Fatigue testing of weldable high strength steels under simulated service conditions

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Fatigue Testing o f Weldable High Strength Steels

Under Simulated Service Condition

by

Natee Tantbirojn

Submitted for the Degree of Doctor o f Philosophy

Department of Mechanical Engineering University College London

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ProQuest Number: U643508

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Abstract

There have been concerns over the effect of Cathodic Protection (CP) on weldable high

strength steels employed in Jack-up production platform. The guidance provided by the

Department o f Energy HSE on higher strength steels, based on previous work, was to

avoid overprotection as this could cause hydrogen embrittlement. However, the tests

conducted so far at UCL for the SE702 type high strength steels (yields strength around

690 MPa) have shown that the effect o f over protection on high strength steels may not be

as severe as previously thought. For this thesis, SE702 high strength steels have been

investigated in more detail. Thick (85mm) parent and ground welded plates were tested

under constant amplitude in air and seawater with CP. Tests were also conducted on

Thick (40mm) T-butt welded plates under variable amplitude loading in air and seawater

with two CP levels (-SOOmV and -lOSOmV). Different backing materials (ceramic and

metallic) for the welding process of the T-butt plates were also investigated.

The variable amplitude sequences employed were generated using the Jack-up Offshore

Standard load History (JOSH). The fatigue results are presented as crack growth and S/N

curves. They were compared to the conventional offshore steel (BS 4360 SOD). The

results suggested that the fatigue life o f the high strength steels was comparable to the BS

4360 50D steels. The effect o f increasing the CP was found to be detrimental to the

fatigue life but the effect was not large. The effect o f CP was less noticeable in T-butt

welded plates. However, in general, the effect of overprotection is not as detrimental to

the Jack-up steels as previously thought.

The load histories generated by JOSH were found to have some unfavourable

characteristics. The framework is based on Markov Chain method and pseudo-random

number generator for selecting sea-states. A study was carried out on the sequence

generated by JOSH. The generated sequences were analysed for their validity for fatigue

testing. This has resulted in recommendations on the methods for generating standard

load histories.

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Acknowledgements

I would like to thank Prof W D Dover, Dr F P Brennan and Dr L S Etube for their

supervision, guidance for the work carried out in this thesis.

I would also like to thank everyone at the NDE centre, Farid, Bijan, Mathews, Han and

Andy for their assistance with the study. Thanks also to my brother, p ’One and Rie for

taking time out to assist with my thesis. The assistance received by the staffs at the

mechanical engineering workshop is also greatly appreciated. I also want to thank Muay

for the friendship and support.

Finally I am most grateful to my family, especially my parents, for giving me the

opportunity and the love and support.

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-Table of Contents

Abstract________________________________________________________________ 1 Acknowledgements______________________________________________________ 2 Table o f Contents________________________________________________________3 T ables_________________________________________________________________ 6 Figures_________________________________________________________________ 7 Nomenclature__________________________________________________________ 10 Chapter 1________________________________________________________________ 11 Introduction and Background Work on Fatigue Testing__________________ 11 Jack-up Steels___________________________________________________ 12 Fatigue Analysis o f Welded Joints___________________________________ 13 Stresses Employed in Fatigue A nalysis__________________________ 14 Fracture Mechanics A pproach__________________________________ 18 Factors Affecting Fatigue L ife __________________________________24 Corrosion Fatigue________________________________________________ 28 Corrosion Process____________________________________________ 28 Cathodic Protection___________________________________________ 31 Variable Amplitude Loading _______________________________________ 39 Crack Retardation and Acceleration_____________________________ 39 Crack Closure_______________________________________________ 41 Fatigue Life Prediction under VA________________________________44 Summary of Chapter______________________________________________ 51 References______________________________________________________ 53 Figures__________________________________________________________5 8 Chapter 2________________________________________________________________ 64 2.0 Variable Amplitude Loading Sequence Generation______________________ 64 2.1 Wave Loading Design______________________________________________ 65 2.1.1 Mathematical Spectrum Models_________________________________ 67 2.2 Service Load Simulation for Fatigue Testing___________________________ 69 2.2.1 W irsching___________________________________________________ 70 2.2.2 Hartt and Lin_________________________________________________ 71 2.2.3 Development o f W A SH ________________________________________72 2.2.4 Jack-up Offshore Standard load History (JO SH )____________________ 77 2.3 Cycle Counting___________________________________________________ 7 8 2.4 Conclusions and Recommendations__________________________________ 79 2.5 References_______________________________________________________ 81 2.6 T ables__________________________________________________________ 83 2.7 Figures__________________________________________________________ 84 Chapter 3________________________________________________________________ 87 3.0 Fatigue Testing o f High Strength Steels_______________________________ 87 3.1 Experimental Programme___________________________________________ 88 3.1.1 Part 1 : Effect of CP on Fatigue in Parent and Ground Welded P lates 88 3.1.2 Part 2: Effect of CP on Fatigue in T-butt Plates____________________ 89 3.1.3 Part 3 : Effect o f Sub-block D uration_____________________________ 89 3.1.4 Part 4 Metallurgical Examination________________________________ 90

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3.1.5 Part 5 Investigation of Different Backing Conditions________________ 90 3.1.6 Long Life Tests (LLT) ________________________________________91 3.2 Experimental Procedure___________________________________________ 91

