HSEHealth & Safety

Executive

Review of low cycle fatigue resistance

Prepared by Failure Control Limited for the Health and Safety Executive 2004

RESEARCH REPORT 207

HSEHealth & Safety

Executive

Review of low cycle fatigue resistance

Failure Control Limited Unit 30

Smithbrook Kilns Cranleigh

Surrey GU6 8JJ

The S-N method of fatigue life assessment is stress based, and is only fully applicable to cyclic stresses in the elastic range. However, in offshore installations the presence of stress concentrations at nodes and other connections can, in some circumstances, lead to cyclic stresses that exceed the yield stress of the material locally. This study is aimed at reviewing the current status of design advice and more recent experimental data with the objective of evaluating the treatment of high stress ranges.

This report and the work it describes were funded by the Health and Safety Executive (HSE). Its contents, including any opinions and/or conclusions expressed, are those of the authors alone and do not necessarily reflect HSE policy.

HSE BOOKS

© Crown copyright 2004

First published 2004

ISBN 0 7176 2830 2

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SUMMARY

The S-N method of fatigue life assessment is stress based, and is only fully applicable to cyclic stresses in the elastic range. However, in offshore installations the presence of stress concentrations at nodes and other connections can, in some circumstances, lead to cyclic stresses that exceed the yield stress of the material locally. This study is aimed at reviewing the current status of design advice and more recent experimental data with the objective of evaluating the treatment of high stress ranges.

Previous reviews of the low cycle fatigue resistance of plate details and tubular connections were supplemented by more recent data identified by a literature search. The present study indicates that a fatigue life less than that predicted by accepted design curves is possible in the high-stress low-life region, but only under laboratory conditions. The very high levels of cyclic stress required to promote this behaviour would not prove acceptable in static strength checks and would not occur in actual offshore installations.

The HSE “Guidance Notes” placed restrictions on the applicability of S/N curves in the high stress – low cycle region because of a lack of experimental data, but did not identify what fatigue damage might be expected when these limits were exceeded. Reanalysis of screened data now suggests that there is no need to impose such a restriction in the future,

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2.0

3.0

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5.0

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CONTENTS

INTRODUCTION 11.1 Background 11.2 HSE Fatigue Guidance 11.3 DNV RP-C203 (2000) and API RP-2A 21.4 Other limit states – members and plates 21.5 Other limit states – tubular joints 21.6 Strain based methods 3

FATIGUE DATA FOR PLATES 42.1 Review of data 42.1.1 Transverse butt welds 42.1.2 Longitudinal fillet welds 52.1.3 Transverse non-load carrying fillet welds 62.1.4 Transverse load carrying fillet welds 6

FATIGUE DATA FOR TUBULAR JOINTS 83.1 Delft compilation3.2 Static modes of failure3.3 Review of data3.3.1 T-joint data3.3.2 X-joint data

DISCUSSION

CONCLUSIONS

REFERENCES

FIGURES

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1. INTRODUCTION

1.1 BACKGROUND The casual paper-clip bender is well aware that that, after a short time, their diversion is terminated by an early failure of the clip. Assessing the likely fatigue life of paper-clips under this extreme loading regime is more demanding than many engineering applications because the clip experiences a combination of elastic and plastic strains. Since the familiar S-N method of fatigue life assessment is stress based, it is only fully applicable to cyclic stresses below the yield point of the material.

In offshore installations the presence of stress concentrations at nodes and other connections can, in some circumstances, lead to cyclic stresses that exceed the yield stress of the material locally. While the strains associated with this post-yield behaviour will be limited by the elastic global response of the structure there is still some uncertainty as to how this behaviour should be accounted for in assessing the fatigue limit state. Most design codes or design guidance either do not consider this question, or set limits on the applicability of the guidance so that the issue is avoided for those regions where experimental evidence is lacking or unreliable.

This study is aimed at reviewing the current status of design advice and more recent experimental data with the objective of updating the treatment of high stress ranges where possible.

1.2 HSE FATIGUE GUIDANCEAlthough no longer actively supported by the HSE, the fatigue sections of “Offshore Installations: Guidance on design construction and certification” were revised extensively in 1995 incorporating the findings of a review panel(1). Advice is given on high stress cycles in fatigue at Section 21.2.13(d).

For welded plates and tubular joints in air or seawater (with or without CP) values of the stress range considered in (b) above greater than twice the yield stress need special consideration. Further information is given in the Background Document. The limits from static strength should be considered (Section 21.2.4). The design tensile stress will be governed by a fracture criterion (Section 21.4.2) or by the tensile limitations on normal member stress (Section 21.2.3). For compressive loading , buckling considerations may be critical and should be included in the analysis.

