spectrum
block i
Nj number of cycles at which design stress range 4Ui,S,d causes failure in
the modified design fatigue resistance S-N curve.
The order of sequence of the blocks has no effect on the results of this calculation.
In some cases it might be convenient to calculate an equivalent constant amplitude
stress range 40E and to compare it directly to the constant amplitude resistance S-N
curve neglecting the constant amplitude fatigue limit.
For the grid of fatigue resistance classes and an initial slope of m=3 predominantly
used in 3.2, the values of the modified characteristic fatigue resistance S-N curves
have been calculated. Stepping down one class corresponds to a division by 1.12. So
different levels of safety 'YM of S-N curve can be achieved (see 6.4.3).
page 97

Tab. {4}-1: Constants, constant amplitude fatige limit and cut-off limits
Values of modified characteristic fatigue resistance S-N curves for Palmgren-
Miner summation. Initial slope IDI =3.0, constant amplitude fatigue limit ,4O"L,k at
5.106 cycles, second slope IDz=5.0, cut-off at 108 cycles.
Class Constant C for S-N constant Constant C for S-N cut-off
curve at N <5e6 amp!. fat. curve at N> 5e6 limit
cycles, m1 =3 limit cycles, m2 = 5
225 2.278e13 166 6.261e17 91.1
200 1.600e13 147 3.474e17 80.9
180 1.166e13 133 2.052e17 72.8
160 8.192e12 118 1.13ge17 64.8
140 5.488e12 103 5.840e16 56.7
125 3.906e12 92.1 3.313eI6 50.6
112 2.810e12 82.5 1.913e16 45.3
100 2.000e12 73.7 1.086e16 40.5
90 1.458e12 66.3 6.411e15 36.4
80 1.012e12 58.9 3.558e15 32.4
71 7.158e11 52.3 1.95ge15 28.7
63 5.oo1e11 46.4 1.078e15 25.5
56 3.512e11 41.3 5.980e14 22.7
50 2.500el1 36.8 3.393e14 20.2
45 1.823ell 33.2 2.004e14 18.2
40 1.280ell 29.5 1.112e14 16.2
36 9.331elO 26.5 6.565e13 14.6
32 6.554elO 23.6 3.643e13 13.0
28 4.390e1O 20.6 1.86ge13 11.3
25 3. 125elO 18.4 1.060e13 10.1
22 2. 130elO 16.2 5.596e12 8.9
20 1.6ooelO 14.7 3.474e12 8.1
18 1.166e1O 13.3 2.052e12 7.3
16 8.192e9 11.8 1.13ge12 6.5
14 5.488e9 10.3 5.840e11 5.7
page 98

4.3.2 Nonlinear Damage Calculation
A nonlinear fracture mechanics damage calculation according to 4.4 is recommended
in cases, where
a) the Miner summation is sensitive to the exact location of the knee point of the
fatigue resistance S-N curve,
b) the spectrum of fatigue actions (loads) varies in service or is changed, and so
the sequence of loads becomes significant or
c) the resistance S-N curve of a pre-damaged component has to be estimated.
Where the parameters for a fracture mechanics fatigue assessment are not known and
only the resistance S-N curve is known, the S-N curve can be used to derive dimen-
sionless fracture mechanics parameters, which allow a damage calculation [31]. The
procedure is based on the "Paris" power law of crack propagation
da
- = C 'AK'"
dN 0
if AK< AKtb then da = 0
dN
where a crack parameter, damage parameter (dimensionless)
N Number of cycles
.AK range of stress intensity factor
.AKu. threshold value of stress intensity factor range
Co, m material constants
The characteristic stress intensity factor range .AKS,k of the fatigue action is calculated
with the stresses of the spectrum .AUi,S,k and the crack parameter a
AKs,1e = AaS,1e ·ra
The characteristic resistance parameters can be derived from the characteristic constant
amplitude fatigue resistance S-N curve: The threshold value corresponds to the fatigue
limit, .AKu.,k=.AUL,R,k' m equals the slope of the S-N curve, and the constant CO,k can
be calculated from a data point (.Aus-N and NS-N) on the S-N curve, preferably from the
fatigue class at 2 .106 cycles
2
CO,1e = -------
(m -2)·Ns_N· AO;-N
The fatigue verification is executed according to 4.4, using an initial crack parameter
~=1 and a final one ar= 00 or a large number e.g. ar=109• The restrictions on life
cycles given in 4.3 are to be considered.
The actual fatigue class of a pre-damaged component is FATact• = FATtVa.
page 99

