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This article was downloaded by: [ABM Utvikling STM / SSH packages]On: 21 July 2009Access details: Access Details: [subscription number 792960683]Publisher Taylor & FrancisInforma Ltd Registered in England and Wales Registered Number: 1072954 Registered office: Mortimer House,37-41 Mortimer Street, London W1T 3JH, UK
Ships and Offshore StructuresPublication details, including instructions for authors and subscription information:http://www.informaworld.com/smpp/title~content=t778188387
Assessment of design criteria for fatigue cracking from weld toes subjected toproportional loadingI. Lotsberg aaDepartment for Offshore Structures, Hvik, Norway
Online Publication Date: 01 June 2009
To cite this ArticleLotsberg, I.(2009)'Assessment of design criteria for fatigue cracking from weld toes subjected to proportionalloading',Ships and Offshore Structures,4:2,175 187
To link to this Article DOI 10.1080/17445300902733998URL http://dx.doi.org/10.1080/17445300902733998
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Ships and Offshore Structures
Vol. 4, No. 2, 2009, 175187
Assessment of design criteria for fatigue cracking from weld toes subjected
to proportional loading
I. Lotsberg
Department for Offshore Structures, DNV, Veritasveien 1, 1322 Hvik, Norway
(Received 20 December 2008; final version received 9 January 2009)
For fatigue design it is necessary to provide guidelines on how to calculate fatigue damage at weld toes based on S-N datawhen the principal stress direction is different from that of the normal direction to the weld toe. Such stress conditionsare found in details in different types of plated structures. Some different fatigue criteria for these stress conditions arepresented in design standards on fatigue design. Criteria used by the International Institute of Welding (IIW), Eurocode,British Standards and in the DNV (Det Norske Veritas) standards have been assessed against some relevant fatigue test datapresented in the literature. Only proportional loading conditions have been considered here. (By proportional loading it isunderstood that the principal stress direction is kept constant during a load cycle.) An alternative equation for calculation
of an equivalent or effective stress range based on stress normal to the weld toe and shear stress at the weld toe has beenproposed. The proposed methodology can be used for nominal S-N curves, which can be used together with a hot spot stressS-N curve with stresses read out from finite element analysis. The different design criteria are presented in this paper togetherwith recommendations on analysis procedure.
Keywords: design criteria; fatigue; welded plate structures; weld toe; principal stress direction; proportional loading
1. Introduction
For fatigue design it is necessary to have guidelines on
how to calculate fatigue damage at weld toes based on S-N
data when the principal stress direction is different from
that of the normal direction to the weld toe. Details of
such stress conditions are found in different types of plated
structures, such as at connections with soft brackets and
at tubulars penetrating plates in ship structures, e.g. DNV
CN 30.7 (2005) and Lotsberg (2004). Some fatigue design
standards have advised to use the largest principal stress
range within 45 to the normal to the weld toe together
with an S-N curve derived for stress ranges normal to the
weld toe for fatigue design. Reference is made to the IIW
(1996), British Standards Institution 5400 (1980), BS 7608
(1993) and DNV CN 30.7 (2005).
The International Institute of Welding (IIW) (2007) de-
cided to change the angle for largest principal stress range
direction from 45 to 60, which is now included in the
present version of the IIW fatigue design guidelines. Thesame revision was also made in DNV-RP-C203 (2005).
During actual design cases it has been found that the new
criterion can have significant impact on the design of some
special details andit is observed that designers have difficul-
ties meeting the required fatigue life at these hot spots when
using this procedure. Therefore, it was decided to make a
further assessment of recommended design criteria based
Email: [email protected]
on a review of some relevant fatigue test data from the liter-
ature. This work is presented in more detail in the following.
2. Fatigue test results for inclined welds
from literature
A literature search has been performed in order to find fa-tigue test data where the principal stress direction relative
to the weld toe has been a varying parameter. It was found
that such fatigue tests have been performed by Kim and
Yamada (2004) using test specimens as shown in Figure 1.
They presentedthe test results at an IIW conference in 2004.
Fatigue test data from specimens shown in Figure 1(b) and
test specimens in Figure 1(a), where fatigue cracking oc-
curred at the straight part of the weld toe, are selected for
assessment. The fatigue test data used in the present assess-
ment are listed in Table 1. The following notations are used
on test specimens in this table: G for gusset specimen and N
for non-load-carrying cruciform specimen. Specimen no-tation also include angle as defined in Figure 2 and the
specimen number.