The review panel referred to above noted that in previous guidance (pre-1995) the T curve for tubular joints was limited to a maximum hot spot stress range equal to twice the material yield stress, 2Sy. This was due to a lack of data in the high stress region and a concern that the fatigue process might change from a stress controlled high cycle fatigue mechanism to strain controlled low cycle fatigue. It was noted that some tubular joints could experience stresses beyond this limit during severe conditions, although this mostly occurred in older platforms constructed from low yield stress steel, or when the joint geometric SCF exceeded 5, or both(2).

Previous Guidance gave a similar restriction for non-nodal connections (classes B-G) where the peak stress was primarily due to bending or a stress concentration. For simple axial stress the maximum stress range was restricted to twice the allowable axial stress of 0.8 Sy for extreme loads and 0.6 Sy for operating loads.

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The review committee concluded that the existing (i.e. pre-1995) combination of limits on stress range and allowable stress for static loading were satisfactory, and that no extrapolation of the design curves to higher stress ranges was justified in the later guidance on the basis of the available data.

1.3 DNV RP-C203 (2000) AND API RP-2A (1991)Neither the DNV recommended practice C203(3), nor API-RP2A(4) require separate consideration of high stress ranges. In both documents the S-N design curves continue in the high stress – low cycle regime without modification. The commentary to DNV RP-C203 presents a plot of the contribution to fatigue damage for different regions of a Weibull distribution when the fatigue design line has slopes of m=3 and m=5 (N>107 cycles) and a total fatigue damage (Miners Number) of 0.5 and 1.0. This showed that the loads comprising the highest 1000 cycles (severe storm loading) made a negligibly small contribution to the overall fatigue damage.

1.4 OTHER LIMIT STATES – MEMBERS AND PLATESIn the assessment of offshore structures, the ultimate/serviceability limit states and the fatigue limit state are normally considered as separate aspects of design. However, practical considerations such as the need to design a member or joint configuration to acceptable limits for extreme loading is likely to influence the magnitude of the cyclic stress in severe sea states.

Large scale plasticity is prevented in offshore structures by limitations on either the allowable stress under extreme loading or suitable application of load and resistance factor design. For example, Section 21.2.3 of the HSE/D.En “Guidance Notes” referred to above. This states:

In no case should the calculated tensile stress in a member exceed 60% of the yield stress under normal operating conditions and 80% of the yield stress under extreme loading conditions.

All members and panels should remain nominally elastic and, even under those storm conditions relevant to fatigue assessment, the nominal tensile stress should be limited to 0.6Fy. Therefore, cyclic plasticity in plate type construction and tubular members will tend to be restricted to localised regions and discontinuities such as member intersection, holes and other sources of stress concentration.

1.5 OTHER LIMIT STATES – TUBULAR JOINTSLimitations on the allowable loads and moments for tubular joint design (tubular intersections) are based on parametric formulae derived from test data under tension, compression or moment loading. The limiting criteria in such tests are normally buckling of the chord or tearing at the brace-chord intersection, both of which may involve significant strain. Therefore, even with the application of appropriate resistance factors to the ultimate load capacity of tubular joints, there is a possibility of cyclic plasticity in normal operation.

In spite of the above, practical upper limits for fatigue loading are set by the partial safety factors applied to the characteristic static strength of the joint. Different codes advocate different approaches, but as an example the HSE “Guidance Notes” recommend partial safety factors of 1.70 for operating and 1.28 for extreme conditions. Therefore as an absolute maximum the stress cycle should be (1/1.28) or 78% of the load causing failure in a single cycle. This will be seen as being an important factor in the assessment of X-joints (section 3.1.2 below).

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1.6 STRAIN BASED METHODS

When the local cyclic stress range in a fatigue test is fully elastic, strains are proportional stresses, and failure can be predicted by reference to an appropriate S-N curve such as the P or T′ curves. The limitation of 2Sy placed on the applicability of the T′ curve is intended to ensure that this is achieved. In medium strength steels such as 355EM, the cyclically stabilised stress­strain curve is expected to be equal to or exceed the monotonic stress-strain curve.

i.e. ∆ε= ∆σ/E

However, for pseudo-elastic stresses that exceed this limit, or in high strength steels that may cyclically soften, the actual level of cyclic strain may exceed that inferred from the for pseudo­elastic stress. Since fatigue crack initiation and fatigue crack growth are largely strain controlled processes, this implies that in the first instance the results of tests should be compared in terms of cyclic strain rather than cyclic stress.

