Fatigue Testing of Full

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Fatigue testing of full-scale girth welded

pipes under variable amplitude loading

Yan-Hui Zhang and Stephen J Maddox, TWI Limited, Cambridge, UK

Paper presented at OMAE 2012 31st Annual Conference on Ocean, Offshore and ArcticEngineering, 1-6 July 2012, Rio de Janeiro, Brazil. Paper No.83054

Abstract

In the fatigue design of steel catenary risers there are concerns regarding the fatigue damage togirth welds from low stresses, below the constant amplitude fatigue limit, in the loading

spectrum and the validity of Miner's cumulative damage rule. In both cases there is increasingevidence that current design methods can be non-conservative. These fundamental issues wereaddressed in a recent JIP. A key feature was development of the resonance fatigue testing rigs to

enable them to test full-scale pipes under variable amplitude loading. Such tests were performed

under a loading spectrum representative of that experienced by some risers, with many testslasting over 100 million cycles to investigate the fatigue damage due to small stresses as well asthe validity of Miner's rule. However, the resonance rigs are only capable of producing spectrum

loading by gradually increasing or decreasing the applied load, whereas more 'spiky' random

load sequences may be relevant in practice. Therefore the programme also included fatigue testsin conventional testing machines on strip specimens cut from pipes to compare the two types of

loading sequence. This paper presents the results of these tests, conclusions drawn and

recommendations for changes to current fatigue design guidance for girth welded pipes regardingthe definition of the fatigue limit, allowance for the damaging effect of low stresses and thevalidity of Miner's rule.

1 Introduction

The majority of structures and components are subjected to variable amplitude (VA) loading in

service. Fatigue design of welded joints in such structures is normally based on data obtainedunder constant amplitude (CA) loading, usually in the form of a design S-N curve, used in

conjunction with Miner's rule to estimate the damage introduced by the different magnitudes of

stress cycles in the service stress history. For a VA stress spectrum consisting of ni cycles at

stress range Si and an S-N curve from which the CA life expected at Si is Ni, where i = 1,2,3, etc,Miner's rule states that Σ (ni/Ni)=1 at failure.

An implicit assumption in Miner's rule is that the fatigue damage due to the application of a particular stress cycle in a VA loading sequence is exactly the same as that due to the same stress

cycle under CA loading. However, there is extensive evidence [Tilly, 1985; Gurney, 2006; Berger et al, 2002;

Zhang and Maddox, 2009] to suggest that it can be higher, with the result that under certain circumstances

Miner's rule is unsafe (ie the Miner's sum Σ (n/N) <1.0 at failure). There is also some

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disagreement in design codes about how to allow for the damaging effect of stresses below the

CA fatigue limit (CAFL). The common approach is to extend the S-N curve beyond the CAFL

but there are various options regarding the definition of the CAFL and the slope of theextrapolated curve. This issue is particularly important in the fatigue design of deepwater steel

catenary risers (SCRs). These can experience very high numbers of low stress cycles due to

vortex induced vibration (VIV) with the result that the choice of method for accounting for themcan be highly significant in terms of the estimated fatigue life.

This situation raises a number of fundamental questions regarding the current fatigue designapproach for girth welds:

Is Miner's linear cumulative damage rule correct?

How should the constant amplitude fatigue limit (CAFL) be defined?

What fatigue damage is induced by stress ranges below the CAFL?

What is the most appropriate fatigue design method for girth welds in which most fatigue

damage is due to low stresses?

These issues were addressed in a recent Joint-Industrial Project (JIP) managed by TWI. This

paper describes the results obtained and the fatigue design guidance proposed for girth welded pipes.

2 Outline of project

The fatigue damage due to low stresses in a VA loading spectrum representative of those

experienced by risers was investigated by fatigue testing full-scale girth welded pipes under

spectrum loading. The testing was carried out using resonance testing rigs, specially modified toallow VA loading

[Zhang and Maddox, 2011]. The high loading frequency achievable in such rigs

enabled very long endurance testing (up to 2x108 cycles) to be performed in a reasonabletimescale, thus facilitating investigation of the damaging effect of small stresses in the spectrum.

A feature of the resonance rigs is that VA loading can only be applied by gradually increasing or

decreasing the applied load, whereas more 'spiky' random load sequences may be relevant in practice. In order to investigate the effect of loading sequence, comparative fatigue tests were

also performed in conventional computer-controlled fatigue testing machines on strip specimens

cut from girth welded pipes under both the sequential spectrum used for the resonance tests onfull-scale pipes and a random version of the same spectrum.

A spectrum representative of that experienced by SCRs and with fatigue damage predominantly

from small stresses was derived to investigate the effect of small stresses. Several sub-spectrawere produced from this spectrum. These had the same peak/maximum stress histograms but

different minimum stresses. Using the same approach as [Gurney, 2006]

, it was anticipated that by

successively adding progressively smaller stress ranges, Miner's sum would increasesignificantly as non-damaging stress ranges were approached. On the other hand, if the lowest

stress in a spectrum produced fatigue damage consistent with the CA S N curve extrapolated

beyond the CAFL without a slope change, Miner's sum would be expected to be almost constant.

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3 Experimental details

3.1 Specimen designs

The full-scale specimens consisted of lengths of steel pipe joined together with either one or two

girth welds located in the middle. For the latter, the welds were about one diameter apart.406mm (16 inch) outside diameter (OD) by 19.1mm wall thickness (WT) seamless steel pipe to

API 5L-X70 specification was used. The yield and tensile strengths of the parent and weldmetals, both determined from tensile tests, are given in Table 1.

Table 1 Tensile properties of the base and weld metals

MetalYield strength,

MPa

Tensile strength,

MPa

Elongation,

%

Reduction area,

%Loading direction

Base 519.0 601.0 24.0 79.0 Parallel to pipe

Weld 728.0 782.1 18.1 64.5 Parallel to weld

The specimens were manufactured by Heerema Marine Contractors (HMC). The girth welds

were made from the outside in the 5G position by PGMAW/GMAW onto copper backing. Theweld cap at each start/stop position was ground flush with the pipe surface. All welds were

shown to be acceptable to a typical SCR specification on the basis of 'Rotoscan' and magnetic

particle inspection carried out in compliance with API 1104.

