HSE Health & Safety

Executive

Variable amplitude corrosion fatigue of jack-up steels

Prepared by University College London for the Health and Safety Executive

OFFSHORE TECHNOLOGY REPORT

2001/079

HSE Health & Safety

Executive

Variable amplitude corrosion fatigue of jack-up steels

N Tantbirojn, R J Bowen, L S Etube and W D Dover University College London

UCL NDE Centre Dept of Mechanical Engineering

Torrington Place London WC1E 7JE

United Kingdom

P J Kilgallon, T Roberts and J Spurrier

SIMS Cranfield University

College Rd Cranfield Bedford

Bedfordshire MK43 0AL United Kingdom

HSE BOOKS

© Crown copyright 2002 Applications for reproduction should be made in writing to: Copyright Unit, Her Majesty’s Stationery Office, St Clements House, 2-16 Colegate, Norwich NR3 1BQ

First published 2002

ISBN 0 7176 2319 X

All rights reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means (electronic, mechanical, photocopying, recording or otherwise) without the prior written permission of the copyright owner.

This report is made available by the Health and Safety Executive as part of a series of reports of work which has been supported by funds provided by the Executive. Neither the Executive, nor the contractors concerned assume any liability for the reports nor do they necessarily reflect the views or policy of the Executive.

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SUMMARY

This report presents the findings from studies carried out on Variable Amplitude Corrosion Fatigue (VACF-T). Two main steels were investigated. These include SE 702 and DILIMAX690E. These steels are both of high strength Jack-up steel grades with yield strengths in the region of 700MPa. The focus of the investigation was twofold. The primary focus was on the investigation of the effect of cathodic protection (CP) on both parent and welded plates of different specimen geometry, with tests conducted both under constant and variable amplitude loading conditions using the Jack-up Offshore Standard load History (JOSH). The second focus of the investigation was to study the effect of transition period in the variable amplitude sequence on the fatigue performance of steels studied.

The overall investigation was conducted in four parts. Part one looked at the effect of CP on thick parent plates and ground welded plates under constant amplitude loading conditions. The plates were tested at two stress ranges, in air and in seawater under CP (-1050mV). Part two looked at the effects of different CP levels on the fatigue performance of T-butt welded plates under variable amplitude loading conditions. In part three, T-Butt welded plates were used to study the effect of transition period on the fatigue performance of high strength steels under variable amplitude loading conditions at two CP levels similar to those used in part two. Part four of the investigation concentrated on the examination of fracture surfaces to identify any significant metallurgical properties relevant to the fatigue performance of the materials examined. A further additional part (Part five) was subsequently conducted and included in this report. This involves corrosion fatigue tests to establish the influence of CP at long life fatigue.

All the results obtained from this investigation are presented in this report and compared with the fatigue performance of conventional fixed platform steels. There is an observable trend for shorter fatigue lives as the CP level is increased from - 800mV to - 1050mV. The reduction in life is of the order of 20% to 50% depending on specimen geometry and applied stress. Metallurgical work on the cracked specimen has found that crack multinucleation occur in most cases with exception to parent plate tested in air. The severity of crack branching also increase with the level of stress intensity factor and was more prevalent under CP than in air. It was observed that the transition period has an effect on the resulting fatigue life for tests conducted under variable amplitude loading conditions but this effect is not significant. Overall, high strength steel plates show satisfactory fatigue performance at high stress and there is no conclusive evidence to show that they are any worse that 50D under CP conditions. The long life corrosion fatigue tests showed this conclusion are not true for all high strength steels and that under some circumstances the effect of overprotection could be substantial at long lives.

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Printed and published by the Health and Safety Executive

C30 1/98

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TABLE OF CONTENTS

SUMMARY iii TABLE OF CONTENTS v 1. INTRODUCTION 1 2 EXPERIMENTAL PROGRAMME 2 2.1 Part 1: Effect of CP on fatigue in Parent and Ground Welded Plates 2 2.2 Part 2: Effect of CP on fatigue in T-butt Plates (VACF-T) 2 2.3 Part 3: Effect of sea state transition period (VACF-T) 2 2.4 Part 4 Metallurgical Examination 3 2.5 Part 5 Long Life Corrosion Fatigue (LLCF) 3 3 EXPERIMENTAL PROCEDURE 4 3.1 Fatigue Testing 4 3.2 Metallurgical Examination 5 4 RESULTS 7 4.1 Results of Metallurgical Examination 7 4.2 Results of Fatigue Tests 7 5 DISCUSSION 10 5.1 Discussion on Metallurgical Examination 10 5.2 Discussion on Fatigue Test Results 14 6 CONCLUSION 19 6.1 Metallurgical Examination (Cranfield) 19 6.2 Fatigue Tests (UCL) 19 7 REFERENCES 21 8 TABLES AND FIGURES 23 9 APPENDIX I: THE CORROSION FATIGUE CRACK GROWTH RATE PLOTS FROM CLI 85 10 APPENDIX II: THE CORROSION FATIGUE BEHAVIOUR OF JACK-UP STEELS 86 11 APPENDIX III: WELDING PROCEDURES FOR FLUSH GROUND WELDED PLATES 89 12 APPENDIX IV: WELDING PROCEDURES FOR T-BUTT WELDED PLATES 93

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1. INTRODUCTION

High strength steels are increasingly being specified for use offshore [1]. However, there are still concerns regarding their corrosion fatigue behaviour, especially with high negative values of applied potential (e.g. –1050mV Ag/AgCl) because of issues regarding their hydrogen embrittlement behaviour.

High strength steels (700MPa) have been used for some time on Jack-up structures but the periodic dry docking of such structures allowed the opportunity for regular inspection and for any necessary repairs to be carried out. However, developments where Jack-ups are designed to operate as fixed installations with long design lives (25 years), such as BP-Harding, have posed new design requirements in respect of corrosion fatigue.

The investigation presented in this report was conducted to study the fatigue performance of high strength steel welded plates. The study was designed to address issues relating to fatigue performance (both initiation and propagation) of high strength steels. In particular the effect of cathodic protection on the fatigue life of welded joints under variable amplitude loading conditions.

The project involved four major studies. The first part of this investigation was aimed at the performance of parent plates and flush ground welded plates under constant amplitude loading conditions and the effects of cathodic protection. This has previously been reported [2].

The second part of the program involved a study of the effect of CP on the fatigue performance of high strength steel T-butt welded plates under variable amplitude loading conditions using the JOSH sequence. The emphasis of this part of the study was to look at the effect of overprotection on the fatigue performance of high strength steels, which are thought to be more susceptible to hydrogen embrittlement under cathodic protection conditions when compared with conventional fixed platform steels. Two CP levels ( - 1050mV and - 800mV ) were therefore used in order to compare performance under an adequate cathodic protection potential (- 800mV ) with performance at an over protected CP level ( - 1050mV ). Some of the results obtained from this part of the study and those from part one have been reported before [3]. Earlier work by CLI reported the Paris law data for SE702 [4]. The resultant plot and the Paris relationship are shown in Appendix I. Work on corrosion fatigue was also carried out by Cranfield, which is shown in Appendix II. This Final report includes all of the results in part two and also brings together results from other parts of the whole programme.

The third part of the investigation concentrated on studying the effects of transitions period in the variable amplitude sequence used. The tests conditions used here are similar to those used in part two and are reported in detail in this report.

Part four involved an extensive programme of metallurgical examination of the fracture surfaces for some of the specimens used during the investigation. This work was undertaken at Cranfield University.

Part five involved an additional series of longlife corrosive fatigue tests on T-butt welded plates at two CP levels ( - 800mV and - 1050mV )

This report presents all the details of the methodology used for this investigation and results obtained. The data generated can be used in the development of suitable guidance on the use of high strength steels Offshore.

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2 EXPERIMENTAL PROGRAM ME

This section of the report gives details of the experimental programme. This includes details of experimental test parameters used for each section of the investigation.

2.1 PART 1: EFFECT OF CP ON FATIGUE IN PARENT AND GROUND WELDED PLATES

This part of the study was carried out on DILIMAX690E. The mechanical properties of this steel are shown in Table 1 while Table 2 shows its chemical composition. The data on chemical composition were obtained from Dillinger Hutte. Two types of plates were tested; thick parent plates and flush ground welded plates. The specimen dimensions are shown in Figure 1. The welding procedures for the ground welded plates are in Appendix III. The objective was to study the fatigue performance of the high strength steels in air and in seawater with cathodic protection and to compare results obtained from the two categories of specimens.

All of the eight tests in this part were carried out under constant amplitude loading conditions. Two stress levels were used, 350 MPa and 412 MPa with a cathodic protection level of - 1050mV for the seawater tests. A summary of the test parameters is shown in Table 3.

2.2 PART 2: EFFECT OF CP ON FATIGUE IN T-BUTT PLATES (VACF-T)

The main objective in part 2 was to study the effects of different levels of CP on the fatigue life and crack growth behaviour in T-butt welded plates under variable amplitude loading conditions.

This part of the investigation was conducted on SE 702, a high strength steel with yield strength of over 700MPa. The mechanical and chemical properties of SE 702 are shown in Tables 1 and 2 respectively. T-butt welded plates made from 40mm thick plates were used. Details of the specimen geometry are shown in Figure 1 and Table 6 shows a summary of the test parameters. The welding procedures for the t-butt welded plates can be found in Appendix IV. All the tests were conducted under variable amplitude loading conditions and two equivalent stress levels, 146 MPa and 172 MPa, were used. Two levels of Cathodic Protection ( - 800mV and -1050mV ) were used for the seawater tests.

2.3 PART 3: EFFECT OF SEA STATE TRANSITION PERIOD (VACF-T)

Part 3 was aimed at studying the behaviour and the effects of different sea state duration and transition periods on the fatigue life and crack propagation of T-butt welded plates tested using the JOSH sequence. Four tests were carried out at two stress levels, 146 MPa and 172 MPa, for two different transition periods of 10 and 30 minutes. The test conditions were otherwise similar to the test conditions used in part 2 with a CP level of -1050mV being applied. The main difference was in the sequence used. Two different variants of the JOSH sequence were generated with the required transition properties. A summary of the test parameters used is given in Table 3.

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2.4 PART 4 METALLURGICAL EXAMINATION

Some of the test specimens from parts 1 and 2 (six) made from Dillimax 690E-Z15 parent plate (85x85mm) or welded plate (85x170mm) were supplied by UCL to Cranfield from the VACF-T programme. These samples were tested in air or in seawater using either 3-point bend (3PB) or 4-point bend (4PB) loading. In addition, T-butt welded joints manufactured from 40mm SE 702 plate were also available for examination.

The six comprised one each of plate, welded plate and T butt joint tested in air and one each tested in sea water under cathodic protection conditions. Stress levels used during the tests and levels of cathodic protection (CP) are given in Table 9.

In outline, the work conducted at Cranfield involved the following. a) The preparation of metallographic sections normal to the crack b) Hardness measurement traverses taken on these sections c) Examination by optical microscopy to reveal the relationship between local

microstructure and the crack path d) Separation of the sections to reveal the full fracture faces e) An examination of the fractography by scanning electron microscopy.

2.5 PART 5 LONG LIFE CORROSION FATIGUE (LLCF)

The objective of this series of tests was to provide Long Life Corrosion Fatigue (LLCF) data for HSS welded plate. A series of tests was carried out using DILIMAX690E-Z15, a High Strength Jack-Up Steel with a yield strength of 690MPa which was available from Part one. The plate thickness was 16mm and the grade is supplied by Dillinger Hutte in a different composition to the 85mm plate. The objective was to gain crack-growth fatigue data from simulated in-service Variable Amplitude Corrosion Fatigue Testing (VACF-T). The Alternating Current Potential Difference (ACPD) technique was used to monitor crack growth behaviour.

