HSEHealth & Safety

Executive

Integrity of Repaired Welds (Phase 1)- Deliverable 5 Summary Report

Prepared by Serco Assurance for the Health and Safety Executive 2004

RESEARCH REPORT 191

HSEHealth & Safety

Executive

Integrity of Repaired Welds (Phase 1)- Deliverable 5 Summary Report

J K Sharples, L Gardner S K Bate, R Charles

Serco Assurance Birchwood Park

Warrington Cheshire WA3 6AT

J R Yates The University of Sheffield

Sheffield S1 3JD

M R Goldthorpe M R Goldthorpe Associates

The Grange 2 Park Vale Road

Macclesfield Cheshire

SK11 8AR

This report summarises work that has been undertaken by Serco Assurance (formerly AEA Technology Consulting), The University of Sheffield and M R Goldthorpe Associates, on behalf of the Health and Safety Executive. It describes Phase 1 of a proposed multi-stage project aimed at

(i) providing general guidance on when welded repairs may or may not be beneficial, and,

(ii) proposing a suitable engineering procedural method for assessing the integrity of repaired welds ona case-bycase basis. Welds considered are appropriate to ferritic material.

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

HSE BOOKS

© Crown copyright 2004

First published 2004

ISBN 0 7176 2800 0

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

Applications for reproduction should be made in writing to:Licensing Division, Her Majesty's Stationery Office, St Clements House, 2-16 Colegate, Norwich NR3 1BQ or by e-mail to hmsolicensing@cabinet-office.x.gsi.gov.uk

ii

CONTENTS

EXECUTIVE SUMMARY v

INTRODUCTION 1

TASK 1 – REVIEW OF CURRENT INDUSTRIAL PRACTICES AND PREVIOUS PROBLEMS AND ASSESSMENT OF INFORMATION CONTAINED IN THE LITERATURE 3

TASK 2 – SCOPING CALCULATIONS TO ESTABLISH MATRIX OF CASES TOCONSIDER 5

TASK 3 – WELD/SPECIMEN MANUFACTURE 7

TASK 4 – MATERIAL CHARACTERISATION TESTS 9

Tensile Tests 9 Fracture Tests 9 Fatigue Crack Growth Tests 10 Metallography And Hardness Testing 10 Microstructural Examination 10

TASK 5 – RESIDUAL STRESS MEASUREMENTS 12

TASK 6 – TESTS INVOLVING PHOTOELASTIC COATING AND THERMAL EMISSIONMETHODS 13

TASK 7 – DEVELOPMENT OF FINITE ELEMENT MODELS 16

Weld Modelling Technique 16 Material Properties 17 Results of Welding Simulations 18 Analyses of Defects In The Simulated Welds 19

TASK 8 – APPLICATION OF FINITE ELEMENT MODELS TO MATRIX CASES 21

Edge Defects in the Welded Plate 21 Equatorial Defects in the Welded Sphere 24 Embedded Defects in the Welded Plate 25

TASK 9 – ASSESSMENT BY ENGINEERING PROCEDURE METHODS 28

General Methodology 28

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38

Edge Cracks 29 Embedded Cracks 36

TASK 10 – PROVISIONAL GUIDANCE ON WELD REPAIRS

Practical Issues 38 Guidance Resulting From The Finite Element Calculations 39

TASK 11 – PROVISIONAL GUIDANCE ON ENGINEERING PROCEDURE METHOD 41

Route for assessing the significance of a flaw in a weld (as-welded, PWHT orrepaired weld) 41 Route for assessing whether repairing a weld Is likely lo be beneficial 41 Critical Crack Size Evaluation 41 Crack Growth Evaluation 42

TASK 12 – RECOMMENDATIONS FOR FUTURE PHASES OF PROJECT 44

REFERENCES

FIGURES

APPENDIX 1 – LITERATURE REVIEW

APPENDIX 2 – MICROSTRUCTURAL EXAMINATION OF WELD SAMPLES

UNDERTAKEN BY SHEFFIELD UNIVERSITY METALS ADVISORY CENTRE (SUMAC)

iv

EXECUTIVE SUMMARY

This report summarises work that has been undertaken by Serco Assurance (formerly AEA

Technology Consulting), The University of Sheffield and M R Goldthorpe Associates, on behalf of

the Health and Safety Executive. It describes Phase 1 of a proposed multi-stage project aimed at (i)

providing general guidance on when welded repairs may or may not be beneficial, and, (ii) proposing

a suitable engineering procedural method for assessing the integrity of repaired welds on a case-by-

case basis. Welds considered are appropriate to ferritic material.

The project has centred on detailed finite element modelling of a matrix of relevant un-repaired and

repaired weld configurations. Development and validation of the finite element models have been

undertaken by way of mechanical testing; involving photoelastic coating and thermal emission

methods. A number of material characterisation tests have been performed and residual stress profiles

have been evaluated experimentally. Metallurgical examination has also has also been carried out in

order to examine the changes in microstructure resulting from the welding process.

The project has involved the following tasks:

Task 1 – Review of current industrial practices and previous problems and assessment of information contained in the literature.

Task 2 - Scoping calculations to establish a matrix of cases to consider.

Task 3 – Weld/specimen manufacture.

Task 4 – Material characterisation tests.

Task 5 – Residual stress measurements.

Task 6 – Tests involving photoelastic coating and thermal emission methods.

Task 7 – Development of finite element models.

Task 8 – Application of finite element models to matrix cases.

Task 9 – Assessment by engineering procedure methods.

Task 10 – Provisional guidance on weld repairs.

Task 11 – Provisional guidance on engineering procedure method.

Task 12 – Recommendations for future phases of project.

The report constitutes the final deliverable (Deliverable 5) of this phase 1 project. The main results of Deliverables 1 to 4 are summarised and the outline of the work and results are presented for Tasks 9,

10, 11 and 12.

v

1. INTRODUCTION

Repair welds are commonly carried out in industry on components where flaws or defects have been

found in weldments during in-service inspection. However, in some cases the process may actually

have a deleterious effect on the residual lifetime of a component. This can be due to metallurgical

changes in the component material in the vicinity of the repair and because of very high residual

stresses which can be introduced in the repaired region.

A Serco Assurance (formerly AEA Technology Consulting) led consortium involving (in addition to

Serco Assurance) The University of Sheffield and an independent consultant, M R Goldthorpe

Associates, has undertaken Phase 1 of a proposed multi-stage project aimed at (i) providing general

guidance on when welded repairs may or may not be beneficial, and, (ii) proposing a suitable

engineering procedural method for assessing the integrity of repaired welds on a case-by-case basis.

Welds considered are appropriate to ferritic material.

The project has centred on detailed finite element modelling of a matrix of relevant un-repaired and

repaired weld configurations. Development and validation of the finite element models have been

undertaken by way of mechanical testing; involving photoelastic coating and thermal emission

methods. A number of material characterisation tests have been perfiormed and residual stress

profiles have been evaluated experimentally. Metallurgical examination has also has also been

carried out in order to examine the changes in microstructure resulting from the welding process.

The project has involved the following tasks:

Task 1 – Review of current industrial practices and previous problems and assessment of

information contained in the literature.

Task 2 - Scoping calculations to establish a matrix of cases to consider.

Task 3 – Weld/specimen manufacture.

Task 4 – Material characterisation tests.

Task 5 – Residual stress measurements.

Task 6 – Tests involving photoelastic coating and thermal emission methods.

Task 7 – Development of finite element models.

Task 8 – Application of finite element models to matrix cases.

Task 9 – Assessment by engineering procedure methods.

Task 10 – Provisional guidance on weld repairs.

Task 11 – Provisional guidance on engineering procedure method.

Task 12 – Recommendations for future phases of project.

1

The various components (i.e. Tasks) of the project, together with their dependencies, are contained in

the flow diagram of Figure 1.

Reports, constituting Deliverables 1 to 4, have previously been issued that outline the work and

results of Tasks 1 to 8.

This report constitutes the final deliverable (Deliverable 5) of this phase 1 project. The main results of

Deliverables 1 to 4 are summarised and the outline of the work and results are presented for

(i) asessment by engineering procedures (Task 9), (ii) provisional guidance on weld repairs (Task 10),

(iii) provisional guidance on weld procedures (Task 11) and, (iv) recommendations for future phases

of the project.

2

2. TASK 1 – REVIEW OF CURRENT INDUSTRIAL PRACTICES AND

PREVIOUS PROBLEMS AND ASSESSMENT OF INFORMATION

CONTAINED IN THE LITERATURE

A draft report of the literature review carried out under Task 1 has been previously issued. An

updated version of this report is included as Appendix 1.

The papers reviewed can be categorised as folllows:

Numerical analysis: These relate to the prediction of residual stresses in weldments.

Case Studies: These papers discuss the metallurgical examination of repair welds and the evaluation

of found defects.

Weld Repair Procedures and Techniques: These papers present weld repair techniques.

Performance of Repair Welds: An assessment of how various weld repairs have performed in service.

The review has indicated that defects in welded structures can occur during the fabrication process

due to ‘workmanship’ or in-service due to working conditions. During fabrication, PD5500 states that

‘unacceptable imperfections shall be either repaired or deemed not to comply with this standard’.

Repair welds have to be carried out to an approved procedure and subjected to the same acceptance

criteria as the original weld. Thus all welds have to satisfy the requirements of the design

specification before acceptance by the purchaser or inspecting authority.

For defects found in-service there are no standard guidelines available for utilities to use to make a

decision on the need to carry out a weld repair. An industrial survey carried out by EPRI for utilities

in the United States has shown that utilities will rely on the original manufacturer or outside vendors

to assist on this decision. However, it is not clear that the assessment procedures used are consistent

or are indeed reliable. In the UK, the repair of welds appears to rely on in-house experience in the

absence of guidelines to follow. However, this review showed that re-cracking of repair welds still

occurs due to lack of understanding on why original defects have occurred and how they should be

repaired.

Whilst the decision to repair a defect may be aided using an assessment procedure, the practical

considerations identified in a paper by Jones could also usefully be considered. These show that

repair welds should be considered on a case-by-case behaviour, therefore a definitive set of ‘rules’ can

not be given. Instead, the guidelines need to be produced which provide good practice in assessing

defects in welds and the requirements for carrying out a ‘safe’ repair.

A number of References were found illustrating the capabilities of performing a repair weld without

the need for PWHT. This was introduced by the half-bead technique defined in ASME XI primarily

for the nuclear industry. This has been superseded by other temperbead techniques, which are all

aimed at improving the properties within the weld HAZ, whilst saving time and costs by precluding

the time for post-weld heat treatment (PWHT). There is evidence that this method is employed by

other industries in the USA, but it is unclear on the use of this practice in the UK.

In the references associated with case studies and the performance of weld repairs, only a few of them

are related to residual stresses. These papers have indicated that the magnitude of residual stresses in

3

repair welds can be of yield magnitude. The most recent advances in welding simulation were

presented at an e IMechE conference in November 1999. The conference demonstrated the

developments that had been made, mainly in the use of finite element analysis to predict residual

stresses. Sufficient confidence in numerical analysis needs to be demonstrated by making comparison

with measurement methods.

When developing guidelines for the assessment of defects in repair welds, sufficient advice needs to

be given to the user as to when residual stresses need to be considered in the assessment. Advice also

needs to be provided on when the user should use simple approximations of the residual stress pattern,

e.g. upper bound profiles given in BS7910, or to use finite element analysis techniques to predict the

complex behaviour of the material during welding.

4

3. TASK 2 – SCOPING CALCULATIONS TO ESTABLISH MATRIX OF

CASES TO CONSIDER

A detailed section on this Task is included in the Deliverable D2 report previously issued (Reference

1).

That section outlines the matrix of analysis cases planned to be undertaken in Task 8. These cases

were developed to illustrate the effect, on fatigue life or load margin, of either leaving a weld defect

in place or repairing it in-situ. Only in-service repairs would be considered. Since such comparisons

were only possible by considering the behaviour of defects, it was assumed that a defect inadvertently

remains in the weld after “repair”. This may or may not grow by fatigue during subsequent service.

The double V-preparation weld in plate (Figure 2) used in the experimental work of the project would

be studied in Task 8. It is a relatively simple weld geometry, but this would enable a large range of

analysis cases to be considered and so illustrate to non-experts, the effects that different parameters

could have on the decision to repair a defective weld.

The repair procedure carried out on the test plate in Task 3 is shown in Figure 3. This procedure was

considered to be representative of an in-situ weld repair. The repair depth is 15 mm in order to

simulate the grinding out of the weld 2 mm beyond an assumed defect with a depth of 13 mm. There

was lower heat input than a shop repair, using no pre-heat and smaller electrodes. Strong-back plates

were used to simulate the restraint on the surrounding structure and no PWHT was carried out.

For Task 8, it was intended to carry out a variety of mainly two-dimensional, plane stress, finite

element analyses. Comparisons would be made between simulations of un-repaired and repaired

situations for a range of different parameters that affect fatigue life or margin on load. The intention

was to illustrate the transition from cases where the defects are best left in place, to cases where repair

is required. Since comparisons would be made between the un-repaired and repaired situations,

simplified, two-dimensional, plane stress analysis would be capable of illustrating the role of different

parameters in the repair decision.

The base case would be a 40 mm thick plate, with an alternative thickness of 20 mm.

For simplicity, defects would be considered to lie in a plane normal to the surface of the plate and

through the middle of the weld. The repair evacuation would be symmetrical with respect to the

middle of the weld.

The base case for the un-repaired condition would be a surface breaking defect in the weld root as

shown in Figure 4. The defect depth would be equal to one third of the plate thickness. Alternative

cases would consider surface breaking weld root defects with different depths, covering the range

from the minimum detectable by NDT (about 3 mm) to one half of the plate thickness.

Embedded defects in the un-repaired condition (Figure 5) would also be considered. The base case

would be an embedded defect having a total height equal to one third of the plate thickness and

symmetrically positioned about the weld throat. Alternative cases would consider different defect

heights and position relative to the weld.