Consequences of using a strain-based approach include the expectation that, for the same high level of pseudo-elastic stress, a different fatigue endurance may be obtained depending on whether the test is conducted under displacement or load control. In the former, the cyclic strain is proportional to the pseudo-elastic cyclic stress irrespective of the actual stress level, while in the latter, cyclic strains can increase rapidly once the yield stress has been exceeded. Similarly, where plastic yielding is limited by the overall elastic response of the specimen (e.g. at an SCF) the cyclic strain may remain proportional to the cyclic pseudo-elastic stress although the yield stress is exceeded locally

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2. FATIGUE DATA FOR PLATES

2.1 REVIEW OF DATA In 1991 The Welding Institute (TWI) was commissioned to undertake a literature review of test results for welded joints subject to high cyclic stress(5). Their findings were mostly confined to plate details, although there were some very limited results for tubular joints. These are considered in Chapter 3. Many of the data identified in the literature were not relevant to any types of offshore structure, and ultimately 10 investigations were selected for detailed study.

The weld types represented in the TWI survey were:

Transverse butt welds (17 data series)Longitudinal non-load carrying fillet welds (10 data series)Transverse non-load carrying fillet welds (7 data series)Transverse load carrying fillet welds (3 data series)

The great majority of the tests were conducted under axial load, with one set of data for transverse load carrying fillet weld being tested under constant displacement cantilever bending. More importantly, none of the investigations involved stress concentrations other than the intrinsic stress concentration of the weld detail itself that is normally accounted for within the weld detail classification.

The TWI survey was conducted some 10 years ago, and the information was supplemented and updated for the current review using a computerised literature search. A number of promising investigations were followed up, but few were found to be relevant once full details had been obtained. Most significant were an investigation by Ferreira et al(6) of cruciform joints welded with full penetration, partial penetration and fillet welds at cyclic strains between 3 and 10 times the yield strain. These tests produced failure in a few tens of cycles. Also there was an investigation by Sakano et al(7) on structural beam-column joints. A French study undertaken by Gauchet and Rabbe(8) was potentially of interests as it contained information on welded joints with endurances as low as 100 cycles. Unfortunately, no actual results were presented, only schematic diagrams, so that it was not possible to include this study in the analysis of data..

2.1.1 Transverse butt weldsThe butt weld tests by Ida(9) and Radziminski(10) at an R ratio in the range 0 to 0.1 resulted in a fatigue endurance that lay above the relevant S-N design curve. However, this finding was not unexpected since there were no tests with failures at less than 104 cycles. These data are on the periphery of the high stress – low cycle regime and the maximum stress did not exceed 0.8Fy. Consequently, they have not been considered further.

Reed(11) studied pipeline girth welds tested under displacement control using API 5L X65 and X70 grades pipe. Welds were either submerged arc or manual metal arc welding from one side only. Excellent results were obtained that easily exceeded the Class F2 design line, and most would have exceeded Class D for two sided welding. This finding was particularly encouraging since the maximum compressive and tensile strains in there R=-1 tests exceeded ±0.8Fy and fatigue lives as low as 128 cycles were investigated.

Similar strain controlled tests at R=-1 were undertaken by Lieurade et al(12) for transverse butt welds in E355 grade steel. The butt welds were manufactured using either manual metal arc

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(MMA) that produced a convex weld profile, or submerged arc welding with a flat weld profile. The submerged arc welds had the same fatigue strength as the base material and lay above the Class E design curve for all instances where the stress range was less than ±0.8Fy. Two tests using MMA welds with the highest endurances “failed” a little below the Class D fatigue curve. However, in this study failure was defined as a crack exceeding 5% of the nominal section area rather than the conventional definition N3, also the two cracks initiated in the base metal rather than the weld zone, which is not typical. All of the other failures were at a stress range greater than ±0.8Fy.

The effects of variable mean stress in transverse butt welds were evaluated by Nihei(13) who conducted fatigue tests at R ratios of 0, -0.33 and –1. The plates were 14mm thick JIS steel grade SM58. None of the tests at R=0 produced failures at a stress range less than 0.8Fy, and none of the tests at R=-0.33 failed below the Class D design curve. A single failure at R=-1 fell below the design curve at a stress range less than ±0.8Fy which was “unsafe”. However, The slope of the S-N lines through the data points was very flat (m>>3) and this result seems untypical.

12 mm thick plates containing transverse butt welds were studied by Trufyakov(14) using 3 grades of steel, M16S mild steel (Fy=271MPa.), 14G2AF-u low alloy steel (Fy=458MPa.) and 13KhGMF high strength steel (Fy=617MPa.). Results for R=0 all lay above the Class D design line except for a few points when the maximum stress exceeded 0.8Fy.