The seven T-Butt-welded components were made from 16mm thick flame–cut ground plate, and contained full-penetration welds at the attachment plate; dimensions of the specimens are shown in Figure 4. A spectrographic metallurgical examination of the steel is shown in Table 10 and the mechanical properties are given in Table 11.

The tests were performed under four-point bend fatigue loading, using a JOSH Variable Amplitude loading sequence. The dimensions of the test rig are shown in Figure 5.

The specimens were tested in air, using a mean frequency 2Hz, (1.6-2.4Hz range); and in a salt-water environment, using two levels of Cathodic Protection (CP); these tests were performed at a mean frequency of 0.2Hz (0.16-0.24 range).

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3 EXPERIMENTAL PROCEDURE

This section of the report gives details of the approach and methodology adopted in conducting the investigation. It outlines both the process of load control and implementation of data acquisition for fatigue testing presented including information relating to test condition monitoring. The procedure adopted during the metallurgical examination is also presented.

3.1 FATIGUE TESTING 3.1.1 Simulation of Seawater environment

The tests are conducted in a simulated seawater environment. The reason for this is that the processes of corrosion fatigue and hydrogen embrittlement of high strength steels are a complex combination of corrosion and embrittlement chemical reactions. The effects of these chemical reactions, service loading and the influences of any cathodic protection system must be modelled in the laboratory to obtain representative results.

Hence, the test specimens were immersed in simulated seawater for the duration of the tests. The simulated seawater environment was achieved by constructing an environment cell around the welded intersection for each specimen tested under corrosion conditions. Artificial seawater made to ASTM D1141 [5] was pumped from a reservoir through a closed loop passing through the environment cell. The fully aerated seawater was maintained between temperature limits of 8oC and 10�C using an external refrigeration system. The pH was also monitored in the course of the test and maintained between 7.8 and 8.2.

Prior to application of fatigue loading, each joint was subjected to a ‘soak time’ of two weeks. During the soaking period, the welded intersection was submerged in seawater with an active CP system applied and maintained at the appropriate level. The soaking period of two weeks is compatible with previous tests conducted at UCL. An earlier study on SE 702 [5] showed that results obtained after a soaking period of two weeks were comparable with those obtained by employing a soaking period of 4 weeks.

All cathodic protection potentials were based on an impressed current method, applied using a Wenking MP81 potentiostat and measured with reference to an Ag/AgCl electrode.

3.1.2 Simulation of environmental loading conditions

All test conducted under variable amplitude loading conditions were conducted using the JOSH sequence [6]. This was implemented by using an advanced fatigue testing software, which allows sequence play back through a DARTEC digital servo-hydraulic control console. The modes of loading employed for the different categories of specimens are shown in Figure 2 and 1.

The variant of JOSH used was generated using 1000 transitions of 12 sea states. The distribution of these sea states in the sequence used in part 2 is shown in Figure 6. The stress range distribution (SRD) curves corresponding to the different equivalent stress levels used are shown in Figure 7 while the relevant time sequence scaled to –100% is shown in Figure 8. Figure 9 and 10 also show the variants of JOSH used to study the sea state duration behaviour for 10 and 30 minute sea state duration respectively. It is important to note that the overall load amplitude in each sequence will depend on the equivalent stress range used for the test. The corresponding SRD curves are compared in Figure 7. The sequence file JOSHTP10.OUT

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was used for the 10 minutes transition period and JOSHTP30.OUT was used for the 30 minutes transition period.

3.1.3 Test condition monitoring and Data acquisition

The servo hydraulic test machines used could be controlled using either position or load feedback control loops. All tests conducted in this programme were performed under load control. In this mode the specimen is subjected to load cycles of a pre-defined amplitude regardless of any change in stiffness of the specimen.

The fatigue testing control software provides the facility for online inspection and data acquisition in the course of a fatigue test. Fatigue crack development was monitored by taking Alternating Current Potential Drop (ACPD) [7] crack depth measurements (Figure 11) at fixed points around the welded intersection at definable inspection intervals.

3.2 METALLURGICAL EXAMINATION

The Cranfield programme was designed to produce metallography support to the VACF-T Phase 2 studies. Microstructure can play an important role in affecting fatigue growth rates particularly in the heat-affected zone of welds. Hardness, grain size, and microstructure as well as other features such as crack branching are important and these aspects were investigated at Cranfield. The procedure adopted is outlined in this section.

3.2.1 Macro-examination

For the metallographic examination, band-saw cuts were made in the larger welded specimens. This enabled metallographic surfaces to be prepared from one original surface and at ¼, ½ and ¾ thickness. A consistent face of each section was ground flat and polished. The polished area was then etched in 2% Nital and photographed directly with a 35mm camera to enable full-size macro-photographs to be produced.

The parent plate samples were not as thick as the welded specimens, and were expected to have a uniform microstructure through the thickness. These therefore were not sectioned, but a metallurgical face was prepared from one of the original surfaces of each specimen.

3.2.2 Hardness measurements

Vickers’ micro-hardness measurements were made on the welded sections with a Matsuzawa Seiki testing machine using a 500g load. (Some verification tests were also made using a lower load of 100g. Differences were negligible.) Traverses were made in two directions,

a) Close to the edge of the crack over the first 6mm or so of crack growth, and b) Parallel to the top surface of the sample on both sides of the crack, at 4mm depth.

Parent plate samples have been treated in a similar way, but using a standard Vickers’ diamond pyramid testing machine with a 5kg load. This was necessary because the micro-hardness machine is not able to accommodate the larger thickness of the un-sliced parent plate samples. (No significant differences are expected in the mean hardness values between these tests and micro-hardness values, although the latter may be expected to show increased scatter due to the smaller sampling volume.)

T butt joints were investigated with the microhardness machine using a 500g load, so that the results could be compared directly with those for welded plates. Traverses were made in two locations, with the objective of giving information on near-surface hardness and on the

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microstructure sampled by the crack path. The indents were positioned at a distance of about 0.5mm from the specimen surface or the crack surface, through weld metal to the crack mouth and along the crack path, and from weld to parent steel around the corner on the opposite weld toe from the crack.

3.2.3 Micro-examination

Small specimens cut from the larger sections were mounted in Bakelite, polished and etched. These specimens were used for determination of the microstructure and grain size, and for examination of the local profile of the crack path using a Nikon microscope and attached camera.

3.2.4 Fracture surfaces

In order to view the fracture surfaces, it is necessary to break the remaining ligament below the tip of the crack, whilst trying to ensure minimum distortion to the profile that has been produced during the test. To achieve this, a significant proportion of the ligament was cut by bandsaw and hacksaw (a region that appears at the bottom of the relevant photographs in this report) and the specimens were then immersed in liquid nitrogen before separation by impact loading. This produces a brittle fracture that can be distinguished readily from the fatigue crack surface, and produces relatively little plastic deformation.

3.2.5 Scanning electron microscopy (SEM)

Selected regions of the fracture surfaces have been examined using an ABT-55 scanning electron microscope. Samples were ultrasonically cleaned and rinsed in distilled water, and dried with propan-2-ol and warm air. The surfaces of the two cathodically protected samples (P4 and W2) had a significant build up of corrosion product, and tended to produce ‘flaring’ and unacceptable contrast levels in the scanning electron microscope. Therefore, attempts were made to clean the surface with a rust-removing agent before vacuum coating with gold/palladium to attempt to improve the uniformity of secondary electron emission.

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4 RESULTS

The results are presented in two sections. The first section relates to results obtained from the metallurgical examination conducted at Cranfield while the second relates to results obtained from the fatigue tests conducted at UCL.

4.1 RESULTS OF METALLURGICAL EXAMINATION

The following results are presented in this report … Table 9 Identification of samples Figure 12 Illustration of sectioning locations for welded plates (showing

W2) Figure 13 separation

Fracture surfaces of welded samples W1, W2 after

Figure 14 Parent plate tests P3, P4 – view of top (tension) surfaces after testing

Figure 15 Fracture surfaces of parent plate samples P3, P4 after separation

Figure 16 & 17 Crack profile sections for welded specimens W1, W2 Figure 18 & 19 Optical micrographs of crack initiation region with respect to

welds Figure 20, 21 & 22 Micrographs of fatigue crack profiles at depths of 10, 30 and

50mm Figure 23 Illustration of hardness traverses with respect to a crack

profile Figure 24 & 25 Hardness traverse values, samples W1, W2, P3 and P4 Figure 26, 27 & 28 Scanning electron micrographs of the fatigue failure surfaces Figure 29 and 30 Photographs of T-butt welds T04 and T08 after failure Figure 31, 32, 33 & 34 Crack profile sections for T-butt welded specimen T04 Figure 35, 36, 37 & 38 Crack profile sections for T-butt welded specimen T08 Figure 39 Hardness traverse values and sketch of locations for T04 and

T08

4.2 RESULTS OF FATIGUE TESTS

The results obtained from the fatigue tests conducted in this study are presented in two forms, the fatigue lives (SN data) of the specimens and the fatigue crack growth data. Both SN data and fatigue crack growth results are given below.

4.2.1 SN Data

The fatigue lives obtained from the tests are presented as SN curves. These data are for the thickness effects for 16mm thick plates. As a comparison to the 50D material, the parent and flush ground welded plates were plotted with the class C 50D mean line. The T-butt welded plates were plotted against class F 50D mean line.

The SN data obtained for the thick parent plate and flush ground welded plates under constant amplitude loading conditions are given in Table 12 while results from variable amplitude loading of the T-butt plates are shown in Table 13. For fatigue tests conducted under variable amplitude

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loading conditions, the stress range shown is the equivalent stress range for the variable amplitude sequence used while the life represents the number of fatigue cycles to failure. In each case, the specimen was considered to have failed once the fatigue crack reached half the plate thickness. The resulting SN curves are shown Figure 40 in to Figure 45. Figure 40 shows the results of tests conducted on parent and flush ground welded parent plates under constant amplitude loading conditions in Part 1 of the study. One of the results obtained for the parent plate P1, was tested at two stress range levels. The first level was 350 MPa, which lasted for 1,000,000 cycles and the second level at 412 MPa for 268,000 cycles. This could explain the shortened fatigue life. The data can be corrected using Miners rule. The new value calculated could be seen in Table 14. The results obtained under variable amplitude loading for Part 2 are shown in Figure 41. Figure 41 shows the comparison between the results with the emphasis on the effects of cathodic protection. Figure 42 shows results obtained from Part 3 for different transition periods. All the data from Parts 2 & 3 are shown in Figure 44. All the SN data points obtained from the study are compared in Figure 44. Results from this study have also been compared with those obtained for other high strength steels. This comparison is shown in Figure 45.

For the SN plots, mean lines were drawn through the data points. These were done in Figure 55 to Figure 61. However, bearing in mind the limited amount of test conducted, these lines were put in to show the trend of the material rather than the actual mean line. The slope of the lines were all corrected to –3. In Figure 59, Figure 60 and Figure 61, the mean minus two standard deviations line, were plotted for the t-butt plates in Air and CP. The standard deviation values were obtained from the results of the tests conducted here. Similar calculation were not done for the parent and ground welded plates, as the numbers of data were too small.

In Table 15, the results of T-butt plates are compared, with reference to different transition period and cathodic protection level. The results for T-plates T01 to T12 were tested under 20 minutes sea state duration.