4.4 FATIGUE ASSESSMENT BY CRACK PROPA-
GATION CALCULATION
The fatigue action represented by the design spectrum of stress intensity factor ranges
tl.Ki,S,d :;:: tl.Ki,s,k·YF
is verified by the material resistance design parameters against crack propagation
CO,d = CO,k 'Y:' = CO,k·rM
using the "Paris" power law
da
- = C -tl.Km
dN 0 '
da
if tl.K<tl.Kth then - = 0
dN
where a crack parameter, damage parameter
N Number of cycles
aK range of stress intensity factor
aKua threshold value of stress intensity factor range
Co, m material constants
At stress intensity factors which are high compared with the fracture toughness of the
material, ~, an acceleration of crack propagation will occur. In these cases, the
following extension of the "Paris" power law of crack propagation is recommended.
In the absence of an accurate value of the fracture toughness, a conservative estimate
should be made.
da C·tl.Km
o
=----
dN (l_R) __tl._K
Kc
where
Kc fracture toughness
R stress ratio
The number of life cycles N is determined by integration starting from an initial crack
parameter ~ to a final one ar. The calculated number of life cycles N has to be greater
or equal to the required number of cycles.
page 100

In general, the integration has to be carried out numerically. The increment for one
cycle is
It is recommended that a continous spectrum is subdivided to an adequate number of
stress range blocks, e.g. 8 or 10 blocks, and the integration performed blockwise by
summing the increments of a and the number of cycles of the blocks. The entire size
of the spectrum in terms of cycles should be adjusted by multiplying the block cycles
by an appropriate factor in order to ensure at least 20 loops over the whole spectrum
in the integration procedure.
4.5 FATIGUE ASSESSMENT BY SERVICE TESTING
4.5.1 General
Components or structures may be tested or verified in respect to fatigue for different
reasons:
a) Existence of a new design with no or not sufficient knowledge or experience of
fatigue behaviour.
b) Verification of a component or structure for a specified survival probability
under a special fatigue action (stress) history.
c) Optimization of design and/or fabrication in respect of weight, safety and
economy.
The fatigue tests should be performed using the data of the fatigue action history (see
3.7), factored by the partial safety factors 'YF and 'YM'
The tests should be performed according to well established and appropriate procedu-
res or standards [32].
The verification or assessment depends of the safety strategy considered (see 5.2). Safe
life strategy on the one hand, and fail safe or damage tolerant strategy on the other,
have to be distinguished.
4.5.2 Safe Life Verification
The number design life cycles of the component or structure should be less than the
factored number of the log mean of test life cycles.
page 101

N
N <2-
d F
where
Nd number of design life cycles, up to which the component or structure
may be used in service
NT log mean value of number of test life cycles of the test specimens or
number of cycles of the first test specimen to fail, whichever is ap-
plicable.
F factor dependent of the number of test results available as defined in
tables {4.5}-1 and {4.5}-2.
Before using the tables, an estimate on standard deviation of log N has to be made.
The standard deviation varies with the life cycles of the component to be assessed, see
fig. (3.7)-1. For geometrically simple stuctures at a number of cycles between IQ4 and
IQ5 a standard deviation of 0.178 may be chosen. For complex structures at cycles up
to 106, 0.25 is more appropriate. For higher cycles near the endurance limit, no
estimate can be given. Here, special verification procedures are recommended, see ref.
[32]
If all components or test specimens are tested to failure, table {4.5}-1 shall be used.
If the tests are carried out until failure of the first test specimen, table {4.5}-2 shall be
used (see also 6.4). The F-factors refer to a 95 % survival probability at a confidence
level of 75% of the mean.
Tab. {4.5}-1: F-factors for failure of all test specimens
I Stdv. n II 2 I 4
I 6 I 8
I 10 I
0.178 3.93 2.64 2.45 2.36 2.30
0.200 4.67 2.97 2.73 2.55 2.52
0.250 6.86 3.90 3.52 3.23 3.18
Tab. {4.5}-2: F-factors for the first test specimen to fail
I Stdv. n
II 2
I 4
I 6 I 8
I 10
I
0.178 2.72 2.07 1.83 1.69 1.55
0.200 3.08 2.26 1.98 1.80 1.64
0.250 4.07 2.77 2.34 2.09 1.85
The factor F may be further modified according to safety requirements as given in
page 102

chapter 5.3. For more details see appendix 6.4.
4.5.3 Fail Safe Verification
Fatigue life verification of fail safe structures depends largely on the design and
operation parameters of a structure. The effectivness of statically over-determined
(hyperstatic) behaviour or redundancy of structural components, the possibility of
detection of failures in individual structural parts and the possibility of repair determi-
ne the level of safety required in the individual structural parts. So, no general recom-
mendation can be given.
It is recommended that 4.5.2 is used as a general guidance and to establish agreement
on the factor F.
4.5.4 Damage Tolerant Verification
The verification is based on crack growth measurements, starting from a crack size,
which can be detected in inspection up to a critical crack size, at which the limit state
of critical safety against brittle or plastic fracture or other modes of failure of the
remaining sectional area is attained.
The criteria for factoring the observed life cycles for the test depend of the application.
It is recommended to establish agreement on the factor F.
page 103