Figures 6 and 7 show the case of ‘wide’ and ‘narrow’ excavations that would be studied. These were

considered to bound the repair procedure specified in Figure 3.

5

Various defects remaining after the repair would be studied. Generally, these defects would be

smaller in height than those in the un-repaired condition. The base case for repaired weld defects

would be an embedded one caused by incomplete excavation, as indicated in Figures 6 or 7. Various

defect heights would be considered, ranging from a minimum of 3 mm to a maximum smaller than the

un-repaired size.

Alternative analysis cases for the repaired condition would consider different surface defects

remaining after improper repair of pre-existing surface defects (Figure 8) and embedded defects

resulting from improper repair of embedded defects (Figure 9). Although in practice the former are

likely to be weld toe cracks, the analyses would consider cracks situated in the middle of the weld.

Figures 10 to 12 show the different defect configurations it was intended to analyse for the 40 mm

thick plate, and Figures 13 to 15 show the defects for the 20 mm thick plate. Table 1 gives a summary

of the un-repaired and repaired defect sizes with a code for each case. The finite element

computations would actually consider a large range of defect sizes in order that calculations of fatigue

crack growth could be undertaken.

In addition to the geometrical parameters referred to above, the planned matrix of cases contained

variations in tensile properties, fracture toughness, residual stresses and service stresses (service

stresses would be simulated in the plate geometry by applying a tensile stress transverse to the weld).

The variations in these parameters are included in Table 1.

As will be seen in Section 9, the finite element analysis covered a good selection of the cases

described above that were proposed under Task 2.

6

4. TASK 3 – WELD/SPECIMEN MANUFACTURE

A detailed section on this Task is included in the Deliverable D2 report of Reference 1.

Motherwell Bridge Group was contracted to prepare a suitable welded steel plate using materials and

welding/repair processes typical of current industrial practice. They used available steel plate of

thickness 40mm to BS1501 490 LT50. The weld procedure qualification record is shown in Figure 2.

An asymmetric double “V” preparation was used with the weld root positioned 2/3 of the plate

thickness from the surface of side 1, which was filled first. Typical pre-heat and interpass

temperatures were used of 75°C and 250°C respectively. No PWHT was carried out. Visual

inspection, Magnetic Particle Inspection (MPI) and ultrasonic testing confirmed that there were no

detectable defects after welding.

The test plate is shown in Figure 16 and comprised two 40 mm thick plates, with length 4000 mm and

width 500 mm, welded together at the long edges. Half of the welded plate, (i.e. a 2000 mm length)

was cut into five sections as shown to provide as-welded material for the experimental work under

project Tasks 4 (material characterisation), 5 (residual stress measurements) and 6 (photoelastic and

thermal emission experiments), along with two blanks for manufacture of further test specimens in a

later phase of the project. Motherwell Bridge Group retained the remaining half of the test plate for

repair weld processing described below.

Strong back plates, made from the same material as the test plate, were used to restrain out of plane

bending during welding. The strong-back plates formed 40 mm thick ribs, 400 mm high, running

across the full 1000 mm width of the test plate on the opposite side to that being welded. Each

strong-back plate was attached to the test plate by fillet welds, which extended for 300 mm from each

end. A central 150 mm cut out was formed to accommodate pre-heaters in the case of the original

weld only. Eight strong-back plates were used for the original welding of the 4000 mm long test

plate, placed at 500mm intervals, commencing 250 mm from the end. The strong-back plates were

fixed to test plate side 2 whilst welding side 1 and vice versa.

The weld repair process carried out on the second 2000 mm length of test plate (Figure 16) was

designed to simulate the site repair of a central root defect in the original weld. This involved typical

grinding out from the narrower side of the weld (side 2) to a depth of 17 mm to ensure removal of a

defect in the original weld root at a depth of 13.5 mm. The weld procedure qualification record for

the repair weld is shown in Figure 3.

To simulate a repair process being applied to a structure on site, rather than under ideal workshop

conditions, some modifications were agreed to the weld procedure. Welding under more difficult

access conditions was simulated by use of smaller electrodes and more rapid passes with less “weave”

than was the case for the original weld. This process (known as “stringer bead” technique) resulted in

a lower heat input than for the original weld. This was exacerbated by the omission of pre-heat for

the repair, simulating a site situation where pre-heat could be difficult to apply effectively. Lower

heat input results in more rapid cooling of the weld metal, which can lead to changes in the material

properties. No PWHT was carried out following the repair welding. Visual inspection, MPI and

ultrasonic testing confirmed that there were no detectable defects after repair.

For the repair weld, four strong-back plates of the type used for the original weld were attached to

simulate structural restraint. These were set at 500 mm spacing on the 2000 mm long test plate, fixed

to side 1 only as the repair was single sided.

7

8

5. TASK 4 – MATERIAL CHARACTERISATION TESTS

A detailed section on this Task is included in the Deliverable D2 report of Reference 1. Initial

material characterisation tests, covered in Reference 1, were those to determine tensile, fracture and

fatigue crack growth properties. The results of metallography and hardness testing are also presented

in Reference 1. Narrow bands of high hardness were measured in the heat affected zone (HAZ) of the

samples (see below). To provide an understanding of the formation of these, it was decided to carry

out a more detailed microstructural examination of the welded regions in samples for both the as-

welded and weld repair specimens.

Results of all the material characterisation tests are summarised as follows:

5.1 TENSILE TESTS

Tensile properties of the weld material in both as-welded and repair-welded conditions at room

temperature were obtained from tests on 3.5 mm diameter round bar specimens.

The tensile test results are listed in Table 2. True stress/true strain data are given in Reference 1.

The results show that, in the as-welded state, the weld was overmatched by 46%, based on the 0.2%

proof stress (PS) values of approximately 512 MPa and 350 MPa for weld and parent plate

respectively. The parent material exhibited typical upper and lower yield point behaviour, which was

not present in the weld metal results. The ultimate tensile stress (UTS) for the weld was 18% higher

than that for the parent material, with average values of 622 MPa and 527 MPa respectively.

For the repair weld material, higher values of 0.2%PS were obtained compared to the as-welded

condition. The near surface average value for repair weld was 540 MPa compared to 512MPa for the

as-welded condition (5% increase), whilst the near root average value for repair weld was 580MPa

(13% increase). The UTS value obtained from near surface repair weld was similar to that for the as-

welded condition (628 MPa against 622 MPa respectively), whilst the value for near root repair weld

was 670 MPa (approximately 13% increase on as-welded). It should be noted that a spurious result

was obtained from repair weld specimen WI12, due to failure outside the gauge length, and this has

therefore been discounted.

5.2 FRACTURE TESTS

Fracture toughness J resistance curves at room temperature were obtained from single edge notch

bend (SENB), side grooved, unloading compliance specimens to BS 7448 Part 4 for the original weld

and the repair weld. Two specimens were tested in each condition. The specimen notch was aligned

centrally in the through-thickness direction. The specimen orientation was selected, and the initial

crack length, after fatigue pre-cracking, adjusted within the standard limits, to ensure that the crack tip

lay in original weld or repair weld, as desired.

The results are shown in the crack growth resistance curves of Figures 17 and 18, for as-welded and

repair-welded material respectively. The results showed that the fracture toughness behaviour was

similar in both the as-welded and repair-welded specimens, with initiation toughness J0.2 values of -2

approximately 105 kJm-2

and 102 kJm respectively (allowing for blunting, J0.2BL of approximately -2

116kJm-2

and 119kJm respectively).

9

5.3 FATIGUE CRACK GROWTH TESTS

Fatigue crack growth properties at room temperature were obtained for the original weld and the

repair weld using Compact Tension (CT) specimens in accordance with ASTM E647. The specimen

notch was aligned centrally in the weld in the through-thickness direction. The specimen orientation

was selected, and the initial crack length adjusted within the standard limits, to ensure that crack

growth was obtained in original weld or repair weld, as desired.

The results of the fatigue crack growth tests on weld metal are shown in the Paris Law plots of Figure

19. The data indicate that similar fatigue crack growth behaviour was obtained with both the as-

welded and repair-welded material. The slopes of the Paris Law plots are very similar, with some

offset giving slightly higher growth rates with the as-welded material. The valid region of stress

intensity factor range, DK, was from approximately 25 MPaÖm to 60 MPaÖm.

5.4 METALLOGRAPHY AND HARDNESS TESTING

Sections from the weld in the as-welded and repaired states were polished and etched to reveal the

welds, macro photographs taken and hardness testing carried out. In addition to examination of

transverse sections, the edges of the samples (i.e. the surface of the test plate) were also prepared by

polishing down to the level of the plate surface. Surface hardness measurements were taken to

compare with the sub-surface values obtained from the transverse sections. The Vickers Hardness

surveys (Hv 10kg load) of the parent materials, welds and HAZs were carried out according to BS EN

288-3.

The original welds had typical, well-defined runs, with HAZs in the order of 2-3mm wide. The area

of weld repair had a less well-defined weld run structure due to the larger number of smaller beads.

The Vickers Hardness survey according to BS EN 288-3 showed no significant hard spots in any of

the samples for the transverse sections. The hardness values in the unaffected parent material were in

the region of approximately Hv140 to Hv180. The highest hardness values were recorded in the

HAZ, as expected. The HAZ on the repair weld was slightly harder than the original weld, with

maximum recorded values of Hv331, and Hv268 respectively. These levels are below the maximum

permitted hardness value of Hv350 stated in BS EN 288-3 for this class of material.

The results for the surface measurements show a similar but less pronounced variation in hardness to

that recorded for the transverse sections. The maximum HAZ hardness values recorded were Hv258

and Hv284 for the as-welded and repair-welded conditions respectively. This gives some confidence

that increased hardness could be indicated by measurements on the accessible surface of a structure,

but suggests that small, isolated areas of peak hardness may not be detected since they may occur sub-

surface.

5.5 MICROSTRUCTURAL EXAMINATION

The more detailed microstructural examination was carried out by the Sheffield University Metals

Advisory Centre (SUMAC). The details of this are given in Appendix 2. The SUMAC work

consisted of examinations on both as-welded and repair-welded samples in terms of microstructural

observations, standard hardness tests, microhardness surveys and microanalysis using dispersive x-

rays.

It was shown that the HAZ microstructure followed the typical pattern of a multi-pass weld with a

zone of grain growth at the fusion line, backed by a band of recrystallization followed by a

spheroidized/tempered zone before the unaffected matrix. Each weld pass imposed a further HAZ on

the underlying weld (and it’s HAZ) leading to a refined microstructure at the overlap. The grain

10

growth and recrystallization zones had a microstructure of grain boundary and Widmanstatten ferrite

(the amount depending on the local austentising temperature and subsequent cooling rate) in a

transformed matrix. In carbon and low alloy steels of this type, the matrix can be a mixture of the

phases ferrite, pearlite bainite and martensite. The root run area was completely refined and tempered

and contained no “hard spots”. The macro and micro-hardness testing indicated that the HAZ of the

‘toe’ welds in weld 2 (the smallest weld on the side containing the repair weld) of both the as-welded

and repair-welded samples had higher hardness values than elsewhere. The microstructure, whilst not

exhibiting defined ‘pools’ of hard phase, did show structural refinement and reductions in pro-

eutectoid ferrite that could explain the increased hardness.

The study concluded that both the as-welded and repair-welded samples passed the hardness

requirement and some potentially high hardness values obtained by microhardness should not detract

from this, particularly as they were in areas where this might be expected and were not found

elsewhere in the weld.

11

6. TASK 5 – RESIDUAL STRESS MEASUREMENTS

The destructive technique of block removal, splitting and layering was used to determine the through

thickness residual stress distribution in the as-welded and repair-welded specimens. Further details of

the procedure and the measured results are contained in Reference 1.

The residual stress results for the as-welded condition are shown in Figures 20 and 21 for the

Y direction (perpendicular to the weld) and X direction (parallel to the weld) respectively. The

stresses in the Y direction are self-balancing through the thickness, with tensile values near the

surfaces and compressive values in the central area. The stresses in the X direction are tensile

throughout the thickness. The distributions are asymmetric as expected, considering the asymmetric

weld preparation, with minimum values occurring at a depth of approximately 25 mm from weld side

1, which corresponds to the location of the weld root. Stress maximum values occur at depths of

approximately 5 mm and 35 mm. The peak tensile stress in the Y-direction (perpendicular to the

weld) is ~220-350 MPa, and in the X-direction (parallel to the weld) ~500-580 MPa.

The residual stress results for the repair-welded condition are shown in Figures 22 and 23 for the

Y direction (perpendicular to the weld) and X direction (parallel to the weld) respectively. The form

of the stress distributions is basically the same as for the as-welded condition (Figures 20 and 21).

The stress minimum values are of similar magnitude to the as-welded but occur closer to the centre of

the plate, corresponding to the location of the repair weld root. Also, the stress maximum values at

depth of 5 mm show a noticeable increase over the as-welded for both the Y and X directions, whilst

the maximum values at depth of 35mm remain at similar levels. The increase in peak tensile residual

stress therefore occurs on the side remote from the weld repair, rather than on the repaired side. The

peak tensile values at depth of 35 mm are 600 MPa and 750 MPa for Y and X directions respectively,

the latter being in excess of the weld metal yield stress measured in the tensile tests. The reason for

this high peak is not clear but the two sets of strain measurements taken in the X direction gave very

similar results, which suggests that it is not due to an experimental error or test equipment fault.

As a further check on the residual stress levels at the surfaces, measurements were made using the

shallow hole drilling technique. This technique involves using a trepanning air-abrasive jet drilling

technique, which has been shown to introduce practically no residual stresses into the component

under test. The technique involves the drilling of a small blind hole (typically 1.8 mm diameter x 1.8

mm deep) in the centre of a special three-element strain gauge rosette. Local strain relaxation is

related to the initial stress state in the specimen and calibration using a known (usually uniform) stress

field allows residual stresses to be calculated.