The fatigue lives are also compared as percentages in Table 17, Table 18, Table 19 and Table 20. Table 17 shows the changes in performance under different testing conditions, relative to other tests in the programme. Table 18, Table 19 and Table 20 show the same fatigue life reduction factors, relevant to Parts 1, 2 and 3 respectively. This relative difference in fatigue life obtained for different test conditions is also shown as a bar chart in Figure 40.

4.2.2 Fatigue Crack Growth data

All the crack growth data is presented in two forms. These include crack shape evolution curves and through thickness crack propagation curves obtained as the number of fatigue cycles is increased for each test. These are given below.

4.2.3 Crack shape evolution data

Crack shape evolution data show crack front profile and how this front advances as the number of fatigue cycles is increased. This is obtained from measurements taken at suitable intervals along the entire length of the weld toe at appropriately defined ACPD inspection intervals. Sample results obtained from tests are shown in Figure 49 and Figure 50 for T09 and T10. Each plot represents only the failed side of the weld but in most cases, there were cracks on both sides of the attachment plate.

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4.2.4 Through thickness Crack growth data

Through crack propagation data is obtained by monitoring the crack advance at a single location of interest along the weld toe. This information is obtained from 19 sites on each side of the weld for every test. Figure 47 and Figure 48 show typical crack growth curves obtained corresponding to T09 and T10 respectively.

Figure 51 to Figure 54 show the comparison in through thickness crack growth for different combinations of environmental and loading conditions. Figure 51 shows the effect of keeping the stress level constant. In this figure, the fatigue crack growth curves for air and two different CP levels are compared at an applied equivalent stress range of 146MPa. A similar comparison is shown in Figure 52 for the higher equivalent stress range of 176MPa. In both of these figures, data from air tests are compared with those obtained from test conducted under a cathodic protection level of - 800mV and - 1050 mV . Figure 53 shows the effect of transition period and compares results obtained for 10 and 30 minute transitions periods. All the results are further compared in Figure 54. The two thick vertical lines seen in each figure represent the repetition of the simulated service loading sequence, where the starting cycles tend to have higher overall amplitude.

4.2.5 Results for Long Life Variable Amplitude Corrosion Fatigue Tests

(Part 5)

Table 16 shows the results for all the fatigue tests conducted for part five. The data is also shown in Figure 62. Table 10 shows the chemical composition of the material used in part five and this must be compared with Table 2 for the material used in part two. It is evident that although the strength levels are very similar, the composition of the two supplied steels are clearly different. It is normal practice for the supplier to change the composition with plate thickness.

Three air fatigue tests were conducted, one of which was a run-out. These can be shown with the earlier air tests conducted on a steel of the same strength level to show a similarity in behaviour. Figure 63 shows all the air data obtained with t-butt welded specimens for DILIMAX690E. Taking a slope of -3 shows that the data for both compositions is very close to the P curve (Class F) mean curve.

Figure 64 shows the long life data for a CP of -800mV and also the data from part two. It would seem that the behaviour for both steels is similar, and all results are better than the BS 4360 50D P Curve (Class F) CP line.

Figure 65 shows the long life data for a CP of -1050mV and also the data from the previous study. In this case the data for long life is closer to the P Curve (Class F) CP line whilst from the previous study the data lies between the air and CP line.

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5 DISCUSSION

5.1 DISCUSSION ON METALLURGICAL EXAMINATION 5.1.1 General

In the VACF-T programme, specimens were fatigued without a notch initiator. An aim was to determine where fatigue might be expected to start in a real structure. Near-surface microstructures and residual stresses associated with welding and processing operations should affect the initiation site. Once a small crack has started to grow, the local stress intensity in the crack tip region is expected to dominate the effect of the applied loading. The crack path therefore may be expected to show deviations associated with variations in microstructure (linked to the plastic zone size at the crack tip). The disadvantage of an un­notched specimen is that once the crack initiation has taken place, there is no control on the types of microstructures being sampled by the subsequent fatigue crack growth. In welded specimens, the sampling is likely to be very dependent on the heat affected zone and weld geometry. In addition, it is understood that the parent plate samples that were sent to Cranfield were manufactured by flame-cutting, and that these edges were deliberately not fully removed. Consequently, local variations in heat-treated microstructure are considerable even in specimens designated here as “parent steel”.

5.1.2 Crack profiles Welded plates

Both welded samples W1 and W2 showed a slightly curved crack mouth across the top (tension) surface, as illustrated by specimen W2 in Figure 12. This suggests small amounts of multi-nucleation as the crack grew across this face. In general, the crack appears to have progressed into the specimen in a relatively uniform way, as shown in Figure 13.

“Parent steel”

The crack in the parent plate sample P3 tested in air appears to have initiated at one edge and carried through to the opposite side and on through the sample.

The cathodically (over-) protected parent plate specimen P4 tested in seawater, however, has initiated three main cracks on the tension face, shown in Figure 14. The overlapping of the cracks suggests they were formed independently, merging to produce the final crack shown in Figure 15. It seems probable that the two outside cracks were initiated at small surface flaws associated with the specimen preparation and corner rounding of the beams. The crack probably initiated in the centre of the sample before the higher stresses produced at the edges caused two further cracks to initiate at machining marks on the surface. It is not clear from the one sample whether the cathodic protection has played a role in promoting multi-nucleation in “parent plate”: especially since the welded specimen W1 which was tested in air without cathodic protection appears to have two main initiation cracks. It is likely that the multi-nucleation is associated with the local changes in microstructure associated with flame cutting of the specimens. This factor is discussed at the end on the next section, when the crack paths are examined by optical microscopy.

T butt joints

The failure surface of the T butt joint sample T04, which was tested in air, exhibited a number of ‘beach marks’ indicating the progress of the crack. It was clear from this that the initial

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crack extension of approximately 2 millimetres had occurred relatively uniformly along the length of the crack from front to back. Marks suggested that some crack deceleration occurred, with a slight tendency for increased bowing after this. After about 4 millimetres, faster growth appears to have resumed. This is broadly in agreement with the ACPD results from UCL on this sample, but the retardation at sensor 9 and the acceleration at sensors 17 and 18 in the ACPD work [3] (Progress Report, VACF-T Phase 2 : Thick Plate Specimens, August 1999, Figure12) did not appear to be as pronounced on the failure surfaces as the readings implied.

The seawater exposure of T08 showed less in the way of beach marks, suggesting more uniform progression of the crack, but again the growth appeared to be relatively uniform along the length of the crack.

It appeared that some dressing of the welds had taken place to facilitate attachment of the ACPD probes. This was probably a light grinding operation with a hand-held rotary grinder. It was not clear if this had any influence on the exact position of the toe of the weld. Figure 29 and 30 show that the initiation line for the crack was relatively straight in both T butt joints, particularly for T08. Viewed at low magnification with a magnifying glass, there were no obvious residual parts of weld metal on the opposite side of the crack mouth, suggesting a close relationship between microstructure and cracking.

5.1.3 Crack paths Welded plates

Photographs of the polished sections (1 to 4) for both the welded specimens W1 and W2 are shown in Figure 16 and 17. The multi-pass weld can be seen clearly to the right of the cracks, together with the heat-affected zone (HAZ), which appears as a darker band a few millimetres wide. Section 1 represents the surface. Dimpling at the crack tip tends to produce a poor polish in this region.

The (dominant) crack appears to have started in the vicinity of the weld toe on the machined top surface of both welded specimens. The four regularly spaced sections typify the crack path, but should not be taken to indicate the exact sites of crack initiation. In general, however, they do show that the cracks propagated into parent plate within 5mm or so of crack growth: but this is principally a function of the geometry of the capping passes. (A slightly different fracture appearance within the HAZ/weld region could be distinguished on the fracture surfaces when the specimens were examined by eye at low magnification.)

Figure 19 shows optical micrographs (at a magnification of x50) taken near the suspected initiation sites. The photographs are from longitudinal sections (similar to Figure 17 and 18) and show the tension surface of the sample at the top, and the initial stages of crack growth down the left-hand side of the figures.

For specimen W1 (in air) the crack appears to have initiated in the intercritical region of the HAZ (ICHAZ) and propagated through the sub critical HAZ into parent plate. The slanted path may be caused by multi-nucleation of crack initiators on slightly different planes, and the deviation that results as these grow together. The rather rough nature of the original surface is revealed at this magnification.

In W2, the crack appears to have initiated at the end of the weld toe and travelled perpendicularly, hence progressing into and across the HAZ because of the weld geometry. The path that the fatigue crack growth takes through the steel specimen will be dependent on the local microstructure and stress distribution. In section 1 of specimen W2 there is some

11

indication that the crack has attempted to deviate towards the HAZ, but this is not seen in the other sections. No clear conclusions can be made regarding the relative susceptibilities of the HAZ and parent plate microstructures as the crack has been forced to grow within the parent plate for the majority of propagation. The inclusion of notched specimens in future work may give a clearer indication of the susceptibility of the HAZ.

“Parent steel”

Optical microscopy on the parent plate samples suggests there is a 5mm deep region at the top of the crack profile where the microstructure is appreciably different. It was at first assumed that this was associated with the manufacture and rolling of the plates. However, the region is broadly similar in appearance to the HAZ of a welded sample, with a near-surface grain-coarsened region followed by a finer microstructure and then some evidence of precipitation at prior austenite grain boundaries as depth below the surface increases. This then led to the realisation that the specimens were flame-cut and tested in that state to give some information of possible consequences of repairs etc in service.

As a consequence of this manufacturing process, the surface layers of the “parent steel” samples did not possess a uniform microstructure, since different amounts of the flame cut surface had been removed. This would produce local surface variations in terms of the ease of initiating a fatigue crack, and explain why there was sufficient multi-nucleation to cause some curvature of the crack mouth on sample P3 tested in air, and to produce the macroscopically stepped crack mouth on sample P4.

T-butt joints

The failure surfaces of both T butt joints were examined for signs of multi-nucleation along the initiation line from front to back. Multiple nucleations appeared quite widespread in both samples, particularly for the T04 test in air (lacking environmental enhancement), but these were all very minor events, probably implying that the nucleation events were relatively close to each other in time as well as position. The fatigue crack front appeared relatively uninterrupted by these events.

On each of the four sections of both joints, the fatigue crack appeared to initiate at the toe of the weld, essentially between weld metal and the coarse-grained heat affected zone. The crack propagation direction was then under stress control, with growth proceeding directly through the thickness towards the loading point. The first 3mm or so of crack growth during the fatigue tests were through the heat affected zone from the CGHAZ towards parent steel.

A short crack at the toe of the opposite weld was noticed on one section from joint T08, also emanating from the weld metal / coarse-grained heat affected zone boundary.

The marked deviation in crack growth direction after 20mm or so crack extension, very marked in sample T04 but also seen in section 1 of T08, is assumed to be associated with movement of the deformed specimen in the fatigue machine, and was not investigated.

5.1.4 Crack branching

Figure 20 to 21 show the fatigue crack profiles for samples W1, W2, P3 and P4, at nominal depths of 10mm, 30mm and 50mm from the tension surface. The crack propagation direction is from right to left in all cases, and the magnification is x200. These metallurgical sections were taken to investigate the changes in surface roughness apparent on the fracture faces (as shown in Figure 13 in particular).

12

There is a significant amount of crack branching in all of the samples. It appears that the direction of the crack branch is influenced by the local microstructure. The prevalence of crack branching appears much higher in the cathodic ally protected samples W2 and P4 than in those specimens tested in air. It is likely that the hydrogen associated with cathodic over­protection is influencing this behaviour. (It may be argued that the length of a crack branch is more closely related to the uniformity of the microstructure. Slight variations in local toughness can encourage continued growth of two equal branches. More noticeable variation can result in one branch becoming longer, and then the difference in applied stress intensity should dominate the growth pattern.) Slight evidence of secondary cracking is visible on the cathodically protected samples. This is cracking that occurs after the main crack has passed, and can be normal to or angled against the main crack propagation direction.