For the purpose of data assessment the cruciform joint data from Ferreira et al has also been included with butt welds. This appeared justified because the non-loaded member was only 7 mm thick compared to the 15 mm thickness of the loaded member. Consequently the weld was relatively small. All of the full penetration welds failed in the parent metal and all except one of the partial penetration welds in the weld itself. There was a degree of buckling in the specimens, although their displacement was measured directly by LVDTs.

A compendium of data for R= 0 and R=-1 is plotted in Figure 1 and 2. All bar two of the results for R=0 and ∆S>±0.8Fy lay above the class D/E design lines as discussed above which supports their use for endurances down to 10,000 cycles, the limit of the data. The results for R=-1 are far more interesting as they contain the Ferreira data with endurances down as low as 26 cycles which all lay on or above the Class D line in spite of their actual classification being closer to F. As noted previously there were a substantial number of data points for S>±0.8Fy that fell below the line, but these were based on a very conservative definition of failure and should be excluded in an analysis of the data. From a consideration of relevant data it can be concluded that the pseudo-elastic stress range provides a suitable basis for assessing the low cycle fatigue life of butt welds.

2.1.2 Longitudinal fillet weldsA more noticeable effect of high stress cycles should be apparent in longitudinal fillet welds than transverse butt welds because of the greater stress concentration of the weld detail at the end of an attachment. This SCF would elevate nominal stress ranges below 0.8Fy to peak stress ranges that exceeded 0.8 Fy where potential non-linearity in the S-N response is most likely.

An extensive evaluation of this weld detail was undertaken by Harrison(15) using three medium strength and two high strength steels. The plates were 12.7 mm thick with 150 mm long fillet welded attachments (i.e. on the border line between fatigue classes F and F2). A summary of the materials is given below.

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MaterialBS968Controlled rolled C-MnFortiweldDucol W30Loycon QT

Yield Stress (MPa) 363 386 435 542 723

For each material fatigue tests were conducted under both load control and strain control, the latter including plastic strain ranges of up to 0.7% (1% total strain range including elastic strains). Under these conditions the endurance was a few hundred cycles. The load controlled tests used an R ratio of 0 while the displacement controlled tests used a variable R ratio between –0.03 at the lowest strain range and –1.1 at the highest. This made interpretation of the results more difficult, but since the highest strain ranges correspond to R=-1 the limiting stress range for practical applications will be taken as 1.6 Fy (±0.8 Fy).

Using load control, R=0, all of the materials gave fatigue endurances that exceeded the class F design line, except for those where the maximum nominal stress exceeded the yield stress (1.0 Fy). Similarly, with displacement control, all of the results for the Fortiweld and higher strength steels exceeded the class F design line even when the maximum stress range was greater than 1.6 Fy.

Data for the lower strength controlled rolled and BS968 steels showed a single point that fell below the class F line when the stress range was equal to 1.5 Fy or less, and several more fell slightly below when the stress range was ≥1.6 Fy. However, this behaviour was marginal and seemed only a small departure from the behaviour of the higher strength steels.

2.1.3 Transverse non-load carrying fillet welds

Lieurade(16) investigated the low cycle fatigue behaviour of cruciform joints welded from 12 mm thick E36 and A70 plates. Displacement controlled tests at R=0 on both materials produced results that lay well inside the Class F design line unless the stress ranges were substantially in excess of 0.8 Fy. Results for the A70 steel under load control and R=0.1 and R=-1 and E36 steel at R=0.1 were similarly encouraging, although the data points R=-1 in E36 steel were “strung out” along the ±0.8Fy horizontal line in a similar manner to the BS968 and controlled rolled steels in ref. 15 above. The results are plotted in Figure 5, although it should be remembered that in this investigation “failure” is defined as cracking that extends to 5% of the specimen area. Consequently the reported endurances will be less than those based on the conventional definition of N3.

12 mm thick plates with transverse fillet welded attachments were also studied by Trufyakov(11)

using M16S mild steel (Fy=271 MPa.) The fatigue lives at R=0 were in the range 3.5x104 to 8x104 cycles, which exceeded the Class F design life even though the maximum stress was in excess of 0.8 Fy. However, because this was a low yield point steel, 0.8Fy corresponds to a relatively low cyclic stress and a longer endurance than in conventional offshore steels.