5 SAFETY CONSIDERATIONS
A component has to be designed for an adequate survival probability. The required
survival probability is dependent on the
a) uncertainties and scatter in the fatigue assessment data,
b) safety strategy and
c) consequences of failure.
The uncertainties of fatigue assessment data may arise from fatigue actions, such as
1. determination of loads and load history,
2. determination of stresses or stress intensity factors from the model used
for analysis, and
3. dynamic response problems.
These uncertainties are covered by an appropriate partial safety factor for the fatigue
actions 'YF' which is not considered here.
Uncertainties of fatigue assessment data arising from fatigue resistance and damage
calculation are:
4. scatter of fatigue resistance data,
5. scatter of verification results of damage calculations.
The sources of uncertainty numbered 4. and 5. are considered here. For normal ap-
plications, they are already covered in the fatigue resistance data given here. For
special applications, the data may be modified by the selection of an adequate partial
safety factor 'YM'
5.2 FATIGUE DESIGN STRATEGIES
Different ways of operation in service require different fatigue design strategies. The
definition of a fatigue strategy refers predominantly to the method of fatigue analysis,
inspection and monitoring in service.
5.2.1 Infinite Life Design
This strategy is based on keeping all fatigue actions under the resistance fatigue limit
page 104

or threshold value. No regular monitoring in service is specified. So a high survival
probability has to be provided. For fatigue actions which are almost uniform and act
at very high cycles this strategy may be adequate.
5.2.2 Safe Life Design
This design strategy is based on the assumption that initially the welded joint is free of
imperfections. No regular monitoring in service is specified, so a high survival
probability has to be provided.
5.2.3 Fail Safe Design
This design strategy is based on statically over-determined (hyperstatic) or redundant
structures. No regular monitoring in service is provided. In case of a fatigue failure,
redistribution of forces provides an emergency life, so that the failure can be detected
and repaired. The welded joints can be designed for a normal survival probability.
5.2.4 Damage Tolerant Design
This design strategy is based on the assumption of the presence of cracks as large as
the detection level of the non-destuctive testing method applied. Fracture mechanics is
used to calculate the life cycles until failure. From the number of life cycles, regular
inspection intervals are derived. A normal probability of survival is adequate.
5.3 PARTIAL SAFETY FACTORS
The required partial safety factor 'YM depends largely on circumstances such as
a) fatigue design strategy
b) consequences of failure
c) practical experience in fields of application.
Thus, no general recommendation can be given. In most cases for normal fabrication
quality and regular inspections in service, 'YM=1 might be adequate.
The safety factors are given in terms of stress. If safety factors are needed in terms of
cycles, rM may be calculated using the slope m of the resistance S-N curve
page 105

It should be recognized that the slope m of the S-N curve varies with the number of
cycles, see fig. (3.7)-1. An example of a possible table of partial safety factors is
given in appendix 6.4.
5.4 QUALITY ASSURANCE
Weld quality is assured preferably by the application of an ISO 9000 quality manage-
ment system or a comparable system.
The weld quality should be equal to quality class B according to ISO 5817. However,
some exceptions may be allowed in the tables given in chapter 3.2.
Besides regulations and quality codes, the general standards ofgood workmanship have
to be maintained.
page 106

6 APPENDICES
The appendices are intended to give special guidances, background information and
additional explanations. They are not normative.
6.1 LOAD CYCLE COUNTING
6.1.1 Transition Matrix
To establish a transition matrix, first the number of different stess levels or classes has
to be defined. 32 stress levels are sufficient. Then the numbers (occurrence) of transi-
tions (reversals) from one extreme value (peak or through) to another are counted and
summarized in the matrix. A number in the matrix element ~,j indicates the number
of transitions from a stress belonging to class i to a stress belonging to class j.
a (1, 1) a(l,j) a (l,n)
stress j
a(i,l)
transition from
min to max i<j
transition from
max to min i>j
a(n,l)
Fig. (6.1)-1 Principle of the transition matrix
a(n,n)
stress i
The data for the transition matrix can be obtained by measurement or by time simula-
tion computations. A time signal for fatigue tests or crack propagation simulations or
cumulative frequency diagrams (stress spectra) for damage calculations can be gene-
rated from the transition matrix by a Markov random draw.
6.1.2 Rainflow or Reservoir Counting Method
The algorithm of rainflow counting method is well explained by using the analogy of
the flow of water on a pagoda roof. The stress signal, looked at vertically, is regarded
as the pagoda roof. A cycle is obtained, when a contour is closed by the drop of the
page 107