3.2.2 Test Monitoring Using A C P D __________________________________ 93 3.3 Metallurgical Examination_________________________________________ 95 3.4 Results o f Fatigue T e sts___________________________________________ 96 3.4.1 S/N D a ta___________________________________________________ 96 3.4.2 Fatigue Crack Growth data_____________________________________ 98 3.5 Discussion on Metallurgical Examination____________________________100 3.6 Discussion on Fatigue Test R esu lts_________________________________ 101 3.6.1 Parent and Ground Welded P la te s ______________________________ 101 3.6.2 T-butt Welded Plates_________________________________________ 102 3.7 Summaries and Conclusions_______________________________________ 108 3.8 References______________________________________________________110 3.9 T ables_________________________________________________________ 112 3.10 Figures_________________________________________________________ 120 Chapter 4 _______________________________________________________________ 150 4.0 Analysis Work on Fatigue Test R esults______________________________ 150 4.1 Revised S/N Plot of Fatigue L ives__________________________________ 150 4.1.1 Discussion on Revised S/N P lo ts_______________________________ 151 4.2 Fatigue Crack Initiation__________________________________________ 153 4.2.1 Crack Initiation Results_______________________________________ 153 4.2.2 Discussion of Crack Initiation Lives_____________________________ 154 4.3 Comparison o f High Strength Steels R esults__________________________156 4.3.1 Discussion on Comparison Study_______________________________ 157 4.4 Crack Growth R ate______________________________________________ 158 4.4.1 Calculation o f Crack growth Rate_______________________________ 160 4.4.2 Discussion o f Crack Growth Rate_______________________________ 161 4.5 Discussion on Welded Joint Classification for T-butt Welded P late_______163 4.5.1 Inclusion of SCF Instead of Using Welded Joint Classification_______ 164 4.6 Summaries and Conclusions_______________________________________ 166 4.7 References______________________________________________________168 4.8 T ables_________________________________________________________ 170 4.9 Figures_________________________________________________________ 174 Chapter 5_______________________________________________________________205 5.0 Analysis o f Standard Load History G enerator________________________ 205 5.1 Standard Load Histories Generator Framework_______________________ 206 5.1.1 Markov Chain______________________________________________ 207 5.1.2 Pseudo Random Binary Shift (PRBS) generator__________________ 211 5.1.3 Sea-States Selection_________________________________________ 212 5.1.4 Operation of WASH and JO S H ________________________________ 213 5.2 Analysis of Generated Sequences___________________________________ 214 5.2.1 Stationary points____________________________________________ 215 5.2.2 Equivalent Stress____________________________________________ 215 5.2.3 Power Spectral Density (PSD) function_________________________ 216 5.2.4 Starting Sea-States__________________________________________ 217 5.2.5 Sea-State Sub-block Duration__________________________________ 218 5.3 Discussion o f Analysed Sequences__________________________________ 219

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-5.4 Conclusions and Recommendations________________________________ 223 5.5 R eference_____________________________________________________ 224 5.6 T ables________________________________________________________ 226 5.7 Figures________________________________________________________ 231 Chapter 6_______________________________________________________________242 6.0 Summary o f Thesis______________________________________________ 242 6.1 Summary o f Conclusions and Recommendations_____________________ 242 Appendix A_____________________________________________________________246 Appendix B _____________________________________________________________280 Appendix C _____________________________________________________________281

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-Tables

Table 2.1: Statistic and fatigue damage calculations fo r simulated North Sea condition proposed by

Wirsching [2 .7 ]...83

Table 2.2: Original sea-state data fo r WASH...83

Table 3 .1: Mechanical properties o f DILLIMAX690E...112

Table 3.2: Chemical composition o f DILLIMAX690E %...112

Table 3.3: Summary o f test param eters fo r p a rt 1...112

Table 3.4: M echanical properties o f SE 7 0 2... 113

Table 3.5: Chemical composition o f SE 702 %...113

Table 3.8: Details o f UCL specimens examined by Cranfield University...114

Table 3.9: Summary o f test param eters fo r p a rt 5 T-butt p la te s...114

Table 3.10: M echanical properties fo r Long Life Tests (LLT)...114

Table 3.11: M etallurgical analysis o f different high strength steels and their so u rces...115

Table 3.12: S/N results fo r Parent and Ground welded plates under CA:...115

Table 3.13: S/N results fo r T-butt plates under Variable Amplitude loadin g...116

Table 3.14: Fatigue results fo r LLT...117

Table 3.15: Calculation o f fatigue life o f parent plate P I...117

Table 3.16: Results f o r different sea-states fo r the T-butt p la te s...118

Table 3.17: S/N data matrix fo r all tests showing % change in fatigue life...118

Table 3.18: S/N data matrix fo r parent plates (Part 1) showing % change in fatigue life...119

Table 3.19: S/N data matrix fo r T-parent plates (Part 2) showing % change in fatigue life...119

Table 3.20: S/N data matrix fo r T-parent plates (Part 3) showing % change in fatigue life...119

Table 4.1: Data o f the initiation life o f the fatigue tests...170

Table 4.2: Data o f the initiation life o f the other fatigue tests...171

Table 4.3: Fatigue life result fo r LLCF....172

Table 4.4: Spectrographic metallurgical analysis o f DILLIMAX690E-Z15 steel...172

Table 4.5: Chemical composition o f DILLIMAX690E % (85mm p late thickness)...172

Table 4.6: Mechanical properties fo r LLCF...173

Table 4.7: Mechanical properties fo r 85mm p la te s...173

Table 4.8: Data o f the fatigue tests...173

Table 5.1: Transition matrix fo r 10 minutes sub-block...226

Table 5.2: Sea scatter diagram fo r number o f occurrence from B ery/F rig...226

Table 5.3: Sea-states selected fo r JOSH...227

Table 5.4: Expected duration time fo r JOSH sea-states...227

Table 5.5: The Transition probabilities fo r different sub-block duration tim e...228

Table 5.6: Number o f occurrence fo r each sea-states with different starting sea-state fo r the 12 sequences (1000 transitions)...229

Table 5.7: % o f occurrence fo r each sea-states with different starting sea-state f o r the 12 sequences (1000 transitions)...229

Table 5.8: % Occurrence o f starting sea-states 1 fo r different number o f transitions...230

Table 5.9: % occurrence fo r different duration length...230

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-Figures

Figure 1.1: Definition o f H ot Spot stre ss...58

Figure 1.2: Co-ordinate f o r stresses fo r the near crack tip region...58

Figure 1.3: Three basic crack modes [ 1 .9 ]...59

Figure 1.4: Paris-Erdogan P lo t...59

Figure 1.5: Stress gradient fo r different plates thickness...60

Figure 1.6: Schematic representations o f the combination o f mechanical fatigue and environmentally assisted crack growth [1 .3 3 ]...61

Figure 1.7: Residual stress left behind due to p la sticity...62

Figure 1.8: Crack Closure due to the residual w ake...62

Figure 1.9: Different Closure Mechanisms roughness induced closure...63

Figure 2.1: Energy spectrum from a North Atlantic storm w ave...84

Figure 2.2: Typical double peak density function [ 2 .1 0 ]...84

Figure 2.3: Schematic representation o f creating response spectral density [2 .1 1 ]...85