Reference is made to Figure 2 for calculation of stresses
in the test specimen. Equilibrium in the loading direction
of the test specimen gives
(//sin + cos )w/ cos = 1w (1)
ISSN: 1744-5302 print/ 1754-212X online
Copyright C 2009 Taylor & Francis
DOI: 10.1080/17445300902733998http://www.informaworld.com
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176 I. Lotsberg
Figure 1. Fatigue test specimens used by Kim and Yamada (2004).
Equilibrium in the transverse direction to the specimen
gives
//cos = sin (2)
From these equations the following stresses are derived:
// = 1sin cos ,
= 1cos2
(3)
The stresses as function of unit stress for varying values of
are shown in Figure 3.
3. Design procedures
3.1. Procedure in Eurocode 3 (2005)
The procedure in Eurocode 3 (2005) is a summation of
calculated fatigue damages from normal and shear stress
ranges at the weld toe. This can be presented in the form of
a design equation as
D+D= 1.0 (4)
3.2. Procedure in IIW (2007)
Two alternative procedures are presented in the latest rec-
ommendations on fatigue design by IIW (2007).
1. Principal stress direction.
2. Quadratic interaction of allowable normal stress range
and shear stress range.
The different methodologies are presented in detail in
the following sections:
Principal stress direction in Section 2.2.3.1 of IIW
(2007):
In the case of biaxial stress state at the plate surface,
it is recommended to use the principal stress which is ap-
proximately in line with the perpendicular to the weld toe,
i.e. within 60. The other principal stress may be anal-
ysed, if necessary, using the fatigue class for parallel weldsin the nominal stress approach. Reference is also made to
Figure 4.
Quadratic interaction of allowable normal and shear
stress ranges of IIW (2007):
The effects of combination of normal and shear stresses
shall be verified by
S,d
R,d
2+
S,d
R,d
2= CV (5)
whereR,dorR,dis the design resistance stress range
for the specified number of cycles and the appropriate FAT
class for normal and shear stresses at the weld toe. S,dorS,dare the corresponding design stress ranges.CVis
a comparison value, which is given as 1.0 for proportional
loading in table 4.1 of IIW(2007).
3.3. Procedure by Kim and Yamada (2004)
Kim and Yamada (2004, 2005) proposed to use the follow-
ing expression for effective stress:
Eff= 1 cos (6)
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Ships and Offshore Structures 177
Table 1. Fatigue test data (from Kim and Yamada (2004)).
Number of Stress rangeSpecimen cycles (MPa) Comments
G0-01 216,000 190G0-02 237,000 190
G0-03 1,564,000 120G0-04 3,428,000 98G30-03 603,000 190G30-04 608,000 190G45-05 1,447,000 190G45-06 735,000 204G45-07 1,278,000 190G45-08 982,000 190G45-09 2,270,000 152 Run-outN0-01 198,000 206N0-02 170,000 203N0-03 470,000 160N0-04 556,000 160N0-05 1,415,000 136N0-06 630,000 136
N0-07 990,000 136N0-08 2,788,000 113N0-09 6,764,000 113 RunoutN15-01 360,000 206N15-02 324,000 203N15-03 479,000 161N15-04 867,000 160N15-05 760,000 160N15-06 1,577,000 136N15-07 1,739,000 136N15-08 984,000 136N15-09 2,366,000 123N15-10 4,860,000 123 Run-outK-30-01 502,000 206K-30-02 389,000 203
K-30-03 1,264,000 174K-30-04 2,053,000 159K-30-05 1,620,000 159K-30-06 6,449,000 138K-30-07 10,000,000 138 Run-outK-30-08 10,000,000 123 Run-out
where the stress 1 and the angle are defined in
Figure 2.
3.4. Alternative procedure
A combined stress range (or effective stress range), takinginto account the stress normal and the shear stress along the
weld toe can be expressed in the following form:
Eff=
2 +
2// (7)
where the stress components are explained in Figure 2. The
S-N category will depend on the type of detail in relation
to the normal stress. This will result in different values
as presented in Table 2. The combined stress range should
Figure 2. Definition of symbols and stress components.
be used together with an S-N curve that is selected as if this
stress was acting normal to the weld toe.