The surface stresses evaluated from the shallow hole drilling technique are as follows:

(1/3 weld side) (2/3 weld side)

As-Welded: Perpendicular Stress (MPa) -94 365 368

As-Welded: Parallel Stress (MPa) 225 138 181

Repair-Welded: Perpendicular Stress (MPa) 34 280 386

Repair-Welded: Parallel Stress (MPa) 181 -27 -162

These values have been included in the residual stress distribution plots of Figures 20 to 23. It can be

seen that the surface stresses obtained from the hole drilling method are generally consistent with the

near-surface stress distributions evaluated from the block removal, splitting and layering technique.

12

7. TASK 6 – TESTS INVOLVING PHOTOELASTIC COATING AND

THERMAL EMISSION METHODS

Detailed information on this Task is contained in Reference 2. The Task focused on (i) quantifying

the fatigue crack propagation rate in welded and repair welded steel plate, (ii) investigating the use of

a full field photoelasticity technique to measure residual stresses in the plates, and (iii) investigating

the use of a thermoelasticity technique to measure the true crack tip driving force (i.e. stress intensity

factor) in the two types of weld.

The specimens used for testing were obtained from the initial test plate as described in Section 4. The

specimens tested were identical for both original and repair welds. The geometry used for the study of

fatigue crack growth was a tension specimen, 41.5 mm wide (W) and 12 mm thick (t) with a 4 mm

initial edge notch (a) spark machined in the side of the original or repair weld, as appropriate

(Figure 24).

Stress intensity factors were calculated using the following equation:

K ID YDs= ap (1)

where 2 3 4

Y 231 0 12 1 ç æ

-= .. a

55 10 ç æ

÷ + ö .

a ÷ ö

72 21 ç æ

- . a ÷ ö

39 30 ç æ

+ . a ÷ ö

(2) è W ø è W ø è W ø è W ø

Such values are referred to as ‘DKI Theory’ so as to distinguish them from values determined by

thermoelastic measurement.

The tests carried out consisted of analysing the crack growth for a tensile edge cracked specimen

using thermoelastic stress analysis. The machine used for this purpose was an ESH 100kN servo-

hydraulic machine which allows the application of a cyclic load to the specimen at the frequency and

load convenient for the thermoelastic test.

Seven fatigue tests were carried out for different load conditions, as detailed below:

Identifier Load range,

kN

R ratio Comments

AEA1 32.4 0.13 Original weld

AEA2 40 0.1 Original weld, 30kN range at R=0.1 applied for 800000

cycles with no growth

AEA3 33.0 0.13 Repair weld

AEA4 39.6 0.1 Repair weld. Subsequently used for J test

AEA_F2 36.9 0.28 Repair weld. Test run to fracture of specimen.

AEA_F3 37.6 0.58 Original weld

AEA_F1 37.6 0.58 Repair weld

Images at different number of cycles during the tests were taken. At the same time, for every picture

captured, the number of cycles and the crack length were noted. A vernier microscope was used to

measure the rate at which the crack length had grown between different images.

13

A non-standard J test was carried out by loading in four-point bending one of the edge cracked tensile

specimen used for crack growth analysis. The results obtained were found to be similar to those

previously obtained that are presented in Figure 17. In addition, one of the fatigue tests was run until

failure. The loads at fracture were Pmin = 16.4 kN, Pmax = 51.6 kN, the final crack length being

34.93 mm, including the initial 4 mm slit. Failure occurred after 536770 cycles.

The stress intensity factor ranges were plotted against the crack length for different R values and for

different specimens (original and repair welds). An example is shown in Figure 25 for the R = 0.13

case of the repaired weld specimen, AEA3. “RAT” and “FGD” referred to in Figure 25 are the initials

of the two different operators who processed the results. In all cases, experimental results were

compared to the range of stress intensity factor calculated by Equations 1 and 2. It was observed that

all experimental data lay below the theoretical values when the crack length is long enough. This is

thought to be due primarily to the crack closure effect (see below), but other factors may also be

influential. In particular, the large displacement of the crack at high stress intensity factors may well

mean that the published stress intensity factor calibrations are erroneous at these levels.

At the same time using information from the tests, the crack growth rate against the stress intensity

factor was plotted for the different experiments. Figure 26 is an example of such a plot whereby the

Paris law is presented using experimental values for the stress intensity factor (identified as

‘Deltatherm data’ in the Figure) and values predicted from Equations 1 and 2 and the crack growth

rates obtained from experimental measurement. The “AEAT growth equation” curve included in

Figure 26 has been derived from the data presented in Figure 19.

Finally, an estimate of the closure level was made from the difference between the theoretical DKI and

the value measured using Deltatherm. The values are shown in Figure 27 plotted against the crack

length.

In considering crack closure effects, it has previously been observed that non-linear crack opening

behaviour results in a region of residual tensile deformation in the “wake” of a fatigue crack. The

resulting permanent contact between the two crack faces results in a lowering of the crack opening

displacement and consequently lower driving force for fatigue crack advancement.

A large amount of research has been carried out on this topic during the last few years and the

mechanisms involved have been described. These mechanisms suggest that several types of closure

affect the rate of fatigue crack advance. The possible sources of crack closure are the following:

- Plasticity induced crack closure, due to residual stress in the wake of the crack.

- Oxide induced crack closure, due to the oxide layers formed inside the fatigue crack.

- Roughness induced crack closure, due to the roughness of the fatigue fracture surface.

- Viscous induced crack closure, due to the penetration of viscous fluids inside of the crack.

- Transformation induced crack closure, due to phase deformations at the crack tip caused by stress

or strain.

In addition, the presence of non-uniform residual stresses in a structure will contribute to the crack tip

driving force in addition to primary loads. These complex stresses may increase the stress intensity

factor above that estimated from the external loading, or may decrease it, thereby having a similar

effect on crack closure.

Looking ahead to Figures 37 and 38, which present the finite element determined values of stress

intensity factor for the residual stress fields (refer to section 9.1), it is evident that the KI values are

positive for all crack sizes considered (crack depth, a, ranging from just over 2 mm to 20 mm). Crack

14

opening, as opposed to crack closure, would therefore be expected to occur from the residual stress

distribution.

Some tests were undertaken using reflection photoelasticity with the intention of measuring residual

stress in welds. Two different specimens were used from the original and the repair welds. The

photoelastic results confirmed the previous measurements referred to in Section 6, whereby very little

difference was observed between the residual stresses in the as-welded and repaired weld conditions.

15

8. TASK 7 – DEVELOPMENT OF FINITE ELEMENT MODELS

Detailed finite element modelling of a matrix of relevant un-repaired and repaired weld configurations

has formed a major part of the project. The work was mainly focussed on the modelling of a plate

geometry but a spherical vessel geometry was also considered. This work (covering Tasks 7 and 8) is

fully described in Reference 3 and summarised in the following sub-sections.

8.1 WELD MODELLING TECHNIQUE

In terms of the development of the finite element models, a weld bead lumping approach was used to

model weldments in which a small number of lumped beads was modelled in both original and repair

welds. A non-linear analysis of the welding process was carried out using a simplified ABAQUS

finite element model of the parent plate and weld. In this analysis, the original weld was built up by

the addition of each lumped weld bead in an incremental manner.

A thermal transient analysis was first conducted in order to establish the temperature history of each

point in the plate or sphere due to the addition of each weld bead. A subsequent elastic-plastic

analysis used an almost identical finite element model to simulate the addition of the weld beads.

This mechanical model was loaded by imposing, at each time increment, the temperature of each node

from the above thermal transient analysis. Like the thermal analysis, the mechanical model was

necessarily simplified, so the complex behaviour of the weld and parent metal near melting point was

not considered. However, approximate temperature dependent mechanical properties were used.

Low values of yield stress and perfectly plastic properties were used at temperatures near the melting

point to reduce the loading on adjacent material. However, this did incur the penalty of producing

unrealistically large plastic strains that cannot be annealed.

After adding the final lumped bead of the original weld, the current state of the mechanical model

(displacements, stresses, elastic and plastic strains etc.) was saved for subsequent restarts. Following

this, the elements in the repaired areas were removed, and the lumped beads of the repair were added.

The required state of the model was again saved for subsequent restarts.

Figure 28 shows a part of the finite element mesh used to model a though-thickness section of the

welded test plate in the region of the weld. For convenience, the mesh is shown rotated by 90o with

respect to Figures 2 and 3. The plate thickness was 40 mm measured in the horizontal direction in

Figure 28. The depth of the repair weld was 15 mm, this being slightly smaller than the 17 mm

actually excavated in the real plate weld. The original weld comprised nine lumped beads, and the

repair weld had four. The weld caps were not modelled. The plate width was measured in the vertical

direction in Figure 28. Due to symmetry about the centre of the weld, only one half of the 1000 mm

plate width was modelled.

To make allowance for later generalisation, the finite element mesh was actually three-dimensional,

but only a single element thickness was used in the plate height direction perpendicular to the plane of

Figure 28. The strong back plates used during the actual welding were modelled as beam elements,

with equivalent section modulus, running vertically along the appropriate side of the mesh in Figure

28.

A sphere was modelled with 40 mm thickness and 20 m diameter. The weld was considered to be a

fully equatorial one, with dimensions and bead lumping exactly as modelled in the plate weld. The

repair lay on the outside of the sphere. Figure 29 shows the axisymmetric finite element mesh used.

Again due to symmetry about the centre of the weld, only one half of the sphere was modelled. No

16

strong back restraints were used in the weld mechanical simulation. A further weld repair situation

was also considered for the sphere. This involved a shallow excavation, with removal of only that

material corresponding to repair weld beads 3 and 4 (see lower part of Figure 28) from the original

weld and its replacement by simulated welding. The repair depth in this case was 6.6 mm instead of

the standard 15 mm.

After the modelling of the original weld in both geometries, a creep analysis was carried out to

simulate a PWHT of four hours at 650oC. Here, the temperature of the model was raised from 20

oC to

650oC in five hours at a uniform rate. The temperature was held at 650

oC for four hours and then

reduced back to 20oC in six hours at a uniform rate. After the PWHT simulation, and also after the

subsequent repair, the state of mechanical model was saved for subsequent restarts. These are, in fact,

the welded plate and sphere states into which defects were inserted as discussed later.

8.2 MATERIAL PROPERTIES

The tensile properties of the parent plate and of the original and repair welds at ambient temperature

are reported in Reference 1. Figure 30 shows the measured results for the parent steel, and in the

centre of the original weld in terms of true stress versus true strain.

The measured results were modified in order to be suitable for the weld simulation analysis. Firstly,

the results were smoothed. Secondly, the Luders strain seen in the parent steel was removed since it

was considered to be a consequence of the tensile test specimen geometry, rather than a real material

phenomenon. The rate of strain hardening seen at the end of the tests, at about 3.5% strain, was

extrapolated to larger strains. Thirdly, trial simulations of the original weld showed that an average

equivalent plastic strain of about 12% remains in the weld. This strain could not be annealed within

the analysis. To take account of this prior welding strain, the weld values in Figure 30 had 12% added

to the plastic strain so that after the weld simulation, the stress versus additional strain followed the

smoothed, measured curve. In Figure 31, the solid blue curve shows the tensile properties at 20oC

used for parent plate in the weld simulation analyses. The solid red curve shows the properties of both

the original and repair weld at this temperature.

The resulting 0.2% and 1% proof stresses of the weld and parent materials are as follows:

0.2% Proof Stress (MPa) 1% Proof Stress (MPa)

Temperature (oC) Parent Weld Parent Weld

20 371 362 418 362

2000 5 5 5 5

Values at intermediate temperatures were obtained by interpolation. Despite the apparent weld under-

match (lower yield strength than parent) the plastic strain accumulated during the weld simulation

analysis raised the yield stress; resulting in behaviour like that measured in Figure 30.

The following Paris Law curve constants were used in the fatigue crack growth that were carried out:

da = 10x 2 .1 -8

DK 77.2 (3) dN

where DK is the stress intensity factor range during cyclic loading, measured in MPaÖm, and da/dN is

in mm/cycle. It may be noted that Equation 3 represents the lower curve presented in Figure 19.

As can be seen from Figures 17 and 18, fracture toughness at ambient temperature corresponding to

0.2 mm of tearing, J0.2, was measured to be equal to 105 and 102 KJ/m2, respectively for original and

repair weld material. These values translate to a toughness of about 150 MPaÖm in terms of KJ0.2. In

17

this study, values of toughness were considered that ranged from 160 MPaÖm down to significantly

lower levels of about 30 MPaÖm in the as-repaired condition. As discussed later, such low values of

fracture toughness can result in repair welds due to a variety of circumstances.

The properties used for the weld simulation thermal analysis and the creep properties used in the

simulation of the intermediate post-weld heat treatment are described in Reference 3.

8.3 RESULTS OF WELDING SIMULATIONS

Figures 32(a-b) compare the through-thickness stress distributions at the middle of the weld with the

measured results presented in Figures 20 to 23. It should be noted that, in these and subsequent

similar Figures, the through-thickness distance is always measured from the non-repaired side 1. The

experimental results are shown as solid lines and the predicted results are dashed lines. The stresses

produced by the original weld are shown in blue, those caused by the repair weld are in red. Predicted

results are in general agreement with the measurements, with tension near the plate surface and

compression at mid-thickness. However, the numerical simulation was unable to predict the precise

magnitudes and positions of stress peaks and troughs. This is not surprising, given the simplifications

and approximations involved. It should also be noted that the predictions and measurements agree that

a higher transverse stress occurs in the repaired weld, but on the un-repaired side 1. Both

measurements and predictions show a similar magnitude of peak transverse stress on the repaired side

2.

In Figures 33(a-b), comparisons are made for through-thickness distributions of transverse and

longitudinal stress across the middle of the weld between the four different numerical simulations.