Figure 2.4:PSD fo r WASH (W) [2 .1 0 ]...85

Figure 2.5: Schematic diagram f o r shift register and shifting feedback fo r the PRBS technique [2.10]... 86

Figure 2.6: Simplified model o f a Jack up [2 .1 1 ]...86

Figure 3.1: Dimension o f specimens used fo r fatigue testing:...120

Figure 3.2: Test set up fo r the ground welded plates and parent p la te s...121

Figure 3.3: Test set-up fo r the T-butt p la te s...122

Figure 3.4: Picture o f T-butt p late set up fo r an air test...122

Figure 3.5: Environment cell used fo r the parent plate tests...123

Figure 3.6: Distribution o f sea-states in JOSH sequence used in P art 2...123

Figure 3.7: Stress Range Distribution (SRD) curves...124

Figure 3.8: Loading Sequence f o r JOSH with a 20 minute sea-state duration...124

Figure 3.9: Loading Sequence f o r JOSH with a 10 minute sea-state duration...125

Figure 3.10: Loading Sequence fo r JOSH with a 30 minute sea-state duration...125

Figure 3.11: ACPD schematic diagram...126

Figure 3.12: Illustration o f spot welded ACPD probes at the w eld to e...126

Figure 3.13: Spot w elded ACPD probes on a welded T-butt jo in t....127

Figure 3.14: S/N p lo t fo r plates under tested under C A...128

Figure 3.15: S/N p lo t o f T-butt plates under VA with the effects o f C P...129

Figure 3.16: S/N p lo t o f T-butt plates with different duration p erio d under C P...130

Figure 3.17: S/N p lo t fo r the Material tested as compared to results from BS 4360 50D steel [3.11] .... 131

Figure 3.18: Mean Lines fo r parents and ground w elded p la te s...132

Figure 3.19: Mean Lines f o r Parent and Ground welded plates together...133

Figure 3.20: Mean Lines fo r T-butt plates with the effects o f C P...134

Figure 3.21: Mean Lines fo r T-butt plates with different duration period....135

Figure 3.22: Mean line fo r Air data with -2 S D lin e s...136

Figure 3.23: Mean line fo r CP data with -2 S D lin e s...137

Figure 3.24: Mean line fo r Air and CP data with -2 S D lin es...138

Figure 3.25: Comparison o f fatigue lif e...139

Figure 3.26: Specimen T09, Seawater test 1 4 6 MPa, CP -SOOmV...139

Figure 3.27: Specimen TIO, Seawater test 1 4 6 MPa, CP -1050m V...140

Figure 3.28: Specimen T09, Seawater test 146 MPa, CP -800m V...140

Figure 3.29: Specimen TIO, Seawater test 1 4 6 MPa, CP -1050m V...141

Figure 3.30: Comparison o f crack growth fo r 146 M P a...141

Figure 3.31: Comparison o f crack growth fo r 172 M P a...142

Figure 3.32: Comparison o f crack growth fo r different JOSH duration p e rio d...142

Figure 3.33: Comparison o f crack growth fo r T-butt p la tes...143

Figure 3.34: Air at 146 M P a...143

Figure 3.35: Crack Growth fo r Ceramic backing T-plate, TCOl at 146 MPa and TC02 at 104M Pa 144 Figure 3.36: Finished CP test fo r parent plate with white oxide surface (calcareous deposit)...144

Figure 3.37: Parent plate, CP (350 M P a )...145

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-Figure 3.38: Ground welded plate, CP (412 M P a )...145

Figure 3.39: Ground welded plate, air (412 M Pa)...146

Figure 3.40: Side profile o f welded plate, CP (412 M P a)...146

Figure 3.41: Close up o f the welded plate, CP (412 M Pa)...147

Figure 3.42: Side profile o f ceramic backing welded T-butt p l a t e...147

Figure 3.43: Side profile o f metallic backing welded T-butt p la te...148

Figure 3.44: Side profile o f cracked metallic backing welded T-butt p la te...148

Figure 3.45: Weld profile o f cracked metallic backing welded T-butt p la te...149

Figure 4.1 : R evised SN curves fo r air te s t...174

Figure 4.2: Revised SN curves fo r -SOOmV test...175

Figure 4.3: Revised SN curves fo r -1050m V test...176

Figure 4.4: Diagram showing single sided and double-sided welded plates...177

Figure 4.5: Comparison o f fatigue initiation lives as bar ch a rt...177

Figure 4.6: Percentage offatigue initiation life to fatigue life...178

Figure 4.7: Fatigue initiation life in air as S/N curve...179

Figure 4.8: Fatigue initiation life in seawater under CP as S/N curve...180

Figure 4.9: Fatigue initiation life fo r other air tests ofH SS at UCL...181

Figure 4.10: Fatigue initiation life fo r HSS tubular join ts tested at U C L...182

Figure 4.11: Dimensions o f HSS T-butt welded specimens (all dimensions in m m )...183

Figure 4.12: Dimensions o f the fou r point bend test rig (all dimensions in mm)...183

Figure 4.13: S/N curve fo r HSS from Lindley [4 .9 ]...184

Figure 4.14: Comparison S/N air curve fo r HSS from Lindley [ 4 .9 ]...185

Figure 4.15: Comparison S/N seawater curve fo r HSS from Lindley [ 4 .9 ]...186

Figure 4.16: S/N curve fo r Long life te s t...187

Figure 4.17: S/N curve fo r all air tests on T-butt p la te s...188

Figure 4.18: S/N curve fo r T-butt plates with CP -SOOmV...189

Figure 4.19: S/N curve f o r T-butt plates with CP -1050m V...190

Figure 4.20: Crack Growth Rate against Stress Intensity Factor R a n g e...191

Figure 4.21: Loading Sequence fo r JOSH.OUT...191

Figure 4.22: Specimen T09, Seawater test 1 4 6 MPa, CP -SOOmV...192

Figure 4.23: Comparison o f crack growth fo r T-butt plates...192

Figure 4.24: Crack Growth fo r Ceramic backing T-plate...193

Figure 4.25: Crack Growth Rate against Time...193

Figure 4.26: Crack Growth Rate against Stress Intensity Factor Range fo r A ir...194