The details tested in Figure 1 are classified as E follow-
ing DNV-RP-C203 for small thicknesses of the attachments
and F for larger thicknesses when = 0. The test results for
= 0 are presented in Figure 5 together with the E-curve.
(Reference is made to Table 3 for relation between notations
on S-N curves used in DNV-RP-C203 (2005), IIW (2007)
and Eurocode (2005)).
For presentation of mean S-N curves it is assumed that
a standard deviation in logarithmic format is 0.20. (The
design curve is defined as mean minus two standard devia-
tions assuming the test data to follow a normal distribution
in a logarithmic format.)
S-N category C2 may be used for continuous shear
stress in a full penetration weld according to Table A.8 in
DNV-RP-C203. Assuming that the shear stress is classified
as C2, the following equation for combined or effective
stress is derived from Table 2 when the effective stress is
combined with S-N curve E:
Eff=
2 + 0.64
2// (8)
The basis for this equation is also illustrated in Figure 6
where 1 = // at = 45. The stress components in
Equation (8) are to be combined with different S-N curves
as shown in Figure 6. (When1is acting normal to a weld
toe, it is classified as an E detail or FAT 80. When the detail
is subjected to shear along the weld, S-N curve C2 or FAT
100 should be applied.)
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178 I. Lotsberg
0.0
0.2
0.4
0.6
0.8
1.0
1.2
0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0 90.0
Angle (deg)
Factor
Tau
Sigma normal
Figure 3. Stress as function of unit stress for varying value of.
4. Comparison of design procedures with fatigue
test data
4.1. Eurocode (2005)
The detail shown in Figure 1 is classified as FAT 80 fol-
lowing Eurocode 3 (2005) and IIW (2007) for = 0. This
is the same as the E-curve in DNV-RP-C203. The test re-
sults for = 0 are presented together with the E-curve in
Figure 5.
The S-N curves for stress range normal to the weld toe
and shear stress can be presented as
N = am
N// = a//m//
(9)
For stress normal to the weld the design S-N curve is FAT
80 withm = 3.0. The design S-N curve for shear stress in
Eurocode 3 and IIW is FAT 100 with m = 5.0.
Figure 4. Figure from IIW (2007) showing stress to be used for fatigue analysis.
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Ships and Offshore Structures 179
Table 2. Values for based on S-N curves in DNV-RP-C203(2005).
Stress direction parallelwith the weld
Stress direction normalto the weld toe C C1 C2
D 0.518 0.646 0.810E 0.409 0.510 0.640F 0.322 0.402 0.504F1 0.254 0.316 0.397F3 0.201 0.250 0.314G 0.160 0.199 0.250
A scaling of the stress ranges in the fatigue test data is
performed such that the test data can be presented forn =
106 cycles. The following scaling of stress are made for
comparison with Eurocode 3 and IIW (m = 3.0 for normal
stress and 5.0 for shear stress):
106 = test
ntest106
1/3.0//106 = //test
ntest106
1/5.0 (10)
From the equation for summation of damages in Equation
(4) and (9) the following expression for shear stress resis-
tance for Eurocode 3 is derived:
R,d= a// 1
n
3.0S,d
a
1/5.0
(11)
Table 3. Relations between notations in DNV-RP-C203and IIW and Eurocode 3.
DNV-RP-C203 IIW and Eurocode 3
B1 160B2 140
C 125C1 112C2 100D 90E 80F 71F1 63F3 56G 50W1 45W2 40W3 36
This equation together with fatigue test data is shown inFigure 7. It is observed that this figure shows a rather small
interaction effect between normal and parallel stress ranges
at = 45.
4.2. IIW(2007)
The IIW quadratic interaction equation on stress reads
S,dR,d
2
+ S,dR,d
2
= 1.0 (12)
10
100
1000
Number of cycles
Stressrange(MPa)
G0
N0
E design
E mean
100,000 1,000,000 10,000,000
Number of cycles
Stressrange(MPa)
G0
N0
E design
E mean
Figure 5. Test data for principal stress normal to the weld.
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180 I. Lotsberg
1
// C2 (FAT 100)
E (FAT 80)1
// C2 (FAT 100)
E (FAT 80)
Figure 6. Illustration of stresses1 = //when = 45 that are
combined with different S-N curves.
