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Overview of recent welding-related pipeline technology advances at TWI

Edited by Christoph Wiesner

Paper presented at International Symposium of Prof Masao Toyoda's retirement from Osaka University: 'From Welding and Fracture Mechanics to Pipeline Technology' - 29 June 2008.

Abstract

There is continuing emphasis on reducing the costs associated with new pipeline developments. Over half the world's undeveloped hydro carbon reserves are remote from potential users and long and large, up to 42 inch diameter,pipelines are required to transport the fuel to market. Further, new reserves are increasingly discovered offshore in deep water and the products recovered frequently contain corrosive components. Much work is therefore underway inoptimising pipeline fabrications, inspection and safety assessment and this paper provides an overview of relevant topics covered by TWI.

1 Introduction

The present paper summarises advances in pipeline technology development at TWI, focusing on developments of welding processes; non-destructive testing methods; and corrosion and structural integrity assessments of pipe welds.

TWI's welding and manufacturing support to the pipeline industry ranges from on-site support for qualification of fabrication processes to long-term strategic development of practical welding processes offering economic advantages,a topic which has been treated before by Toyoda (2001). TWI's recent main advances in welding process development have been related to the development of power beam processes for welding ofpipelines. The focus has been on the development of both high-power, fibre-delivered hybrid laser-arc and electron beam welding for girth welding of pipelines. However, other novel fabrication techniques are also being explored such asTIG welding with activated fluxes, underwater welding techniques, radial friction welding and other friction-based processes.

The focus of TWI's NDT activities for pipelines is to understand the inspection requirements of pipeline operators which are currently not being met and to provide novel solutions to meet these needs. TWI was the initiator of thedevelopment of guided wave technology for long-range inspection of pipelines. The technology is now in use worldwide and has been applied by many oil majors and pipeline operators for examination of both in-plant and transmissionpipelines.

Identification of pipe weld microstructures and their relationships to material and corrosion/stress corrosion properties and avoidance of fabrication cracking, with subsequent definition of appropriate welding procedures is anotherarea of interest to this sector. Pipeline integrity assurance is a further important issue for this sector. TWI pursues fundamental and applied pre-competitive research to solve problems of common interest and to improve methods ofassessing the integrity of pipelines. This paper provides examples of current welding-related pipeline technology developments at TWI.

2 Pipeline welding process developments

2.1 Fibre laser-arc hybrid processing for welding of pipelines

The welding process used to make on-site pipeline girth welds has a significant bearing on the total cost. Current practice is to use mechanised metal active gas (MAG) welding. However, this process, requires a high manning leveland the costs of providing this and the necessary support in remote regions are a significant component of the overall costs. Recent developments to improve productivity include the use of novel tandem MIG equipment and also a renewedinterest is spirally welded pipe. This section, based on the work by Howse et al (2005), focuses on the progress achieved using laser welding processes.

Laser welding, and in particular fibre-delivered laser welding, has now been developed to a stage where it presents opportunities for cost savings, which arise from reductions in labour content, despite perceived high capital costs.It has been demonstrated (Howse et al, 2002; Booth et al, 2002) that the concept of high-power laser welding of land pipelines is entirely feasible.Welding procedures have been developed that produce good quality welds with satisfactory tolerance to joint fit-up. Additionally, techniques have been developed for welding around 360° and for ensuring a good weld at thestart/stop weld overlap position.

Until recently, there have been two main types of industrial laser used at high powers for deep penetration keyhole welding. These were CO2 gas lasers and Neodymium-doped Yttrium, Aluminium Garnet (Nd:YAG)lasers. In terms of materials processing, the principal difference between Nd:YAG lasers and CO2 lasers is the difference in wavelength of the light emitted. Nd:YAG lasers produce light that can be transmitted tothe workpiece by a fibre optic cable. This is a much more flexible system of beam delivery than for CO2 lasers which require more cumbersome reflective or transmissive optical systems.