In all of the samples, the amount and size of crack branching increases with increasing applied stress intensity range ÄK. Previous work at Cranfield [8 and [9] has indicated the importance of crack branching on fatigue crack growth rates, and has shown that the amount of branching increased with increasing ÄK values regardless of the strength and microstructure of the steel.

Crack branching was not investigated in detail for the T butt joints.

5.1.5 Hardness values Welded plates

Figure 23 is included to illustrate the hardness traverse directions (a and b) with respect to the crack profile for the plate samples.

Micro-hardness values for samples W1 and W2 are plotted and compared in Figure 24. These values confirm a higher hardness in the HAZ, where the maximum recorded was 351HVN for W1 and 342HVN for W2. These values for Dillimax 690E-Z15 may be compared with the mechanical testing requirements for weldability testing of offshore structural steels, where HV10 values of around 350 are usually recommended to avoid hydrogen cracking problems. The 1989 British Standard for weldable offshore steels [10], for steel strengths up to Grade 450, recommends an upper limit to the hardness of 325 at HV10. The requirements for Grade 690 were for PHWT, aimed for a maximum value of 370HV10, where values in excess of 400HV10 are acceptable.

“Parent steel”

Vickers hardness measurements for P3 and P4 are plotted in Figure 25. There is little variation in the hardness values taken across the crack at constant depth, as can be expected for parent plate material. Values were 221-241HVN for P3 and 239-268HVN for P4. However, the hardness of both P3 and P4 increases towards the top surface (and crack mouth) to a maximum of 260HVN for P3 and 386HVN for P4. The microstructural variations in these regions are noted in section above.

T-butt joints

Figure 39 includes a sketch showing the locations of the hardness traverses taken on the T-butt welded joints. The first part of traverse A samples near-surface weld metal hardness, with the toe of the weld positioned at the ‘5mm’ reading. From the ‘5mm’ to ‘8mm’ locations, the microhardness values reflect the material along the crack path. For traverse B, taken on

13

the opposite side, the values show near-surface hardnesses from weld metal around the toe and through the HAZ into parent plate.

Both joints T04 and T08 showed relatively uniform weld metal hardnesses, in the range 320­350 (Vickers’ HVN) for T04 and 330-370 for T08. A softer region of microstructure was found at the uncracked weld toe in T04. This was not investigated further, but the microstructure there could have undergone slight tempering from finish-grinding operations. HAZ hardness values were appreciably higher, reaching 414-416 in all four of the HAZ regions examined. Parent plate hardness values were around 278 at the end of all the traverses on the T butt joints.

The hardness profiles from traverse A show that there was a distinct hardness change between weld metal and HAZ, and the geometry of the specimens provided a significant stress concentration there. These factors appear to have combined to generate the relatively consistent crack initiation line along the T joints. Because no single initiation site was detected, multiple nucleation occurring along the weld toe, it is suggested that the values observed from the (arbitrary) polished sections used in this work can be considered representative.

Hardness measurements were taken on both the parent plates and weld metal in part five. The parent plate hardness was about 298VPN, HAZ 285VPN and the weld metal was 205VPN. General purpose welding rods were used in the fabrication, as it was not considered to be significant for the fatigue tests.

5.1.6 Fractography

Surface oxidation and contamination was unacceptably high for electron microscopy of the fracture surfaces when the samples arrived at Cranfield. Attempts were made to examine all four plate specimens, and typical micrographs (after cleaning) are shown in Figure 26 to 28. The data along the bottom of the photographs are, from left to right,: the accelerating voltage, the nominal on-screen magnification (and micrographs here are produced at that size), the scale bar (ì = micrometre), and a sequence number.

The main features visible are the mainly ductile surfaces for the air tests W1 and P3. There is some evidence of crack branching into grain boundary fissures, but the surfaces show signs of ductile striations (especially P3).

Conversely, the seawater tests with cathodic protection, W2 and P4 show enhanced grain boundary attack. The fatigue surfaces of these two specimens have regions where the crack appears to follow grain boundaries, either along the boundary or in the adjacent material. The fatigue separations appear to be under ductile control, but probably at much reduced energy levels, implying a lower local toughness value. (Additional crack branching, however, may delay the macroscopic rate of fatigue crack growth [7, 8].

5.2 DISCUSSION ON FATIGUE TEST RESULTS Fatigue Life

In general, the results obtained have shown that the T-plates have a lower fatigue life when compared to ground welded plates. This could be due to the weld toe geometry and hence the low initiation life of the welded T-plates. From the SN data presented in Figure 44, it can be seen that there are differences in the performance of T-plates when compared with parent and ground welded plates. Looking at the performance in air for example, increasing the stress level had different effects on the three categories of specimens. On average, a 30% reduction in life was observed for welded T-plates. The corresponding reduction in life for welded

14

ground and parent plates is 11% and 80% respectively. This difference in fatigue life could be down to the shorter initiation lives of the T-plates as compared to the parent and ground welded plates.

The behaviour in air and in seawater is markedly different for parent plate and flush ground weld specimens but far less so for the T-butt. In practice, the seawater behaviour for the three types of specimen is quite close, thus supporting the suggestion of the initiation of the initiation life being the distinguishing feature amongst the test results.

The comparison of fatigue lives for T plates and parent plates cannot be straight forward, however as the T plates were loaded under variable amplitude, while the parent and ground welded plates were tested under constant amplitude loading conditions. In addition, the weld toe stress concentration factor (about 2.4) was not included in the presentation of these T-butt results. In most cases, under variable amplitude loading, the higher peak loads may lead to early crack initiation. Alternatively severe plastic deformation at the crack tip may lead to plasticity induced crack closure, which may have an influence on the crack propagation under variable amplitude loading conditions. The resulting fatigue lives therefore, for tests conducted under variable amplitude loading conditions will depend largely on the balance between crack advance and crack retardation mechanisms. Where there is this balance, one may expect the life under variable loading condition to be comparable to those under constant amplitude conditions. This will be particularly applicable at stress levels where the clipping ratio is significant as is typical for offshore applications where structures are exposed to multi Sea State loading. Under this type of loading conditions, where high clipping ratios are involved, there is a higher probability of higher than expected plastic deformation at the crack tip and this may lead to significant differences between constant and variable amplitude behaviour.

With respect to the effect of cathodic protection, results show that the cathodic protection level has a detrimental effect on the fatigue life for both constant and variable amplitude tests. The main factor involved in shortening the fatigue life of specimens under higher cathodic protection potentials can be put down to hydrogen. The mechanism of hydrogen embrittlement is known to increase the crack propagation rate possibly through enhanced grain boundary attack. But under variable amplitude loading, where there is a combination of crack retardation and acceleration, the effects of hydrogen is not known fully. In addition, the rate of charging at the crack tip with hydrogen will depend largely on the different level of protection used. At the higher level of protection more hydrogen will be generated from the electrochemical reactions involved and this may explain the higher crack growth rates observed for higher cathodic protection potentials when compared with lower potentials.

For the constant amplitude loading of the parent and flush ground welded plates, the plot shows a distinct difference for air and seawater tests under cathodic protection. This clear distinction is not seen for tests conducted on T-butt welded plates.

The percentage changes in fatigue life for different loading and CP conditions are given in Table 17. Table 18, Table 19, and Table 20 show the relevant SN data matrices for parts 1, 2, and 3 respectively. Table 19 shows that a CP level of -1050mV leads to a reduction in life of 87% and 56% respectively for tests conducted at 350 MPa and 412 MPa for the parent plates. For ground welded plates, the effect of stress range seems to be less significant when compared to the effect of CP since tests conducted at 172 MPa and 146 MPa show a reduction in life of 83% and 88% respectively.

Table 19 shows that the effect of CP seems to be less severe on the fatigue behaviour of thick T-butt welded plates. For tests conducted at an equivalent stress range of 146 MPa and a CP level of - 1050mV , the observed reduction in fatigue life ranges from 48% to 57% when

15

compared to tests conducted in air. Table 17 also shows that the corresponding reductions in life observed for tests conducted at the lower CP level of - 800mV falls in the range of 20% to 33%. At the higher equivalent stress range of 172 MPa, the reduction in life due to cathodic protection falls between 52% and 70% for tests conducted at - 1050mV . On the other hand, tests conducted at the lower CP level of - 800mV show reductions in the range of 42% to 64%. The results obtained from this study show that there is a higher percentage reduction in life for the higher CP level of - 1050mV when compared with tests conducted at a CP level of - 800mV . Looking at tests conducted on T-butt welded plates, increasing the CP level from - 800mV to - 1050mV leads to, on average, a 40% reduction in life for 146 MPa. An average reduction of about 30% seems to be applicable for tests conducted at 172 MPa apart from test T07 which shows an increase in life of about 40% when compared with T12 carried out under similar conditions. These results are consistent with results obtained from a previous study [11] on large-scale tubular welded joints, where a reduction factor of 30% was observed for cases where the CP level was increased from - 800mV to - 1000mV for hot spot stresses in the range of 200–225MPa.

As for the fatigue life for different transition period, the effect is not so clear cut. From the SN plot shown in Figure 43 it can be seen that the fatigue lives are not so distinct for different transition periods. The results are also similar to those obtained in part 2 of the study and the relevant reduction factors are shown in Table 19. However, there are differences between the two transition periods examined especially in terms of crack propagation. The results obtained for the 30 minute transition period show a tendency for higher crack propagation when compared with results obtained for tests conducted with a transition period of 10 minutes. The sequence used for most of the tests is the 20 minutes transition period. However, the results in Figure 58 do show that the 30 minutes period is more similar to the 20 minutes than the 10 minutes period.

The transition period is how quick the Sea State moved to the next Sea State. The longer the transition period, the more likely it is that once a more damaging sea state with a lower probability of occurrence is reached, it will persist in the sequence for longer leading to more fatigue damage. However, the less damaging sea state will also last longer, but overall, the longer transition period is viewed as being potentially more damaging.

Overall, the results show that the steels perform satisfactorily as can be seen in the SN curve where their performance is compared with results from other tests (Figure 45). The fatigue data from this programme are within the clusters of the other results tested for other high strength materials.

Part five showed that for the thinner plates 690MPa yield steel; the effect of CP at long life can be quite severe. The low-alloy steel tested in the current programme had an average fatigue life of around 500,000 cycles at a stress range of 108MPa with a CP of -1050mV. In comparison, the tests at -800mV at the same stress had a run-out at 3,800,000 cycles, a factor of 7·.

In contrast, the tests at higher stresses in Part two on the thicker steel showed a factor of only 1.42·. This difference is either a feature of life time and/or composition and it would be valuable to conduct further tests in the long life region with the thicker plate steel that the same minimum yield stress of 690MPa.

Figure 66 shows all the data for T-butt welded specimens from both Part two, three and five. Of concern is the short fatigue life of the long life tests, which are close to and below the P curve (F Class) CP line. It would be desirable to test further specimens the thicker steel to establish the behaviour of the long life tests with a CP of -1050mV more clearly.

16

Figure 66 also shows that the behaviour at a CP of -800mV is clearly superior and possible indicates a change to a curve of gradient -5. Table 10 shows the spectrographic and manufacturers composition of the material used in this and previous studies.

It would appear that steel manufacturers supply a range of steels under the general classification of high steel with a yield stress of 690MPa, depending on plate thickness. The long life fatigue tests were done on a thinner plate with lower alloy additions. All the steels used in the study and the previous tubular joint studies are shown in Table 21.