From Equations (9) and (12) the following expression for
shear stress resistance for IIW is derived:
R,d=a//
n
1/5 1 2S,d
n
a
2/3(13)
This equation together with test data is shown in Figure 8.
It is observed that the fatigue test data, also at = 45,
are in good agreement with the mean line for quadratic
interaction on stress as shown.
4.3. Present proposal
For comparison of Equation (8) with fatigue test data, the
fatigue test data are scaled with respect to stress range to
correspond to 106 cycles. The following scaling of stress
range is made for comparison with DNV-RP-C203 (inverse
negative slope of S-N curve m = 3.0):
106 = test
ntest106
1/3.0//106 = // test
ntest106
1/3.0 (14)
The test data for different principal stress range directions
are presented in Figure 9. It is observed that there is a good
correspondence between the test data and the proposed de-
sign equation for effective stress.
4.4. Kim and Yamada (2004) and the present
proposal
A comparison using different equations for effective stress
is presented in Figure 10. The effective stress from Equation
(6) (Kim and Yamada 2004) is compared with the present
proposal from Equation (8). As the present proposal fits the
test data well, it may be concluded that the procedure by
Kim and Yamada (2004) is slightly on the conservative side.
From Figure 10 it is observed that the effective stress is
reduced bya factor0.63at anangle =45. This means that
using theprincipal stress withinan angle45 to thenormal
to the weld toe becomes conservative for large angles.
0
20
40
60
80
100
120
140
160
0 50 100 150 200 250
Normal stress (MPa)
Para
llelstress(MPa)
Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01N30-01
0
20
40
60
80
100
120
140
160
0 50 100 150 200 250
Normal stress (MPa)
Para
llelstress(MPa)
Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01N30-01
Figure 7. Test data presented in format of interaction equation in Eurocode (2005).
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Ships and Offshore Structures 181
0
20
40
60
80
100
120
140
160
0 50 100 150 200 250
Normal stress (MPa)
Parallelstress(MP
a)
Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01
N30-01
0
20
40
60
80
100
120
140
160
0 50 100 150 200 250
Normal stress (MPa)
Parallelstress(MP
a)
Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01
N30-01
Figure 8. Test data presented in format of IIW quadratic interaction equation on stress range components.
5. Comparison of design procedures
and recommended approach
A review of fatigue test data considering principal stress
direction relative to the weld toe geometry has been per-
formed. Based on this assessment one may reconsider the
text related to Figure 4 from IIW (2007).
The IIW method with the calculation of allowable stress
ranges for stress normal to the weld and shear stress sep-
arately and using a quadratic interaction equation on these
is considered to fit test data very well. A methodology with
adding the damages from these stress components together,
which is used by Eurocode (2005), is not that good.
0.00
20.00
40.00
60.00
80.00
100.00
120.00
140.00
160.00
180.00
0 50 100 150 200 250
Normal stress (MPa)
Parallelstress(MPa) Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01N30-01
0.00
20.00
40.00
60.00
80.00
100.00
120.00
140.00
160.00
180.00
0 50 100 150 200 250
Normal stress (MPa)
Parallelstress(MPa) Design
Mean
G0-01
G30-03
G45-05
N0-01
N15-01N30-01
Figure 9. Test data for principal stress having different angles with the normal to the weld toe compared with proposed effective stressrange.
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182 I. Lotsberg
0.0
0.2
0.4
0.6
0.8
1.0
1.2
0 10 20 30 40 50
Angle between normal to the weld and principal stress direction
Effectivestress
Present proposal
Kim and Yamada (2004)
Figure 10. Comparison of effective stress.
An alternative design equation has been proposed that
combines the stress normal to the weld and the shear stress
at the weld toe into an effective stress range that can be
entered into a single S-N curve for calculation of number
of cycles to failure. This design approach is considered to
be efficient for use together with stresses read out from
finite element analyses.
The general expression for effective stress is derived
from Equation (7) where values are derived from Table 2.
If the hot spot stress is derived by extrapolation of stresses to
the weld toe or to the intersection line from read-out-points
t/2 and 3t/2 as explained in DNV-RP-C203, this hot spot
stress should be combined with S-N curve D. This means
that = 0.81 in Equation (7) for calculation of effective
stress range. If the hot spot stress is based on a read-out-
point at t/2, the hot spot stress should be combined with the
E-curve and = 0.64 in Equation (7).