Nevertheless, Nd:YAG lasers have a significant drawback in that they are inefficient, only converting around 3% of the input energy to produce the laser beam power. This is not a major issue for some manufacturing applications, butthe use of lasers for welding cross-country pipelines relies on the portability of the process, and the Nd:YAG laser process is therefore very difficult to justify economically. One of the major advances in laser technology in recenttimes is the introduction of Ytterbium (Yb) fibre lasers. The lasing medium for these lasers is contained within a fibre and individual units generating 200-300W can be combined to produce single lasers with up to 10kW power andbeyond. The wavelength of light generated is similar to that from Nd:YAG lasers and hence fibre delivery of the energy is used.

TWI has recently added a 7kW Yb fibre system to its laser processing facilities. The laser is capable of delivering power up to 7kW via a 0.3mm diameter, optical fibre cable and has the capability of producing a focused spot with apower density of 5.6 x 106 W/cm2. It operates at a wavelength with similar good material interaction to Nd:YAG lasers. The laser is also relatively compact with length, width and heightdimensions of 0.8, 1.2 and 1.6m, respectively. This laser is a high-efficiency power source which is capable of being containerised and used for pipeline applications.

With this in mind, TWI carried out welding trials in API 5L X80 linepipe material. The work built on early studies which investigated autogenous Nd:YAG laser welding. Whilst welding was possible in wall thicknesses up to 12.7mm, thespeeds were slow. Higher speeds needed for the process to maintain productivity, were used but the autogenous process was not tolerant to variation in joint gap. In addition, the autogenous process produced welds with very low impacttoughness. In order to solve the problems with poor tolerance to fit-up at high speeds and the low toughness, hybrid Nd:YAG laser-MAG welding processes were developed as reported by both Howse et al(2002) and Booth et al (2002). The work showed that it was possible to use high-power Nd:YAG laser welding combined with the MAG process to produce deep penetration welding passes incommercially available pipeline steels that met the requirements of pipeline specification such as BS 4515 and API 1104 in terms of acceptance criteria for imperfection limits. The welds also showed acceptable hardness values and goodlow temperature toughness.

The fibre laser-arc hybrid welding capability has therefore been developed using the Yb fibre laser in combination with a programmable MAG arc power source. The processing heads used held a 250mm length focusing lens producing a0.6mm diameter focused spot and a power density of 2.5 x106 W/cm2. An example of a cross section of the weld is shown in Figure 1 (note: pipewall was bevelled to achieve desired wallthickness).

ΔK, where the cyclic frequency is varied in blocks to determine how frequency affects the fatigue crack growth rate, at a particular value of ΔK. By monitoring the crack growth rate over a relatively short crack increment, it is possible to determine da/dN for much lower frequencies than would be possibleusing conventional test techniques. Figure 13 shows a typical set of data for X65 pipeline steel tested in 3.5% NaCl at an applied cathodic protection potential of -1050mV. It can be seen that, as the frequency decreases, themeasured crack growth rate increases until a plateau is reached, where it appears a further decrease in frequency has no further effect. At ΔK=400N/mm3/2, the plateau occurs at approximately 0.1-1Hz. At a higher value of ΔK, the observed increase in growth rate was initially similar, but aplateau was not reached until a much lower frequency, at least as low as 0.001Hz.

√m for shallow-cracked specimen compared with maximum values of around 500MPa√m for deeply cracked specimens).The HAZ of the girth weldrepresentative of the third pipe test, showed instances of more discernible increases in fracture toughness with reduction in notch depth, the maximum measured values increasing from around 350MPavm to around 690MPa√m.