flow from a peak to a slope of the roof [26 and 27]. The range is then equal to the
difference between the extreme values of the contour. Later the smaller included cycles
can be determined the same way. The non closed contour from the extreme of the
entire signal leads to a half cycle. Reservoir counting is similar.
19· ...........~
12
4
11
1
Fig. (6.1)-2 Illustration of rainflow counting
6.2 FRACTURE MECHANICS
13
Cycles:
2 - 3
5 - 6
4 - 7
8 - 9
11 - 12
open:
1 - 10 - 1
6.2.1 Rapid Calculation of Stress Intensity Factors
A simplified method may be used to determine Mk-factors [19]. Here, the Mk-factors
are derived from the non-linear stress peak distribution Unlp(x) along the anticipated
crack path x assuming no crack being present. Hence, the function of the stress con-
centration factor ~n1p(x) can be calculated. The integration for a certain crack length
a yields:
For different crack lengths a, a function Mk(a) can be established, which is preferably
presented in the form:
COnst
aCXl'
page 108

6.2.2 Dimensions of Cracks
Tab. {6.2}-1: Dimensions for assessment of crack-like imperfections (example)
Idealizations and dimensions of crack-like imperfection for fracture mechanics
assessment procedure (t = wall thickness).
~_=----'(V I. 2e .1 t
I. 2• •1
f.Y
}
- ~ Jt
~
page 109

6.2.3 Interaction of Cracks
Adjacent cracks may interact and behave like a single large one. The interaction
between adjacent cracks should be checked according to an interaction criterion.
There are different interaction criteria, and in consequence no strict recommendation
can be given. It is recommended to proceed according to an accepted code, e.g. [24].
6.2.4 Formulae for Stress Intensity Factors
Stress intensity factor formulae may be taken from literature, see references [13] to
[20]. For the majority of cases, the formulae given below are sufficient.
Tab. {6.2}-2: Stress intensity factors at welds
Surface cracks under membrane stress
c c
~
)
b¢lt<
Q7'-
The formula for the stress in-
tensity factor KI is valid for
a/c < 1,
b I
for more details see ref. [14]
I
b =distance to ......- edge t =distonce to nearest sutfac:e
KI = a V( n . a / Q) F.
Q = 1 + 1.464 (a/c)J.6.S
F. = [MI + Mz' (a/t)2 + M3 ' (a/t)41·g·f·fw
~ = 1.13 - 0.09 (a/c)
M2 = -0.54 + 0.89 / (0.2 + a/c)
M3 = 0.5 - 1 I (0.65 + a/c) + 14 (1 - a/c)~
fw = [sec(n'c v(a/t) 1(2'b»] 112
9 and f are dependent to direction
Halt-direction: 9 = 1
ItcH-direction: 9 = 1 + [0.1 + 0.35 (a/t)2]
page 110
f = 1
f = v(a/c)

Embedded cracks under membrane stress
c c
tensity factor K( is valid for
a/c < 1,b . It )
for more details see ref. [14]~
I
u~r
b
~I
b =distance II> neare5t edge t =distance II> nmrest surface
Kl1 Q, F., f ... as given in A.1.1.1, but:
M( = 1
~ = 0.05 I (0.11 + (a/c)312)
M3 = 0.29 / (0.23 + (a/c)312)
"a"-direction: 9 = 1
·c"-direction: 9 = 1 - (a/t)4 I (1 + 4a/c)
f = 1
f = v(a/c)
Surface cracks under shell bending and membrane stress
~
)
tensity factor K( is valid for
a/c < 1,
for more details see ref. [14].
I---b--I
b =distance II> nearest edge t =distance II> nmrest surface
K( = (omcm + Ho0bcn ) v(7r°a / Q) F.
Q = 1 + 1.464 (a/c)l.65
F. = [M( + ~o (a/t)2 + ~. (a/t)4]ogofof...
M( = 1.13 - 0.09 (a/c)
M2 = -0.54 - 0.89 I (0.2 + a/c)
M3 = 0.5 - 1 / (0.65 + a/c) + 14 (1 - a/c)24
f ... = [sec(7r°c v(a/t) /(2 ob»]
(12
"a"-direction: 9 = 1
"c"-direction: 9 = 1 + [0.1 + 0.35 (a/t)2]
f = 1
f = v(a/c)
The function H is given by the formulae:
"a"-direction: H = 1 + Gda/t) + G2(a/t)2
where G( = -1.22 -0.12o(a/c)
G2 = 0.55 - 1. 05 ° (a/c) 0.75 + 0.47 (a/ c) 1.5
"c"-direction: H = 1 - 0.34 (a/t) - 0.11 (a/c) (a/t)
page 111