Figure 4.27: Crack Growth Rate against Stress Intensity F actor Range fo r C P...194

Figure 4.28: AK v d a/dN for TCOI...195

Figure 4.29: AK v da/dN fo r T C 02...195

Figure 4.30: AK v da/dN for T03...196

Figure 4.32: AK v da/dN for T09...197

Figure 4.33: AK v da/dN for TIO...197

Figure 4.34: AK v d a/dN for T07...198

Figure 4.36: Crack growth and crack growth rate fo r TC02...199

Figure 4.37: S/N Curves fo r class C, F and F 2...200

Figure 4.38: S/N curve fo r all air tests on T-butt p la te s...201

Figure 4.39: S/N curve fo r T-butt plates with C P...202

Figure 4.40: Comparison p lo t fo r all the T-butt tests...203

Figure 4.41: S/N p lo t with SCF....204

Figure 5.1: Amplitude p lo t o f the sequence generated by JO SH...231

Figure 5.2: Fatigue Crack growth fo r a T-butt welded p late in air...231

Figure 5.3: Rainfiow count fo r JOSH 20m in...232

Figure 5.4: Monte Carlo simulation f o r sea-state transition [5 .2 ]...232

Figure 5.5: % error o f equivalent stress range fo r one JOSH 20 min sequence...233

Figure 5.6: % error o f the equivalent stress range fo r eight repeated sequences...233

Figure 5.7: Comparison o f PSD obtained from spectral analyser and F F T...234

Figure 5.8: Another comparison PSD against FFT....234

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-Figure 5.9: Different starting sea-state...235

Figure 5.10: Plot o f number o f occurrence fo r different starting sea-state...236

Figure 5.11: PSD f o r one JOSH 20 min sequence...236

Figure 5.12: PSD f o r different length o f JOSH 20 min sequence...237

Figure 5.13: % changes fo r one frequency o f PSD JOSH 20 m in...237

Figure 5.14: Comparison o f the % error fo r PSD and Equivalent stre ss...238

Figure 5.15: Number o f Occurrence fo r different number o f transition...239

Figure 5.16: Plots o f sequence generated from different length sub-block duration...240

Figure 5.17: PSD f o r sequence generated from different length sub-block duration...240

Figure 5.18: % Occurrence o f sea-state fo r different duration len gth...241

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-Nomenclature

ACPD Alternating Current Potential Difference

ACFM Alternating Current Frequency Measurement

Ag/AgCl Silver/Silver Chloride reference electrode

CA Constant Amplitude

e u

Creusot Loire Industrie

CP Cathodic Protection

HAZ Heat Affected Zone

HSE Health and Safety Executive

JOSH Jack-up Offshore Standard load History

NDE Non Destructive Evaluation

PSD Power Spectral Density

SCF Stress Concentration Factor

SRD Stress Range Distribution

SRPD Stress Range Probability Distribution

UCL University College London

VA Variable Amplitude

VACF Variable Amplitude Corrosion Fatigue

UEO Underwater Engineering Group

WASH Wave Action Standard load History

a Crack Depth

c Crack half length

C Paris Law constant

A Probe gap employed by ACPD technique

£ Spectral bandwidth parameter

F Force

f Frequency

Hs Significant wave height

K Stress intensity factor

m Paris Law constant

N Number of cycles

R Loading ratio

S Stress range

^yy(f) Wave loading spectral density function

Y Stress intensity correction factor

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Chapter 1 Introduction and Background Work on Fatigue Testing

Chapter 1 1.0 Introduction and Background Work on Fatigue Testing

Recent years have seen a growing trend in the offshore industry to use Jack-up platforms

for production purposes due to the economic depression in the 1990’s. The harsh

economic environment meant instead of just employing these mobile platforms for

exploration purposes, they are also utilized for production purposes. Consequently, Jack-

up platforms are deployed for service that was for which they were not intended.

Jack-up platforms are mobile structures and therefore high strength steels are used in their

construction to economise on weight. The design o f these structures is predominantly

based on static strength considerations. For exploration, such an analysis is judged

sufficient since the role of deployment and operation are generally restricted to short

period. However, for production purposes, the long-term life or service life o f these

structures needs to be assessed. This assessment cannot be based on the existing

understanding o f medium strength steels, which are the more conventional platform

material. Even though high strength steels exhibit better performance under static loading,

there is no guarantee of improved performance under cyclic loading. There is also concern

on the brittle nature o f Jack-up steels caused by hydrogen.

Extensive studies have been conducted on the fatigue performance o f medium strength

steels, but relatively little is known about higher strength steels. The main concern

associated with these high strength steels, is the tendency for embrittlement in the

presence o f a hydrogen rich environment. This is particularly important as the corrosion

prevention method, Cathodic Protection (CP), is viewed as damaging to the higher

strength steels by encouraging the hydrogen activities.

Most o f the fatigue studies are usually carried out under constant amplitude loading, but

random loading is more realistic of the actual service condition. This is particularly

critical in the design o f mobile offshore platform, where the steels are susceptible to the

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-C hapter 1 Introduction and Background Work on Fatigue Testing

environment and the loading condition. The offshore fatigue guidance under random

loading is limited, particularly for corrosive environment. This chapter reviews the current

practice in the field of fatigue testing including the variable amplitude loading and its

effects, with emphasis placed on high strength Jack-up steels and corrosion fatigue.

1.1 Jack-up Steels

Jack-up platforms are mobile offshore units, the design o f which can vary considerably.

Over the years. Jack-up rigs have had different numbers o f legs; nowadays most Jack-up

designs are based on three legs. They are usually towed to the desired location with their

legs raised. After reaching their destination, the legs are jacked down into the water to

make contact with the sea floor. The jacking process is powered by either electric motors

or hydraulic jacks, which are locked into the rack teeth that are welded to the side o f each

leg. The Jack-up leg foundations usually comprise spud cans, which are large steel plates

that penetrate the sea floor. The platforms are consequently jacked up above the water and

prepared for drilling.