Each girth weld was instrumented with eight foil resistance strain gauges, 45° apart, to allow

measurement of axial strain and its variation around the circumference on the outside surface of

pipe. The gauge centres were approximately 60mm from the weld cap toe. The nominal strains

on the inside surface were calculated by simply considering a linear stress distribution across the pipe wall thickness under the applied bending.

Strip specimens were extracted from 0.8m long ring samples containing identical girth welds to

those in the full-scale pipe fatigue test specimens. Figure 1 shows the design for the strip

specimens.



Figure 1. Resonance fatigue testing of the full-scale pipes

3.2 Fatigue tests under constant amplitude loading

3.2.1 Full-scale girth welded pipes

Full-scale girth welded pipes were fatigue tested in resonance rigs, as shown in Figure 2, in order

to establish the CA S-N curve. An axial tensile mean stress of 150MPa was applied by pressurising the pipes internally with tap water at a pressure of 34.1MPa (4,950psi). All the testswere conducted in air at ambient temperature and loading frequencies between 25 and 30Hz.

During each test, the motor speed, which controls the stress amplitude, was kept constant until

the test pipe failed or ran out.



Figure 2. Strip specimen dimensions and strain gauge locations

The nominal strains measured by the strain gauges were used in conjunction with measurementsof axial misalignment of the joint, made after testing, to calculate the local stress range in the

region of fatigue crack initiation. Since, as expected, this was the weld root, the local stress was

defined as the nominal stress range on the inside of the pipe multiplied by the stress

concentration factor (SCF), K m, due to joint misalignment adjacent to the failure position. K m was calculated using the analytical solution given in

[BS 7910 (2005)].

For those pipes containing two test welds, the first to fail was repaired by welding in order toallow testing of the second weld to continue. However, if the repair weld then failed, no further

repair was attempted and, as long as there was no evidence of fatigue cracking, the second of the

two original girth welds was declared a run-out.

3.2.2 Strip specimens

CA fatigue tests were also performed on strip specimens to establish their S-N curve. The

specimens were tested under axial loading in fatigue testing machines. Two pairs of curved

packers, with the same radius as the pipe, were used in gripping the ends of the specimen in

wedge jaws. Specimen edges were dressed round and smoothed to prevent premature fatiguefailure from an edge. As for the full-scale specimens, the initial strip specimens were tested at a

constant tensile mean stress of 150MPa. However, higher mean stresses between 230 and 380MPa were also tried as the strip specimens exhibited a higher CAFL than the full-scale pipes, as

discussed later.

The local stress range near the failure site, allowing for any misalignment-induced secondary

bending, was determined for each specimen using the measured strains. By assuming that stress

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is proportional to strain, the SCF at the weld root due to misalignment was calculated by the

following equation:

Eqn 1where εr is the strain measured at the weld root of interest and εt is the strain at the opposite weld

cap toe. The stress range at the weld toe is the product of K m and the nominal stress range, whichwas defined as the applied load divided by the cross-sectional area of specimen adjacent to the

weld.

3.3 Fatigue tests under variable amplitude loading

3.3.1 Loading spectrum

A literature review and discussions with offshore operators was first carried out to determine the

loading spectra representative of those experienced by SCRs. With the additional requirement to produce conservative design data, a spectrum with the following characteristics was proposed[Zhang and Maddox, 2011]

:

Narrow-band loading, with gradual increase and decrease of the stress range for the full-

scale fatigue tests.

High tensile mean stress, around 150MPa. This would be conservative when the testresults are applied to the touch-down position (TDP) of risers.

Maximum stress range of around 180MPa.

An approximately linear stress distribution, as a compromise between the concave-up

distribution for the top of risers and a convex-up distribution for the TDP of risers.

A long block length (i.e. the number of cycles applied before the sequence is repeated).

The proposed stress distribution was checked against a 'generic' West Africa case and showedgood agreement.

Several sub-spectra with the same maximum stress histograms but different minimum stresseswere produced from this spectrum. The characteristics of the five initial sub-spectra, VT-1 to

VT-5, as well as their stress distributions are shown in Figure 3. Subsequently, an intermediate

sub-spectrum VT-4/5, with a minimum stress range of 45MPa, was introduced. In each case, the

exceedence at the minimum stress is the block length for that sub-spectrum.






Figure 3. Stress distribution and definition for each sub-spectrum

All the sub spectra exhibit a linear stress distribution but contain different minimum nominal

stress ranges. For example, the minimum and maximum stress ranges for sub-spectrum VT-5 are

40MPa and 180MPa, respectively. Furthermore, the relative fatigue damage, defined as the percentage of fatigue damage at a stress level ΔSi against the total fatigue damage in the

spectrum,

, where m was assumed to be 3.0, gradually increases with decreasing stress range until a valueof about 28MPa. This spectrum was intended to make the results of the present VA tests

sensitive to the introduction of smaller and smaller stresses. The fatigue damage sum (ΣΔS3n)

increased with decreasing minimum stress range in the sub-spectrum, so that it increased

progressively from VT-1 to VT-5 [Zhang and Maddox, 2011]

.

3.3.2 VA tests for full-scale pipes

Details of the development of the full-scale fatigue test under VA loading have been reported[Zhang and Maddox, 2011]

. In summary, the implementation of a VA test involved several proceduresincluding establishing a polynomial relation between the motor speed in RPM (revolutions per

minute) and strain range, dividing each block (a whole sub-spectrum) into a number of sub-

blocks to simulate the sea states, establishing the relation between RPM and time for each sub-










block, data recording and processing. In each sub-block, the stress range increased from the

minimum value to a peak value and then decreased to the minimum value. The minimum value

was the same for each sub-block, but four different peak values were used, namely 140, 180, 110and 80MPa for sub-blocks A to D respectively. An example of the resulting sequence of sub-

blocks, for sub-spectrum VT-3 (minimum stress range of 60MPa), is shown in Figure 4. Sub-

spectra VT-2 to VT-5 were all divided into six sub-blocks in this way. In these, the upper parts(minimum stress range=80MPa) of the sub-blocks A - C were identical to those in sub-spectrumVT-1 as described below. The three sub-blocks D in each sub-spectrum were identical and the

upper parts of sub-block D were identical in all sub-spectra. For sub-spectrum VT-1, there were

only three sub-blocks and they were identical to the A-C sub-blocks in other sub-spectra.