Surface crack in cylinder under internal pressure
e e
~
I--b--i
b~lt(
~I
tensity factor KJ is valid for
a/c < 1,
for more details see ref. [15],
where D is the diameter in mm
and P is the internal pressure
in N/mm2 •
b =dis!ance lI> .....rest edge t= distance II>.-.est surface
S = p"D~ / (2t)
Q = 1 + 1.464 (a/c)l.6S
F. = 0.97" [MJ + M2" (a/t)2 + M3" (a/t)4] "C"g"f"fw
MJ = 1.13 - 0.09 (a/c)
M2 = -0.54 - 0.89 / (0.2 + a/c)
M3 = 0.5 - 1 / (0.65 + a/c) + 14 (1 - a/c)~
fw = [sec(1l'"c v(a/t) /(2"b»] 112
C = [(Do} + Din2) / (D"",2 - Din2) + 1 - O.S"V aft ] "2t/Din
g and f are dependent to direction
"an-direction: g = 1 f = 1
nCB-direction: g = 1 + [0.1 + 0.35 (a/t)2] f = v(a/c)
Through the wall cracks in curved shells under internal pressure
where
and
with
K = 0mcm V(1l' " a ) Mk
Mk = 1.0
In sphere and longitudinal
cracks in cylinder loaded by
internal pressure. Mk covers in-
crease of stress concentration
factor due to bulging effect of
shell. For details see ref.
[15,18].
for x < 0.8
Mk = v(0.9S + 0.6S"x - 0.03S"x1.6 ) for x > 0.8
and x < 50
x = a / v(r"t)
a half distance between crack tips of through the wall crack
r radius of curvature perpenticular to the crack plane
t wall thickness
page 112

Root gap crack in a fillet welded cruciform joint
.-~l
tensity factor K is valid for
j t
H/t from 0.2 to 1.2 and for a/w
H "d jP
from 0.0 to 0.7. For more de-
SR -I 1.., ~ I-~ tails see ref. [17].
~LDTOECRACK I ~~ I
r. ~ ROOT CRACK
L-
a" (AI + Az " a/w) v (n"a sec(n"a/2w) )
K =
1 + 2"H/t
where w = H + t/2
a = nominal stress range in the longitudinal plates
and with x = H/t
Al = 0.528 + 3.287"x - 4.361"xz + 3.696"x3 - 1.875 "X4 + 0.415 "Xs
Az = 0.218 + 2.717"x - 10.171"xz + 13.122 "X3 - 7.755"x4 + 1.783·xs
page 113

For a variety of welded joints parametric formulae of the Mk functions have been
established and published [18,19]. For the majority of cases, the formulae given below
are sufficient [20].
Tab. {6.2}-3:
Weld local geometry correction for crack at weld toe
n ~I
jrf-:I'~l~IJttvtI
L= toe distance
Applicable for transverse
full penetrating or non-
loadcarrying welds, e.g.
butt weld, transverse
attachment, cruciform
joint K-butt weld. For
more details see ref.
[20 J•
Stress intensity magnification factor Hie > 1 for membrane stress:
for 1ft !S 2:
Hie = 0.sl-(lft)o.27-(a/t)~31 , for (aft) !S O.Os-(lft)o.ss
Hie = 0.83 - (aft)~·IS(lIt)).46 , for (aft) > 0.05 - (lft)o.ss
for 1ft> 2:
Hie = 0.615-(aft)~31, for (aft)!S 0.073
Hie = 0.83- (a/t)~·2 , for (aft) > 0.073
Stress intensity magnification factor Hie > 1 for bending stress:
for 1ft !S 1:
Hie = 0.45-(lft)0.21-(aft)~·31 , for (aft) !S 0.03-(lft)o.ss
Hie = 0.68- (a/t)~·19(lI1)).21 , for (aft) > 0.03- (lft)o.ss
for 1ft > 1:
Hie = 0.45- (aft)~.31 , for (aft) !S 0.03
Hie = 0.68' (aft)~·19 , for (aft) > 0.03
page 114