The Jack-up steels need to have high strength in order to save on weight and bear the

required load. The weldable high strength steels associated with Jack-up construction

could be defined as having yield strengths in the region of 450 to 800 MPa. A steel

commonly used in the production of new Jack-up platform designs is SE702. This is a

weldable high strength steel with a yield strength o f about 690 MPa. The basis for using

the word weldable is because in the past, the welding process limited the use o f higher

strength steels. There was concern about the toughness of the material. However, by

lowering the carbon content o f steels, such as SE702, and increasing the toughness, they

become more weldable. The high strength o f the steels is achieved by quenching,

tempering and alloying processes. Thus, the modem quenched and tempered high strength

steels have the strength, toughness and weldability required for the constmction of a Jack-

up platform.

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-Chapter 1 Introduction and Background Work on Fatigue Testing

The susceptibility o f higher strength steels to hydrogen embrittlement is of particular

concern. The induction o f hydrogen is usually caused by corrosion protection system such

as CP. Problematic hydrogen assisted cracking o f the high strength steels used in earlier

Jack-ups was highlighted in a report by Davey [1.1] for the UK Health and Safety

Executive, Offshore Safety Division. This report showed that cracks could occur in the

chord o f Jack-up legs under North Sea conditions and that the cracks all initiated in the

Heat Affected Zone (HAZ) o f the base material. The study concluded that the problem

was due to hydrogen assisted cracking, which was caused by cathodic overprotection.

Following the study on high strength steels using slow strain rate tests, it was

recommended that a Cathodic Protection (CP) level on Jack-up platforms should not

exceed -825mV (Ag/AgCl). The study concluded that any steel with a yield strength

above 600 MPa should be considered susceptible to hydrogen induced stress corrosion

cracking unless tests showed otherwise.

Due to the findings o f Davey, the 1990 UK HSE guidance on design and construction of

offshore structures [1.2] requires all Jack-up production platforms to be fitted with CP

limiters. However, the employment of CP limiters can be a very costly and unnecessary

process. The introduction of improved material properties and manufacturing processes

has altered the fatigue performance of new generations of Jack-up steels different to that

of old generations. Previous corrosion fatigue studies at UCL on higher strength steel

tubular joints by Etube [1.3] and Myers [1.4] indicated that overprotection is not as

damaging to the fatigue performance as first thought. The publish fatigue data for newer

generations o f Jack-up steels is limited, particularly under corrosive environments, and it

is only possible to determine the performance o f a particular high strength steels via

laboratory tests, as discussed in section 1.3 under corrosion fatigue.

1.2 Fatigue Analysis o f Welded Joints

Welding is the preferred joining technique used in the fabrication o f offshore structures.

The welded offshore structures are subjected to variable loading conditions during their

service lives. Thus, rigorous stress analysis and fatigue studies are carried out during the

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design o f the offshore structures identifying and analysing critical areas. This section

reviews topics associated with stress analysis and fatigue assessment o f welded structures.

This includes the main controlling parameter in fatigue crack initiation and propagation

mechanisms. The design guidance and recommendations used for the design of offshore

structures are reviewed with emphasis on the fracture mechanics and stress/life (S/N)

analysis techniques. The effects of corrosion fatigue and variable amplitude loading are

discussed in this section.

1.2.1 Stresses Employed In Fatigue Analysis

In stress analysis studies, stresses within welded joints are identified for design and

fatigue life calculations. As there are numerous definitions o f stress used for fatigue

analysis, this section presents the main definitions used throughout this thesis.

Nominal stress, cTnom, is the calculated stress remote from any discontinuities using basic stress formulae. The stress calculated does not take into account any influence of the

welded attachment, which would normally alter the stress distribution. An example for

tensile loading is shown in equation (1.1), where F is the force applied and A is the cross sectional area.

^nom - —

The concept o f a nominal stress avoids the need for complicated analysis. In complex

welded structures, nominal stress is often used to identify the loading of the structure

since the actual stress experienced is too complex to calculate.

Another form o f stress usually employed in fatigue calculations is the hot spot stress. This

is a form of geometric stress. Hot spot stress is a term used to describe the stress at a

critical point in the structure. This is usually located at a geometric discontinuity or a

notch. The hot spot stress ignores the stress concentration effect due to the weld toe. It can

be obtained by measuring the strain via strain gauges or calculated using parametric

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equations. The hot spot stress approach is used for components where the crack is

assumed to grow at the weld toe. Thus, it is not suitable for crack growth from embedded

defects or the root o f a fillet weld.

Most hot spot stresses can be determined from experiments by measuring the strain near

the weld toe at two specified locations from the weld toe. The stresses are then linearly

extrapolated to work out the stress at the weld toe. Usually the strain gauges, employed

for stress measurement, should be sufficiently far away from the weld toe, so that the

weld toe notch does not affect the measurements as represent in Figure 1.1. Various

methods for obtaining hot spot stress via strain gauges for tubular joints can be found in

guidance such as published by UEG offshore research [1.5].

1.2.1.1 Stress Concentration Factor

While nominal stress is rather easier to calculate for a structure, more detail is required for

a more complex analysis. The maximum stress for a particular joint can be obtained from

multiplying the nominal stress to the Stress Concentration Factor (SCF). The stress

concentration factor is the ratio of maximum stress at the point o f interest, Cmax to the nominal stress, anom- The equation (1.2) shows the relationship between SCF and cTmax’

^m ax (1’2)

This is particularly useful for calculating the stress in welded connections such as tubular

joints. Usually the SCFs can be obtained experimentally or by employing parametric

equations for welded offshore structures. Most o f the experimental methods are very

similar, involving the use o f strain gauges. For comparison of methodology, the

Underwater Engineering Group [1.5] discusses and compares the existing SCF design

codes for offshore structures. The usual practice for experimental measurement is to place

two strain gauges close to the weld toe so that the stress gradient can be measured. The

maximum stress is usually taken as the hot spot stress and is measured by linear

extrapolation to the weld toe. Strain gauges are often used to measure the hot spot stress

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on the actual structure. Another method for obtaining or predicting the SCF is to use

empirical formulae. Parametric equations based on experimental work or from finite

element analysis could be used to calculate SCF. The parametric solutions tend to be

conservative. For tubular joints, finite element analysis on thin shells can be used for

obtaining parametric equations for calculating SCFs [1.6]. For the analysis work on

welded T-butt plates, the parametric method for obtaining SCFs can be found from

published work based on 2D finite element models [1.7].