These cases are:

(i) as originally welded (blue diamonds),

(ii) as originally welded followed by post-weld heat treatment (green diamonds),

(iii) as originally welded, followed by partial weld removal and repair welding (red circles),

(iv) as originally welded, followed by post-weld heat treatment, partial weld removal and finally

repair welding (orange circles),

In case (ii), the effect of heat treating the original weld is apparent, with a large reduction of both

components of stress compared with the as-welded case (i). In Figure 33(a) it is seen that the through-

thickness transverse stresses in the weld for the two repair cases (iii) and (iv) are similar. The repair

of the PWHT weld thus re-establishes a pattern of stress as if the original PHWT had not been carried

out. Furthermore, close to the surface of the un-repaired side 1, the repair causes an increase in

transverse stress to a higher peak level than the un-heat treated original weld (compare the orange

with blue curves). Figure 33(b) shows that the longitudinal stress is affected by repair mainly on the

repaired side 2 itself.

Figures 34(a-b) compare the predicted residual stress results for the four simulation cases carried out

on the sphere. The general pattern of results is similar to that of the plate in Figures 33(a-b).

Figure 35(a) compares transverse stresses for cases (ii) and (iv) between the plate (open symbols) and

sphere (filled symbols). For case (ii), the original PWHT weld shown in green, the peak transverse

tensile stresses predicted in the sphere are about half those in the plate on the last welded side 2. This

situation is reversed on the first welded side 1. The sphere therefore appears to have a component of

through-wall bending stress. For case (iv), repaired stresses shown in orange, the sphere has higher

18

values than the plate at the un-repaired side. Figure 35(b), showing longitudinal stresses, also

illustrates slightly lower predictions in the sphere than the plate in respect of the PWHT original weld,

case (ii).

Comparisons of residual stresses for ‘deep’ and ‘shallow’ repairs in the sphere are shown in Figures

36(a-b). These graphs show results for the original PWHT weld, the standard simulated repair of

depth 15 mm and also for the shallower repair with a depth of 6.6 mm. On the repaired side of the

weld, the shallow repair promotes peak values of transverse and longitudinal stress similar to the

deeper repair. On the un-repaired side of the weld, the shallow repair gives peak stresses lying

between the un-repaired PWHT cases and the deep repair case. Thus, shallow weld repairs can

promote high local residual stresses if the component is not heat treated.

8.4 ANALYSES OF DEFECTS IN THE SIMULATED WELDS

Crack-like defects were inserted into the plate and sphere weld cases (ii) and (iv) of the previous

section. Additional loads were applied to the models to give stresses on the defective section

typically experienced by engineering structures, and crack driving forces (CDFs) were calculated.

These parameters were then used to determine limiting or critical defect sizes for various values of

weld fracture toughness in the two welded states. Comparisons were made between limiting defect

sizes for defects in these heat treated and as-repaired situations.

Using the CDFs, fatigue crack growth calculations were also carried out to determine the number of

loading cycles required to reach the limiting condition for a range of initial defect sizes. Comparisons

were made between fatigue lives of defects in the heat treated and as-repaired states for a range of

initial defect sizes and fracture toughness.

Some modelling simplifications were made in these analyses of defects in welds and these are

explained in Reference 3.

Defects were inserted into the plate model on the plane through the middle of the weld. The two

configurations considered in the welded plate are actually those shown in Figures 4 and 5. In Figure 4,

a surface defect of depth a is shown in the weld. In some cases, the tip reaches into the original weld

(for the weld repair cases). In Figure 5, an embedded or internal defect is considered in the weld. In

some of the weld repair cases, this also reached into the original weld. As for the surface defect, this

defect was also considered as fully extended along the whole length of the weld. The defect is

characterised by its depth, 2a, and the distance of its nearest tip from the repaired surface, p.

In the welded sphere, surface defects were considered in the middle of the repair weld, like Figure 4.

Since the repair was considered to lie on the outside of the sphere (Figure 29) and the finite element

model was axisymmetric, this corresponds to a fully extended outer surface defect of depth a along an

equatorial weld.

Modelling of the defects was accomplished by removing the symmetry boundary conditions along the

line of the defect. These restraints were replaced by equivalent forces that were reduced to zero in

several subsequent elastic-plastic increments of the analysis. The created defect usually opened

under the influence of the residual stress field. In some circumstances however, the defect closed over

at least part of its depth due to a predominantly compressive residual stress. In such cases, the contact

of the opposing faces of the defect was not modelled, so the defect was allowed to ‘over-close’.

Simultaneous introduction of the entire crack surface is mechanistically different to the modelling of

slow, sub-critical crack growth where the crack is introduced progressively. In the former, a zone of

plastic deformation appears at the crack tip(s) only. In the latter, a wake of plastic deformation

develops on the crack flanks as (each) crack tip moves forward.

19

In terms of the primary loading, a remote, uniformly distributed, tensile load was applied to the top of

the modelled plate, 500 mm away from the defect plane. This represents loading in the weld

transverse direction normal to the plane of the defect; causing it to open further, or to open if closed in

the residual stress field acting alone. Various magnitudes of remote membrane load were applied,

with a maximum of 225 MPa. This load was considered to be the occasional ‘overload’ condition for

which the possibility of ductile crack initiation or cleavage fracture was assessed. A remote load of

180 MPa was considered to be the cyclic ‘operating’ load that causes fatigue crack growth. This value

of nominal stress is about 50% of the 0.2% proof stress and 34% of the UTS of the parent plate, and

so is typical of an engineering structure.

An internal pressure was applied to the sphere. This results in an equi-biaxial stress in the spherical

shell that acts to open the defect. Various magnitudes of pressure were applied, with a maximum of

1.8 MPa corresponding to a meridional stress of 225 MPa according to thin shell theory. Again, this

was considered as the overload condition. The operating condition was a repeatedly applied pressure

of 1.44 MPa, causing a nominal stress of 180 MPa in the shell.

Crack driving force was evaluated in terms of stress intensity factor. This parameter was evaluated

both elastically (designated K) and from an elastic-plastic analysis (designated KJ). Because of the

complexity of the finite element analyses, the conventional J-contour integral option with ABAQUS

could not be accurately employed to evaluate K and KJ. The primary reason for this is that the

contour integral calculation of J requires that significant unloading of the material does not take place.

This was not the case in the present finite element analyses that simulated welding, heat treatment,

material removal and repair welding. An alternative calibration approach, based on the crack opening

displacements at the node immediately behind the crack tip, was therefore used as a proxy for J. Full

details of this calibration procedure are contained in Reference 3. It may be noted that J was 0.5

converted to K by the usual equation, K = [(EJ)/(1-n 2)] where E is Young’s modulus (taken as 200

GPa and n is Poisson’s ratio (taken as 0.3).

20

9. TASK 8 – APPLICATION OF FINITE ELEMENT MODELS TO

MATRIX CASES

This Task is associated with applying the finite element models and methodology referred to in

Section 8 above to a matrix of cases. It should be noted that because of previously unforeseen

complexities of the analyses (e.g. the requirement to develop the calibration method used to evaluate

crack driving force), it was not possible to include all the cases that had originally been suggested

under Task 2 (Section 3). A good selection of the cases was included in the analyses however.

9.1 EDGE DEFECTS IN THE WELDED PLATE

Figures 37(a-b) show results for elastic stress intensity factor K for various defect depths and levels of

primary load in the welded plate in the un-repaired heat treated and the as-repaired states,

respectively. The magnitude of primary load is indicated in the legends. 0 MPa corresponds to

residual stress only. The stress intensity factors for the defect in the repaired weld are obviously larger

than in the un-repaired PWHT case. The two curves for residual stress only show a tendency to rise

with increasing defect depth and then gradually fall, reaching a maximum K for about 11 mm defect

depth. This is a consequence of the residual stress fields presented in Figure 33(a) whereby the

stresses are shown to start decreasing in magnitude after reaching tensile peak values at a distance of

about 10 mm from the appropriate side of the plate. The other curves simply show that the additional

stress intensity factor is proportional to the primary load applied.

Figures 38(a-b) show results for KJ calculated from J obtained from elastic-plastic analyses. The

curves for zero primary load are unchanged from Figures 37(a-b). With increasing crack depth and

load, the value of KJ becomes larger than the corresponding value of K in Figure 37 due to plasticity

effects. The KJ results in the as-repaired state are higher than in the PWHT state, particularly for

intermediate defect depths and loads. For deeper defects and higher loads, the residual thermal strains

arising from welding are reduced by the mechanical plastic strains, and so the difference in CDFs

between the two welded states is reduced.

Repeated loading and unloading between zero and 180 MPa was considered. Fatigue crack growth

predictions are made using the Paris law, Equation 3, but with the more representative parameter

DKJ=KJmax-KJmin used in preference to DK. Here KJmin is the crack driving force for the appropriate

residual stress acting alone and KJmax is the total CDF for combined residual stress plus 180 MPa

applied stress. Both these parameters are available in Figures 38(a-b). For each updated crack depth

the value of KJ for an occasional 225 MPa applied stress was also available. This KJ was required to

assess when the critical crack size had been reached during the fatigue crack growth calculations (i.e.

fatigue crack growth was based on an applied stress range of 180 MPa and critical crack size was

based on an overload stress of 225 MPa).

Results of fatigue crack growth predictions are illustrated in Figures 39(a-b). These graphs show

crack depth, a, as a function of the number of loading cycles, N, between zero and 180 MPa for the

different initial defect depths indicated in the legends. Defects in the as-repaired weld, Figure 39(b),

need fewer cycles to grow to a given depth compared with the PWHT state, Figure 39(a), since the

value of DKJ is generally lower for the PWHT state (Figure 38).

Ductile crack initiation, or cleavage failure in the ductile-to-brittle transition region of ferritic steels,

is considered to occur when KJ is equal to a given fracture toughness KJc. No differentiation is drawn

between these types of failure and the term ‘limiting condition’ is used hereafter. In Figures 40(a-b)

21

results are presented for fracture toughness, KJc, versus the number of 0-180 MPa loading cycles, Nf,

required to cause the limiting condition due to an occasional 225 MPa overload. Curves are shown for

different initial defect depths. A comparison of the two graphs shows that for a given fracture

toughness and initial defect, fewer cycles are required to grow to the limiting condition in the as-

repaired weld.

Figure 41 shows the relationship between critical defect depth, ac, at the limiting condition and

fracture toughness in the two weld states. For a given toughness, the critical defect depth is smaller in

the as-repaired weld. The difference in critical defect depth between the two welds depends on

toughness. For example, for a weld toughness of 160 MPaÖm, the critical defect depth is about 17.5

mm in the PWHT weld and 16.7 mm in the as-repaired case. This difference in depth is not

significant. However, for a lower fracture toughness of 100 MPaÖm, the respective critical defect

sizes are about 13.5 mm and 9.5 mm. This difference is more significant.

Figure 42 shows curves of the ratio of the number of loading cycles to the limiting condition for a

defect in the repair, Nf(repaired), to the number of cycles in the un-repaired PWHT state, Nf(un-

repaired). These curves assume the same initial defect depth in both weld states. Each curve

represents a different fracture toughness that is also assumed to be the same in both welds. So in this

graph, a comparison is made of the fatigue life of the same size defect and same fracture toughness in

the repaired and un-repaired welds. Values less than unity imply a worse life for the repair. Of course,

in the majority of cases, this is the case due to the higher repair residual stresses. Some results are

greater than unity for initial defects between 8 mm and 14 mm deep for high toughness. This occurs

because of high values of KJ at zero load in the as-repaired state, Figure 38(b), giving lower values of

DKJ in the as-repaired weld compared with un-repaired and so reduced fatigue crack growth rates.

The series of graphs in Figures 43(a-e) also illustrate the ratio of operating cycles required to reach

the limiting condition for repaired and un-repaired cases. These take account of different initial defect

depths and fracture toughness in the two weld states. The scenarios are either: an edge defect is left in

the (un-repaired) weld, or a repair is carried out that leaves the same size or shallower edge defect

located in material with the same or reduced local fracture toughness. The trade-off is thus explored

between introducing the same or shallower defect in the repair and higher levels of residual stress and

lower fracture toughness in that weld.

Firstly, Figure 43(a) shows comparisons between leaving un-repaired a 5 mm deep edge defect and

inadvertently introducing either 5 mm, 4.2 mm or 3.3 mm deep defects in the as-repaired weld. Curves

are shown of the ratio of operating cycles to reach the limiting condition in the repaired and un-

repaired weld versus the percentage reduction in repaired fracture toughness from the original PHWT

value. Each curve represents a combination of repair defect depth and original toughness. The highest

values of PWHT fracture toughness are represented by blue curves, and the lowest by red. For

example, the blue squares show the effect of leaving in the repair the same size 5 mm deep defect for

an original PWHT fracture toughness of 160 MPaÖm; slightly greater than the initiation toughness of

the plate test welds. The operating life of the repair is always lower than the un-repaired life (ratio of

cycles to the limiting condition is less than unity). Repair life gets comparatively worse as the

repaired toughness reduces. So, a 40% reduction of the repaired toughness, compared with the

original PWHT value, leads to a halving of the repaired life compared with the life if left un-repaired.

The open blue diamonds show the effect of introducing into the repair a 4.2 mm defect compared with

leaving un-repaired the PWHT weld containing a 5 mm defect. The repaired life slightly exceeds the

un-repaired life by only a small margin, though if the repaired toughness drops more than 20% below

the original 160 MPaÖm, the life of the repair becomes less than the un-repaired life. The blue

triangles show the comparison between having a 3.3 mm defect in the repair and leaving un-repaired

22

the 5 mm defect. The repaired exceeds the un-repaired life until the repaired toughness drops below

about 43% of the PWHT level.