6.3 FORMULAE FOR MISALIGNMENT
Tab. {6.3}-1: Formulae for assessment of misalignment
1# ITYPE OF MISALIGNMENT I
1 Axial misalignment between flat plates
e'l
t le km = 1+,t, 1
t(ll +12)
- --tt-=----f
I~ 11 ~IE 12 ~I
Ais dependent on restraint, A=6 for unrestrained joints,
For remotely loaded joints assume I.= 12,
2 Axial misalignment between flat plates of differing thickness
t1 1e
"
km
6e t1
= 1+-,--
--~~ tl tl"+~"
T t2
t2 ~ t1
Relates to remotely loaded unrestraint joints,
The use of n = 1.5 is supported by tests,
3 Axial misalignment at joints in cylindrical shells with thickness chan-
ge
t1 t 1e 6e t"
k = 1+
1,
--
--~
m
tl (1-v2) tt+~"
t 1 ~
t2
f2~h
n= 1.5 in circumferential joints and joints in spheres,
n=O.6 for longitudinal joints,
page 115

4 Angular misalignment between flat plates
Assuming fixed ends:
with ~ = 21~ 3".
t E
2 I k = 1+ 3y . tanh(P/2)
m t P/2
altern.: k = 1+~. «oZ. tanh(P/2;
m 2 t P/2
assuming pinned ends:
k = 1+ 6y . tanh(P)
m t P
altern.: k = 1+ 3«oZ. tanh(P)
m t P
The tanh correction allows for reduction of angular misalignement due to
the straightening of the joint under tensile loading. It is always < 1 and it
is conservative to ignore it.
5 Angular misalignment at longitudinal joints in cylindrical shells
Assuming fixed ends:
with ~ = 21~ 3(1-.2)."
t E
2/
k = 1+ 3d . tanb(P!2)
m t(1-v2) P/2
assuming pinned ends:
k = 1+ 6d tanh(P)
m t(1-v2) P
d is the deviation from the idealized geometry
page 116

6 Ovality in pressurized cylindrical pipes and shells
Wei
l.S(Dmax -Dmin)-cos(2<P):
k = 1+.....,.. ~ ...... m
t (1+ O.5P~(1-V2). ( ~)']
/'
/
I'
__t
I

D......
I
OIMX. /
...... /'
- .,..
7 Axial misalignment of cruciform joints (toe cracks)
~+~
e·l
k = l. 1
m t(ll+12)
IE 11 "1< 12 ~I A is dependent on restraint
f1 ~ 12
A varies from A=3 (fully restrained) to A=6 (unrestraint). For unrestrai-
ned remotely loaded joints assume: 11=12 and A=6
8 Angular misalignment of cruciform joints (toe cracks)
~-&~
11.12
k = l+locx·
~
-:f m t(ll +12)
l1"IE 12 "I
A is dependent on restraint
IE
If the inplane displacement of the transverse plate is restricted, A varies
from >"=0.02 to A=O.04. If not, A varies from A=3 to A=6.
page 117

9 Axial misalignment in rillet welded cruciform. joints (root cracks)
~ refers to the stress range in weld throat.
page 118
e
k = 1+-
m t+h

6.4 STATISTICAL CONSIDERATIONS ON SAFETY
6.4.1 Statistical Evaluation of Fatigue Test Data
The different methods as described in 3.7 consider different statistical effects, when
evaluating a set of fatigue data. Ideally, all effects have to be considered, e.g.
a) Variance of data
b) Variance of the mean value
c) Difference of the distribution of the whole set of data (population) and
the distribution of the sample (Gaussian versus t-distribution)
d) Deviation from the assumed Gaussian distribution
The values given are so called characteristic values xk• These are the values at a 95 %
survival probability at a confidence level of 75 % of the mean Xm • The general formula
for the characteristic value is given by:
x =x - k·Stdvk m
The factor kJ considers the effects a) to d). It is calculated by
where t
n
~
tk = (0.875,n-l) +<1>-1 •
1 .;n (0.95)
n-l
2
X(O.l2S, n-l)
value of the t-distribution, here for a one sided probability of 0.875 or
a two sided probability of 0.75 at n-1 degrees of freedom
number of thest pieces
distribution function of the Gaussian normal distribution, here 1.645 for
a survival probability of95% (superscript -1 indicates inverse function)
Chi-sqare, here for a probability of 0.125 at n-1 degrees of freedom
The Chi-square correction covers a possible deviation of test data from the assumed
Gaussian normal distribution. If there is evidence about the distribution from other
tests, or if a large number of test specimens is available, this correction may be
dropped, giving:
z, = 1(0.875, n-l) +<1>-1 _ 1(0.875, n-l) + 1.645
~ .;n (0.95) - .;n
At a higher number of test specimens, the t-distribution may be replaced by the
Gaussian normal destribution. Hence
-1
z, = <I>(0.875) + <I>-1
"'3 .;n (0.95)
= 1.15 + 1.645
rn
page 119