1.2.1.2 SN Approach

The SN approach is widely used in the design and assessment o f offshore structures. Its

practicality lies in its simplicity and that the results are based on experimental fatigue

data. For the S/N plot, the stress range, S, is conventionally taken as the hot spot stress range for tubular joints and nominal stress range for welded plates and N is the fatigue life

of the structures or the number of cycles until failure.

The plots for S/N design curves employed for the design of offshore platforms are usually

based on the BS 4360 50D type steels. The S/N curves can be obtained from published

literature standards such as the 1990 HSE Fatigue Guide [1.2]. The basic design S/N

curves are usually represented in the form of equation (1.3) and plotted on a log-log scale.

l-ogio W = Logi^

Z .% (5^ )

Where

N is the predicted number to failure under stress range Sb

K] is a constant relating to mean S/N curve

d is the number of standard deviations below mean cris the standard deviation of Logio N

m is gradient o f the inverse slope of the curve.

Different welded specimens can be represented by various S/N curves. These are

categorised into different classes for different loading types and welding geometry. For

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example, the tubular joints are represented by the T curve. For welded plates, the curves

are divided into many classes, where the joint types are described in most fatigue

literatures such as the 1990 HSE Fatigue Guidance.

The data employed for the construction o f the fatigue S/N curves were obtained

experimentally. The best-fit line or the mean line can then be drawn through the points.

The standard deviation is calculated from the scatter o f the data. For the tubular joint

standard, the guidelines are for plate thickness o f 16mm and 32 mm. For safety reasons,

the design line is twice the standard deviation below the mean line.

The S/N curve from the guidance should be applied with caution to other grade of steels,

as it is only a guide for medium strength steels, such as BS 4360 50D. The data are also

obtained primarily under constant amplitude conditions. Thus, it is not necessarily

representative of a variable amplitude fatigue test. Constant amplitude tests for steels

usually display a fatigue limit, where any stress range below a certain level would have no

effect on the fatigue life. Under variable amplitude loading, the sequence consists o f a

range o f stress amplitudes. The larger stress ranges can propagate the crack, resulting in a

different behaviour nearer the fatigue limit. The S/N curve can be employed for designing

structures [1.8] that will undergo variable amplitude loading, but this could lead to errors,

since the data were not obtained from such tests. Thus more emphasis should be made

toward testing under realistic loading condition, since there are limited fatigue life data

available for tubular joints and welded plates tested under variable amplitude loading.

Studies carried out under variable amplitude loading are difficult to compare due to the

difference in the loading sequences and the method employed for equivalent stress range

calculation. Some older work employed the root mean square stress range method to

calculate the equivalent stress range while more recent studies would employ root mean

cube stress range. Root mean square method is usually associated to the energy for most

signal analysis. However for cumulative fatigue damage calculations, root mean cube is a

better fit for the slope of the crack growth curve o f welded joints [1.9]. For unprotected

joints in seawater, the 1990 HSE Fatigue Guidance suggested that the basic S/N curve be

reduced by a factor o f 2 for all joint classes. However, there is no clear guideline on

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higher strength steels (yield higher than 400 MPa) and the guidance suggests that the

reduction factor o f 2 might be inadequate.

1.2.2 Fracture Mechanics Approach

The fracture mechanics approach is more powerful than the SN approach, in the sense that

it can be used to model and predict the propagation rates of cracks. The idea is based on

using stress analysis to calculate the stress intensity factor. Most o f the early studies of

fracture mechanics were related to the energy-based theory by Griffith [1.10]. This was a

linear approach and more suitable for brittle fracture. His theory stated that a crack would

propagate when the decrease in elastic strain energy is at least equal to the energy required

to create the new crack surface. Griffith’s work was mainly used to show the importance

o f a notch and brittle fracture. This work ignored plastic deformation in the crack region,

which made it unsuitable for ductile metals and limited its use for most engineering

application. Most o f the calculations in fracture mechanics nowadays are based on stress

intensity factor (SIF). Irwin [1.11] first thought up this idea of stress intensity factor in

1957. He expressed the stress singularity ahead o f the crack tip in terms of the scalar

quantity or SIF, also known as K. This led to Linear Elastic Fracture Mechanics (LEFM), which attempted to explain fatigue crack propagation in terms of stress intensity factor.

Paris and Erdogan [1.12] suggested the relationship between fatigue crack advance per

cycle (da/dN) and the stress intensity range (AK). Though this assumption lacked scientific basis, it has found wide applications in modem fracture mechanics due to its

simplification of fatigue crack growth. This came to be known as the Paris law or the

power law.

For relatively ductile materials, LEFM is not applicable due to the occurrence of plastic

deformation. Irwin [1.11] recognised this and estimated the plastic zone ahead of the

crack tip. As long as the plastic zone size is small compared to the geometry of the

specimen, LEFM can be employed. The plastic zone correction is applied by increasing

the crack length to include the plastic zone. However, for larger plastic zone, other

methods are required instead.

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In 1961 Wells [1.13] suggested the idea of measuring the crack tip opening to overcome

the limitation of LEFM. He noted that during plastic deformation, the crack faces move

apart, which led to the work in crack tip opening displacement (CTOD). Wells assumed

that there must be a critical crack tip opening size for fracture to occur. For this same

problem. Rice [1.14] in 1968 developed the J integral idea, as the energy release rate o f a

non-linear material. CTOD is mostly used and developed in the UK, while J is developed in the US. Most o f the fatigue work carried out in this thesis was more concerned with

less ductile material like high strength steels and low stress failure by fatigue. This allows

LEFM to be applied to ductile metals. In this section different fatigue mechanisms and

analysis techniques associated with fatigue testing for offshore structures are reviewed.

1.2.2.1 Different Stages o f Fatigue Crack Growth

The fatigue life can be divided into three different stages, the initiation stage, crack

propagation stage and the final failure of the component. In this section, the fatigue

behaviours are discussed for the different stages.

Initiation is linked to microscopic material behaviour. Cracks are initiated by localised

cyclic plastic deformation, as a result of dislocation motion under cyclic strain. Since

dislocation mobility is greater at a free surface than in the bulk o f the material, in general,

fatigue initiation is a surface phenomenon. Normally cracks will initiate at locations with

the largest bending stress or strain range.

Under successive load cycles, dislocation motion may occur in different adjacent slip

planes in the crystalline material, resulting in extrusions and intrusions on the surface.