Two of the three full-scale pipe tests with 1.6 and 4mm deep flaws survived strains greater than 2.6% at temperatures down to 0°C without brittle fracture occurring. Ductile behaviour, in terms of the strains achieved andinitiation of ductile tearing, was demonstrated. This contrasts with the behaviour of the deeply notched SENB fracture toughness specimens which failed by cleavage at minimum K values of around 60MPa√m in the HAZ and 110MPa√m in the weld metal. The results from deeply notched fracture toughness specimens are thereforeunrepresentative of full-scale pipe with shallow notches. Two major factors contribute to this behaviour: differences in constraint between fracture toughness test specimen and pipe, and residual stresses. The effect of constraint (orstress triaxiality at the crack tip) is indicated by the fracture toughness tests on the shallow-notched bend tests which can show a significant increase in toughness compared with deeply notched specimens. However, one of the sixshallow-cracked fracture toughness specimens fractured at K values similar to the deeply notched specimens. The reason for this may be that the constraint levels in the shallow-cracked specimen is still higher than in the pipe which isexpected since the applied through-thickness stress distribution in the pipe is essentially tensile, whilst the fracture toughness test is specimen tested in bending.

The second factor is the beneficial role played by welding residual stresses in the full-scale pipe tests. Numerical analysis was used to predict the transverse residual stresses through the thickness of the RPEB girth weld. For thepipe sizes, strength grade and welding conditions employed, compressive residual stresses were predicted at both the pipe internal and external surface. These are balanced by tensile residual stresses within the central two thirds ofpipe wall thickness. The brittle behaviour of the second full-scale pipe test is consistent with the above hypotheses. The pipe contained a deep notch/crack almost half way through the pipe wall, so fracture toughness determined usingdeeply notched specimens would be appropriate. Furthermore, most of the crack and especially the deepest part of the crack tip region was located in a tensile residual stress field.

The full-scale pipe bend tests and fracture mechanics tests have shown that contrary to the expectations from conventional, deeply notched fracture toughness bend specimens, girth welds made by the RPEB process in X65 pipe canbehave in a ductile manner at temperatures down to 0°C when shallow surface flaws (<4mm deep) are present. The reasons for this are attributed to the beneficial effects of low constraint in the pipe (which is loadedpredominantly in tension) containing shallow circumferential flaws, compared with standard (SENB) fracture toughness tests, and compressive transverse residual stresses near to the surface. These benefits are lost if the girth weldscontain deeper flaws because of the increase in constraint and presence tensile residual stresses within the central two thirds of pipe wall thickness. However, the on-line seam tracking system, which is an integral part of the RPEBprocess, should avoid the creation of large welding flaws. If small flaws should be present, these should be readily detected by non-destructive testing. Such flaws (provided they are less than 4mm high) do not therefore present aparticular risk of fracture for strains expected during pipelay.

5.2 Evaluation of weld metal strength mismatch in X100 pipeline girth welds

Use of grade X100 strength pipeline steels for transmitting gas in harsh and environmentally sensitive regions, poses particular challenges for weld metal design. The usual means of increasing strength is through increasing alloylevels. But as strength levels increase, weld metal properties become more sensitive to cooling rate variations than for lower strength weld metals. Furthermore, alloying tends to increase the risk of weld metal hydrogen cracking andtends to limit the fracture toughness achievable. Welding consumable development therefore requires parallel development of corresponding welding processes and procedures but these tend to be to less user-friendly, necessitating morecare in implementation on the part of fabricators to ensure the necessary weld properties and quality are achieved.

In practice, it is difficult to achieve the desired balance of weld metal strength and toughness, particularly without reducing weldability and usable procedure ranges in practice. Consequently, it is important to take a moreholistic view of specification requirements in order to contain costs and minimize risks while ensuring that failure will not occur under envisaged loading conditions. For example, it may be more cost effective to re-design weld jointgeometries and welding processes to ensure higher levels of weld quality than to increase fracture toughness requirements. However, in order to avoid unnecessary delays and to maintain an economic welding process, it is unrealistic torequire repair of every flaw indication reported by non-destructive testing.