2.1.4 Transverse load carrying fillet welds

Load carrying fillet welds are not normally used in situations where they are likely to be subject to cyclic loading, full penetration welds being preferred for this detail. However, the data that does exist is summarised in Figure 6.

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The first series of tests considered was conducted by Dunn and Anderson(17) using 12.7 mm plate to BS1501-224-26A LT30 with a low yield stress of 265 MPa. The specimen was a plate containing a transverse attachment subject to fully reversed cantilever bending displacements (R=-1). The resulting failures lay well above the relevant Class F2 design line, and in fact performed better than Class F. In these tests the pseudo-elastic stress ranges corresponding to the imposed displacements approached 10 Fy, indicating substantial reversed plasticity. Also, the resulting fatigue lives were as short as 300 cycles.

A second series of tests were undertaken by Trufyakov using the same M16S mild steel discussed in Section 2.1.1, but in a double lap joint configuration. All of the failure points lay above the Class F2 design line.

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3. FATIGUE DATA FOR TUBULAR JOINTS

3.1 DELFT COMPILATION In 1989 the Technical University of Delft prepared a review(18) of 228 low cycle fatigue results on tubular T and X joints. They applied the results in a general investigation of SCF equations, the thickness effect, and weld improvement techniques rather than a study of low cycle fatigue behaviour in particular. Since that time, design guidance for fatigue resistance in tubular joints has improved considerably (for example revisions to the HSE “Guidance Notes” in 1995) and much of the Delft work now seems dated. However, it still serves as a convenient source of data with 26 references to experimental studies, SCF equations and design guidance. Moreover, a literature survey to update this work did not identify any new information for tubular joints in the most relevant area of the high stress – low cycle regime (typically 10-1000 cycles).

For the present study, the raw data for T and X joints contained in the Delft review has been re­analysed in terms of the hot spot stress range. The Efthymiou equations were selected since they are the most widely accepted. In many instances, the cyclic strain was measured, but the measurement location was too far from the intersection to characterise the hot spot strain field. Consequently, these strain data have not been used in assessing fatigue performance.

Some of the joints had a ground weld profile, and these were excluded from the data analysis because the fatigue endurance would not be representative of normal joints with an as-welded profile. Ideally the data would also have been screened on chord or brace thickness since some of the chords were as thin as 1.2 mm and were little more than models. Unfortunately there is insufficient data, particularly for X-joints, to restrict the data to practical thicknesses (for example 12.7 mm minimum), and it was decided to overlook the significance of thickness unless the data fell below the T’ design line or its extension to higher stresses. The HSE does not recommend taking advantage of a potentially increased fatigue life for chords thinner than

(19).the reference thickness of 16 mm

3.2. STATIC MODES OF FAILURESome of the joints tested failed after a single cycle. This type of failure cannot be described as fatigue, and the failure mode was presumed to be either buckling of the chord under a compressive part of the load cycle, or ductile tearing of the brace-chord intersection.

In the design of offshore jacket structures, the loads applied to joints during extreme events are compared with their “characteristic strength”. This assumed strength is less than the mean strength of the joint, and its use is intended to take account of uncertainty in joint load resistance. Therefore joints that fail in one cycle are likely to have been subject to loads well in excess of the characteristic strength. Also, applied loads for extreme events are a combination of environmental, live and dead loads containing load factors and this combination would normally exceed the maximum or minimum load in a fatigue cycle. Therefore tubular joints in practical designs would not experience environmental load cycles of a magnitude that could cause single cycle failure.

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3.3 REVIEW OF DATA

3.3.1 T-Joint data S-N results for T-joints are shown in Figure 7. These show the T’ curve to be a satisfactory lower bound to the data except for two points at endurances of 4.4x104 and 5.0 x104 cycles. On closer examination, these two joints were found to have an extreme geometry (a=1, b=1, g=11.6) so that the stress range derived using the Efthymiou equation may have been an under-prediction. Also, the endurances at failure are in the region where the T’ curve is well established rather than the high stress – low cycle end of the data where its applicability is in greater doubt. Consequently, it has not been necessary to consider whether the loading was in excess of the joint characteristic strength.

3.3.2 X-Joint data Figure 8 shows the compilation of data for X-joints, which extends to much shorter endurances than that for T-joints. Some specimens failed at only a single load cycle as discussed above, and this load level is likely to equal to or exceed the characteristic strength of the joint. Consequently, these data, included in Figure 8, are not all relevant to an assessment of fatigue strength.

Figure 9 has been prepared with the data points removed that correspond to these excessive load levels, it can be seen that almost all of the “unsafe” values belonged to this deleted category.