With this simplification, the evaluation is equal to the conventional one at about 10 test
specimens.
Tab. {6.4}-1: k-values for the different methods
I n
I t I x2
I kl I
2 2.40 0.028 11.53
3 1.61 0.270 5.41
4 1.44 0.69 4.15
5 1.32 1.21 3.58
10 1.24 4.47 2.73
15 1.21 8.21 2.46
20 1.20 12.17 2.32
25 1.19 16.26 2.24
30 1.18 20.45 2.17
40 1.18 29.07 2.09
50 1.17 37.84 2.04
100 1.16 83.02 1.91
6.4.2 Statistical Evaluation at Component Testing
Testing all test specimens to failure
Starting from the formula in 4.5.2, there is
N
N <2
d F
logNT-logF> logNd
k2 I k3 I
3.34 2.45
2.57 2.30
2.36 2.22
2.23 2.15
2.04 2.00
1.96 1.93
1.91 1.90
1.88 1.87
1.86 1.85
1.83 1.82
1.81 1.80
1.76 1.75
Taking the acceptance criterion from chapter 3.7 Xm - k Stdv > Xk the factor F can
be received:
logF ;:: k·Stdv
With the formula for k the different values of F can be calculated, depending on
number of test specimens n and on the assumed standard deviation Stdv of the test
pieces in terms of 10gN.
k = t(O.87S, n-1) ..4..-1 _ t(O.S7S, n-1) + 1.645
/it + 'I' (0.95) - /it
page 120

Tab. {6.4}-2: k-values for 75 % confidence of the mean
181~.34 I~.37 I~.19 I~.1O I~~04
Testing all test specimens simultaneously until iIrst failure
100
1.76
When considering statistical evaluation, account must be taken of additional effects:
1)
2)
3)
Distribution of the lIn-th extreme
value
Distribution of the sample be-
tween lIn-th extreme and mean
Safety margin for the characteris-
tic value
first failure
Frequency
NT
------ 12 _ _ __
Nk __O - - - - _
N (log)
I
mean of the sample
charactristic value
design value Fig. (6.4)-1 Distribution of action and
resistance
With these definitions the criterion for
assessment will be:
logNT - k1' (X 'Stdv +k,.'Stdv - ~'Stdv > 10gNd
with Stdv standard deviation of the sample
ex from table of variance order statistics
k% from table of expected values of normal order statistics
For more details see ref. [35, 36].
With the formula or the table {6.4}-3 for k the different values of F can be calculated,
depending on number of test specimens n and on the assumed standard deviation Stdv
of the test specimens in terms of log N.
logNT-k'Stdv > 10gNd
k = ~+(X·kl-k,.
logF = k'Stdv
Tab. {6.4}-3: Values k for testing until first failure
B ~.44 I~.77 I~.48 I~.28
page 121
10
1.07
II

6.4.3 Statistical Considerations for Partial Safety Factors
No general recommendations on partial safety factors are given. For special fields of
application, tables of safety factors may be established. Table {6.3}-4 shows a possible
example which may be adjusted according to the special requirements of the individual
application.
Tab. {6.4}-4: Possible example for partial safety factors 'YM for fatigue resistance
Partial safety factor 'YM - Fail safe and damage Safe life and infinite life
Consequence of failure tolerant strategy strategy
Loss of secondary struc- 1.0 1.15
tural parts
Loss of the entire struc- 1.15 1.30
ture
Loss of human life 1.30 1.40
page 122

7 REFERENCES
General:
[1] ISO 2394
General principles on reliability for structures.
Second edition 1986-10-14
[2] Niemi E.
Recommendations concerning stress determination for fatigue analysis of
welded components.
IIW doc. XIII-1458-92/XV-797-92
[3] Gurney T.R.
Fatigue of Welded Structures.
Cambridge University Press, UK, 1978
[4] Maddox, S.l.
Fatigue Strength of Welded Structures.
Abington Publishing, Abington UK, 1991
[5] Radaj D.
Design and analysis of fatigue resistent welded structures
Abington Publishing, Abington Cambridge, U.K. 1990
[6] Hobbacher A. et al.
Design recommendations for cyclic loaded welded steel structures
IIW doc. XIII-998-811XV-494-81; Welding in the World, 20(1982), pp. 153-
165
Geometric stress procedure:
[7] Huther M. and Henry l.
Recommendations for hot spot stress definition in welded joints.
IIW doc. XIII-1416-91
[8] Huther M, Parmentier G. and Henry l.
Hot spot stress in cyclic fatigue for linear welded joints.
IIW doc. XIII-1466-92/XV-796-92
page 123