Figure 4. An actual cyclic loading sequence (peak stresses in each loading and unloading

cycle) that illustrates the good repeatability of the loading block in the VA testing

3.3.3 VA tests for strip specimens

As noted previously, the significance of the practical limitation that resonance rigs can only

apply VA loading spectra in a sequential order, with gradual changes in applied stress range

( Figure 4), was checked by performing comparative tests on strip specimens in testing machinescapable of applying both sequential and random VA loading.

Sub-spectra VT-2 (with a minimum nominal stress range of 70MPa) and VT-3 (with a minimum

nominal stress range of 60MPa) were chosen for this investigation. However, to compensate for

the higher CAFL associated with the strip specimens (described later), all stress ranges in thesespectra were doubled. Thus, the maximum and minimum nominal stress ranges were increased to

360 and 140MPa for sub-spectrum VT-2 and 360 and 120MPa for sub spectrum VT-3,



respectively. The applied mean stress in all tests was 230MPa, the same as that used for most CA

tests.

The tests were performed in a computer-controlled testing machine that was programmed to

apply the stress cycles in each block from the same basic loading spectrum in either a sequential

or a random order. For the former, the loading sequence was the same as that used in the full-scale tests. For the latter, it was achieved by selecting the relative stress ranges, Pi (ΔSi/ΔSΔmax),

using a random number generator. When the whole of the first block had been applied the

process started again and subsequent blocks were applied in the same sequence. This processwas repeated until the specimen failed or achieved run-out. A segment of the loading pattern

from both sequential and random loading is shown in Figure 5. It should be emphasised that this

was an exploratory investigation simply comparing 'sequential' and 'random' loading. Precise

details of the random loading sequences seen by SCRs are not available and the present one maynot be truly representative.

Figure 5. Loading patterns of the sub-spectrum VT-2, used for VA testing of the strip

specimens:

(a) Sequential loading (same sequence as that used for the full-scale pipes);



(b) Random loading.

3.4 Post-test examinations

Comprehensive post-test examinations were carried out for all full-scale girth welds tested. For

full-scale pipes, the examination included measurements of wall thickness, weld root hi-lo, weld

root bead height (WRBH) relative to the neighbouring pipe surface at each gauge location, aswell as dye penetrant inspections of the test welds. When a test weld ran out, a macro-section

was taken at a location where the weld root bead profile was considered to be the most

unfavourable (ie highest stress concentration, generally where the hi-lo and WRBH were largest)

along the circumference of that weld.

4 Test results

4.1 Fatigue tests under CA loading

4.1.1 Full-scale pipes

Six pipes, with a total of eleven girth welds, were tested to establish the CA S-N curve. The test

results are plotted in Figure 6 where they are seen to be generally better than the [BS 7608 (1993)]

Class D mean S-N curve. Examination of the failed girth welds confirmed that fatigue cracking







always initiated at the weld root. Noting that the lowest stress that produced failure, with an

endurance of 4.4x107 cycles, was 55.4MPa, it appears that the CAFL for these girth welds may

correspond to an endurance between 5x107 and 10

8 cycles, which considering the two unbroken

welds that survived for 108 cycles, would be between 45 and 55MPa. The same was found in

relation to CA tests on girth-welded 20 inch pipe [Maddox and Zhang, 2008]

. It is apparent that the Class

D mean fatigue limit of 74MPa, corresponding to 10

7

cycles, over-estimates the value suggested by the present tests. This argument was further supported by the results of the post-testexamination on two unbroken welds which had both been tested under relatively low stresses.

One was tested under a local stress range of 54.4MPa and declared a run-out after 5.16x107

cycles while the other was tested under a local stress range of 45.2MPa and declared a run-outafter 10

8 cycles. The post-test examination revealed fatigue cracking from the weld root for both

welds, 0.26mm deep in the first and 0.14mm deep in the other.

Figure 6. Comparison of the constant amplitude fatigue performance of the full-scale and

strip specimens with the Class D mean curve.

Assuming that all the results from failed welds lie on the same S-N curve, regression analysis

produced the mean S- N curve ΔS3.25

N = 2.02 x 1013

, slightly shallower than the Class D curve,for which m = 3. However, the value of m = 3 is within the 95% confidence estimates of the

slope and therefore it would be legitimate to assume this value when fitting the S-N curve. For

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convenience this assumption was made and the resulting S- N curve became ΔS3 N=6.174 x 10

12.

This curve was used to calculate N values for the application of Miner's rule in the analyses of

the VA test results from the full scale pipes. Furthermore, since there were no failures beyond5x10

7 cycles, it was assumed that the CAFL corresponded to this endurance, giving a value mean

of 49.8MPa, 42% lower than the stress corresponding to N=107 cycles, the conventional CAFL.

4.1.2 Strip specimens

As for the full-scale specimens, initially the strip specimens were tested at a constant tensilemean stress of 150MPa. However, although the two results obtained at relatively high stress

ranges were in agreement with those obtained from the full-scale pipes, the test at a stress range

of 120MPa was a run-out at 2.67x107 cycles, which was significantly greater than the fatigue

endurance of the full-scale pipe tested at this stress level. It was thought that this difference

might be due to relaxation of welding-induced residual tensile stresses during extraction of strip

specimens. Therefore, the mean stress was increased for the additional tests at this and lower

stress levels. However, these specimens, tested under mean stresses of 250 and 380MPa, still ran

out, indicating a CAFL more than double that for the full-scale pipes.

Subsequently, further CA tests were carried out at higher stress ranges and a mean stress of230MPa. Furthermore, as post-test examinations of the tested welds suggested that the fatigue

performance of girth welds strongly depended on weld root bead profile, one strip specimen with

a hi-lo of 0.48mm (resulting in a very poor profile) was tested at a nominal stress range of110MPa. This specimen failed in a life close to the lower bound of the data scatter for the strip

specimens. Recalling that some tests ran out at higher stress ranges and higher mean stresses, it

appears that the higher fatigue limit associated with the strip specimens, as compared with the

full-scale, can be at least partly attributed to their favourable weld root profiles.