There are more interesting consequences for lower PWHT fracture toughness. Consider a toughness

of 100 MPaÖm in the PWHT state; the three sets of orange curves and symbols in Figure 43(a). A

defect in the repair having a depth of either 5 mm or 4.2 mm always has a shorter operating life than

the 5 mm deep defect in the PWHT weld. A 3.3 mm deep repair defect, shown by orange triangles,

gives a slightly longer life than the un-repaired 5 mm case for no reduction of toughness. However, a

mere 10% or so reduction of toughness due to the repair results in a shorter operating life. For the

lowest 80 MPaÖm PWHT toughness (red curves and symbols), all repaired defects from 3.3 mm to 5

mm depth imply an inferior fatigue life even if the repaired toughness does not change. These results

therefore demonstrate that repairing a shallow surface defect by re-welding is likely to result in a

shorter operating life if it leaves a defect and reduces the fracture toughness. This is particularly

apparent for materials with low original toughness. Although the repair surface defects considered

here could be detected visually or by Magnetic Particle Inspection, it is considered that a defect about

3 mm deep cannot be sized accurately by Ultrasonic Techniques.

Figure 43(b) shows similar sets of predictions for a 6.7 mm deep original defect. Here a defect of

depth 6.7 mm, 5 mm or 3.3 mm is considered left in the repair. The trend of the predictions is similar

to the 5 mm case discussed above, but a larger reduction of toughness is needed to obtain a shorter life

in the repaired situation. For example, the orange triangles show that for 100 MPaÖm toughness in the

PWHT weld, a 45% reduction due to repair is required to give a shorter life for a 3.3 mm deep repair

defect.

Figures 43(c-e) however provide more support for repairing deeper surface defects. Figure 43(c)

compares an un-repaired 9.2 mm defect with repaired defects of 6.7 mm, 5 mm or 3.3 mm. Note that

not all symbols in the legend are seen on the graph because some initial defect/toughness

combinations considered meet the limiting condition and so imply zero operating life (see Figure 41)

or the repair life exceeds twice the un-repaired. The steeper angle of the curves suggests that, for these

deeper initial defects, the effect of toughness reductions due to repair can be more severe. For

example, the red diamonds compare the un-repaired 9.2 mm defect with 5 mm in the repair for a low

PWHT toughness of 80 MPaÖm. If the toughness reduces by up to 10%, the life of the repair is still

over twice the life if un-repaired. However, a toughness reduction of 30% due to repair causes the

repaired life to drop drastically to about one quarter of that if the weld was left un-repaired.

Figure 43(d) shows a comparison of the 10.8 mm deep un-repaired defect with 9.2 mm, 6.7 mm or 5

mm defects in the repair. Since it is unlikely that a 9.2 mm defect is left in a repaired weld, the

shallower depths are perhaps more feasible. Considering 160 MPaÖm PWHT toughness and a 5 mm

defect in the repair (blue triangles) then a 60% toughness reduction due to repair (down to about 64

MPaÖm) is required to obtain a shorter fatigue life in the repair. If the PWHT weld has a lower 80

MPaÖm toughness (red triangles) then only a 30% reduction down to about 56 MPaÖm will give a

worse or even no repair life.

Finally, Figure 43(e) compares the un-repaired 13.3 mm deep defect with 9.2 mm 6.7 mm or 5 mm in

the repair. Many of the ratios are zero or unreported because there is no un-repaired or repaired life:

the initial defect is at or beyond the limiting condition. Obviously, this original 13.3 mm defect is

more likely to warrant repair than the shallow ones discussed above. However, onerous welding

conditions, giving the likelihood of poor toughness and a remaining defect, can result is a worse life.

Leaving a mere 5 mm deep defect, and reducing an original toughness of 100 MPaÖm (orange dotted

curve) by 45%, will lead to little or no life of the repair. It will reach the limiting condition on first

overload. As seen in Figure 41, a 5 mm deep defect with a toughness of about 60 MPaÖm is near the

limiting condition in the repair weld.

23

9.2 EQUATORIAL DEFECTS IN THE WELDED SPHERE

The stress intensity factors for primary load alone are in good agreement with standard results for

extended edge defects in spheres. Figures 44(a-b) give the results for the parameter KJ from the

welded sphere simulations. It is apparent that the crack driving forces are lower in this geometry than

previously seen for the plate. Due to algebraically more compressive residual stress at the defective

side 2 of the PWHT sphere weld (Figure 35a) the crack driving forces are negative for the shallowest

and also for the deepest defects in the absence of primary load, as seen in Figure 44(a).

Repeated loading and unloading between nominal biaxial stresses of zero and 180 MPa was again

considered, with fatigue crack growth predictions made using Equation 3 and DKJ=KJmax-KJmin. Only

the range over which KJmin is positive contributes to fatigue since the crack is actually closed if KJmin is

negative according to Figure 44(b). The value of KJ was calculated for an assumed overloading to a

nominal stress of 225 MPa at each crack depth and associated number of cycles. Fatigue crack growth

predictions are shown in Figures 45(a-b). As seen in the welded plate case earlier, defects in the as-

repaired state need fewer cycles to grow to a given depth compared with the PWHT state.

Figures 46(a-b) show results for fracture toughness versus the number of loading cycles required to

cause the limiting condition at the 225 MPa load. As with the welded plate, fewer cycles are required

in the as-repaired sphere weld to grow the defect to the limiting condition. Note that the range of

toughness is shifted to lower values compared with the welded plate because of the lower crack

driving forces in the sphere.

Figure 47 plots critical defect depth at the limiting condition as a function of fracture toughness in the

two weld states. Again for a given toughness, the critical defect depth is smaller in the as-repaired

weld. However, the difference between the two cases is more significant due to the generally lower

levels of toughness illustrated. For example, for a weld toughness of 100 MPaÖm, the critical defect

depth is about 19 mm in the PWHT weld and 10.5 mm in the as-repaired case. The respective critical

depths for the welded plate (Figure 41) are about 13.5 mm and 9.5 mm. Thus the difference between

critical depths in the welded sphere is clearly more significant than for the plate. The green curve in

Figure 47 suggests that for PWHT toughness close to 60 MPaÖm there is a large change in critical

crack depth. This is due to the flat or falling CDF in Figure 44(a) arising from compressive PWHT

residual stress at distances from side 2 greater than about 10 mm, see Figure 35(a).

Figure 48 compares the fatigue life of the same initial size defect and fracture toughness in the

repaired and un-repaired welds. Results are always less than unity; implying a worse life for defects in

the repair.

Figures 49(a-e) illustrate, for edge defects in the welded sphere, the trade-off between introducing the

same or shallower defect in the repair and higher residual stress and lower toughness there. These

graphs are similar to Figures 43(a-e) for the welded plate discussed earlier, with the exception that the

maximum toughness examined here is lower due to smaller crack driving forces in the sphere.

Figure 49(a) contrasts leaving un-repaired a 5 mm deep defect with introducing either 5 mm, 4.2 mm

or 3.3 mm deep defects in the as-repaired weld. With few exceptions, the operating life of the repair is

always lower than the un-repaired life for PWHT fracture toughness up to 110 MPaÖm. Repairing a

long 5 mm deep surface defect in this weld geometry by re-welding without heat treatment is not

beneficial if it is likely that a mere 3 mm or so deep surface defect can remain undetected after repair.

The probability that the toughness will be reduced by a non-heat treated repair reinforces this

conclusion.

24

Figure 49(b) shows comparisons between leaving un-repaired a 6.7 mm deep defect and introducing

6.7 mm, 5 mm or 3.3 mm deep defects in the repair. For the highest PWHT toughness of 110 MPaÖm

and the smallest repair defect of 3.3 mm (blue triangles), the defective repair has a longer life unless

the repair causes a toughness reduction of about 35% to 72 MPaÖm. For the lowest PWHT toughness

examined of 70 MPaÖm, a mere 12% toughness reduction will give a lower life for a 3.3 mm deep

defect in the repair (red triangles).

Looking ahead to Figure 49(d), compares leaving un-repaired a 10.8 mm deep defect in the PWHT

weld with having 9.2 mm, 6.7 mm or 5 mm defects in the as-repaired state. For 110 MPaÖm PWHT

toughness and leaving the 5 mm defect after repair (blue triangles), a lower life is achieved by the

repair should the toughness fall by more than 35% to about 72 MPaÖm. For the lowest considered

PWHT toughness of 70 MPaÖm, only a 12% or so reduction in toughness will give a lower fatigue

life for the 5 mm repair defect (red triangles).

Finally, Figure 49(e) compares a 13.3 mm un-repaired defect with 9.2 mm, 6.7 mm and 5 mm defects

in the repair. For 110 MPaÖm PWHT toughness and a 5 mm defect in the repair (blue triangles) then

about 43% toughness reduction, to 63 MPaÖ m. is required to obtain a shorter life in the repair.

9.3 EMBEDDED DEFECTS IN THE WELDED PLATE

This section explores the behaviour of embedded defects in both the un-repaired and repaired weld in

the plate. Two initial types of defect configuration were considered. In the first, labelled ‘p+2a=16.7

mm’, the upper defect tip closest to side 1 (see Figure 5) lies at a depth of 16.7 mm from the repaired

side 2 of the plate. Various initial defect heights 2a were examined. In the second configuration,

‘p+2a=10.8 mm’, the upper defect tip is 10.8 mm from side 2. Again, various initial defect heights

were studied. In all cases examined here, only the lower tip of the defect closest to repaired side 2

was considered. This necessary simplification meant that fatigue crack growth was not considered at

the upper tip closest to side 1. This is not as approximate as it might at first appear, particularly for

p+2a=16.7 mm, since the upper tip lies far from side 1 and generally experiences lower crack driving

forces (and ranges) than the lower tip of the defect. Given the power law dependence of the Paris law,

Equation 3, this leads to much lower rates of fatigue crack growth than experienced by the lower tip.

Figures 50(a-b) gives some KJ crack driving force results for increasing height of an embedded defect

in the un-repaired and repaired weld. These relate to the case p+2a=16.7 mm. A comparison of

Figures 50(a-b) and Figures 38(a-b) shows that, for the same defect height/depth, the CDFs for the

embedded cases are comparable to the edge cases, particularly for higher/deeper defects. At first

sight, this appears to be inconsistent with what is generally understood: that edge cracks have higher

CDFs than embedded cracks of the same depth. However, the embedded defect tip is developing

towards the repaired surface and so experiencing an increasing tensile nominal stress field. By

contrast, the edge defect results relate to the (only) tip of the defect in the ‘deep’ position which

develops towards a more compressive stress field at plate mid-thickness. Should the 2a=14 mm high

embedded defect break through the 2.7 mm remaining ligament to the repaired surface, it is re-

characterised as a 16.7 mm edge defect. In both PWHT and as-repaired welds, the CDF will increase

at the 220 MPa maximum applied load plotted: compare Figures 38(a-b) for a=16.7 mm with Figures

50(a-b) for 2a=14 mm.

Figures 51(a-b) plot fracture toughness versus number of loading cycles to the limiting condition for

the case p+2a=16.7 mm. As expected, for the same toughness fewer cycles are achieved in the as-

repaired weld. Figures 52 and 53 compare critical crack depths for un-repaired PWHT and as-repaired

welds for the two embedded cases p+2a=16.7 mm and p+2a=10.8 mm, respectively. A lower range

of toughness is displayed in these cases compared with the edge defects (Figure 41) consistent with

the generally smaller crack driving forces obtained. The rapid change of critical defect height with

25

toughness in Figure 53 compared with Figure 52 is due to the defect tip in question being closer to,

and so more sensitive to, the tensile part of the residual stress field near the plate surface.

Figure 54 plots the ratio of cycles to limiting condition in the as-repaired weld to cycles in the un-

repaired condition as a function of toughness and initial defect height for the case p+2a=16.7 mm.

This shows that, for the same initial defect size and toughness in both welds, a shorter life is generally

obtained in the repair. A longer repair life is however seen for the very short initial defects examined.

This behaviour is due to the defect tip of interest lying in the compressive part of the repair residual

stress field, resulting in low rates of fatigue crack growth compared to the un-repaired case. Results

for the case p+2a=10.8 mm are seen in Figure 55. There is a more restricted range of toughness to

show here due to the low CDFs for this shorter defect.

Figure 56(a) contrasts leaving un-repaired a 5 mm high defect with introducing either 5 mm, 4.2 mm

or 3.3 mm high defects in the as-repaired weld for p+2a=16.7 mm. The squares show the effect of

having the same size 5mm deep defect in the repair. Obviously the repaired life is always lower than

the un-repaired life, and gets comparatively worse as the repaired toughness reduces. If the repaired

defect is 4.2 mm high (diamonds), a reduction in toughness is needed to get a worse life out of the

repair. The triangular symbols for the shallow 3 mm high defect in the repair are well over unity due

to a large life of that repair. This is due to the defect tip in question lying well inside the compressive

region of the repair residual stress field, giving low initial crack growth rates. However, the fatigue

lives are generally very long for this un-repaired defect (see rightmost curve in Figure 51(a). Leaving

un-repaired small height defects near the middle of the plate is therefore likely to be a reasonable

course of action.

Results comparing a 6.7mm high defect in the un-repaired weld with 6.7 mm, 5 mm or 4.2 mm high in

the repair are illustrated in Figure 56(b). The diamonds show that introducing a smaller 5 mm defect

in the repair always gives a shorter life. The triangles start to appear, showing the smallest repaired

defect of 4.17 mm where the fatigue life ratios remain well above unity.

Figure 56(c) compares the 9.2 mm high un-repaired defect with 9.2 mm, 6.7 mm or 5 mm in the

repair. The diamonds have moved up slightly compared with the previous graph, but the squares have

shifted downwards. This is an interaction between tip position and the associated residual stress field.

The defect tip is growing towards the repaired surface, so initially higher (longer) defects experience

more strongly the tensile region of the residual stress near the repair surface.

In Figure 56(d), a 10.8 mm high un-repaired defect is compared with smaller 9.2 mm, 6.7 mm or 5

mm defects in the repair. Leaving the same size in the repair (squares) always gives a shorter

operating life, particularly for lower toughness. The 6.7 mm repair defect (diamonds) needs a large

reduction in toughness to give shorter life than the un-repaired 10.8 mm defect.