It is realised that the present classification of details
with guidance on S-N curve is not refined enough for some
special details in DNV-RP-C203 (2005). Therefore, a more
detailed classification is proposed as shown in Figure 11
and Table 4.
Figures 11(a and b) are intended to be used for nominal
stress analyses. The selection of the E and F curves depends
on the thickness of attachment as presented in Table A.7 of
DNV-RP-C203.
Figure 11(c) is intended to be used in special cases
when using the hot spot stress methodology as presented in
DNV-RP-C203 (2005). Figure 11(c) can be used together
with the hot spot stress methodology in general.
The stress range in both the two principal directions
should be assessed with respect to fatigue. Here a design
criterion for within an angle 45 to the normal to the
weld has been assessed against fatigue test data. For a prin-
cipal stress direction 45 < 90, anS-N curve for stress
direction parallel with the weld can be used due to the ef-
fective stress reduction factor of 0.63 at = 45 as was
given in Section 4.4.
Table 4. Classification of details and selection of S-N curve.
Anglein Figure 11
Detail classified asF for stress direction
normal to the weld
Detail classified as E forstress direction normal
to the weld
S-N curve when usingthe hot spot stress
methodology
030 F E D3045 E D C24560 D C2 C26075 C2 C2 C2
7590 C2 C2 C2
A higher S-N curve may be used in special cases. See Table A-3 in DNV-RP-C203 for further information. http://webshop.dnv.com/global/category.asp?c0=2624&c1=2627.
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Ships and Offshore Structures 183
C2C2
F FEE
DD
Principal stressdirection
Weldtoe
Section
C2C2
F FEE
DD
Principal stressdirection
Weldtoe
Section(a)
(b)
(c)
C2C2
E EDD
Principal stressdirection
Weldtoe
Section
C2C2
E EDD
Principal stressdirection
Weldtoe
Section
C2C2
D D
Principal stressdirection
Weldtoe
Section
C2C2
D D
Principal stressdirection
Weldtoe
Section
Figure 11. Classification of details and selection of S-N curve: (a) Detail classified as F for stress direction normal to the weld, (b) Detailclassified as E for stress direction normal to the weld, (c) S-N curve when using the hot spot stress methodology.
Different design criteria and interaction equations are
presented in Figure 12 and Figure 13 for comparison of
design criteria at 106 and 107 cycles respectively.
6. Derivation of hot spot stress using finite element
analysis
Two alternative methods can be used for hot spot stress
derivation in the revised DNV-RP-C203 (2008). These are
described as follows:
Method A:
For modelling with shell elements without any weld
included in the model, a linear extrapolation of the stresses
to the intersection line from the read-out points at 0.5t and
1.5t from the intersection line can be performed to derive
hot spot stress.
For modelling with three-dimensional elements with
the weld included in the model, a linear extrapolation of the
stresses to the weld toe from the read-out points at 0.5t and
1.5t from the weld toe can be performed to derive hot spot
stress.
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184 I. Lotsberg
0
20
40
60
80
100
120
140
0 20 40 60 80 100 120
Normal stress (MPa)
Parallelstress(MPa)
Present proposal
IIW (2007)
Eurocode (2005)
Figure 12. Comparison of design equations at 106 cycles.
Thenotations for stress components are shown in Figure
14 and Figure 15.
The effective hot spot stress to be used together with
the hot spot S-N curve D (FAT 90) is derived as
Eff= max
2 + 0.81
2//
1 |2|
(15)
where
= 0.90 if the detail is classified as C2 with stress
parallel to the weld at the hot spot (ref. Table A-3 in DNV-
RP-C203 (2008)).
= 0.80 if the detail is classified as C1 with stress
parallel to the weld at the hot spot (ref. Table A-3).
= 0.72 if the detail is classified as C with stress
parallel to the weld at the hot spot (ref. Table A-3).
0
10
20
30
40
50
60
70
80
0 10 20 30 40 50 60
Normal stress (MPa)
Parallelstress(MPa)
Present proposal
IIW (2007)
Eurocode (2005)
Figure 13. Comparison of design equations at 107 cycles.
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Ships and Offshore Structures 185
Principal stressdirection
Weld
toe
Section
Fatigue crack
//
//
Principal stress
direction
Weld
toe
Section
Fatigue crack
//
//
//
//
Figure 14. Fatigue cracking along weld toe.