17

Fatigue crack growth

From the results obtained for the JOSH.OUT sequence, the pattern is for high crack propagation rates at the start of the sequence. The simulated sequence contains high stress ranges at the beginning and these become smaller towards the end. This can be observed by looking at the plot of the sequence in Figure 8. This distribution represents long term loading conditions commencing from more severe loading representing stormy conditions. This is one of the reasons why there is a high crack growth rate at the beginning of each repetition followed by almost constant growth rate and subsequent fall. This illustrates the potential significance of sequence effects under typical multi sea state loading conditions and the mechanisms of crack growth acceleration and retardation involved under these conditions need to be modelled for adequate fracture mechanics analysis under variable loading conditions.

The classic example of a typical crack growth for this can be seen in Figure 47 for specimen T09. The crack seems to have initiated quite quickly and propagate at a steady rate. At around 130,000 cycles, the crack growth seems to have changed and become more rapid, the slope becoming steeper until failure. In addition, if one looked at the point where the sequence started again, that is the point where the crack growth increased as indicated by the red line in Figure 51. This can be seen in all the crack growth curves under the JOSH.OUT sequence, that when it reaches this repetition of cycles, the crack growth rate increases.

Comparison between the air curve to the seawater curve is not very simple. Some of the specimen under cathodic protection did not get through the whole sequence thus a test at lower stress range would have revealed more in terms of crack propagation behaviour under variable amplitude loading. However, from what was observed, Figure 54 shows a couple of interesting points. One obvious point is that the crack initiated quicker under higher stress range. The effect of cathodic protection is very detrimental, with early initiation and has a more aggressive crack growth rate. The interesting point is that slopes for air normally becomes less steep until the sequence is repeated, but for the tests conducted with cathodic protection, the slope continued at its usual rate. The next point of observation is that if the crack growth is below 2mm when the sequence is repeated, the specimen will usually survive until the next repetition of the cycles. Anything above 2mm will result in a correspondingly high growth rate resulting in failure.

18

6 CONCLUSION

6.1 METALLURGICAL EXAMINATION (CRANFIELD)

• The parent material, which were flame cut, retained some of the heat-affected layer, with microstructure and properties similar to the HAZ of welded joints. It may be expected to show variable initiation behaviour in this condition

• Under CP conditions, both the parent and welded specimens exhibited some degree of crack multinucleation and crack branching linked to the local microstructure.

• The severity of observed crack branching increased with the applied stress intensity factor (Dk).

• For tests in air, crack growth appeared to have been slower. Initiation in the plate samples showed frequent local multinucleation of cracking.

6.2 FATIGUE TESTS (UCL)

• The SN curves obtained for the high strength steels showed that for SE702 there was a better fatigue life for air and under CP than that shown for 50D material. By comparing the results obtained with the mean SN curve for 50D data, the data obtained from the study for parts 1, 2 and 3 respectively show that all the data points lie above the mean C and F class curves.

• The results from part two clearly show that the CP level is significant and that higher CP potentials will have a detrimental effect on the fatigue performance of high strength steels. It was observed that, by increasing the CP level, the fatigue life is shortened. However, this effect is not as severe it was originally thought for SE702. The effects are less obvious in T-butt plates as compared to the parent and ground welded plates and this is most likely due to differences in the microstructure and surface finish for the two specimen categories examined.

• Regarding the effects of different transition periods (sea state duration), the results suggest that the fatigue lives for 20 and 30 minutes period are more similar, while 10 minutes period seems to show a longer life. This observation is very important and it illustrates that the inaccurate modelling of sea state duration can lead to large differences in the resulting fatigue life. This effect is very important for service applications and needs to quantify by undertaking a more detailed further study on transition behaviour.

• The results are satisfactory and show a good agreement with the data obtained for SE702 tubular joints although fatigue performance at high equivalent stresses will not necessarily be identical to behaviour at lower stress levels typical for service applications. More data on high strength steels is required within the low stress regime, which is more relevant to service applications.

• Air fatigue tests at long life for two steels in Part two and five with yield strength of 690Mpa show similar fatigue behaviour to that of type.

19

• Tests at low stress ranges in Part five showed a considerable difference between CP levels of -800 and -1050mV. The -800mV tests had lives in excess of 7· greater than the -1050mV specimens.

• There would appear to be a need to establish whether the SE702 material retains a good corrosion fatigue performance at long life.

20

7 REFERENCES

[1] Webster S. “Structural Materials for Offshore Structures – Past, Present and

Future” IBC Conference on Safe Design and Fabrication of Offshore Structures,

London, 1993.

[2] Etube LS, Tantbirojn N, Brennan FB and Dover WD, ‘Progress report on

Variable Amplitude fatigue of Jack-up Steels (Phase 2:Thick Plate Specimens),

UCL November 1998’

[3] Tantbirojn N, Etube LS, and Dover WD, ‘Progress report on Variable Amplitude

fatigue of Jack-up Steels (Phase 2:Thick Plate Specimens), UCL August 1999’.

[4] Coudert E and Renaudin C, “Variable amplitude corrosion fatigue behaviour and

hydrogen embrittlement of high strength steels for off-shore applications”,

Proceedings of the International Offshore and Polar Engineering Conference,

Vol. 4, pp116-122, 1998.

[5] American Society for Testing and Materials, “Specification for Substitute Ocean

Water.”, ASTM D 1141-75, 1980.

[6] Etube L S, Brennan F P and Dover W D, “Variable Amplitude Corrosion Fatigue

(VACF) of High Strength Steels Used in jack-up Structures”, Final Project

Report, UCL NDE Centre, 1998.

[7] Technical Software Consultants Ltd., ACFM Crack Microgauge – Model U10

User Manual, April 1991, Milton Keynes.

[8] Drury J. “An Investigation into the Fatigue and Corrosion Fatigue Properties of

Two High Strength Low alloy Steels and their HAZs” PhD Thesis, Cranfield

University, 1992.

[9] Maden G.C. “Corrosion Fatigue of High Strength Low alloy Steel under

Conditions Likely to Promote Hydrogen Embrittlement” MPhil Thesis, Cranfield

University, 1994.

[10] British Standard Specification for Weldable Structural Steels for Fixed Offshore

Structures, BS 7191:1989, British Standards Institution, London, 1989

21

[11] Etube L S, Myers P, Brennan F P, Dover W D and Stacey A, “Constant and

variable amplitude corrosion fatigue performance of a high strength steel jack-up

steel”, Proceedings of the International Offshore and Polar Engineering

Conference, Vol. 4, pp123-130, 1998.

[12] Stacey A, Sharp A and King R N, “High Strength Steels used in Offshore

Installations” OMAE96, Vol. III, ASME 1996

22

8 TABLES AND FIGURES

Table 1: Mechanical properties of DILLIMAX690E

Material Steel grade

Material Specification

Minimum Specified YS UTS Tests Used

DILLIMAX690E-Z15 DILLING-E06/97 690 814 Parent and Ground Welded Plates

Requirement to UTS from yield/tensile<0.95

Table 2: Chemical composition of DILLIMAX690E % (85mm plate thickness)

C Si Mn P S Al Mo Ni Cr V Nb 0.155 0.328 1.43 0.01 0.0008 0.079 0.42 0.81 0.86 0.03 0.001

B FO-02 FO-31 N Cu Ti Co As Sn O2 N2 0.0017 0.71 0.34 0.005 0.032 0.005 - - - - -

Table 3: Summary of test parameters for part 1

Specimen Category

Specimen ID

Load Ratio R

Type of load

Stress level

Environ ment

CP level

Freq. Hz

Parent plates P1 0.013 CA 412 Air - 2 P2 0.052 CA 412 Seawater -1050 0.3 P3 0.013 CA 350 Air - 2 P4 0.052 CA 350 Seawater -1050 0.3

Welded Plates W1 0.065 CA 412 Air - 1 W2 0.0625 CA 412 Seawater -1050 0.3 W3 0.065 CA 350 Air - 1 W4 0.0625 CA 350 Seawater -1050 0.3

Table 4: Mechanical properties of SE 702 (40mm thick)

Material Steel grade

Material Specification Yield stress UTS Tests Used

SE 702 SUPERELSO 702 732 809 T-Butt Welded Plates

Table 5: Chemical composition of SE 702 % C Mn P S Al Mo Ni Cr V B

0.126 1.253 0.006 0.001 0.051 0.585 2.233 0.547 0.003 27ppm

FO-02 FO-31 N Cu Ti Co As Sn O2 N2 0.129 0.003 0.018 0.017 0.007 7ppm 71ppm

23

Table 6: Summary of test parameters for part 2

Specimen Category

Specimen ID

Clipping Ratio

Type of load

Stress level

Environ ment

CP level

Freq. Hz

T-Butt Plates T01 9.91 VA 146 Air - 1 T02 9.91 VA 146 Air - 1 T03 9.91 VA 172 Air - 1 T04 9.91 VA 172 Air - 1 T05 9.91 VA 146 Seawater -800 0.2 T06 9.91 VA 146 Seawater -1050 0.2 T07 9.91 VA 172 Seawater -800 0.2 T08 9.91 VA 172 Seawater -1050 0.2 T09 9.91 VA 146 Seawater -800 0.2 T10 9.91 VA 146 Seawater -1050 0.2 T11 9.91 VA 172 Seawater -800 0.2 T12 9.91 VA 172 Seawater -1050 0.2

Table 7: Summary of test parameters for part 3

Specimen Category

Specimen ID

Clipping Ratio

Type of load

Stress level

Environ ment

CP level

Freq. Hz

T-Butt Plates T13 8.13 VA 10min 146 Seawater -1050 0.2 T14 8.12 VA 10min 172 Seawater -1050 0.2 T15 6.80 VA 30min 146 Seawater -1050 0.2 T16 6.79 VA 30min 172 Seawater -1050 0.2

Table 8: Summary of test parameter for part 5

Plate Number

Test type Frequency Hz

CP level mV

Equivalent stress level MPa

HSS1 Air 2 NA 83.33

HSS2 Sea Water 0.2 -800 104.17

HSS3 Sea Water 0.2 -800 83.33

HSS4 Sea Water 0.2 -1050 104.17

HSS5 Sea Water 0.2 -1050 104.17

HSS6 Air 2 NA 83.33

HSS7 Air 2 NA 104.17

24

Table 9: Details of UCL specimens received at Cranfield. (Cathodic protection CP is with respect to Ag/AgCl)

Specimen No Material Loading Stress (MPa) Environment

P3 Parent 3PB 350 Air

W1 Welded 4PB 412 Air

P4 Parent 3PB 350 Sea water with

CP @ -1050 mV

W2 Welded 4PB 412 Sea water with

CP @ -1050 mV

T04 T Butt Weld 4PB VA 172 Air

T08 T Butt Weld 4PB VA 172 Sea water with

CP @ -1050 mV

Table 10: Spectrographic metallurgical analysis of DILIMAX690E-Z15 steel (16mm thick)

Element Fe Mn Si C Cr Mo Al V P Ni S Ti Cu Pb % composition 97.85 1.12 0.285 0.177 0.096 0.095 0.083 0.04 0.021 0.021 0.008 0.006 0.003 0.0002

Table 11: Mechanical properties for Part five

Material Steel grade

Material Specification

Measured Yield

Measured UTS

DILLIMAX690E-Z15 DILLING-E06/97 790 840

Table 12: SN results for Parent and Ground welded plates under CA:

Test No. Type of Loading

Plate Type Stress Range(MPa)

CP Level (mV)

N (cycles)

P1 CA Parent 412 - 307,000 * P2 CA Parent 412 -1050 132,000 P3 CA Parent 350 - 1,532,000 P4 CA Parent 350 -1050 195,000 W1 CA Welded 412 - 938,000 W2 CA Welded 412 -1050 159,000 W3 CA Welded 350 - >2,500,000 W4 CA Welded 350 -1050 296,000

* Calculated from testing at two stress levels

25

Table 13: SN results for T-butt plates under Variable Amplitude loading (Parts 2 &3)

Test No. Type of Loading Nominal Stress Range Corrected to 16mm Thick

Plate (MPa)

CP Level (mV)

N (cycles)

T01 VA 20min 192 - 327,000 T03 VA 20min 192 - 271,000 T02 VA 20min 226 - 263,000 T04 VA 20min 226 - 165,000 T05 VA 20min 192 -800 218,000 T06 VA 20min 192 -1050 140,000 T07 VA 20min 226 -800 96,000 T08 VA 20min 226 -1050 79,000 T09 VA 20min 192 -800 299,000 T10 VA 20min 192 -1050 155,000 T11 VA 20min 226 -800 161,000 T12 VA 20min 226 -1050 134,000

T13 VA 10min 192 -1050 299,000 T14 VA 10min 226 -1050 143,000 T15 VA 30min 192 -1050 152,000 T16 VA 30min 226 -1050 141,000

Table 14: Calculation of fatigue life of parent plate P1 Stress Range with

corrected thickness effects for16mm plate (MPa)

Number of cycles done

n

Predicted life form mean line N

% of Fatigue life used

340 1,000,000 9,610,000 10.4 680 268,000 1,213,000 22

The Life is calculated using Miners rule below.