This section, based on the work by Pisarski et al, (2004), therefore gives an example of how fracture mechanics can be used to examine the effects of such competing demands on girthweld performance. Specifically, it examines the effects of strength mismatch on fracture resistance of girth welds in a X100 pipe. A number of papers have been published to assess these effects, often based on advanced finite elementanalysis; although it may be difficult for practioners to assess the effect of the changes in key input parameters on the outcome without further analyses (Wang and Horsey, 2004). Analternative is to evaluate effects by experiment using large-scale tests such as curved wide plates (Denys et al, 2004), but in practice, only a limited set of parameters can be tested andmaterial/test variability can disguise important trends. The analytical method described here is based on generally recognised procedures and evaluates the effect on fracture toughness requirements for girth welds in X100 strengthpipeline of weld metal strength under and overmatching and the interplay with weld width for different applied axial stress regimes.

The assessments are based on the methods described in BS 7910, modified for strength mismatch effects in accordance with the R6 procedure. The main modification is the introduction of an 'equivalent' stress-strain curve to generatethe failure assessment diagram (FAD) and to define an equivalent yield strength. The equivalent stress-strain curve is derived from a weighted average of the weld metal and parent pipe stress-strain curves. The weighting is provided bythe ratio of the mismatch limit load to the limit load for homogeneous material.

The mismatch ratio M is defined as Mx = σYw/σYb, where σYw or σYb areflowstresses corresponding to the same amount of plastic strain of the weld metal and base metal, respectively. The sub script x defines the plastic strain at which mismatch is defined. Since the weld metal and base metal will, ingeneral, work harden differently, the degree of mismatch will vary with plastic strain. Mismatch limit loads are given in R6 for fully circumferential internal cracks in cylinders. This geometry was considered representative of a longsurface flaw in a pipeline girth weld. The mismatch limit load solutions consider parallel sided welds and are for weld metal centreline flaws and flaws located at the interface (in this paper, the interface is considered to berepresentative of the HAZ/fusion line). The choice of limit load depends on the expected deformation pattern as described by Pisarski et al, (2004).

Figure 18 shows how the allowable axial stresses (represented as the ratio of allowable stress, Pm, to the specified minimum yield strength (SMYS) of X100 strength pipe changes with mismatch level indifferent weld widths. The assessments were conducted assuming that a minimum CTOD of 0.25mm at the minimum design temperature is achieved. The first analysis was conducted assuming that the flaw is in homogeneous material (no strengthmismatch) with tensile properties equal to the lowest strength present. This is the recommended procedure when mismatch limit load solutions are not available. As can be seen, this provides, as intended, a conservative estimate ofallowable axial stress. If mismatch and weld width effects are specifically considered, benefits to the allowable stress are realised as the weld width is reduced from 20 to 5mm. Although the allowable stress increases for undermatchedwelds (M<1), it always remains below that for matched and overmatched welds. When overmatching is present and the flaw is located in weld metal, increasing the weld width increases the allowable stresses. However, when thedifference in weld metal and parent pipe strength is small (M ≈ 1), the best strategy is to design a narrow weld. For narrow welds, if slight undermatching occurs,the reduction in allowable stresses minimised, and there is still a benefit in allowable stress if overmatching occurs. Further analyses show that for a given level of undermatching, HAZ flaws are more sensitive to weld width effectsthan weld metal flaws. Nevertheless, there is a clear benefit of using narrow welds if undermatching is likely.

spcswjune08f18.gif
Fig.18. Ratio of allowable axial stress to SMYS of X100 pipe for 3mm deep circumferential weld metal flaws at different levels of strength mismatch and weld

Figure 19 shows how the CTOD requirement varies with applied axial stress for different weld widths and mismatch levels. Also included are analyses for homogeneous material. These provide bounds to the analyses andillustrate the benefits of allowing for mismatch effects. The figure shows that CTOD requirements increase rapidly when the applied stress approaches yield strength magnitude. A limit is reached when further increases in applied stresscannot be accommodated by further increases in fracture toughness. This condition is obtained when failure is predicted to occur by plastic collapse rather than fracture. In practice, crack growth by stable ductile tearing will occurbefore plastic collapse and this will affect the analyses at higher CTOD values (say above about 0.3mm when initiation by ductile tearing is likely to have taken place).