The test results are shown in comparison with those for the full scale pipes in Figure 6 . It will beseen that failed strip specimens tested under applied mean stresses of both 150 and 230MPa gaveresults in agreement with one another and with those of the full-scale pipes. The best-fit curve to

the strip specimen results had a slope of 2.47. By fixing the slope at m = 3.0, it became

ΔS3 N=6.275 x 10

12, very close to the S-N curve for the full-scale pipes. This S-N curve was used

to calculate N values for the application of Miner's rule in the analyses of the VA test results

from the strip specimens.

4.2 Fatigue tests under VA loading

4.3 Full-scale pipes

The VA test results obtained from full-scale pipes were analysed in terms of both Miner's sums

and, for graphical presentation, the equivalent constant amplitude stress range, ΔS eq. For a

particular endurance, ΔSeq is the constant amplitude stress range which, according to Miner's

linear cumulative damage rule, is equivalent in terms of fatigue damage to the VA stressspectrum. It relates to the constant amplitude S- N curve for the detail under consideration, ΔS

m N

= A (a constant), as follows:



Eqn 2

where m is the slope of the relevant S- N curve, i.e. 3 in the present case, and ΔSi is the stressrange that is applied ni times in a spectrum. Note that (Σni) is the spectrum block length N b. It

may be noted that ΔSeq values below, on or above the CA S-N curve correspond to Σ(ni/Ni)

values <1, =1 or >1, respectively.

An implicit assumption in Eq. [2] is that stresses below the CAFL are as damaging as implied by

the S-N curve extrapolated beyond the CAFL at the same slope. However, it is common practiceto assume that they are less damaging by adopting a bi-linear S-N curve with a slope change

from m to m+2 at the CAFL. As noted above, BS7608 assumes that the CAFL is the stress range

corresponding to N=107 cycles, but the present full-scale pipes suggest a lower value,

corresponding to N=5x107 cycles. To allow for this, ΔSeq was also calculated using the bi-linear

S-N curve for which the slope changed from m=3 to m=5 at 5x107, on the basis of the equivalent

fatigue damage concept expressed below:

Eqn 3

where m1 = 3 and m2 = 5. Selection of m and A on the right hand side depends on whether Seq isabove or below the slope transition point in the S-N curve. The left hand side of the equation is

the sum of the fatigue damage ratios due to stresses above (subscript 1) and below the fatigue

limit (subscript 2), respectively. Parameters A1 and A2 correspond to the S-N curve above and

below the fatigue limit, respectively.

The VA results obtained from full-scale pipes are summarised in Table 2. This includes the sub-spectrum used, its block length, the fatigue endurance, in terms of number of cycles and number

of blocks, and the calculated Miner's sum Σ(n/N). The mean S-N curve for the full-scale pipes,

ΔS3 N=6.174x10

12, was used to calculate N. However, two values of Σ(n/N) are given, one based

on N values derived from the S-N curve extrapolated beyond the CAFL without a slope change

and the other for N derived from the bi-linear S-N curve that changed slope from m=3 to 5 at

N=5x107 cycles, for stress ranges below 49.8MPa. The results are also presented in Figure 7 in

comparison with the CA results, with ΔSeq calculated using Eq.[3].



Figure 7. Comparison of variable amplitude test results for the full-scale pipes with bi-

linear versions of constant amplitude S-N curves, with VA data presented in terms of

equivalent stress range calculated using a bi-linear S-N curve with slope change from 3.0 to5.0 at 5x107 cycles.

Table 2 Summary of the VA test results for the full-scale pipes

Pipe

ID

Sub-

spectrum

Min

nominal

stress

range,

MPa

Block

length,

cycles

EnduranceAverage Miner's

sum, Σ(n/N)

Number of

cycles

Number of

blocks

No slope

change

Bi-

linear1

S23 VT-1 80 28,237 9.76x10 345.2 1.59 1.59

S24 VT-1 80 28,196 7.30x10 260.4 0.90 0.90S28 VT-2 70 85,073 3.15x10 >370.2 >2.90 >2.90

S25 VT-3 60 215,400 3.50x10 162.5 1.55 1.55

S26 VT-3 60 216,100 >1.95x10 >900.9 >10.33 >10.33

S27 VT-3 60 217,100 >1.52x10 >700.0 >7.23 >7.23

S14-W1

VT-4 50 601,600 >6.06x107 >101.4 >1.90 >1.87



S14-W2

4.22x107 70.3 1.15 1.11

S17-

W1VT-4 50 603,330

1.89x107 31.3 0.78 0.78

S17-

W2 >2.85x10

7

>47.3 >1.03 >1.02

S22-

W1VT-4 50 602,220

2.30x107 38.2 1.05 1.05

S22-

W2>1.30x10

8 >216.2 >4.72 >4.70

S29 VT-4 50 600,850 >1.59x10 >263.8 >4.30 >4.14

S20-W1

VT-4/5 45 1,106,950 >1.71x108 >155.4

>3.26 >2.89

S20-

W2>3.36 >2.99

S31-W1

VT-4/5 45 1,110,000 >1.32x108 >120.0

>4.12 >4.04

S31-

W2>4.32 >4.26

S15-W1

VT-5 40 1,671,000 >1.96x108 >115.3

>3.61 >3.16

S15-

W2>3.29 >2.81

S16-W1

VT-5 40 1,672,300 >1.43x108

>85.7

>2.77 >2.44

S16-

W2>3.27 >2.98

S30 VT-5 40 1,677,500 8.27x10 49.3 1.18 0.96

Notes:1. Slope change from 3.0 to 5.0 at the assumed slope transition point which corresponds to 5x10

7

cycles with a stress range of 49.8MPa. Symbol '>' indicates that the weld was unbroken.

2. Pipe S28 failed from corrosion pitting on the inside of the pipe remote from the weld. Theweld profile was comparatively good.

3. Pipe S30 failed from an undetected 0.65mm deep lack of fusion defect at the w

eld root.

Referring to Table 2, it will be seen that Σ(ni/Ni) depends on the sub spectrum. It varied from

0.78 to 1.59 for pipes that failed under the spectra with a minimum stress range ≥50MPa (VT-1

to VT-4),. However, as seen in Figure 7 , these values correspond to fatigue strengths that are

within the range expected considering the scatter in the CA data. Thus, Miner's rule can be

considered to be accurate for these sub-spectra. However, it may be noted that this would not be



true if the slope change in the CA S-N curve was assumed to correspond to N=107 cycles, as

indicated in Figure 7 , thus further confirming that the current design approach of locating the

slope change at 107 cycles is non-conservative. Miner's rule proved to be conservative for spectra

with the minimum stress range <50MPa (VT-4/5 and VT-5), with eight out of nine welds being

unbroken after endurances considerably more than predicted by Miner's rule (Σ(ni/Ni) > 1).