� N340 n 680 �N680 =��

N 340 - n 340 ��

Ł ł

N680 = 307,000 cycles

26

Table 15: Results for different sea states for the T-butt plates

412 MPa 146 MPa

40

30

n)

20

on

(mi

du

rati

10

Sea

sta

te

0

0 50000 100000

N

150000 200000

350 MPa 172 MPa

40

in)

30

ion

(m

20

du

rat

10

Se

a s

tate

0

0 100000 200000

N

300000

Stress Range (Mpa)

CP Levels(mV)

Sea State Duration (min)

10 20 30 146 -1050 299,000 140,000

155,000 151,000

172 -1050 143,000 79,000 134,000

141,000

146 -800 218,000 299,000

172 -800 96,000 161,000

146 Air 327,000 271,000

172 Air 263,000 165,000

350 MPa

412 MPa

N

0

50000

100000

150000

200000

250000

300000

350000

-1050mV -800mV

172 MPa

Air

0

50000

100000

150000

200000

250000

N

-1050mV -800mV

146 MPa

Air

27

Table 16: Results for part five

Note: ‘+’ signifies a run-out.

Plate Number

Test type Frequency Hz

CP level mV

Stress level MPa

Cycles to failure

HSS1 Air 2 NA 83.33 1,550,000 (A)

HSS2 Sea Water 0.2 -800 104.17 2,600,000+

HSS3 Sea Water 0.2 -800 83.33 1,900,000 +

HSS4 Sea Water 0.2 -1050 104.17 690,000 (B)

HSS5 Sea Water 0.2 -1050 104.17 340,000 (B)

HSS6 Air 2 NA 83.33 3,800,000 +

HSS7 Air 2 NA 104.17 1,220,000 (B)

Table 17: SN data matrix for all tests showing % change in fatigue life P2 P4 W2 W4 T05 T06 T07 T08 T09 T10 T11 T12 T13 T14 T15 T16

Specimen Cycles

132,

000

195,

000

159,

000

296,

000

218,

000

140,

000

95,8

25

79,1

55

299,

361

155,

477

161,

400

134,

301

299,

000

142,

914

151,

774

141,

000

P1 (412, air) 301,000 -56 -35 -47 -2 -28 -53 -68 -74 -1 -48 -46 -55 -1 -53 -50 -53 P2 (412, -1050) 132,000 0 48 20 124 65 6 -27 -40 127 18 22 2 127 8 15 7

P3 (350, Air) 1,532,000 -91 -87 -90 -81 -86 -91 -94 -95 -80 -90 -89 -91 -80 -91 -90 -91 P4 (350, -1050) 195,000 -32 0 -18 52 12 -28 -51 -59 54 -20 -17 -31 53 -27 -22 -28 W1 (412, air) 938,000 -86 -79 -83 -68 -77 -85 -90 -92 -68 -83 -83 -86 -68 -85 -84 -85W2 (412, -1050) 159,000 -17 23 0 86 37 -12 -40 -50 88 -2 2 -16 88 -10 -5 -11 W3 (350, Air) 2,500,000 -95 -92 -94 -88 -91 -94 -96 -97 -88 -94 -94 -95 -88 -94 -94 -94W4 (350, -1050) 296,000 -55 -34 -46 0 -26 -53 -68 -73 1 -47 -45 -55 1 -52 -49 -52 T01 (192, air) 327,000 -60 -40 -51 -9 -33 -57 -71 -76 -8 -52 -51 -59 -9 -56 -54 -57 T02 (226, air) 263,000 -50 -26 -40 13 -17 -47 -64 -70 14 -41 -39 -49 14 -46 -42 -46 T03 (192, air) 271,000 -51 -28 -41 9 -20 -48 -65 -71 10 -43 -40 -50 10 -47 -44 -48 T04 (226, air) 165,000 -20 18 -4 79 32 -15 -42 -52 81 -6 -2 -19 81 -13 -8 -15 T05 (192, -800) 218,000 -39 -11 -27 36 0 -36 -56 -64 37 -29 -26 -38 37 -34 -30 -35 T06 (192,-1050) 140,000 -6 39 14 111 56 0 -32 -43 114 11 15 -4 114 2 8 1 T07 (226, -800) 95,825 38 103 66 209 127 46 0 -17 212 62 68 40 212 49 58 47 T08 (226, -1050) 79,155 67 146 101 274 175 77 21 0 278 96 104 70 278 81 92 78 T09 (192, -800) 299,361 -56 -35 -47 -1 -27 -53 -68 -74 0 -48 -46 -55 0 -52 -49 -53 T10 (192, -1050) 155,477 -15 25 2 90 40 -10 -38 -49 93 0 4 -14 92 -8 -2 -9 T11 (226, - 800) 161,400 -18 21 -1 83 35 -13 -41 -51 85 -4 0 -17 85 -11 -6 -13 T12 (226, -1050) 134,301 -2 45 18 120 62 4 -29 -41 123 16 20 0 123 6 13 5 T13 (192, -1050, T10) 299,000 -56 -35 -47 -1 -27 -53 -68 -74 0 -48 -46 -55 0 -52 -49 -53 T14 (226, -1050, T10) 142,914 -8 36 11 107 53 -2 -33 -45 109 9 13 -6 109 0 6 -1 T15 (192, -1050, T30) 151,774 -13 28 5 95 44 -8 -37 -48 97 2 6 -12 97 -6 0 -7 T16 (226, -1050, T30) 141,000 -6 38 13 110 55 -1 -32 -44 112 10 14 -5 112 1 8 0

28

Table 18: SN data matrix for parent plates (Part 1) showing % change in fatigue life

P2 P4 W2 W4

Specimen Cycles

132,

000

195,

000

159,

000

296,

000

P1 (412, air) 301,000 -56 -35 -47 -2 P2 (412, -1050) 132,000 0 48 20 124 P3 (350, Air) 1,532,000 -91 -87 -90 -81 P4 (350, -1050) 195,000 -32 0 -18 52

W1 (412, air) 938,000 -86 -79 -83 -68 W2 (412, -1050) 159,000 -17 23 0 86 W3 (350, Air) 2,500,000 -95 -92 -94 -88 W4 (350, -1050) 296,000 -55 -34 -46 0

Table 19: SN data matrix for T-parent plates (Part 2) showing % change in fatigue life

T05 T06 T07 T08 T09 T10 T11 T12

Cycles

218,

000

140,

000

95,8

25

79,1

55

299,

361

155,

477

161,

400

134,

301

T01 (192, air) 327,000 -33 -57 -71 -76 -8 -52 -51 -59 T02 (226, air) 263,000 -17 -47 -64 -70 14 -41 -39 -49

T03 (192, air) 271,000 -20 -48 -65 -71 10 -43 -40 -50 T04 (226, air) 165,000 32 -15 -42 -52 81 -6 -2 -19 T05 (192, -800) 218,000 0 -36 -56 -64 37 -29 -26 -38 T06 (192,-1050) 140,000 56 0 -32 -43 114 11 15 -4 T07 (226, -800) 95,825 127 46 0 -17 212 62 68 40 T08 (226, -1050) 79,155 175 77 21 0 278 96 104 70

T09 (192, -800) 299,361 -27 -53 -68 -74 0 -48 -46 -55 T10 (192, -1050) 155,477 40 -10 -38 -49 93 0 4 -14 T11 (226, - 800) 161,400 35 -13 -41 -51 85 -4 0 -17 T12 (226, -1050) 134,301 62 4 -29 -41 123 16 20 0

Specimen

Table 20: SN data matrix for T-parent plates (Part 3) showing % change in fatigue life

T13 T14 T15 T16

Specimen Cycles

299,

000

142,

914

151,

774

141,

000

T13 (192, -1050, 10min) 299,000 0 -52 -49 -53 T14 (226, -1050, 10min) 142,914 109 0 6 -1 T15 (192, -1050, 30min) 151,774 97 -6 0 -7 T16 (226, -1050, 30min) 141,000 112 1 8 0

29

Table 21: Metallurgical analysis of different high strength steels and their sources

1 2 3 4 5 6

Element Fe Mn Si C Cr Mo Al V P Ni S Ti DILIMAX690E-Z15 (1) 97.85 1.12 0.285 0.177 0.096 0.095 0.083 0.04 0.021 0.021 0.008 0.006 DILIMAX690E-Z15 (2) 1.43 0.328 0.155 0.86 0.42 0.079 0.03 0.01 0.81 8E-04 0.005 SE 702 (3) 1.253 0.126 0.547 0.585 0.051 0.003 0.006 2.233 0.001 0.003 SE 702 (4) 0.9 0.3 0.14 0.7 0.55 0.05 0.01 1.5 0.004 SE 702 (5) 1.1 0.256 0.125 0.467 0.474 0.069 0.008 0.007 1.404 5E-04 0.003 SE 702 (6) 1.05 0.25 0.12 0.51 0.48 0.08 0.02 0.009 1.34 0.001 0.01

1 2 3 4 5 6

Element Cu Pb Nb B Co As Sn FO-02 FO-31 N O2 N2

DILIMAX690E-Z15 (1) 0.003 2E-04 0.001 0.002 0.71 0.34 0.005 DILIMAX690E-Z15 (2) 0.032 SE 702 (3) 0.129 27ppm 0.018 0.017 0.007 7ppm 71ppm

SE 702 (4) 0.003 SE 702 (5) 0.185 0.003 0.004 0.001 0.011 0.007 0.003 SE 702 (6) 0.19 0.01 0.01

1. Spectographic results for DILIMAX690E-Z15 steel (Part 5).2. Chemical composition for DILIMAX690E-Z15 steel (Part 1).3. Chemical composition for SE702 steel, (Part 2).4. Quoted chemical composition for SE702 steel, maximum values (used in Tubular

Y joint tests).5. CLI Independent chemical composition for SE702 steel (used in Tubular Y joint

tests).6. UCL Independent chemical composition for SE702 steel (used in Tubular Y joint

tests).