Even this sub-set of the data contains joints with loading that would not be permissible from a static strength viewpoint. In section 1.5 it was noted that different codes advocate different approaches, but as an example the HSE “Guidance Notes” recommend partial safety factors of 1.70 for operating and 1.28 for extreme conditions. Therefore as an absolute maximum the stress cycle should be (1/1.28) or 78% of the load causing failure in a single cycle. Figure 10 shows the joints that satisfy this more stringent condition where only a single data point falls below the T’ design line.

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4.0 DISCUSSION

Many laboratory investigations are designed to research into the mechanisms of material behaviour or show changes in failure mechanisms under extreme conditions. Although interesting from a scientific viewpoint, not all of these investigations will be relevant to the design of offshore structures or their design codes.

The tubular members that make up jacket structures or the plates in box girders and columns are subject to either conservative levels of allowable stress, or alternatively load and resistance factor design. This ensures that in normal operation the members remain elastic and the basic working stress even under heavy weather loading should not exceed 80% of the yield stress. In practical terms many members are subject to compressive loads/buckling and the extreme stress is likely to be well under 80% of the yield stress. Therefore, those tests in which the nominal stress exceeds this level are not relevant to practical designs. A useful background to the design provisions for the static strength of tubular joints is given in the HSE report OTO 2001/082(20).

As a second consideration, offshore structures are subject to wide band variable amplitude loading from wind and wave action. It is well established that the great majority of fatigue damage comes from small and medium sized waves that have a large number of occurrences, rather than the much smaller number of storm waves. Therefore uncertainties in the precise response of welded joints to loading in the high-stress, low-cycle regime are much less significant than uncertainties in the low-stress-high-cycle regime, where variations in predicted fatigue life of 8 times were found depending on the assumed form of the S/N curve(21). In any event, the primary concern that has been expressed related to the possibility of tubular joints suffering high levels of cyclic stress compared to their yield point was in respect of older platforms constructed from low yield stress mild steel, or when the joint geometric SCF exceeded 5, or both. The current emphasis on weight reduction (so that jackets may be installed by lifting) has encouraged the use of considerable quantities of 450 grade steel with 350 grade material being the minimum strength used. Also, the availability of node quality steel in plates up to 120 mm thick and a better understanding of stress concentration factors has reduced the SCFs in more recent designs compared to less efficient configurations. Consequently the concerns expressed for platforms designed in the 1970’s will have little relevance to new designs.

For tubular joints that satisfy the static strength criteria, the relevant experimental evidence appears to endorse the use of the T’ curve without restriction. For plates the evidence is less clear since no tests seem to have been done for those welded details where stress concentrations are likely to occur (ratholes, changes in flange width etc), although the work on beam to column intersections (ref 7) addresses this comment in a limited manner. A review of fatigue results in those welded plates that satisfy the static strength criteria shows that the accepted fatigue detail classifications remain valid.

The tests on transverse butt welds at R=0 contained few points below 104 cycles and were of limited interest, while those tested under fully reversed loading (R=-1) showed conflicting results. Firstly, the data for full penetration cruciform joints (likened to butt welds, see 2.1.1) and partial penetration cruciform joints showed a high fatigue endurance that exceeded the Class D design curve even with endurances as short as a few tens of cycles and pseudo-elastic stress ranges (E∆ε) that were many times the yield stress of this low yield material (250-300 MPa). Conversely, data from Lieurade (ref 12) with pseudo-elastic stress ranges exceeding 0.8 times the yield stress of E355 steel fell below the Class E line. This may explained by the

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authors use of the number of cycles to first appearance of a crack as the endurance (N1) rather than specimen separation (N3). Unfortunately the authors did not provide the more conventional N3 data so it was not possible to plot all the data on the same basis.

Figures 3 and 4 show data for longitudinal fillet welded attachments under load and displacement control. All “valid” data points except one lay above the relevant Class F design line.

Figure 5 shows data for transverse load carrying fillet welds that is principally from the work of Lieurade and Maillard-Salin (ref 16), with a few points from Trufyakov at relatively high endurances. The French work appears to show that at endurances less than 2x103 cycles the results fall well below the Class F design line, although the corresponding pseudo-elastic stress ranges are excess of 0.8Fy. Unfortunately the results are difficult to interpret because once again N1 has been taken as the failure criterion rather than N3 and the slope of the data is very shallow (m≈6). Consequently, it was decided to ignore these data as the failure criterion was unconventional and the stress levels outside the range of practical interest.

Endurance data for transverse load-carrying fillet welds is shown in Figure 6. All of the data lay well in excess of the Class F2 design curve apart from a singe point at 2x104 cycles. This is considered atypical and lays within the high cycle part of the curve where the fatigue resistance of such details is well established.