26

The final graphs, Figures 57(a-b) show similar comparisons for the embedded cases in which

p+2a=10.8 mm. There is a more restricted range of defect heights and toughness to consider and so

fewer points are plotted than in Figure 56. Also, the growing defect tip of interest lies at a

comparatively shallow depth in the repair, so it tends to experience more tensile repair residual

stresses and so has a comparatively shorter fatigue life. In Figure 57(a) the life ratio of many of the

triangular symbols is below unity, indicating that repairing the 5 mm high defect but leaving a 3.3 mm

one gives a shorter life. The highest 60 MPaÖm PWHT toughness considered needs only 16%

reduction in the repair to give a shorter life. As noted earlier, the defect tip considered is now shifted

towards the tensile part of the repair residual stress, so the repair has a relatively shorter life. In fact,

for a large range of toughness the initial repair is at a limiting condition with respect to the overload

considered, so the repair life is zero cycles. Figure 57(b) compares a 6.7 mm high un-repaired defect

with 6.7 mm, 5 mm or 4.2 mm repaired. Many repaired cases have no life for the range of toughness

considered. The triangles show that having a 4.2 mm defect in the repair gives a lower life than the

un-repaired 6.7 mm defect if repair causes a modest reduction in toughness.

27

10. TASK 9 – ASSESSMENT BY ENGINEERING PROCEDURE

METHODS

10.1 GENERAL METHODOLOGY

British Standard BS 7910:1999, "Guide on Methods for Assessing the Acceptability of Flaws in

Metallic Structures" (Reference 4), contains three levels for the assessment of fracture resistance.

The materials involved, the input data available and the conservatism required are the factors which

determine the chosen level. For the purpose of the calculations carried out in this task, Level 2, the

normal assessment route was applied. This involves values of the plastic collapse parameter, Lr and

fracture mechanics parameter, Kr to be evaluated and plotted on the appropriate fracture assessment

diagram (FAD).

Three residual stress conditions were assumed. They were (i) as-welded, (ii) as-welded followed by

post weld heat treatment (PWHT) and (iii) weld repair. In case (i), the transverse residual stress

distribution shown in Figure Q.1(a) of BS 7910 was considered. It was judged that this distribution

could be approximately represented by a through-wall bending stress equal to the material yield stress.

Case (ii) assumed the residual stresses to be a membrane stress equal to 20 % of the yield stress

(Section 7.2.4.2 of BS 7910). In case (iii), the transverse residual stress distribution shown in Figure

Q.1(d) of BS 7910 was considered. It was judged that this distribution could be approximately

represented by a membrane stress equal to the material yield stress. In these calculations, the yield

stress was taken as the lowest value of 0.2% proof stress given in Table 2 (i.e. 345 MPa). To

summarise therefore, the residual stress conditions assumed were:

(i) As-welded condition – through-wall bending stress (+ 345 MPa at the surfaces).

(ii) As-welded followed by PWHT condition – membrane stress of 69 MPa.

(iii) Weld repair condition – membrane stress of 345 MPa.

The term r is included in the evaluation of Kr in order to cover interaction between the primary and

secondary stress systems. The procedure used to determine r was as detailed in Annex R.2 of BS

7910 as follows:

a) Determine KIs, the linear elastic stress intensity factor for the flaw size of interest, using the

elastically-calculated secondary stresses. KIs is positive when it tends to open the crack.

If KIs is negative or zero, then r is set to zero and the remainder of this procedure does not

apply.

b) Determine the ratio KIp/Lr.

s sc) Determine KI /(KI

p/Lr) from the result of a) and b). If KI /(KI

p/Lr) > 4 then Annex R.3 of BS

7910 should be used to evaluate r. This is a more detailed procedure for calculating r and the

steps involved are:

si. Calculate the parameters KI

s and Kp . Advice on determining the effective (elastic-

plastic) stress intensity factor Kps

is given in Annex R.4 of BS 7910. In these

calculations, Kps, was evaluated by the route given in Annex R.4.3 of BS 7910 which is

based on the small-scale yielding correction to KIs.

28

sii. Determine the ratio Kp /(KI

p/Lr) where KI

p and Lr are calculated as in Sections 7.3 and

7.4 of BS 7910.

iii. Obtain the parameter y from the table in Annex R.1 of BS 7910 in terms of Lr and the s

parameter Kp /(KIp/Lr) calculated in step (II). Linear interpolation should be used for

svalues not given in the table. If Kp

s = KI , then r is set equal to y and the remainder of

this annex does not apply.

iv. Obtain the parameter j from the table in Annex R.2 of BS 7910 in terms of Lr and the s

parameter Kp /(KIp/Lr) from step (II). Linear interpolation should be used for values not

given in the table.

v. Determine r from the following equation.

K s ö æ

I -1r -Y = jç

ç è

÷÷ øK s

P

If this results in a negative value for r, then r is re-defined to be zero.

The detailed procedure, described above, for the determination of r was automated in the s

calculation if KI /(KIp/Lr) > 4. A visual basic program was used to search two tables to find

s sappropriate values of y and j as functions of Kp /(KI

p/Lr) and Lr, where Kp is as described in

sAnnex R4.3 of the procedures. If KI /(KI

p/Lr) < 4, r was evaluated following the simplistic

route of steps d) and e) below.

a) Determine r1 from Figure R.1 of BS 7910.

b) Determine r.

r = r1 Lr <= 0.8

r = 4r1(1.05 - Lr) 0.8 < Lr < 1.05

r = 0 1.05 <= Lr

10.2 EDGE CRACKS

10.2.1 Available Solutions

and s

In the case of an edge crack, two possible solutions are available in BS 7910 for the calculation of KI

Reference. Figures M.6 and M.10 of BS 7910 show a long surface flaw and an edge flaw geometry

respectively. The solutions described for the long surface flaw geometry, Figure M.6 of BS 7910,

were used rather than those for the edge crack flaw geometry. The reason for this is that the axis of

the plane of bending (as required for the residual stress case (i) ) is not correct in the case of the edge

flaw geometry.

However, two KI solutions were calculated to compare the difference between the two crack

geometries for just membrane loading. This clearly showed that there was a negligible difference

between the two solutions.

The stress intensity factor solution (Eq. M.1 of BS 7910) is:

29

K I = (Ys ) pa

where, for Level 2 assessments

Ys = (Ys )p + (Ys )s (Eq. M.4 of BS 7910)

where (Ys )p and (Ys )s represent contributions from primary and secondary stresses respectively.

They are calculated as follows:

(Ys )p = Mf [k M M P + k M M {P + (km - 1)Pm }] (Eq. M.5 of BS 7910)w tm km m m tb kb b b

(Ys )s = M + Q M b (Eq. M.6 of BS 7910)mQm b

where, for the case under consideration, M, ktm, Mkm, ktb, Mkb, fw = 1, and Mm and Mb are given below

for a/B <= 0.6 (Section M3.3 of BS 7910):

4Mm = 1.12 - 0.23(a/B) + 10.6(a/B)2 - 21.7(a/B)3 + 30.4(a/B)

4Mb = 1.12 - 1.39(a/B) + 7.32(a/B)2 - 13.1(a/B)3 + 14(a/B)

The reference stress for a long surface flaw in flat plates is as follows (Section P.3.2 of BS 7910

assuming normal bending restraint):

5.0

P + [Pb

2 + 9P 2

(1 - a ' ' ) 2 ]b m

s = ref 2 (1 3 - a ' ' )

where a'' = a/B

10.2.2 Results

Figure 58 contains values of crack depth versus number of fatigue cycles for initial crack depths of

3.33 mm, 10 mm and 16.67 mm in the plate subjected to 180 MPa cyclic primary stress for the un-

repaired PWHT case. The calculated values using the BS 7910 K solutions are compared to the FE

solutions of Figure 39(a). As has been explained previously, the FE calculations were carried out in

terms of the elastic-plastic DKJ as opposed to the elastic DK. Since the BS 7910 calculations were

evaluated in terms of DK, such evaluations have also been carried out based on the FE K solutions

(using the data given in Figure 37) and the results of these have been included in Figure 58. It should

be noted that since Kmin in the DK calculations is based on the weld residual stress alone and Kmax is

based on the primary stress plus the residual stress, then the actual value of the residual stress is not

sensitive to such calculations. It can be seen from Figure 58 that the BS 7910 based crack depth vs

cycles results are very similar to those based on the FE elastic solutions. This implies of course that

the BS 7910 K values were very similar to those of the elastic FE K values. This aspect is considered

further in section 10.2.3 below.

Figure 59 contains the same type of information as for Figure 58 but for the as-repaired condition.

Again, the BS 7910 based crack depth vs cycles results are shown to be very similar to those based on

the FE elastic solutions.

30

Figure 60 contains fracture toughness vs critical crack depth curve evaluated by BS 7910 for the

weldment material states (i.e. residual stresses) (i), (ii) and (iii). The calculations have been carried

out for a 225 MPa overload stress along with the various residual stress assumptions. As would be

expected, the PWHT state results in the largest critical crack depth sizes (the residual stress being 69

MPa membrane), the as-welded state results in the second largest values (the residual stress being 345

MPa through-wall bending) and the PWHT state results in the lowest values (the residual stress being

345 MPa membrane).

Figure 61 shows comparisons of the BS7910 evaluated and FE calculated values of critical crack

depth for the un-repaired PWHT condition. It can be seen that the critical crack depths based on the

BS 7910 calculations are somewhat lower than those obtained by the FE analysis. This aspect is

considered further in section 10.2.3 below.

Figure 62 contains the same type of information as for Figure 61 but for the as-repaired condition.

For this state, the critical crack depths based on the BS 7910 calculations are seen to be significantly

lower than those obtained by the FE analysis. Again, this is further considered in section 10.2.3.

10.2.3 Refined Calculations

In the previous sub-section, with reference to Figures 61 and 62, it was shown that the critical crack

sizes based on the BS 7910 calculations were lower than those obtained by the FE analysis. It was

considered that the differences in residual stress distributions between those of the BS 7910 document

and those evaluated by finite elements was likely to be the main reason for these differences in critical

crack sizes. The residual stress distributions evaluated by the finite element techniques were

therefore considered in refined BS 7910 calculations. Since the residual stress fields are of a

sinusoidal nature and since no solutions are available in BS 7910 for evaluating stress intensity factor

(KI) solutions for such stress distributions, alternative K solutions were employed as follows.

KI, for an edge crack in a plate or a cylinder can be represented (Reference 5) as:

0.5 2 3KI = (pa) [F0A0 + 2(a/t)/p F1A1 + (a/t) /2 F2A2 + 4(a/t) /(3p) F3A3] (4)

where, a is crack depth and t is wall thickness.

A0 to A3 are constants in the cubic polynomial equation representing the through-wall stress

distribution, s, over the depth (a) of the crack (but prior to the presence of the crack), i.e.:

2 3 s = A0 + A1(x/t) + A2(x/t) + A3(x/t)

x is the distance into the plate thickness

For a flat plate, the Fn functions are given by:

F0 = [1.148 – 0.9913 a/t + 3.076(a/t)2] / [1 – a/t]

F1 = [1.077 – 0.8345 a/t + 1.543(a/t)2] / [1 – a/t]

F2 = [1.007 – 0.7007 a/t + 0.781(a/t)2] / [1 – a/t]

F3 = [1.015 – 0.7296 a/t + 0.446(a/t)2] / [1 – a/t]

Solutions for Fn at the deepest point of the crack are also available for cylinders.

It is evident that for applied membrane loading, only the first term in Equation 4 will be required

since s = A0.

31

For through-wall bending, the first two terms in Equation 4 are required since s=A0+A1a/t where A1=-

2A0 with A0 being the stress at the surface.

The above equations were used to evaluate values of KI for the various residual stress distributions

modelled in the finite element analyses of Tasks 7 and 8. In preliminary calculations, the “true” finite

element calculated distributions were considered. However, since these generally consisted of

compressive stresses at the surface of the plate where the crack was situated, the calculational route

described above resulted in compressive values of KI, i.e. crack closure, for all lengths of crack. A

study of all the residual stress distributions (a. As-welded, b. As-welded and Repair, c. As-welded and

PWHT, d. As-welded and PWHT and Repair) indicated that the peak tensile stress occurred at a

distance of approximately 7.1 mm from the surface. As a compromise in modelling the residual stress

distributions to evaluate KI, it was assumed that the peak tensile stress occurred over the first 7.1 mm

of the plate. These modified residual stress distributions are shown in Figure 63 over approximately

one half of the plate, i.e. for distances of up to approximately 20 mm from one side of the plate.

When evaluating KI for these stress distributions, different cubic polynomial equations were fitted for

different crack sizes since it was not possible for one equation to accurately represent all crack sizes

in the range being considered (i.e. 0 to 20 mm). Once values of KI had been evaluated for the

different crack sizes, a cubic polynomial equation was fitted for each of the four residual stress cases,

represented by the equation:

2 3KI = C0 + C1a + C2 a + C3 a

where a is crack depth (in mm) and KI is in MPaÖm.

For completeness, values of KI were also obtained for a sinusoidal distribution (tensile at both

surfaces and compression in the middle region of the plate) and for membrane and bending (tensile

stress on one side of the plate and compressive stress on the other). Values for C0 to C3 for the

different stress cases considered are as follows:

32

Stress Case C0 C1 C2 C3

As-Welded

(Stress = 312.5 MPa for a = 0 to 7.1 mm)

4.9282 11.388 -0.5382 0.008

As-Welded then Repair

(Stress = 316.2 MPa for a = 0 to 7.1 mm)

4.5909 11.788 -0.5944 0.013

As-Welded then PWHT

(Stress = 110.3 MPa for a = 0 to 7.1 mm)

3.3661 3.495 -0.1406 0.0012

As-Welded then PWHT then Repair

(Stress = 317.6 MPa for a = 0 to 7.1 mm)

9.7278 10.043 -0.4285 0.0085

Sinusoidal Distribution

(Stress peaks at 110.3 MPa at Surfaces)

1.9775 4.0770 -0.3384 0.00773

Sinusoidal Distribution

(Stress peaks at 317.6 MPa at Surfaces)

5.6939 11.740 -0.9744 0.02226

Bending

(Stress = 110.3 MPa to –110.3 MPa)

2.2364 3.5034 -0.2544 0.009

Bending

(Stress = 317.6 MPa to –317.6 MPa)

6.4394 10.0877 -0.7327 0.0259

Membrane

(Stress = 110.3 MPa)

1.8927 4.0597 -0.2658 0.0128

Membrane

(Stress = 317.6 MPa)

5.4498 11.6897 -0.7655 0.0368

Values of KI plotted against crack depth are presented in Figures 64 to 66.