30

85mm

85mm Parent Plate

1m a)

170mm

85mm Welded Plate

1m b)

40mm

200mm wideT-butt Plate

550mm c)

Figure 1: Dimension of specimens used for fatigue testing: a) Parent Plate b)Flush ground welded plate c) T-butt plate

31

390mm 390mm

Parent Plate

I beam

300mm 300mm

Ground Welded Plate

I beam

237mm 237mm

Figure 2: Test set up for Ground Welded Plates and Parent Plates

32

125mm 125mm

I-Beam

40mm

T-butt40mm

111mm 128mm 111mm

Figure 3: The Rig set-up for T-butt Plates

Figure 4: Dimensions of HSS T-butt welded specimens (all dimensions in mm)

33

Figure 5: Dimensions of the four point bend test rig (all dimensions in mm)

Nu

mb

er

of

tra

ns

itio

ns

450

400

350

300

250

200

150

100

50

0

Distribution of sea states in the sequence (JOSH-VACFTP2)

1 2 3 4 5 6 7 8 9 10 11 12

Sea state number

Figure 6: Distribution of sea states in JOSH sequence used in part 2

34

Stress Distribution N

um

ber

of O

ccu

ran

ce

7000

6000

5000

350 JOSH.OUT 4000 412 JOSH.OUT

350 10min Tran

412 10min Tran

3000 350 30min Tran 412 30min Tran

2000

1000

0 0 200 400 600 800 1000 1200 1400 1600 1800

Stress Range (MPa)

Figure 7: Stress Range Distribution (SRD) curves

Figure 8: Loading Sequence for JOSH.OUT

35

Figure 9: Loading Sequence for JOSHTP10.OUT

Figure 10: Loading Sequence for JOSHTP30.OUT

36

a=56o

r=3.8mm

Spot Welded ACPD Probes

Connected to U-10

Figure 11: Illustration of spot welded ACPD probes at the weld toe

Section 1 Section 2 Section 3 Section 4

Locations of band-saw cuts

Figure 12: Position of sections through fatigue specimen W2 (x0.57) (Photograph reproduced from Ref 2)

37

W1

W2

4 3 2 1

4 3 2 1

Figure 13: Fracture surfaces of the welded plates

38

P3

P4

Figure 14: Parent plate samples, P3 and P4, following fatigue testing (x0.7)

39

P3

P3

P4

Figure 15: Fracture faces of the parent plates

40

Figure 16: Photographs of polished and etched sections of specimen W1 (x1)

41

3 4

2 2

Figure 17: Photographs of polished and etched sections of specimen W2 (x1).

42

Figure 18: Micrograph of the weld, HAZ and crack initiation for W1 (x50)

Figure 19: Micrograph of the weld, HAZ and crack initiation for W1 (x50)

43

W1

W2

P3

P4 Figure 20: Fatigue crack profiles at a depth of 10mm (x200)

44

W1

W2

P3

P4 Figure 21: Fatigue crack profiles at a depth of 30mm (x200)

45

W1

W2

P3

P4

Figure 22: Fatigue crack profiles at a depth of 50mm (x200)

46

30

Traverse b

25 20 15 10 5 0

Distance (mm)

400

350

300

250

200

HV

N

HVN

250 300 350 400

Dis

tanc

e (m

m)

200 0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

b a

Section 1 (Front face x1.15)

Figure 23: Hardness measurements on specimen W2, Section 1

47

Figure 24: Microhardness values for samples W1 and W2

48

Figure 25: Vickers hardness values for P3 and P4

49

Figure 26: Fractographs of W1

50

Figure 27: Fractographs of W2

51

P3

P4 Figure 28: Fractographs of P3 and P4

52

Figure 29: Photograph of T Butt Weld T04

Figure 30: Photograph of T Butt Weld T08

53

Figure 31: Photograph of polished and etched Section 1 of specimen T04

Figure 32: Photograph of polished and etched Section 2 of specimen T04

54

Figure 33: Photograph of polished and etched Section 3 of specimen T04

Figure 34: Photograph of polished and etched Section 4 of specimen T04

55

Figure 35: Photograph of polished and etched Section 1 of specimen T08

Figure 36: Photograph of polished and etched Section 2 of specimen T08

56

Figure 37: Photograph of polished and etched Section 3 of specimen T08

Figure 38: Photograph of polished and etched Section 4 of specimen T08

57

Traverse A

Traverse B

Microhardness values for T-butt joints : Traverse A

200

250

300

350

400

450

0 2 4 6 8

Distance (mm)

HV

N T04

T08

Microhardness values for T-butt joints : Traverse B

200

250

300

350

400

450

0 2 4 6 8

Distance (mm)

HV

N T04

T08

Figure 39: Microhardness values for T butt joints T04 and T08

58

SN Curve for Parent and Ground Welded tests with Thickness correction to 16mm thick plate

S (

MP

a)

1,000

100

10

P Curve (Class C) Air

P Curve (Class C) CP

Parent Plates Air

Parent Plates CP -1050 mV

Welded Plates Air

Welded Plates CP -1050 mV

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

Figure 40: SN plot for plates under CA with comparison to the 50D

59

60

SN Curve for All T-butt Welded Plates with Thickness correction to 16mm thick plate

for Different CP Levels

10

100

1,000

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

S (

MP

a)

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates Air

T-Plates CP -800

T-Plates CP -1050

Figure 41: SN plot of T-butt plates under VA with the effects of Cathodic Protection

SN Curve for T-butt Welded Plates with Thickness correction to 16mm thick plate

for Different Transition Period

S (

MP

a)

1,000

100

10

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates Air

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

Figure 42: SN plot of T-butt plates with different Transition period

61

SN Curve for All T-butt Welded Plates with Thickness correction to 16mm thick plate

S (

MP

a)

1,000

100

10

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates Air

T-Plates CP -800

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

Figure 43: SN plot of all the tests for T-butt Plates

62

SN Curve for all the tests with Thickness correction to 16mm thick plate

1,000

100

10

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

P Curve (Class F) Air

P Curve (Class F) CP

mean P Curve (Class C) Air

mean P Curve (Class C) With CP

T-Plates Air T-Plates CP -800

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

Parent plates Air

Parent plates CP -1050 Welded Plates Air

Welded plates CP -1050

S (

MP

a)

Figure 44: SN plot all the tests done

63

Comparison of SE 702 Plates with other high strength steels tested in air.

10

100

1000

1,000 100,000 10,000,000

Number of Cycles

448-5011MPa [11]

540-586MPa [11]

730-797MPa [11]

830MPa [11]

SE 702

Ground Welded plates

mean P Curve (Class C) Air

mean P Curve (Class C) With CP

mean P Curve (Class F) Air

mean P Curve (Class F) With CP

Parent and Ground welded plates(DIL) Air

Parent and Ground Welded Plates CP

T-butt Plates(SE702) Air

T-butt Plates CP

Str

ess

ran

ge

(MP

a)

Figure 45: SN plot for the Material tested as comparison to other high strength

steels tested

64

Figure 46: Fatigue Life of the test done Plotted against each other

Cra

ck D

epth

Crack Profile ACPD09 Specimen N2.4, Sea water test 350 MPa, CP -800mV

20.00

18.00

16.00

14.00

12.00

10.00

8.00

6.00

4.00

2.00

0.00 1 2 3 4 5 6 7 8 9 10 11

Site Number

12 13 14 15 16 17 18 19

Fatigue Life

Air -1050 mV Air -1050 mV 500,000

450,000

350M

Pa

412

MP

a

350

MP

a

412

MP

a

146

MP

a

172

MP

a

146

MP

a

172

MP

a T-butt Plates Air 146 MPa 172 MPa

400,000

350,000 146MPa 172 MPa -800 mV -1050 mV -800 mV -1050 mV

300,000

250,000

200,000

150,000

100,000 Parent Plates Welded Plates

50,000

0 P3 P1 P4 P2 W3 W1 W4 W2 T01 T03 T02

Plate No. T04 T05 T09 T06 T10 T07 T11 T08 T12

Figure 47: Specimen T09, Seawater test 146 MPa, CP -800mV

65

Figure 48: Specimen T10, Seawater test 146 MPa, CP -1050mV

Crack Profile ACPD10 Specimen N3.9, Sea water test 350 MPa, CP -1050mV

20.00

18.00

16.00

14.00

12.00

10.00

8.00

6.00

4.00

2.00

0.00 1 2 3 4 5 6 7 8 9 10 11

Site Number

12 13 14 15 16 17 18 19

Cra

ck D

epth

Crack Growth

T09 Sea water test 350 MPa, CP -800mV

20.00

18.00

16.00

14.00

12.00

10.00

8.00

6.00

4.00

2.00

0.00

0 50000 100000 150000 200000 250000

Number of Cycles

300000

Cra

ck D

epth

Figure 49: Specimen T09, Seawater test 146 MPa, CP -800mV

66

Figure 50: Specimen T10, Seawater test 146 MPa, CP -1050mV

Crack Growth T10 Sea water test 350 MPa, CP -1050mV

20.00

18.00

16.00

14.00

12.00

10.00

8.00

6.00

4.00

2.00

0.00

0 20000 40000 60000 80000 100000 120000

Number of Cycles

140000 160000

Cra

ck D

epth

Crack Growth 350 MPa: AIR: T01, T03(9); -800mV: T05(4), T09(2.4) ; -1050mV: T06(5), T10(3.9); Tran30: T15(3.5)

18

20 T06 T10 T05 T03 T09 T01

16

14

12

10

8

4

6

T15

2

0 0 50000 100000 150000 200000

Number of cycles

250000 300000 350000

Cra

ck d

epth

Figure 51: Comparison of crack growth for 146 MPa

67

Figure 52: Comparison of crack growth for 172 MPa

T07T07 T07T07

T07

Crack Growth 412 Mpa: AIR: T02, T04(8); -800mV: T07(N3.7), T11(N3.4 ); -1050mV: T08(N3.8)

Tran10: T14(N2.5),Tran30: T16(2.7)

0

2

4

6

8

10

12

14

16

18

20

Cra

ck d

epth

T02

T04

T07T08 T11T14

T16

0 50000 100000 150000 200000 250000 Number of cycles

T07

T07

T07

Crack Growth 350 MPa: T01, [T06(5), T10(N3.9), T15(10 min): (-1050mV )]

412 Mpa: T02, T04(8), [T08(N3.8), T14(10 min), T16(30 min): (-1050mV )]

0

2

4

6

8

10

12

14

16

18

20

Cra

ck d

epth

T06

T08

T10

T14T15

T16

0 20000 40000 60000 80000 100000 120000 140000 160000

Number of cycles

Figure 53: Comparison of crack growth for different Transition period

68

T07T07

T07

Cra

ck d

epth

Crack Growth 350 MPa: T01, T03(9), [T05(4), T09(2.4): ( -800mV )], [T06(5), T10(N3.9): ( -1050mV )]

412 Mpa: T02, T04(8), [T07(N3.7), T11(N3.4): ( -800mV )], [T08(N3.8), T12() : ( -1050mV )]

16

18

20

T08 T07 T06

T04

T10

T11

T05

T02

T03 T09

T01

14

12

10

8

6

4

2

0 0 50000 100000 150000 200000 250000 300000 350000

Number of cycles

Figure 54: Comparison of crack growth for T-butt plates

69

SN Curve for Parent and Ground Welded tests with Thickness correction to 16mm thick plate

S (

MP

a)

10

100

1,000

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

P Curve (Class C) Air

P Curve (Class C) CP

Parent Plates Air

Parent Plates CP -1050 mV

Welded Plates Air

Welded Plates CP -1050 mV

Mean Line Parent Air(m=-3)