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5.0 CONCLUSIONS

The HSE “Guidance Notes” placed restrictions on the applicability of S/N curves in the high stress – low cycle region because of a concern that plasticity effects might cause a reduced life, but were unable to offer any definite advice because of the lack of suitable data. This concern was particularly focussed on high SCF tubular joints in low yield steel. As far as we are aware, this restriction was not implemented in the fatigue analysis software used by jacket designers because (a) there was no alternative given, (b) other design codes did not contain the restriction, and (c) it would have made very little difference to the calculated fatigue life.

A review of available data now suggests that a fatigue life less than that predicted by accepted design curves is possible in the high-stress low-life region, but only under laboratory conditions. The very high levels of cyclic stress required to promote this behaviour would not prove acceptable in static strength checks and would not occur in actual offshore installations.

Consequently there appears to be no practical need to impose such a restriction in code revisions, and other design codes are not open to criticism for failing to include a restriction.

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6.0 REFERENCES

1 “Offshore Installations: Guidance on design, construction an d certification”, Fourth Edition (including Amendment 3, 1995), Department of Energy, publ. HSE Books

2 “Twice Yield Stress as a Fatigue Criterion”, Lloyd’s Register of Shipping, Report OSG/TR/84003, November 1984.

3 “Fatigue Strength Analysis of Offshore Steel Structures”, Recommended Practice RP-C230, Det Norsk Veritas, 2000

4 “Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms – Load and Resistance Factor Design”, API Recommended Practice 2A-LRFD, American Petroleum Institute, 1991

5 TD ROSENBURG, RM ANDREWS and TR GURNEY “A compilation of fatigue test results for welded joints subjected to high stress / low cycle condition – Stage 1”, Dept. of Energy Report OTI 91 552.

6 J FERREIRA, CA CASTIGLIONI, L CALDO and MR AGATINO "Low cycle fatigue strength assessment of cruciform welded joints", J. of Constructional Steel Research, 1998, vol 47, pp 223-244

7 M SAKANO, I MIKAMI, E YONEMOTO and Y MOMO "Low cycle fatigue test of steel beam-column joint”, Technology Reports of Kansai University, March 1993 vol 35, pp 285-290

8 A GAUCHOT and P RABBE "Comportement en fatigue d’aciers à haute limite d’élasticité", European Commission Report EUR 8061 FR, 1983.

9 K IlDA. et al. "Low cycle fatigue test of hemispherical pressure vessel with misaligned welded joint", Proc. 3rd Intl Conf. P.V. Tech., Tokyo April 1977, Part II, ASME, pp.733-739

10 JB RADAMINSK and FV LAWRENCE "Fatigue of high strength steel weldments" Welding J. Research Supplement, August 1970, pp.365s-374s

11 RP REED "Fitness for service criteria for pipeline girth weld quality" WRC Bulletin No.296, July 1984

12 HP LIEURADE and C MAILLARD-SALIN "Low cycle fatigue behaviour of of welded joints in high strength steel" ASTM STP 770, 1982, pp 311-336.

13

13 M NIHI et al "Effect of programmed mean stress on fatigue strength of welded joints for SM58 steel”. Trans NRIM, Vol 23, No. 3, 1981, pp33-44.

14 VI TRUFYAKOV and VV YAKUBOVSKII “Low cycle fatigue of welded joints under zero tension loadings” Automatic Welding, Vol 34, No. 10, 1981

15 JD HARRISON “Further low cycle fatigue tests on butt and fillet welded high strength steel” Welding Institute Report n. C215/11/69

16 HP LIEURADE "The low cycle fatigue behaviour of cruciform welded joints in high yield stress steels” Soudages et Techniques Connexes, Vol 32, Nos. 11-12, 1978, pp 405-418

17 JW DUNN and J ANDERSON "An assessment of the over-conservatism of the ASME III low cycle fatigue predictions for fillet welds”. 2nd Int. Conf. on Fatigue and Fatigue Thresholds. Birmingham, Sept. 1984, Vol III, pp1753-1761.

18 GJ van der VEGTE, J de BACK and J WARDENIER "Low cycle fatigue of welded structures – analysis of low cycle fatigue tests on tubular T and X joints”, Stevin Laboratory report 25.6.89.11/A1, Delft university of Technology, April 1989.

19 R. KING "Background to new fatigue guidance for steel joints and connections in offshore structures”, HSE Report OTH 92 390.