Figure 64 contains the evaluated KI distributions for the As-welded, As-welded-Repair, As-welded-

PWHT and As-welded-PWHT-Repair cases. As would be expected by consideration of the stress

distributions given in Figure 63, the As-welded-PWHT-Repair case gives the highest KI values and

the As-welded-PWHT case gives the lowest.

Figure 65 contains the evaluated KI distributions for the As-welded-PWHT case together with the

evaluated KI distributions for the sinusoidal stress distribution case (the stress peaking at 110.3 MPa

at the surface), the bending stress distribution case (stress = 110.3 MPa on one surface and –110.3

MPa on the other surface) and the membrane stress case (=110.3 MPa). As would be expected by

consideration of the respective stress distributions, the KI distribution for the As-welded-PWHT case

is similar to that for the membrane case for crack depths of up to approximately 7 mm after which the

KI values for the latter increase significantly. It may be noted that for the BS 7910 calculations

covered in section 10.2.2, the stress distribution for the As-welded-PWHT case was taken to be 69

MPa membrane. The KI vs crack depth curve for this distribution would therefore lie some 40%

lower than the membrane curve shown in Figure 65 (see below with reference to Figure 68).

Figure 66 contains the evaluated KI distributions for the As-welded-PWHT-Repair case together with

the evaluated KI distributions for the sinusoidal stress distribution case (the stress peaking at 317.6

MPa at the surface), the bending stress distribution case (stress = 317.6 MPa on one surface and –

317.6 MPa on the other surface) and the membrane stress case (=317.6 MPa). Again, as would be

expected by consideration of the respective stress distributions, the KI distribution for the As-welded-

PWHT-Repair case is similar to that for the membrane case for crack depths of up to approximately 7

mm after which the KI values for the latter increase significantly. For the BS 7910 calculations

covered in section 10.2.2, the stress distribution for the As-welded-PWHT-Repair case was taken to

be 345 MPa membrane. The KI vs crack depth curve for this distribution would therefore lie just

about 10% above the membrane curve shown in Figure 66 (see below with reference to Figure 69).

33

Figure 67 contains a comparison of the various KI solutions considered in this work for an edge

cracked plate subjected to a membrane stress of 220 MPa. The “handbook” solutions of Rooke and

Cartwright (Reference 6) for both restrained and un-restrained bending have also been included. It

can be seen that the values obtained from the BS 7910, finite element, Sharples et. al. solutions

(Reference 5) described above and the Rooke and Cartwright un-restrained bending solutions are all

very similar to one another, with the FE values lying slightly below the others. The Rooke and

Cartwright restrained bending solution values lie significantly below those of the other solutions.

This confirms that the solutions considered in this work are relevant to the un-restrained bending case.

Such conditions are relevant for application to the applied membrane stress case. However, since

residual stresses arise from a displacement control mechanism, restrained bending conditions may be

more appropriate when evaluating values of KI for such stresses. The use of the BS 7910 and

Sharples et. al. solutions may therefore result in over-estimates of KI values for residual stresses.

Figure 68 contains a comparison of the KI values evaluated from the various solutions for the As-

welded-PWHT residual stress case. The BS 7910 and Rooke and Cartwright values are for a

membrane stress of 69 MPa. For completeness, the Rooke and Cartwright restrained bending values

have been included. Again, the BS 7910 and Rooke and Cartwright un-restrained bending values are

identical to one another. As suggested above with reference to Figure 65, the Sharples et. al. solution

values are somewhat higher than the BS 7910 (and Rooke and Cartwright un-restrained bending)

solution values up to a crack depth of just under 15 mm. The finite element solution values are fairly

close to the BS 7910 values up to a crack depth of approximately 10 mm after which they start to

diverge and become considerably lower.

Values of critical crack depth obtained by using the Sharples et. al. residual stress KI solutions have

been included in Figure 61 (referred to as “BS 7910 refined”). It can be seen that the critical crack

depth values for this case are actually lower than those of the original BS 7910 calculations. This of

course can be explained by the difference shown in Figure 68 between the BS 7910 and Sharples et.

al. values of KI for the residual stress. Although such calculations have not been performed, it is

evident from Figure 68 that critical crack depth values closer to those obtained from the finite element

analysis may be obtained by using the Rooke and Cartwright restrained bending solutions for

modelling the residual stress field as 69 MPa membrane.

Figure 69 contains a comparison of the KI values evaluated from the various solutions for the As-

welded-PWHT-Repaired residual stress case. The BS 7910 and Rooke and Cartwright values are for

a membrane stress of 345 MPa. Again for completeness, the Rooke and Cartwright restrained

bending values have been included. The BS 7910 and Rooke and Cartwright un-restrained bending

values are of course identical to one another. As suggested above with reference to Figure 66, the

Sharples et. al. solution values are higher than the BS 7910 (and Rooke and Cartwright un-restrained

bending) solution values. It is also interesting to note that the Sharples et. al. solution values are very

similar to those of the Rooke and Cartwright restrained bending solution. The finite element solution

values are significantly lower than the other values.

Values of critical crack depth obtained by using the Sharples et. al. residual stress KI solutions have

been included in Figure 62 (referred to as “BS 7910 refined”). It can be seen that the critical crack

depth values for this case are slightly higher than those of the original BS 7910 calculations. This of

course can be explained by the difference shown in Figure 69 between the BS 7910 and Sharples et.

al. values of KI for the residual stress. This time, it is evident that even using the Rooke and

Cartwright restrained bending solutions for modelling the residual stress field as 345 MPa membrane

would not result in critical crack depth values closer to those obtained from the finite element

analysis.

34

The significance of the BS 7910 calculation results in terms of the fatigue life for the repaired

condition compared to the fatigue life for the un-repaired (i.e. as-welded-PWHT state) condition can

be understood from the information presented in Table 3. This table gives values of Nrepaired/Nun-repaired

for initial crack depths (in both the repaired and un-repaired conditions) of 3.33 mm, 10 mm and

N

16.67 mm for fracture toughness values (KIC) ranging from 60 to 160 MPaÖm. The values in the table

have been compiled from the crack depth versus number of cycles, N, data presented in Figures 58

and 59 and from the fracture toughness versus critical crack depth data of Figures 60 and 61. The

finite element values based on KJ, given in Table 3, are of course those plotted in Figure 48. An

explanation of the values shown in brackets in Table 3 is given in the next paragraph. With the

exception of four cases, It can be seen that the finite element Nrepaired/Nun-repaired values based on elastic

K are somewhat higher than those based on elastic-plastic KJ. It can also be seen that the BS 7910

refined calculation values are somewhat higher than the BS 7910 original calculation values. It is also

noticeable that for the lower fracture toughness values of 100 and 120 MPaÖm, the BS 7910

repaired/Nun-repaired values are considerably lower than the finite element values. However for the higher

fracture toughness values of 140 and 160 MPaÖm, the BS 7910 Nrepaired/Nun-repaired values are closer to

the finite element values. Where direct comparisons between the BS 7910 and the finite element

results are available within Table 3, the same trends are shown for both methods, i.e. that the fatigue

life for the repaired state is lower than that for the un-repaired state. It is shown however, that the

finite element margins of Nrepaired/Nun-repaired are not produced by the BS 7910 calculations.

The values contained in brackets in Table 3 have been derived as follows:

The Paris Law equation, as in Equation 3 is of the form:

dN da = CDKm (5)

Rearranging this equation gives the incremental cycle, dN , as

da dN = (6)

CDKm

da or dN µ (7)

DKm

5.0 (Now DK µ p s D a) (8)

and since in the work being considered here, sD for the repaired case is taken to be the same as that

for the un-repaired case then:

DK µ a 5 .0 (9)

For the increment cycle being considered in these calculations, crack depth a, grows from the initial

size, aI, to the limiting size, aL. It is reasonable to assume therefore, that the average value of a

(i.e. (aI+aL)/2) can be used in Equation 9. Therefore:

5.0

æ a + a ö DK µ ç i L

÷ (10) è 2 ø

Setting da to (aL-aI) and the value of m to 2.77 (Equation 3) and substituting Equation 10 into

Equation 7, results in:

35

a - a dN µ L i (11)

385.1 (ai + a )L

dN 385.1

repaired (aL - ai )repaired

(ai + aL )unrepaired or = (12)

385.1 dN unrepaired (aL - ai )unrepaired (ai + aL )repaired

The values in brackets in Table 3 were therefore obtained from Equation 12.

It can be seen from Table 3 that these values are very close to those (without the brackets) obtained

from the detailed fatigue crack growth calculations for when the elastic K’s are used in the

calculations. As may have been expected, the values are not as close to those when the inelastic K’s

(i.e. KJ’s) are used in the calculations.

Based on the elastic route therefore, Equation 12 seems to be a reliable and relatively easy route for

determining values of Nrepaired/Nun-repaired values once the critical or limiting crack length, aL, has been

evaluated for both the repaired and un-repaired cases.

10.3 EMBEDDED CRACKS

10.3.1 Available Solutions

In the case of an embedded crack, a solution is available appropriate to Figure M.1 of BS 7910 for the

calculation of KI and sReference. The geometry for this solution again presented problems in that the axis

of the plane of bending is not appropriate for this particular case and it assumes that the crack is in the

centre of a flat plate.

Another solution, appropriate to Figure M.7 of BS 7910, was then considered. In this case the crack

geometry does not satisfy conditions set. The geometry is not identical to the embedded crack case in

that Figure M.7 has an elliptical crack of length 2c whereas the problem has a crack length of the

same magnitude as the width of the specimen, W. Therefore, 2c/W > 0.5 and not < 0.5 as specified in

the conditions.

After considering the two representations as described above, It was decided that the geometry in

Figure M.1 and associated KI (section M.3.1 of BS 7910) and sReference (section P.3.1 of BS 7910)

solutions should be used, even though it was not exactly like the embedded through-wall crack case in

question.

Membrane stress could be represented in the normal way.

Bending stress (actually not used in the current calculations) could be represented by specifying a

relevant component of membrane stress and a relevant component of bending stress to allow for the

fact that the crack is not in the centre of the plate specimen.

10.3.2 Results

Some comparisons were made between the BS 7910 calculated stress intensity factor values and those

obtained from the finite element analyses. These comparisons are presented in Figures 70 to 72 (for

the position of the crack, p+2a, being 16.7 mm).

36

Figure 70 compares the KI values for an applied membrane stress of 220 MPa. Values obtained from

the appropriate Rooke and Cartwright (Reference 6) solutions have also been included as a check.

These solutions are shown to be almost identical to those of BS 7910. The finite element values are

shown to be very close to those of the BS 7910 calculations up to a crack depth (2a) of approximately

8 mm. For crack depths greater than 8 mm, the FE values then increasingly become higher than the

BS 7910 values. Fatigue crack growth calculations have not been performed using the BS 7910 KI

solutions for the embedded crack cases. However the KI comparisons of Figure 70 tend to suggest

that fatigue crack growth evaluations for the BS 7910 KI solutions would be identical to those of the

FE solutions up to a crack depth of 8 mm after which the former solutions would show a reduced

crack growth rate. Based on just these considerations therefore, the comparisons would tend to

suggest that the BS 7910 KI solutions may be non-conservative for fatigue crack growth for the

particular case being considered.

Figure 71 compares the BS 7910 and FE KI values for the As-welded-PWHT residual stress case. As

well as the elastic FE values (denoted K), the elastic-plastic FE values (denoted KJ) have also been

included in the figure. It can be seen that there is practically no difference between the elastic and

elastic-plastic FE values. The BS 7910 values (based on a membrane stress of 69 MPa) are shown to

be higher than the FE values up to a crack depth of approximately 11 mm, after which the opposite is

true. Critical crack calculations have not been performed using the BS 7910 KI solutions for the

embedded crack cases. However, the information contained in Figures 70 and 71 suggests that the

critical crack sizes (for different fracture toughness value assumptions) for the as-welded-PWHT case

would be underestimated (compared to the FE calculated values) for crack depths up to approximately

8 mm. For higher crack depths, the critical crack sizes may be overestimated which is non-

conservative. The BS 7910 critical crack height curve could therefore be on the right of the

corresponding FE curve presented in Figure 52.

Figure 72 compares the BS 7910 and FE KI values for the As-welded-PWHT-Repair residual stress

case. This time the elastic-plastic FE values start to diverge from the elastic FE values at a crack

depth of approximately 8 mm. The BS 7910 values (based on a membrane stress of 345 MPa) are

shown to be higher than the FE elastic values by margins greater than 20 MPaÖm. The information

contained in Figures 70 and 72 suggests that the critical crack sizes (for different fracture toughness

value assumptions) for the As-welded-PWHT-Repair case would be underestimated (compared to the

FE calculated values) for all crack depths considered, with the possible exception of the largest cracks

(i.e. possibly overestimated for 2a = 12 mm to 14 mm say). The BS 7910 critical crack height curve

would therefore likely be on the left (for crack sizes up to approximately 12 mm) of the corresponding

FE curve presented in Figure 52.

37

11. TASK 10 – PROVISIONAL GUIDANCE ON WELD REPAIRS

Under this Task, provisional guidance on weld repairs has been developed. The resulting guidance

firstly focuses on practical issues that have been highlighted from the review carried out under Task 1

and from other relevant information. Secondly, guidance resulting from the finite element

calculations of the matrix of cases considered is presented.