Mean Line Ground Air(m=-3)

Mean Line Parent CP(m=-3)

Mean Line Ground CP(m=-3)

Figure 55: Mean Lines for parents and ground welded plates

70

SN Curve for Parent and Ground Welded tests with Thickness correction to 16mm thick plate

with mean lines

10

100

1,000

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

S (

MP

a)

P Curve (Class C) Air

P Curve (Class C) CP

Parent Plates Air

Parent Plates CP -1050 mV

Welded Plates Air

Welded Plates CP -1050 mV

Parent and Ground Welded Air(m=-3)

Parent and Ground Welded CP-1050(m=-3)

Figure 56: Mean Lines for Parent and Ground welded plates together

71

SN Curve for T-butt Welded Plates with Thickness correction to 16mm thick plate

for Different CP levels

1,000

100

10

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

P Curve (Class F) CP

T-Plates Air

T-Plates CP -800

T-Plates CP -1050

Mean Line T-butt air(m=-3)

Mean Line T-butt CP-800(m=-3)

Mean Line T-butt CP-1050(m=-3)

S (

MP

a)

Figure 57: Mean Lines for T-butt plates with the effects of CP

72

SN Curve for T-butt Welded Plates with Thickness correction to 16mm thick plate

for Different Transition Period

S (

MP

a)

1,000

100

10

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

P Curve (Class F) CP

T-Plates Air

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

Mean Line T-butt air(m=-3)

Mean Line T-butt CP-1050(m=-3)

Mean Line T-butt CP-1050 Tran10(m=-3)

Mean Line T-butt CP-1050 Tran30(m=-3)

Figure 58: Mean Lines for T-butt plates with different transition period

73

SN Curve for T-butt Welded Plates with comparison for Air CP with -2SD lines

S (

MP

a)

1,000

100

10

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates Air

Mean Line T-butt air(m=-3)

T-butt air(m=-3) -2 SD(0.0954)

10,000 100,000 1,000,000 10,000,000

N (number of cycles)

Figure 59: Means line for Air data with –2SD lines

74

SN Curve for T-butt Welded Plates with comparison of CP with -2SD lines

S (

MP

a)

1,000

100

10

10,000 100,000 1,000,000 10,000,000

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates CP -800

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

T-butt CP(m=-3)

T-butt CP(m=-3) -2 SD(0.1329)

N (number of cycles)

Figure 60: Means line for CP data with –2SD lines

75

SN Curve for T-butt Welded Plates with comparison for Air and CP with -2SD lines

S (

MP

a)

1,000

100

10

10,000 100,000 1,000,000 10,000,000

P Curve (Class F) Air

P Curve (Class F) CP

T-Plates Air

T-Plates CP -800

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

Mean Line T-butt air(m=-3)

T-butt air(m=-3) -2 SD(0.0954)

T-butt CP(m=-3)

T-butt CP(m=-3) -2 SD(0.1329)

N (number of cycles)

Figure 61: Means line for Air and CP data with –2SD lines

76

SN Curve for part 5 T-butt Welded Plates S

tre

ss

ss M

Pa

1,000

100

10

10,000 100,000 1,000,000 10,000,000

Mean P Curve (Class F) Air

HSS Air

HSS -800

HSS -1050

Design P Curve (Class F) Air

Design P Curve (Class F) CP

Number of cycles

Figure 62:SN data for part five

77

SN Curve for T-butt Welded Plates tested in air with Thickness correction to 16mm thick plate

Str

es

s ss

MP

a

1,000

100

10

Mean P Curve (Class F) Air

Design P Curve (Class F) Air

T-Plates Air

T-Plates Ceramic Air

T-Plates Metallic Air

HSS Air

10,000 100,000 1,000,000 10,000,000

Number of cycles

Figure 63: SN data for all T-butt plates in air

78

SN Curve for CP at -800mV T-butt Welded Plates with Thickness correction to 16mm thick plate

Str

es

sss

MP

a

1,000

100

10

T-Plates CP -800

T-plate Ceramic CP -800

HSS -800

Design P Curve (Class F) Air

Design P Curve (Class F) CP

10,000 100,000 1,000,000 10,000,000

Number of cycles

Figure 64: SN data for T-butt plates tested at –800mV

79

SN Curve for CP at -1050mV T-butt Welded Plates with Thickness correction to 16mm thick plate

Str

es

s ss

MP

a

1,000

100

10

10,000 100,000 1,000,000 10,000,000

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

T-plate Ceramic CP -1050

HSS -1050

Design P Curve (Class F) Air

Design P Curve (Class F) CP

Number of cycles

Figure 65: SN data for T-butt plates tested at –1050mV

80

SN Curve for All T-butt Welded Plates with Thickness correction to 16mm thick plate

Str

es

s ss

MP

a

1,000

100

10

10,000 100,000 1,000,000 10,000,000

Number of cycles

Mean P Curve (Class F) Air T-Plates Air

T-Plates Ceramic Air

T-Plates Metallic Air T-Plates CP -800

T-Plates CP -1050 T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

T-plate Ceramic CP -800 T-plate Ceramic CP -1050

HSS Air

HSS -800 HSS -1050

Design P Curve (Class F) Air Design P Curve (Class F) CP

Figure 66: SN for all the T-butt test

81

SN Curve for T-butt Welded Plates tested in air with Thickness correction to 16mm thick plate

1,000

100

10

10,000 100,000 1,000,000 10,000,000

Design P Curve (Class F) Air

T-Plates Air

T-Plates Ceramic Air

T-Plates Metallic Air

HSS Air

mean Air

mean Air-2(0.1757)

Number of cycles

Str

es

s ss

MP

a

Figure 67: SN air data with trend line

82

SN Curve for CP at -800mV T-butt Welded Plates with Thickness correction to 16mm thick plate

Str

es

sss

MP

a

1,000

100

10

T-Plates CP -800

T-plate Ceramic CP -800

HSS -800

Design P Curve (Class F) CP

mean CP-800

mean CP-800 -2(0.2136)

10,000 100,000 1,000,000 10,000,000

Number of cycles

Figure 68: SN CP –800 with trend line

83

SN Curve for CP at -1050mV T-butt Welded Plates with Thickness correction to 16mm thick plate

Str

es

s ss

MP

a

1,000

100

10

10,000 100,000 1,000,000 10,000,000

T-Plates CP -1050

T-Plates CP -1050 Tran 10min

T-Plates CP -1050 Tran 30min

T-plate Ceramic CP -1050

HSS -1050

Design P Curve (Class F) CP

mean CP-1050

mean CP-1050 -2(0.2400)

Number of cycles

Figure 69: SN CP –1050 with trend line

84

9 APPENDIX I: THE CORROSION FATIGUE CRACK GROWTH RATE PLOTS

FROM CLI

The relationship in the Paris Regime: Seawater, Base metal: da = 1.73 ·10 -8 DK 3.03

dN Seawater, HAZ: da = 1.63 ·10 -8 DK 2.88 dN

85

10 APPENDIX II: THE CORROSION FATIGUE BEHAVIOUR OF JACK-UP

STEELS

P. Kilgallon, J. Healy & J. Billingham Marine Technology Centre Cranfield University

Introduction The work carried out at Cranfield University was conducted as a parallel study to a larger research programme headed by Professor Dover at University College London (UCL) [1]. The main aim of the UCL programme was to study the corrosion fatigue behaviour of high strength steel tubular welded joints as used in the construction of jack-up platforms. Included in the UCL work was the development of an elementary service load history, analysis of generated constant amplitude and variable amplitude corrosion fatigue crack growth data for tubular welded T joints and assessment of the effects of cathodic protection (CP) on corrosion fatigue crack growth rates.

Mechanical Properties The steel used in the study was Creusot-Loire Industrie SE702 which is representative of the high strength steels with yield strengths >690MPa used in the construction of jack-ups. The mechanical properties of the SE702 parent plate was determined using five standard tensile specimens of diameter 5.5mm according to BS EN 10002 [2]. The mean values obtained from these tests are given in Table 1. Charpy V notch impact tests were also carried out on duplicate specimens, the mean values are listed in Table 2 and show that the SE702 parent plate easily meets the usual offshore specification criteria of impact values of 55J at -40 C. Metallurgical examination of two welded joints was carried out and hardness surveys in the parent plate, weld metal and HAZ were conducted. Table 3 gives the mean and range of the hardness values measured.

Fatigue Testing A standard three-point edge notch bend specimen (130x16x24mm) according to BS7448 [3] was used for the fatigue testing. Prior to testing the specimens had a fatigue precrack grown in-air at a frequency of 10Hz using a reducing load method. Initially the corrosion fatigue test specimens in this programme were precharged for 8 weeks in seawater (8 C) at the test potential prior to testing as this was deemed to be the worst case. To quantify the effects of shorter precharging times a couple of fatigue tests were carried out at -850mV(Ag/AgCl) after only two weeks precharging. Tests were conducted using constant amplitude cycles of high mean stress ratio, R = 0.6, maximum load 8.5kN, and test frequency of 0.5Hz in-air and in artificial seawater [4] at 8 C with cathodic protection levels of -830 and -1080mV(Ag/AgCl). Crack growth was monitored optically and with an alternating current potential drop measuring system [5]. Test data in the form of crack length versus number of test cycles, N, was recorded and converted into the format da/dN versus stress intensity range, ÄK, using a computer spreadsheet developed at Cranfield which incorporates the ASTM 7 point polynomial analysis technique [6]. The da/dN versus ÄK curves obtained from the fatigue tests are summarised in Figure 1.

86

References 1. WD Dover - ‘Corrosion fatigue fracture mechanics of jack-up steels’ NDE Centre,

University College London. 2. British Standards Institute, BS EN 10002:1990, ‘Tensile testing of metallic materials’ ­

Part 1 - ‘Method of test at ambient temperature’, 1990. 3. British Standards Institute, BS 7448:1991, Part I - ‘Fracture mechanics toughness

tests’, 1991. 4. ASTM D1141 ‘Standard specification for substitute ocean water’ 1992.5. Kilgallon P J, ‘The effect of sulphate reducing bacteria on the hydrogen absorption

of cathodically protected high strength low alloy steels’, PhD Thesis, Cranfield University, 1994.

6. Spurrier J, Private communications, 1995, Cranfield University.

Table 1. Mean Values Obtained from Tensile Tests on SE702 Parent Plate

Yield Strength (MPa) UTS (MPa) Reduction in Area (%) Elongation (%)

748 815 63 20

Table 2. Charpy V Notch Impact Data for SE702 Parent Plate

Test Temperature (oC) Charpy Impact Energy (J)

Ambient 151

-40 120

-60 86

Table 3. Mean and Range of Vickers Hardness Numbers (10kg) Measured in SE702 Parent Plate, Weld Metal and CGHAZ of Two Welded T-joints

Measure Parent Weld Metal CGHAZ

Cap Root Cap Root

Average 257 305 251 391 306

Range 242-274 260-336 245-258 363-409 262-363

87

11 APPENDIX III: WELDING PROCEDURES FOR FLUSH GROUND WELDED

PLATES

The welding report obtained from Statoil

89

90

91

92

12 APPENDIX IV: WELDING PROCEDURES FOR T-BUTT WELDED PLATES

The welding report obtained from CLI (Cresusot-Loire Industrie)

93

94

95

Printed and published by the Health and Safety Executive

C30 1/98Printed and published by the Health and Safety Executive

C0.35 4/02

OTO 2001/079

£25.00 9 780717 623198

ISBN 0-7176-2319-X