20 BOMEL LIMITED "Comparison of tubular joint strength provisions in codes and standards”, HSE Report OTO 2001/082.

21 FAILURE CONTROL LIMITED "Fatigue design curves for welded joints in air and seawater under variable amplitude loading”, HSE Report OTO 1999/058.

14

Stre

ss ra

nge

(MPa

.)

FIGURES

10000

Class E design line

Class D design line

Ref 9 Smax<0.8 Fy

Ref 9 Smax>0.8 Fy

Ref 10 Smax<0.8 Fy

Ref 10 Smax>0.8 Fy

Ref 13 Smax>0.8 Fy8

Ref 14 Smax<0.8 Fy

1000 Ref 14 Smax>0.8 Fy

Stre

ss ra

nge

(MPa

.)

1001.000E+01 1.000E+02 1.000E+03 1.000E+04 1.000E+05

Cycles to failure (N3)

Figure 1: Endurance data for transverse butt welds (R=0)

10000 Ref 11 ∆S > ±0.8Fy

Ref 12 ∆S < ±0.8Fy

Ref 12 ∆S > ±0.8Fy

Ref 13 ∆S < ±0.8Fy

1000

100

Class D design line

Class E design line

Ref 6 ∆S > ±0.8Fy

1.0E+01 1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06

Cycles to failure (N3)

Figure 2: Endurance data for transverse butt welds (R=-1)

15

100

1000

Stre

ss R

ange

(M

Pa.)

1.0E+02 1.0E+03 1.0E+04 1.0E+05

Max stress > 0.8Fy

BS 968 steel

Controlled rolled C-Mn

Fortiweld

Ducol W30

Loycon QT

Class F design line

Cycles to failure N3

Figure 3: Endurance data for longitudinal fillet welded attachments(displacement control)

100

1000

Stre

ss R

ange

(M

Pa.)

1.0E+01 1.0E+02 1.0E+03 1.0E+04 1.0E+05

Max stress > 0.8Fy

Fortiweld

Ducol W30

Loycon QT

Class F design line

Cycles to Failure (N3)

Figure 4: Endurance data for longitudinal fillet welded attachments(load control)

16

10000

Stre

ss R

ange

(MPa

.)

Class F design line

A70, Load control, R=-1

A70, Load control, R=0.11

A70, Disp. control, R=0

A36, LC, R=-1

A36, LC, R=0.11

A36, DC, R=0

1000 M16S, DC, R=0

A70, Disp. control Smax>0.8 Fy

A36, Disp. control Smax>0.8 Fy

100 1.0E+01 1.0E+02 1.0E+03 1.0E+04 1.0E+05

Cycles

Stre

ss ra

nge

(MPa

.)

Figure 5: Endurance data for transverse non load-carrying fillet welds

10000

Class F2 design line

Ref 17 R=-1

Ref 14 R=0

Ref 6 R=-1

Ref 7 R=-1

1000

100 1.0E+01 1.0E+02 1.0E+03 1.0E+04 1.0E+05

Cycles to failure (N3)

Figure 6: Endurance data for transverse load-carrying fillet welds

17

100

1000

10000

Hot

spo

t st

ress

ran

ge (

MPa

.)

1E+02 1E+03 1E+04 1E+05

Nf (R=-1)

Nf (R=0)

T' curve

Hot

spo

t st

ress

ran

ge (

MPa

.)

Cycles to failure (N3)

Figure 7: Delft Endurance Data for T-Joints

10000

1000

100 1E+00 1E+01 1E+02 1E+03 1E+04 1E+05

Nf (R=0.1 - 0.25)

T' curve

Cycles to failure (N3)

Figure 8: Delft Endurance Data for X-Joints

18

100

1000

10000

Hot

spo

t st

ress

ran

ge (

MPa

.)

Valid results

T' curve

1E+00 1E+01 1E+02 1E+03 1E+04 1E+05

Cycles to failure (N3)

Figure 9: Delft Endurance Data for X-Joints - excluding those with the same stress range that caused single cycle failures (see text)

100

1000

10000

Hot

spo

t st

ress

ran

ge (

MPa

.)

∆S<78% single cycle failure

T' curve

1E+00 1E+01 1E+02 1E+03 1E+04 1E+05

Cycles to failure (N3)

Figure 10: Delft Endurance Data for X-Joints - excluding those within 78% of the stress range causing single cycles failures (see text)

19

Printed and published by the Health and Safety ExecutiveC30 1/98

Printed and published by the Health and Safety ExecutiveC0.06 03/04

9 78071 7 628308

ISBN 0-7176-2830-2

RR 207

£10.00