The formation of these intrusions is the initiation stage. Microscopic slip may occur in

single grains at a stress level below the general yield stress of the material. Under

continued load cycling, one or more intrusions may develop into what is commonly

denoted as stage I crack growth. In this first stage, the crack grows in a shear mode, the

same crystallographic plane upon which slip originally took place, i.e. in a plane oriented

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at 45 degrees to the maximum principal stress direction. Commonly these stages are

referred to as initiation or crack nucléation. In most welded structures, small flaws are

normally expected in the welded region, unless the surface has been treated, hence the

initiation period is short. Thus, the initiation period could be improved by introducing

compressive residual stress by shot peening or removing o f small flaws by grinding.

Therefore, in the case o f T-butt welded plates, the initiation period of the crack is

expected to be short when compared to the crack propagation period. However, for plain

surface structures, where the surface has been smooth-grounded and has no stress

concentration, the initiation stage would dominate the fatigue life. It is difficult to define

exactly when the initiation period ends and the crack propagation period commences. In

the case o f the fatigue study carried out in UCL, the initiation period is normally taken as

the number o f cycles required until a fatigue crack can be detected. This method is very

dependent on the operator and the technique employed in detecting the crack. From the

past work [1.4] carried out on tubular joint with Alternating Current Potential Drop

(ACPD) technique at UCL, a crack depth of 0,1 mm is usually taken as the initiation

point.

The crack propagation stage is of most interest, because many o f the offshore structures

are welded and are affected by geometric stress, and usually the propagation period is the

more dominant in the fatigue life of the structures. As the crack increases in depth from

the initiation stage, the crack growth direction will be in a direction perpendicular to the

largest tensile stress. From this point onward, the maximum principal stress and the

associated crack tip stress intensity factor range will be the principal driving force for the

crack growth.

In the final stage, the crack growth rate is rapid. The main factors affecting the crack

growth rate are the mean stress and the microstructure o f the material. However, the more

likely failure scenario for the specimen is due to the static load as the remaining static

strength o f the structure would have been sufficiently low for it to fail under a single

fatigue load.

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1.2.2.2 LEFM

The basis for fracture mechanics is in understanding the stress distribution around the

crack tip. The stress field near the crack tip is normally described by the stress intensity

factor, K. Although K is known as the stress intensity factor, it is not a factor as unlike Stress Concentration Factor (SCF), it has units. The stress intensity factor is only suitable

for small-scale yield linear elastic material, and the method is known as Linear Elastic

Fracture Mechanics (LEFM). This method does break down, when there is a large amount

of plasticity or non-linearity. However, as most of the studies for offshore structures are

subjected to relatively low stress levels, the plastic zone is likely to be small, making

LEFM adequate for most fatigue calculations.

Fracture mechanics calculations rely on using the Stress Intensity Factor (SIF) to describe

the crack tip condition. SIF is a function of crack geometry and the applied stress. The

stress field at the crack tip can be described by the equation (1,4). Figure 1.2 describes the notations used.

Where is the stress field component, r and 6 are the polar co-ordinates and fxy is the function o f 9.

For Mode I loading, the crack tip stress field on the expected fracture plane can be

represented as:

yy

r

The use o f the notation / in Æ}, this refers to Mode I (one) loading type for the crack

specimen. The crack surface can be displaced in three loading modes or in any

combination o f the three as represented in Figure 1.3. Mode I is tensile opening mode,

where the crack faces are separated normal to the plane o f the crack. Mode II is in plane

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sliding mode and Mode HI is tearing shear mode. Mode I is the most common mode and

the main factor in most fatigue crack growth situations. Mode II and IQ tends to be less

significant when compared to Mode I in most practical applications.

In order to be able to use K for normal engineering applications, modification is required for equation (1.5). This is written in the form o f equation (1.6) for Ki. Y is placed in the equation, as the finite correction factor for the plate and other properties like environment

and loading conditions. However, Y is normally a fimction o f the crack and plate geometry and can be found by analytical or experimental methods.

Kj = Yg^ [ ^ (16)

y = /W

0

-

7

)

Where for single edge notch t is the specimen thickness and a is the crack depth.

1.2.2.3 Fatigue Crack Growth

The fatigue crack growth rate can be divided into three different regions. Figure 1.4 shows

the curve o f crack growth rate against the stress intensity factor range. In region I, the

crack growth rate is low. Below a certain SIF range, there is no longer any crack growth.

This minimum SIF range is called the threshold value or AKth. The crack growth mechanism within this stage is most likely to be influenced by microstructural effects.

Mean stress is also highly influential as it can shift the value o f the threshold.

Region II is characterised by the linear relationship in the Log-Log curve. This is the

region where the Paris law may be applied. For air tests on mild steel under constant

amplitude loading, the microstructure and the mean stress level do not have a significant

effect on this region. However, under a corrosive environment, the influence of mean

stress can be very significant. The majority of the fatigue life crack growth falls into this

region, thus most crack modelling is done for this region. Region IH shows a high

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acceleration of crack growth rate, as this is a region o f high stress or low cycle loading.

Mean stress is highly influential, as is the material fracture toughness.

There have been attempts to characterise fatigue crack propagation. Paris and Erdogan

[1.12] were the first to suggest the characterisation o f crack growth under cyclic loading

by a power law relationship for region H crack growth. They related the AK value to da/dN as represented in equation (1.8).

— = c (a a :)'” (1.8)

dN

This equation became more widely known as the Paris and Erdogan law, where da/dN is the crack growth rate, AK is the stress intensity factor range, C and m are material constants. The stress intensity range can be defined as the difference between the

maximum SIF and the minimum SIF; AK = K ^ ^ ^ the original work, Paris based

his findings work on aluminium alloy and gave m values o f between 2 and 4. The amount of data available was limited hence the power law is open to many different

interpretations. This relationship also suggested that the crack growth rate in this region is

only dependent on the stress range and the effect o f load ratio, R, is insignificant. However, this is a break through in relating the fatigue crack propagation to the applied

stress and has found universal use in most fatigue calculations.

The power law can also be used to estimate the number of cycles required for a given

changes o f crack propagation. By integrating the equation (1,8), the number of cycles, Np, which is required to propagate the crack depth of a, to a/, can be calculated, as shown in

(L9).