11.1 PRACTICAL ISSUES

Types of Defect

Weld defects can be categorised as follows:

· Category 1 - cracks or crack-like defects (planar discontinuities) such as incomplete fusion or

penetration.

· Category 2 - geometric defects, including undercut, misalignment and incorrect weld profile.

· Category 3 - slag and porosity, they are the most innocuous type of weld defect but the most

frequently repaired

Causes of Defects

During fabrication, defects may arise due to problems with the fabrication procedure which must be

addressed quickly. Poor joint design and weld misalignment are frequent causes of fabrication

defects. During service, the cause and nature of the cracking must be investigated and action must be

taken to prevent recurrence. Otherwise, the same type of defect may develop. Incorrect joint design

and unforeseen service conditions are commonly cited causes of defect development.

Significance of Defects

Many repairs can be unnecessary and sometimes reduce the integrity of the structure. This is because

there is a gap between the high integrity resulting from the weld quality associated with workmanship

standards and the often much lower level of quality required to satisfy a fitness-for purpose

assessment.

Necessity of Repair Welding

Repair welding is not always necessary, particularly for shallow defects that may be removed by an

alternative method, for example by grinding out.

Problems Associated with Repair Welding

There are several problems associated with repair welding which may lead to a reduction of the

structural integrity of a weld compared with its original defective condition. These include:

· inadequate removal of the original defect; for example incomplete excavation of a crack,

· introduction of new defects; hydrogen cracking is a likely source of new defects in repair welds,

38

· unfavourable site conditions for re-welding including: poor access, the inability to apply

sufficient preheat and poor weld positioning,

· unfavourable conditions for inspection and testing of the repaired weld,

· repairs of restrained welds in structures may have a higher risk increased residual stress or

distortion,

· post-weld heat treatment of a site repair is often unfeasible, giving a higher probability of

increased residual stress and lower toughness.

In this regard, repairing a weld can lead to inferior or inappropriate properties of the repair due to a

degraded microstructure. Poor fracture toughness is of particular concern. Inappropriate properties

can arise from a number of different reasons including:

- inappropriate welding consumables,

- insufficient pre-heat,

- inappropriate (usually too low) heat input since a different welding process may be used in the

repair compared with the original fabrication site (repair welds are invariably made manually but

the original fabrication weld may have been an automatic process),

- inappropriate composition and weldability of the local parent material if a repair is being carried

out in a previously unwelded region of a component.

In terms of fracture toughness, there is strong evidence to suggest that too high or too low heat input

can have a deleterious effect on toughness. Reductions of CTOD toughness in Heat Affected Zones

(HAZ) by a factor of between 5 and 8 have been reported for carbon and carbon-manganese steel

welds (corresponding to reductions in KJ toughness by factors of approximately 2.2 and 2.8). On the

other hand, high preheat and inter-pass temperatures and post-weld heat treatment help to increase

HAZ toughness. These latter improvements are unlikely to be obtained in difficult on-site working

conditions however. For carbon-manganese steel welds, lower bound CTOD toughness of the HAZ

can be increased by PWHT by a factor of 10 compared with as-welded toughness. Mean CTOD

toughness of the HAZ can increase by a factor of 5. It should be noted though that in some modern

thermo mechanically rolled or TMCP steel, HAZ toughness may not improve or actually reduce due

to PWHT. To summarise, In standard carbon and carbon-manganese steels PWHT obviously

improves the toughness of HAZ in particular. This, combined with the benefit derived from using

proper pre-heat, correct heat input, a satisfactory welding position and other factors, must mean that

on-site repairs are likely to be produce lower values of fracture toughness than original shop welds.

11.2 GUIDANCE RESULTING FROM THE FINITE ELEMENT CALCULATIONS

In this study of crack-like defects in welds, comparisons have been made between the behaviour of

various sizes of defects in a heat treated weld and in a non-heat treated repaired weld. The main

parameters considered are those that are quantifiable and can be used within a fracture mechanics

framework. These are: weld residual stress pattern; defect depth or height and local fracture

toughness. The initial defects in the two weld cases are assumed to grow by fatigue due to a cyclic

operating load. Occasional overloads at each resulting defect depth or height are assessed to see

whether this causes a limiting condition for given values of fracture toughness. Of course, it seems

rather pessimistic to consider that a defect always remains after repair, and it is difficult to place

39

limits on its possible size. However, limits of sizing in ultrasonic testing provide a useful lower bound

to the repair defect position and height.

The main findings are as follows:

· For a given fracture toughness the critical defect depth at the limiting condition is smaller in the

as-repaired weld than the PWHT weld. As the fracture toughness reduces the relative difference

in critical defect size between the two cases becomes more significant. This phenomenon is

particularly apparent for the welded sphere studied.

· Predictions of fatigue crack growth in both the plate and the sphere show more rapid rates of

crack growth in the repair compared with the original PWHT weld. Fewer loading cycles are

required in the repair to cause a limiting condition for the same initial defect size and fracture

toughness.

· A graphical framework is presented to compare fatigue lives of defects in original heat treated

welds with the same or shallower defects in as-repaired welds (Figures 42, 43, 48, 49, 54, 55 and

56). From this information, it can be concluded that:

- Weld repairing shallow defects and low toughness parent or weld materials is more likely to

give a shorter fatigue life than leaving the weld un-repaired.

- A relatively shorter fatigue life of the weld repair compared with leaving un-repaired is more

likely in the sphere than the plate. The welded sphere geometry is more sensitive to reductions

in fracture toughness in both PWHT and conditions.

- For embedded defects in double ‘V’ preparation butt welds, the case for whether to repair or

not depends on the depth and height of the defect.

- Short embedded defects near the middle of the plate are likely to experience low or

compressive levels of compressive stress and so low crack driving forces and relatively long

fatigue lives. The best course of action is likely to be to leave these defects in place.

40

12. TASK 11 – PROVISIONAL GUIDANCE ON ENGINEERING

PROCEDURE METHOD

Section 10 above has reported on Task 9 of the project dealing with assessment by engineering

methods. This task has focused on the BS 7910 methodology. From the experience gained in

undertaking Task 9, provisional recommendations are now made on the use of this methodology to

firstly assess the significance of flaws in weldments (as-welded, PWHT or repaired weld) on a case-

by-case basis, and secondly, to assess as to whether repairing a weld is likely to be beneficial or not.

These provisional recommendations are derived from the work relating to the edge-cracked plate

configuration considered in this study. However, many of the aspects given below may be applied in

a generic sense.

12.1 ROUTE FOR ASSESSING THE SIGNIFICANCE OF A FLAW IN A WELD (AS-

WELDED, PWHT OR REPAIRED WELD)

The significance of a flaw known or postulated to occur in a weldment (or indeed in parent material)

can be evaluated from the following three steps:

1. Evaluate the critical crack size.

2. Evaluate the operating time or cycles to grow (e.g. by fatigue) the flaw to the critical crack

size.

3. Use the information obtained from 2 to decide as to whether continued operation in the

current state is possible, what the future inspection frequency should be or plan for repair or

replacement.

12.2 ROUTE FOR ASSESSING WHETHER REPAIRING A WELD IS LIKELY TO BE

BENEFICIAL

Assessing whether repairing a weld is likely to be beneficial or not can be evaluated from the

following four routes:

1. Evaluate the critical crack sizes for the un-repaired weld and for the repaired weld states.

2. For the un-repaired weld state, evaluate the operating time or cycles for the known flaw to

grow (e.g. by fatigue) to the critical crack size.

3. For the repaired weld, evaluate the operating time or cycles for the maximum size of flaw

that could be missed by the relevant detection techniques to grow (e.g. by fatigue) to the

critical crack size.

4. Compare the results of 2 and 3 to conclude as to whether it is likely to be beneficial or

detrimental to go ahead with the weld repair.

12.3 CRITICAL CRACK SIZE EVALUATION

It is recommended that critical crack size should be evaluated by following the Level 2 procedures of

BS 7910 section 7 (Assessment for Fracture Resistance).

For the edge crack plate configuration under consideration here, the stress intensity factor (KI) and

reference stress (sref) for the primary Stress (membrane) should be evaluated from BS 7910 with

reference to Figures M.6 and M.10 of that BS.

41

For the secondary residual stresses, ideally, elastic FE solutions for KI of the appropriate residual

stress profile should be obtained (as for the FE calculations presented above). In practice though,

such solutions are usually not available and the time and effort required to produce them is likely to

be restrictive. The guidance given in BS7910 is therefore likely to produce conservative (i.e. under-

estimates) values of critical crack size (Figures 61 and 62) whereby residual stress for the various

conditions can be represented as follows:

As-welded - + 0.2% proof stress, Through-wall Bending

PWHT - 20% of 0.2% proof stress, Membrane

Weld Repair - 0.2% proof stress, Membrane

KI solutions for the membrane residual stress assumptions can be evaluated with reference to Figure

M.6 and M.10 BS 7910. KI solutions for the residual stress assumption of through-wall bending can

be evaluated with reference to Figure M.6 of the BS.

(Note: the refined calculations referred to in section 10.2.3 for evaluating KI from a simplified

distribution of the FE residual stress profiles were shown to produce only a marginal benefit from

those of the BS 7910 route as described above.)

It is important that the tensile and fracture toughness materials properties relevant to the appropriate

weld state are taken into account in the calculations. This is particularly true in calculations for

assessing whether repairing a weldment may be beneficial since the fracture toughness in the repaired

state may be significantly different from that in the un-repaired state due to different heat treatments

being evident between the two cases.

12.4 CRACK GROWTH EVALUATION

If fatigue crack growth is the relevant mechanism, then the procedures of BS 7910 section 8

(Assessment for Fatigue) should be employed. It may be noted that other likely crack growth

mechanisms are corrosion fatigue (covered to a certain extent in section 8 of BS7910), creep (covered

in section 9 of BS7910), stress corrosion cracking (mentioned in section 10 of BS7910 but essentially

a specialist topic) and creep/fatigue (combined procedures of sections 8 and 9 of BS7910 but

essentially a specialist topic).

In the simplified fatigue crack growth route considered here, the R ratio (defined in BS7910 as

minimum stress/maximum stress but in reality is minimum KI/maximum KI) has not been taken into

account and therefore only the primary stress needs to be considered in the fatigue crack growth

calculations. More accurate fatigue crack growth calculations can be obtained by taking the R ratio

into account in the Paris Law relationship as outlined in BS7910 together with threshold

considerations. The use of the R ratio would necessitate the residual stresses to be taken into account.

Furthermore, in line with R5 procedures (Reference 7), inelastic values of K (i.e. KJ), instead of

elastic values, should really be used for evaluating DK as was done in the calculations reported in

section 9 above.

For assessing whether repairing a weld is likely to be beneficial for a flawed component subjected to

fatigue loading, the route described in the relevant sub-section above, together with the considerations

described in the paragraph above, should ideally be pursued.

However, as has been shown in section 10.2.3 above, the use of equation 12, i.e.

42

385.1 dN repaired

(aL - ai )repaired (ai + aL )unrepaired

= 385.1 dN unrepaired (aL - ai )unrepaired (ai + aL )repaired

is a reasonable approximation to establish whether repairing a defect is beneficial or not in terms of

improved fatigue life. It may be noted that the verification of this equation has been undertaken

(Table 3) by considering initial defect size (ai) to be the same in both the un-repaired and repaired

states. However, there is no reason why the relationship should not be equally valid when aI for the

repaired state is different (usually smaller) than for the un-repaired state.

43

13. TASK 12 – RECOMMENDATIONS FOR FUTURE PHASES OF PROJECT

Recommendations for work to undertake in future phases of this project are given below:

1. Repair Length. This first phase of the project has focussed on a weld repair configuration

extending along the full width of the plate (i.e. the original wide plate used for the experimental

programme) and the geometry has been simplified to a two-dimensional situation (i.e. an edge-

cracked or through-cracked plate). In reality though, the length of a defective weld, and hence

the subsequent repair weld, is likely to be more limited. It is therefore important to study the

influence on repair length on the residual stress distribution, as well as depth. At the same time,

it is necessary to consider realistic semi-elliptical surface or elliptical embedded cracks as

opposed to the simpler edge or through-wall cracks considered here.

2. Defect Position in Weld. For simplicity, the flaws in this initial study have been assumed to

occur at the centre of the weld. In reality however, it is common for flaws to occur in the HAZ.

Further phases of the programme should therefore address HAZ flaws and their subsequent

repair.

3. Development of Defect. In this study fatigue crack growth has been simulated as a post-

processing operation; that is, gradual growth of the crack tip is not explicitly simulated in the FE

models. Other studies have concluded that crack driving forces can be reduced due to the build-

up of the plastic wake behind of the growing crack. This is an effect that has been examined

experimentally in Task 6. It could also be usefully explored in future numerical models.

4. Residual stress relief due to mechanical loading (either operation or proof testing) has not been

taken into account in the present study, but it could be incorporated in future developments.

5. Fatigue Simulation. The load-unload-reload sequence is likely to be predominantly elastic. That

is, due to strain hardening there is unlikely to be an effect of reverse plastic straining in real

defective structures. Thus, crack driving force ranges may be over-estimated in the calculation

of fatigue crack growth rates. This aspect could usefully be investigated further.

6. Other Sub-Critical Crack Growth Mechanisms. Alternative forms of crack growth could be

considered using the crack driving forces obtained in this work. Stress corrosion cracking is a

common sub-critical mechanism that can be strongly affected by welding residual stress. Given

material properties, it would be a relatively simple task to make alternative predictions of

operating life in un-repaired and repaired situations.

7. Effect of Weld Process Control on Fracture Toughness. Guidance is required on how the various

welding parameters (e.g. heat input) effect the material fracture toughness.

8. Residual Stress Distributions. Improvements need to be made on providing more realistic

residual stress distributions for as-welded, PWHT and repaired conditions.

44