Attention is drawn to two tests that gave results that were not consistent with the others. Pipe S28

failed from a corrosion pit on the inside of the pipe with no evidence of fatigue cracking from the

weld root. Although the weld root bead profile was good it was no better than that in other weldsthat did fail from the root. Thus, its exceptionally long life could not be explained. In contrast, it

was clear that the single pipe failure under VT-5 (pipe S30) occurred because the fatigue crack

initiated at a 0.65mm deep lack of fusion defect at the weld root.

Since two full-scale pipes tested under CA loading ran-out at 108 cycles, another possibility is to

introduce a slope change at this endurance. Although not shown here, the ΔSeq values

corresponding to this bi-linear S-N curve were also calculated. Similar to Figure 7 , the ΔSeq

values were also compared with the CA S-N curve with a slope change at 10

8

cycles. It wasfound that all the VA data for endurances greater than 2x107 cycles were above the mean curve,

while those for endurances greater than 108 cycles were above the upper bound S-N curve. In

view of this, it is considered that the case for adopting the slope change at 108 cycles, rather than

5x107 cycles, is not strong and is unduly conservative. Meanwhile, using the slope change at 10

8

cycles is clearly the more conservative of the two approaches and so it might be considered

preferable for particularly critical cases.

4.4 Strip specimens

Fourteen strip specimens were tested under sub-spectrum VT-2 and four under spectrum VT-3,

half under the sequential version of the spectrum and half under random loading. The results are presented in Figure 8 in terms of ΔSeq, calculated using Eq.[2]. No attempt was made to use

Eq.[3], with a bi-linear S-N curve, because although the CAFL was not well defined it appearedas if all stress ranges in VT-2 and VT-3 were above it. Figure 8 also shows the results obtained

under sequential loading from the full-scale pipes and the S-N curves based on the combined CA

data from full-scale pipes and strips. Although not conclusive, it appears that the random loadingwas the more damaging, with most of the results lying below the mean CA S-N curve while most

sequential loading results were above it. Although there is no overlap of the results from failed

specimens it will be evident that the VA data from both the strips and full-scale pipes follow theCA S-N curve. It will also be evident that apart from one of the results for failure under the

sequential loading and three run-outs, which were all beyond the upper bound of the CA scatter-

band, all the VA data from strip specimens are within the CA scatter-band. This suggests thatMiner's cumulative damage rule was reasonably accurate for calculating the fatigue lives of thestrip specimens under VA loading, as was found for the full-scale pipes tested under sub-spectra

with all stress ranges above the CAFL.

The relatively large scatter in the strip specimen results in Figure 8 probably reflects variations

in the weld root bead profiles. However, it was observed that two strip specimens extracted from

the same original girth weld had very similar relatively poor profiles, with both a hi-lo and



WRBH of about 0.45mm. One was tested under the sequential version of VT-2 and the other

under random loading. This provided a good opportunity to investigate the effect of loading

sequence independently of a possible effect of a difference in weld root bead profile. The lifeobtained under sequential loading was ~1.3 times greater than that obtained under random

loading.

Figure 8. Variable amplitude fatigue data obtained from full-scale pipes and strip

specimens illustrating effect of loading sequence ('sequential' or 'random') on fatigue

performance.

It was also apparent that weld root bead profile was significant in the welds tested under VT-3.

Only two of the four failed, the unbroken welds surviving for lives more than 5 times those predicted by Miner's rule. Both had exceptionally good profiles, as indeed did the single

unbroken weld tested under VT-2, with both the hi-lo and WRBH being less than 0.1mm.

However, it is probably also significant in the case of VT-3 that stresses below 140MPa are

unlikely to be as damaging as implied by the CA S-N curve in the region close to the CAFL.

5 Post test examination

5.1 Failure location for the full-scale pipes



Of the 33 welds tested major fatigue cracks occurred in 17 of them. All but three failed from the

weld root bead toe, with two cap failures and one failure from a corrosion pit. All root failures

originated within 10° of the weld start position or 4° of the weld stop position. The failurelocations were often characterised by relatively poor weld root bead profiles, with the hi-lo and

WRBH ≥0.5mm.

5.2 Fracture surfaces and macro-sections

All the girth welds tested were examined in detail, either from macro-sections or directly on the

fracture surfaces of failed welds. Cracking was revealed in many run-out welds. Results from theexaminations of four weld macro-sections are described below.

Pipe S12, which contained two welds and was tested under CA loading, was unbroken after 108

cycles. No fatigue cracking was found at the weld start position in weld W1, but a 0.14mm deep

weld root crack was found at the weld stop position, where the WRBH was 0.25mm, Figure

9(a). Two sections from weld W2 of pipe S12 did not reveal any indication of fatigue cracking.

One of these sections was made at the position with the poorest weld root bead profile (WRBHand hi-lo of 0.47mm), Figure 9(b).



Figure 9. Macro-sections of the two welds in pipe S12 after being tested under CA loading

for 108 cycles:

(a) Weld W1, showing a 0.14mm deep crack at the toe of the weld root bead;



(b) Weld W2, with no evidence of cracking in the region of poorest profile (WRBH=

0.47mm). The definitions for hi-lo and WRBH are shown in the figure

Both weld W1 and 2 of pipe S15, tested under sub-spectrum VT-5, were unbroken after 1.96x108

cycles but macro-sections revealed fatigue cracking from the weld root bead toe in both. Twocracks were found in W1, one just 67µm deep at the weld start position but the other 0.19mm

deep at the weld stop position, Figure 10. A section of W2 at a location away from the weld

start/stop positions but where the WRBH was 0.58mm revealed a 93µm deep crack, Figure 11.

Therefore, the remaining endurances of the two welds were significant even though their Miner'ssums were already respectively greater than 2.8 and 3.1 (assuming a bi-linear S-N curve with a

slope change from 3 to 5 at 5x107 cycles).