Effective notch stress procedure:
[9] Petershagen H.
A comparison of approaches to the fatigue strength assessment of welded
components
IIW document XIII-1208-86, 1986
[10] Petershagen H.
Experiences with the notch stress concept according to Radaj (transl.)
15. Vortragsveranstaltung des DVM Arbeitskreises Betriebsfestigkeit, Ingol-
stadt 18.-19.10.1989
[11] Olivier R., Kottgen V.B., Seeger T.
Welded connection I: Fatigue assessment of welded connections based on local
stresses (transl.)
Forschungskuratorium Maschinenbau, Bericht No. 143, Frankfurt 1989 (143
pages)
[12] Kottgen V.B., Olivier R., Seeger T.
Fatigue analysis of welded connections based on local stresses
IIW document Xm-1408-91, 1991
Fracture mechanics:
[13] Murakami Y.
Stress Intensity Factors Handbook
Pergamon Press, Oxford U.K. 1987
[14] Newman J.C. and Raju LS.
Stress intensity factor equations for cracks in three-dimensional finite bodies.
ASTM STP 791 1983, pp. 1-238 - 1-265.
Stress intensity factors for internal surface cracks in cylindrical pressure ves-
sels.
Journal of Pressure Vessel Technology, 102 (1980), pp. 342-346.
An empirical stress intensity factor equation for the surface crack.
Engineering Fracture Mechanics, vol 15. 1981, No 1-2, pp. 185-192.
[17] Frank K.H. and Fisher J.W.
Fatigue strength of fille welded cruciform joints.
J. of the Structural Div., Proc. of the ASCE, vol 105 (1979) pp. 1727-1740
page 124

[18] Folias E.S.
Axial crack in pressurized cylindrical shell.
Int. J. of Fracture Mechanics, vol 1 (1965) No.2, pp 104
[19] Hobbacher A.
Stress intensity factors of welded joints.
Engineering Fracture Mechanics, vol 46 (1993), no 2, pp. 173-182, et vol 49
(1994), no 2, p. 323.
[20] Maddox S.J., Lechocki J.P. and Andrews R.M.
Fatigue Analysis for the Revision of BS:PD 6493: 1980
Report 3873/1186, The Welding Institute, Cambridge UK
Fatigue strength modifications:
[21] 0rjasreter, O.
Effect of plate thickness on fatigue of welded components.
IIW doc. xm-1582-95 1XV-890-95
Weld imperfections:
[22] IIW guidance on assessment of the fitness for purpose of welded structures.
IIW doc. SST-1157-90
[23] Hobbacher A. et al.
Recommendations for assessment of weld imperfections in respect of fatigue.
IIW doc. XIII-1266-88/XV-659-88
[24] Guidance on some methods for the derivation of acceptance levels for defects
in fusion welded joints.
British Standard Published Document 6493: 1991
[25] Ogle M.H.
Weld quality specifications for steel and aluminium structures.
Welding in the World, Vol. 29(1991), No, 11112, pp. 341-362
Stress spectrum:
[26] Endo T. et al.
Fatigue of metals subjected to varying stress - prediction of fatigue lives
(transl.) Kyushu District Meeting of the JSME, Nov. 1967.
also: Rain flow method - the proposal and the applications.
page 125

Memoir Kyushu Institut of Technical Engineering, 1974.
[27] Standard Practice for Cycle Counting in Fatigue Analysis.
ASTM E 1049-85
Damage calculation:
[28] Palmgren, A.
On life duration of ball bearings (transl.).
VDI-Z. vol. 68(1924), pp 339-341
[29] Miner, A.M.
Cumulative damage in fatigue.
J. Appl. Mech. September 1945. pp 151-164.
[30] Haibach E.
Modified linear damage accumulation hypothesis considering the decline of the
fatigue limit due to progressive damage (transl.)
Laboratorium fUr Betriebsfestigkeit, Darmstadt, Germany, Techn. Mitt. TM
50/70 (1970)
[31] Hobbacher A.
Cumulative fatigue by fracture mechanics.
Trans. ASME Series E, J. Appl. Mech. 44(1977), pp. 769-771
Fatigue testing:
[32] Lieurade H.P.
Fatigue Testing of Welded Joints
IIW doc. XIII-1516-93 (ISO porposal)
Quality and safety considerations:
[33] ISO 6520: 1982 (EN 26520: 1982)
Weld irregularities
[34] ISO 5817: 1992 (EN 25817: 1992)
Quality groups of welds
[35] Ruther M.
Uncertainties, Confidence Intervals and Design Criteria
IIW dec. XIII-1371-90
page 126

[36] Maddox S.l.
Statistical Analysis of Fatigue Data Obtained from Specimens Containing many
Welds
IIW doc. JWG-XIll-XV-l22-94
page 127

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