The principal stresses are calculated as
1 = +//
2+
1
2
//
2+ 4 2//
2 = +//
2
1
2
//
2+ 4 2//
(16)
The first equation for effective stress (Equation (15)) is
made to account for the situation with fatigue cracking
along a weld toe as shown in Figure 14 and the second and
third equations are made to account for fatigue cracking
when the principal stress direction is more parallel with
the weld toe as shown in Figure 15.
Principal stressdirection Weld
toe
Section
Fatigue crack
//
//
Principal stressdirection Weld
toe
Section
Fatigue crack
//
//
//
//
Figure 15. Fatigue cracking when principal stress direction is more parallel with weld toe.
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186 I. Lotsberg
Method B: For modelling with shell elements without
any weld included in the model the hot spot stress is taken
as the stress at the read-out point 0.5t away from the inter-
section line.
For modelling with three-dimensional elements with the
weld included in the model the hot spot stress is taken as
the stress at the read-out point 0.5t away from the weld toe.The effective hot spot stress to be used together with
the hot spot S-N curve D (FAT 90) is derived as
Eff= max
1.12
2
+ 0.81 2
//
1.12 1
1.12 |2|
(17)
where,1and2are explained under Method A.
The first equation for effective stress (Equation (17))
is made to account for the situation with fatigue cracking
along a weld toe as shown in Figure 14 and the second andthird equations are made to account for fatigue cracking
when the principal stress direction is more parallel with the
weld toe as shown in Figure 15.
7. Conclusions
The purpose of the present assessment has been to arrive
at guidelines on how to calculate fatigue damage at weld
toes based on S-N data when the principal stress direc-
tion is different from that of the normal to the weld toe.
Some different fatigue criteria have been assessed together
with fatigue test data from the literature. Only proportionalloading has been considered here.
The method used by Eurocode (2005) is to calculate the
fatigue damage due to stress range normal to the weld toe
and the damage due to the shear stress at the weld toe and
then adding the damages together. This sum should be less
than 1.0. It is observed that this method shows somewhat
low interaction effect between normal stress andshear stress
when compared with the test data.
IIW (2007) presents two methods for proportional load-
ing. The first one is to calculate the principal stress at the
weld toe on a nominal basis. If the angle between the prin-
cipal stress and the normal to the weld toe is less than 60,
this principal stress is used together with S-N curves for a
detail with stress acting normal to the weld toe.
The second method in IIW (2007) is to calculate allow-
able stress ranges for stress normal to the weld and shear
stress separately and use a quadratic interaction equation
on these.
From the present assessment it is found that the first
method is considered to be conservative and it should be
explained that this approach is conservative for large angles
and that the document includes more accurate alternatives
that can be recommended to be used. The second method is
found to be in good agreement with fatigue test data and is
the preferred methodology based on comparison with test
data.
An alternative equation for calculation of an effective
stress range based on stress normal to the weld toe and shear
stress has been proposed. The equation for effective stress
range reads
Eff=
2 +
2// (18)
where
= stress normal to the weld,
// = stress parallel with the weld,
= factors from Table 5.
The -factor is derived from the S-N curve constants
in the design standard such that the calculated fatigue life
using this equation for effective stress equals that using
nominal S-N curves for stress parallel with the weld toe.
Equation (18) is considered to be efficient for calcula-
tion of fatigue life when used together with the hot spot
stress concept (or structural stress concept) with stresses
read out from finite element analysis. This methodology
can also be used for nominal stress S-N curves.
The new alternative design approaches are included in
a revision of DNV-RP-C203 that was issued in April 2008.
Table 5. Recommended -factor for design.
-factor
S-N curve for stressnormal to the weld
S-Nclassification
C2 (FAT100) for pureshear stress
S-Nclassification
C1 (FAT112) for pureshear stress
D (FAT 90)When used as nominal S-N curve
or hot spot S-N curve byextrapolation of stress to the
weld toe or the intersectionline from read-out points 3t/2andt/2 (t= plate thickness)
0.81 0.64
E (FAT 80)When used as nominal S-N curve
or hot spot S-N curve by stressfrom read-out pointst/2 (t=plate thickness) from the weldtoe or the intersection line
0.64 0.50
F (FAT 71) 0.50 0.40F1 (FAT 63) 0.40 0.31F3 (FAT 56) 0.31 0.25G (FAT 50) 0.25 0.20
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