Figure 10. Macro-sections of the two welds in pipe S15 after being tested under VA-5

loading (minimum stress range of 40MPa) for 1.96x108 cycles, showing fatigue cracking at

the toe of weld root bead:

(a) W1 weld, 0.19mm deep crack;



(b) W2 weld, 93µm deep crack that initiated in a region of poor profile (WRBH = 0.58mm)



Figure 11. Dependence of the ratio of the calculated fatigue life based on a bi-linear S-N

curve to that based on a single slope S-N curve on the minimum stress range in the applied

loading spectrum

6 Discussion

6.1 Failure locations

In the full-scale CA and VA tests, a total of seventeen major fatigue cracks, which either led to

failure or were more than 1mm deep, were observed, most near the weld start position and a few

near the weld stop position. No failures initiated at locations away from these two positions,

although indications of small cracks were found. For those welds which ran out, macro-sectionsrevealed that cracking often initiated at the toe of a weld root bead with a poor profile.

In the present context, it seems that the best indication of a poor weld root bead profile is the

combined weld root hi-lo and WRBH, especially the former. Not only does this lead to a sharp

corner, and hence more severe stress concentration than would arise from a smooth transitionfrom pipe to weld surface, but any associated axial misalignment would increase the local stress

by introducing secondary shell bending. From the practical viewpoint, this highlights the

importance of achieving as near perfect inner pipe wall alignment as possible when making girth

welds. If perfect alignment between pipe wall surfaces cannot be achieved due, for example, to a



difference in the wall thicknesses of the pipes, it would be preferable to place the thickness

transition on the outside where the weld toe stress concentration is less severe and can be

reduced by flush grinding the weld cap to achieve a smooth transition. This approach isrecommended in

[DNV 2010].

Although it is unquestionable that weld root bead profile has a significant influence on thefatigue performance of girth welds, other factors may also play a part. This was evident from the

fact that in some girth welds the worst weld root profile was not always at the weld start or stop

position. However, major cracking still occurred at the weld start/stop positions. Other possiblekey factors could be welding imperfections and residual stresses.

6.2 Constant amplitude fatigue limit

The fatigue limit is generally defined as the maximum stress range at which fatigue failure will

never occur under CA loading. However, as suggested by [Pyttel et al, 2011]

, a fatigue limit may not

exist. They found that even smooth specimens in various metals can fail after more than 108

cycles and concluded that a more useful property was the fatigue strength at a defined number ofcycles. The present CA test results add some support to this view in that fatigue failure occurred

after 4.4x107 cycles and fatigue cracking was found in several welds tested up to 10

8 cycles, all

far greater than the 107 cycles commonly specified by many international codes as the endurance

corresponding to the CA fatigue limit. Therefore, in this paper, fatigue limit means fatigue

strength at a particular endurance. On this basis, defining the CAFL as the fatigue strength at 108

cycles, the value for the present girth welds is between 45 and 50MPa. The S-N curve started tochange slope as this limit was approached at an endurance of ~5x10

7 cycles, similar to that

obtained previously from girth-welded 20 inch OD pipes with a fatigue performance close to

Class E [Maddox and Zhang, 2008]

, lower than that for the present 16 inch pipes.

In the present work the CAFL is significant only as a guide to the choice of location for thetransition point at which the S-N curve used in conjunction with Miner's rule changes slope. As

noted previously, for convenience the value finally adopted for the present welds was the fatiguestrength at N=5x10

7 cycles, namely 49.8MPa.

6.3 Comparison of the fatigue performance of full-scale pipes and strip

specimens cut from pipes

The fatigue performance of the present strip specimens was comparable with that of the full-

scale pipes at high applied stress ranges but superior at low (stress ranges ≤~120MPa), even

when the strip specimens were tested under very high tensile mean stress conditions. This

finding was in agreement with the results in previous work by the present authors [Maddox and Zhang,

2008] but at odds with the claim that fatigue data from strip specimens can be correlated well with

those from full-scale pipes as long as the strip specimens are tested under high tensile meanstress conditions to compensate for the loss of tensile residual stresses (Salama, 1999). It is

recalled that one strip specimen, with a poor weld root bead profile (WRBH=0.48mm), was

deliberately chosen for the CA fatigue test at a nominal stress range of 110MPa. Althoughseveral strip specimens ran out at 120MPa under a high tensile mean stress, this specimen failed,

with an endurance close to the lower bound to the combined strip specimens and full-scale pipe

http://www.twi.co.uk/technical-knowledge/published-papers/fatigue-testing-of-full-scale-girth-welded-pipes-under-variable-amplitude-loading/?vAction=swpTextOnly#dnv2010



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data, Figure 6 . Clearly, the weld root bead profile of the strip specimens also has a very

significant effect on the difference in fatigue performance between the strip specimens and full-

scale pipes in the long-life regime.

At high applied stresses, crack initiation would be easier and the fatigue life would be dominated

by fatigue crack growth. As a result, the fatigue endurance would not be so sensitive to the weldroot bead profile and the fatigue performance of the strip specimens would be comparable with

that of the pipes, as seen in the present project. Therefore, when strip specimens have to be used

to characterise the fatigue performance of full-scale girth welded pipes and the weld root bead profile of these strip specimens is favourable when compared to the worst profile of a full-scale

weld, they should be tested under fully tensile high applied stress ranges where fatigue life is

dominated by fatigue crack growth.

6.4 VA results from full-scale pipes and validity of Miner's rule

The VA test results obtained from the full-scale pipes (Table 2) indicate that, for the basic

spectrum used, Miner's rule is reasonably accurate if the minimum stress range ≥50MPa (sub-spectra VT-1, VT-2, VT-3 and VT-4). The results obtained under sub-spectra VT-3 and VT-4

were the most widely scattered, with many welds remaining unbroken. However, this is not

surprising since the fatigue damage in these spectra came mainly from stress ranges thatcorrespond to the transition regime as the CA data approach the fatigue limit, where the CA data

are also widely scattered. It will be noted that the choice of S-N beyond the CAFL, extrapolation

with or without a slope change, had little or no effect on the life calculated by Miner's rule. Thisis reasonable since all applied stress ranges were above the estimated CAFL. However, this is

not the case using a bi-linear S-N curve with the slope change at 107 cycles, as currently

recommended in many design rules. The S-N curve would then be as shown in Figure 7 .