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1.2.3 Factors Affecting Fatigue Life

The fatigue life o f a specimen can be affected by many factors, although the real driving

force is the stress range. Thus, when comparing the fatigue lives for the different

specimens in a test series, the only difference should be the stress range applied. Fatigue

life can be affected by factors such as the mean stress ratio, residual stress, specimen

dimension, material properties and environmental effects. The major points that will be

discussed in this section are the thickness effect, the mean stress level and the corrosion

effect.

1.2.3.1 Thickness Effects

When dealing with fatigue crack propagation in a specimen, it is important to identify the

size o f the specimen. For offshore structures, thicker plates are usually employed thus

plates o f 16 mm or more are used for fatigue testing. Thick specimens, however, are

found to have a shorter fatigue life. This could be due to a number o f factors, for example

the processing and manufacturing methods of the specimen and also the differences in the

stress distribution through the plate thickness.

The thickness o f the wall can have considerable effects on the fatigue life of a structure.

The crack growth rate is often higher in thicker specimens. In welded structures, where

the crack growth is the dominant part o f the fatigue life, the stress gradient is lower for a

thicker specimen. Thus, the same size flaw in a thick plate would be subjected to a higher

stress level for a larger region than that o f a thin plate as demonstrated in Figure 1.5. It is

important when comparing fatigue performance for different specimens that the

component size is taken into account.

The correction factor recommended by the 1990 HSE fatigue guidance [1.2], which was

based on the study by Gurney [1.15], is given in the form of equation (1.10),relating the corrected stress range to plate thickness.

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s = s.

( i . w )

Where S is the stress range at the plate thickness, t. Sb is the stress range at the reference

plate thickness, ts and y is the thickness correction exponent. The value o f y employed depends on the reference plate thickness, tg, and is quoted as y = 0.3 for 16 mm thick plates in the 1990 HSE fatigues guidance [1.2]. This is particularly useful when

comparing data for different tests as all the results can be plotted on one S/N curve.

There is currently no specific thickness correction factor study for fatigue lives under

variable amplitude loading. A study on the thickness effect could be carried out, by noting

from studies for different variable amplitude fatigue specimens. However, these

specimens would have to be identical and tested under a similar loading spectrum. The

offshore technology report on variable amplitude loading published by the HSE [1.8]

recommended that there was no fundamental reason why the thickness correction factor

employed under constant amplitude would not be applicable to variable amplitude

loading.

1.2.3.2 Mean Stress

Mean stress is the average stress applied during cyclic loading. This can be introduced

into the structure either by a residual stress due to welding or post weld heat treatment or

static load. Therefore, for most welded offshore structures in the as-received condition,

the joints are likely to be under a certain amount of mean stress already.

Under constant amplitude cyclic loading, the maximum stress and the minimum stress are

normally used to calculate the mean stress. The preferred notation used for the mean

stress is in the form o f the load ratio, R. This is defined by the equation (1.11). This relationship represents the ratio o f the minimum load or stress, <Tmin, over the maximum load or stress,

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(1.11)

^max

Under variable amplitude loading, the load ratio cannot be represented by the ratio o f the

maximum to the minimum load. The method usually employed, is the clipping ratio.

Clipping ratio is the ratio between the maximum stress against the root mean stress of the

sequence. The ratio can also be represented by the ratio o f Kmean over Krms and is known as g r [1 16] as represented below.

Q ^ ^ ^ j n e a n _ ( IJ2)

^rms

In air, the effects o f mean stress on offshore steels may not be significant under constant

amplitude loading. However, under variable amplitude loading, the load interaction

effects may affect the crack initiation phase as well as the crack propagation mechanism.

Also, the stress level or the mean stress directly affects the corrosion fatigue mechanism.

Austin [1.17] carried out fatigue tests under variable amplitude loading on tubular joint.

He stated that the effect of mean stress on fatigue lives under variable amplitude loading

in air is less than under constant amplitude loading. The reason could be due to

“shakedown of crack tip stress field by large stress peaks in most severe sea states”. It is

to be expected that the effect o f mean stress could still be significant under variable

amplitude loading. Thus the “shakedown” could be referring to the breaking down of the

mean stress due to residual stress rather than static load.

1.2.3.3 Effects o f Residual Stresses

In engineering structures where cold working or welding of steels are involved, residual

stresses will usually exist. Residual stresses are locked in stresses within the structure.

Welding process can create residual stresses, which usually have a detrimental effect on

the fatigue life o f a structure, due to the tensile residual stress on the surface. However,

when used correctly, residual stresses can also be beneficial. Post weld treatment methods

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like shot peening, deliberately introduce surface compressive residual stresses, which in

turn improve fatigue life. Residual stresses can be described as static stresses caused by a

static load. They do not contribute to the stress range, but can contribute towards the mean

stress or stress ratio resulting in crack closure, thus affecting the fatigue performance.

Due to the welding process, a Heat Affected Zones (HAZ) is introduced into the welded

structure. The residual stresses are also introduced during the welding process due to local

beating and different cooling rates. This can shorten the fatigue life by creating surface

tensile residual stresses. The surface tensile stress can have a detrimental effect as far as

fatigue life is concerned. The tensile residual stress near the surface shortens fatigue life

by increasing the mean stress especially during the crack initiation stage. For crack

growth, once the crack propagates from the surface, the residual stresses are redistributed

and are unlikely to affect the propagation stage of the crack. However, it is difficult to

calculate and exactly measure the residual stresses. The effect o f residual stresses is

usually included into the fatigue calculation by increasing the mean stress. The residual

stresses can be reduced by post weld beat treatment. This involves beating o f the welded

region and cooling the specimen more uniformly, bi thicker specimen, it is advisable to

perform a post weld beat treatment to improve resistance to brittle fracture and also the

joint is more likely to be affected by the residual stresses due to different cooling rate

caused by the welding process.

Beneficial, compressive residual stress at the surface may induce the opposite effect and

prolong the crack initiation stage by reducing the effective tensile stress. The compressive

residual stress can be introduced by a shot peening method or by similar surface treatment

processes.

1.2.3.4 Environmental Effects

The environmental factors that could affect fatigue mechanism in most offshore structures

are the temperature and chemical environment. The most obvious factor for offshore

structures is the marine environment. The main concerns for the marine environment are