Although the ΔSeq values corresponding to this S-N curve are not shown, they would be either

the same (tested under higher minimum stresses) or lower (tested under lower minimum stresses)than those obtained for a S-N curve with a slope change at 5x107 cycles. Since many of the data

points already lie below the mean S-N curve with a slope change at 107 cycles, the current design

approach of locating the slope change at 107 cycles would be non-conservative.

Sub-spectra VT-4/5 and VT-5 included stress ranges down to 45MPa. Consequently, shorterlives were estimated assuming the bi-linear rather than single-slope S-N curve. However, these

were still considerably shorter than the actual endurances suggesting that stress ranges below the

estimated CAFL are less damaging than implied by the bi-linear S-N curve with the slope changeat 5x10

7 cycles. Again referring to Figure 7 the same may not be the case for a slope change at

107 cycles; ideally more data, especially from welds that fail, are required to confirm this.

Meanwhile, noting from Figure 7 that the data obtained under sub-spectra VT-4/5 and VT-5 areclearly biased above the bi-linear S-N curve slope change at 5x107 cycles, it seems reasonable to

adopt bi-linear CA S-N curves with the slope change at 5x107 cycles as a safe but not unduly

conservative approach for cumulative damage calculations. The one factor that might be seen to

contradict this conclusion is the test result obtained from pipe S30 under VT-5. This failed at alower Seq value than the deduced CAFL as a result of a 0.65mm deep weld root defect. If

inspection of girth welds for offshore pipelines cannot be relied upon to detect such small defects

it would clearly be prudent to adopt a lower CAFL, corresponding to N closer to 108 cycles.



6.5 Effect of loading sequence

The VA test results from the strip specimens appear to indicate that random loading was more

damaging than the sequential loading used for testing the full-scale pipes. A comparison of thefatigue endurances between two specimens with similar weld root bead profiles suggested that

the fatigue endurance under sequential loading was about 1.3 times longer than that obtainedunder random loading.

In terms of the impact of these findings on the fatigue design of SCRs, it should be born in mind

that the random sequence was an artificial one. It was not based on actual observations of thesequence actually experienced by risers but simply randomisation of the stress spectrum. In

practice, SCRs may not see such a sequence. Indeed, it could be argued that, although actual

SCRs may experience random loads, the stress changes would be gradual as they are controlled

by environmental loading. Then the random loading sequence used in the present tests may not be truly representative and the deduced life reduction factor may be too high. However, if actual

SCRs do experience such random loading or when it is uncertain, it would be prudent to apply a

life reduction factor of 1.3 to account for the possible loading sequence effect.

6.6 Implication of current results to fatigue design of girth welds

The fatigue test results obtained from the girth welds tested under sub-spectra with different

minimum stress ranges were assessed by Miner's rule using two different CA S-N curves, one

with no slope change and the other a bi-linear curve with a slope change from 3 to 5 at 5x107

cycles (the assumed CAFL). From Table 2, it will be seen that the introduction of a slope changefrom 3 to 5 at 5x10

7 cycles made little difference to the Miner's sums for sub-spectra VT-1 to

VT-3. The effect was small for sub-spectrum VT-4 and increased for sub-spectra VT-4/5 and

VT-5. The results indicated that it would be over-conservative to adopt a single-slope S-N curve

in design for loading spectra with minimum stress ranges <50MPa. Assessments based on a bi-linear S-N curve were more accurate, although introduction of the slope change at 10

7 cycles, as

currently recommended in many design rules, would be potentially unsafe. A slope change in the

S-N curve is also in compliance with the possibility that the proportion of life spent initiating afatigue crack becomes more significant in the long endurance regime, hence producing a

shallower S-N curve.

As an illustration of the benefit of adopting a bi-linear S-N curve, Figure 11 shows the variation

in ratio of fatigue life calculated using Miner's rule in conjunction with a bi-linear S-N curve,

with the slope change at 5x107 cycles, and that obtained using a single slope curve for each sub-

spectrum. It will be seen that the relative fatigue endurances increase with decreasing minimum

stress range in the spectrum. Thus, adoption of the bi-linear curve justifies a significant increasein estimated fatigue design life, or corresponding increase in allowable stress range, when

compared with that estimated on the basis of a single-slope S-N curve. For example, for aloading spectrum with a linear stress distribution and a minimum nominal applied stress range of

10MPa, the calculated design fatigue life can be increased by a factor of 1.76 using the bi linear

S-N curve.



Although the recommendation to retain the bi-linear S-N curve approach for cumulative damage

calculations remains, the case of failure from an undetected 0.65mm deep weld root defect in

pipe S30 mentioned earlier leaves a question mark over the choice of the point at which the slopechange should be introduced. If inspection of girth welds for offshore pipelines cannot be relied

upon to detect such small defects it would clearly be prudent to extend the S-N curve to perhaps

10

8

cycles before the slope change.

One outcome from the failure of pipe S30 from a defect that could be seen as positive is that it

provides an indication of the acceptance level for weld root defects in girth welds. As seen in Figure 7 , the test result agreed very closely with the mean S-N curve obtained from defect-free

welds which itself was above the Class D mean curve. Clearly further confirmation is needed

from fatigue tests over a wide range of applied stress levels but, at this stage, it appears possible

that Class D, but even more certainly Class E, fatigue performance could be achieved from girthwelds with root defects approaching 1mm in depth. At the same time it would have to be

recognised that it may be necessary to assume a lower CAFL than that for defect-free welds.

6.7 Design considerations and recommendations for full-scale pipes underspectrum loading

The present work suggests that the fatigue performance of girth welded pipes under spectrum

loading containing significant proportions of low stresses should be evaluated using Miner's rulein conjunction with a bi-linear S-N curve with the slope change at 5x10

7 cycles. This is in

agreement with the results obtained in a previous JIP [Maddox and Zhang, 2008]

from lower design class

girth welds.

Equation.4.

When facilities for full-scale testing under VA loading are not available, the following designapproach is proposed:

1. Establish the constant amplitude S-N curve using strip specimens extracted fromrepresentative girth welds in the pipe of interest.

2. Conduct variable amplitude tests, using a loading spectrum representative of the actual

service condition, to check the 'severity' of the spectrum by establishing the validity of

Miner's rule on the basis of Equation.4.values at failure.3. Establish the constant amplitude S-N curve and CAFL for full-scale specimens.

Experience from tests on girth welds suggests that the CAFL should be defined as the

fatigue strength at which the CA data start to deviate from the linear logS-logN curvefitted to data obtained at higher stresses. This should not be assumed to correspond to an

endurance <107 cycles. In fact, it may be necessary to continue testing to endurances

approaching 108 cycles to be sure of reaching the CAFL.

4. Calculate the fatigue life of the full-scale specimen using Miner's rule, in conjunction

with the 'spectrum severity' (ie Equation.4. value at failure) obtained in step (b) and the







S-N curve obtained in step (c). A slope change from m to m+2 at the STP of the full-scale

specimen is recommended for use in the calculation of Equation.4.

With respect to the fatigue tests, the strip specimens should be selected to include the worst weld

root bead profiles (highest hi-lo and WRBH) found in the actual girth welds, they should be

conducted under fully tensile loading, preferably at a high tensile mean stress, and they shouldutilise spectra that contain only stress ranges that are well above the apparent CAFL.

With reference to the difference between sequential and random loading, the present endurancetests on strip specimens indicate there can be a difference in fatigue performance. Therefore,

ideally step b) above should be conducted using the actual loading sequence anticipated in

service. If this is not known the Miner's summation values obtained in the variable amplitudeloading sequence used in the tests may need to be reduced; the present limited study suggested a

reduction factor of 1.3. More study is required, including checks to determine if a 'spiky' random

loading sequence of the type used here can ever actually arise in a SCR where factors such as

inertia may limit the rate of change of stress, to establish the general applicability of such a

factor.

7 Conclusions

The following conclusions refer to the results of fatigue tests on full-scale 16 inch OD steel pipes

and strip specimens extracted from them, both containing girth welds with fatigue strengths

better than UK Class D, carried out under constant and variable amplitude loading, the latter based on a load spectrum considered to be relevant to SCRs.

1. Fatigue failures were obtained from full-scale girth welded pipes at stress levels up to42% lower than the conventional fatigue limit (corresponding to N=10

7 cycles), such that

the fitted S-N curve extended to 5x107 cycles. Consequently, the CAFL was defined asthe stress range corresponding to N = 5x10

7 cycles on the fitted S-N curve.

2. The fatigue performance of the full-scale and strip specimens agreed for applied stresses

above 110MPa but the CAFL for the strips was around double that for the full-scale

pipes.3. Miner's rule proved to be reasonably accurate for the loading spectrum used.

4. The VA results suggested that it would be non-conservative to adopt a bi-linear S-N

curve with a slope change at 107 cycles as recommended in most fatigue design codes.

However, introducing the slope change at 5x107 cycles (the new CAFL) was

conservative, but not unduly so.

5. Limited data from comparative VA fatigue tests on strip specimens indicated that a

'spiky' random loading sequence was more damaging than the 'sequential' loadingconditions possible with a resonance test rig, with a reduction factor of about 1.3 onendurance. However, further work is needed to assess the relevance of the 'spiky'

sequence to actual SCRs and, if necessary, perform further tests on more realistic

sequences.6. The fatigue performance of the girth welds was strongly influenced by the weld root bead

profile. The weld root bead height (WRBH) and hi-lo, sources of stress concentration and

secondary bending, were identified as significant indicators of fatigue performance.



7. Based on the present test results, as well as others from a previous JIP, a change to the

current fatigue design guidance for girth welded pipes is proposed. In particular, Miner's

rule should be used in conjunction with the appropriate constant amplitude S-N curve (eg. a BS 7608 design curve or the lower bound curve established by special testing) with the

slope change from 3 to 5 introduced at N= 5x107 cycles. For greater accuracy, a

procedure was proposed involving special VA fatigue tests on strip specimens under trulyrepresentative spectrum loading conditions, including a representative sequence ofloading if known, to establish the appropriate Miner's rule sum for design, together with

CA tests on full-scale pipes to establish the appropriate design S-N curve.

8. One reservation regarding the above recommendations was the case of a girth welded pipe failing from an undetected 0.65mm deep weld root defect to give a lower fatigue life

than that expected from the S-N curve obtained from the defect-free welds. If defects of

this size cannot be detected reliably or indeed for particularly critical cases, it would be

prudent to perform cumulative damage calculations using S-N curves that extended to perhaps 10

8 cycles before introducing the slope change.

8 Acknowledgements

The authors are pleased to acknowledge financial and technical support of the sponsors for this

work: BP Exploration and Production Ltd, Chevron Energy Technology Company, ExxonMobilUpstream Research Company, Heerema Marine Contractors, Petrobras, Saipem SpA, Tenaris

Tamsa and the UK Health and Safety Executive. Help from Nick Zettlemoyer of ExxonMobil

regarding the derivation of the effective stress range ΔSeq for the bi-linear S-N curve is also

acknowledged.

9 References

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BS 7608, 1993: 'Fatigue design and assessment of steel structure', British Standards Institution,

London.

BS 7910, 2005: 'Guide on methods for assessing the acceptability of flaws in metallic structures',

British Standards Institution, London.

Det Norske Veritas (DNV), 2010: 'Fatigue design of offshore steel structures', DNV-RP-C203,

Norway.

Gurney T R, 2006: 'Cumulative damage of welded joints', Woodhead Publishing, Cambridge.

Maddox S J and Zhang Y H, 2008: 'Comparison of fatigue of girth-welds in full-scale pipes and

small-scale strip specimens', Proceedings of the 27th Int. Conf, on Offshore Mechanics and

Arctic Engineering, Paper No.OMAE2008-57103.



Pyttel B, Schwerdt D and Berger C, 2011: 'Very high cycle fatigue - is there a fatigue limit?', Int.

J Fatigue 33 (2011) 49-58.

Salama M M, 1999: 'Fatigue design of girth welded pipes and the validity of using strips',

Proceedings of the 18th Intl Conf on Offshore Mechanics and Arctic Engineering,

OMAE99/MAT-2003, June 11-16.

Tilly G P, 1985: 'Fatigue of land-based structures', Int. J Fatigue, 7(2):67-78.

Zhang Y H and Maddox S J, 2009: 'Investigation of fatigue damage to welded joints under

variable amplitude loading spectra', Int. J Fatigue, 31, 138-152.

Zhang Y H and Maddox S J, 2011; 'Development of fatigue testing of full-scale girth welded

pipes under variable amplitude loading', Proceedings of the 30th Int. Conf, on Offshore

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