spcswjune08f19.gif
Fig.19. Relation between CTOB requirements and ratio of axial stress (Pm) to Grade X100 SMYS for HAZ flaws (3mm high0 with different mismatch levels (M) and weld widths (2H)

The analyses show the clear benefit of even small amounts of weld strength overmatch. For the weld metal flaw case considered, axial stresses above the pipe SMYS are acceptable with a 3mm deep weld metal flaw present provided thatCTOD is greater than about 0.1mm. The results also show how fracture toughness requirements need to be increased if undermatching is present to achieve the same flaw tolerance. However, the increase would not be as great as predictedby an analysis based on the least strong material. For example, Figure 19 shows that for M = 0.83, an analysis based on the GMAW weld metal would require a CTOD of 0.25mm for an axial stress 85% of SMYS. By specificallyanalysing for the undermatched condition (M = 0.83), the CTOD requirement can be significantly reduced, to about 0.05mm.

It should be noted that if the HAZ strength undermatches the strength of both weld metal and parent pipe, the likelihood of cleavage can be increased. This can result in very low fracture toughness values being measured in the HAZ(Pisarski and Harrison, 2002). In addition, there will be effects of internal pressure acting on top of axial loading, because of effects on constraint. Finally, residual stresses need tobe considered in deriving practically applicable fracture toughness requirements. Nevertheless, the relatively simple stress-based procedures described here are useful to illustrate the effects of mismatch and weld width on fracturetoughness requirements to avoid fracture in girth welds.

In conclusion, a simple method based on recognised flaw assessment procedures has been used to illustrate the effects of weld width and strength mismatch on the CTOD requirements for girth welds in X100 strength pipeline materialsubjected to axial stresses. It has been shown that CTOD requirements derived from analyses based on homogeneous material (i.e. making no allowance for mismatch) are conservative and can be reduced if mismatch effects arespecifically considered.

Analyses show that the adverse effects of weld strength undermatching can be mitigated by reducing weld width and this can be demonstrated by specifically considering mismatch in the calculations. The results also show the benefitsof even small amounts of weld metal overmatching on the maximum tolerable axial stress. The results illustrate the use of fracture mechanics to examine the effects of the competing effect of weld metal strength and toughness on girthweld performance.

5.3 Development of flaw assessment procedures for girth welds subjected to plastic straining

The definition of rational flaw acceptance criteria for girth welds in pipelines subjected to axial straining in the context of codified assessment procedures is problematic since these are essentially stress-based. Although thereis no fundamental problem in using such procedures, the solutions provided are not always the most suitable for strain-based assessments. Nevertheless, with appropriate modifications, assessments based on BS 7910 procedures have beenused successfully for a number of years to set acceptance criteria for pipeline installation methods involving plastic straining. The flaw acceptance criteria provided by these methods have, in many cases, enabled larger flaws to beacceptable than the limits in workmanship standards.

Another benefit to industry of using fracture mechanics assessments is that flaw size information provided by automated ultrasonic testing can be assessed properly, since workmanship acceptance criteria are based on flaw length, notheight. Nevertheless, for certain situations involving plastic straining, flaw sizes predicted by fracture mechanics procedures can be smaller than those based on workmanship standards. Although it could be argued that there is nothinginherently wrong with such a conclusion, because the workmanship criteria are intended for installation methods not involving plastic straining, industry experience indicates that flaw tolerance is better than predicted by fracturemechanics analyses.

A previous joint industry programme, conducted by DNV, TWI and SINTEF, developed guidelines for fracture control of pipeline installation methods involving plastic straining. The work provided the basis for the published RecommendedPractice DNV-RP-F108. The flaw assessment procedure developed, referred to hereafter as the reeling procedure, is based on BS 7910 but with adjustments to make it suitable for plastic straining conditions. Novel features of theprocedure include the use of single edge notched tension (SENT) specimens to define fracture toughness and the use of sub-scale segment specimens to validate, by experiment, the procedure used to generate acceptable flaw sizes. Theprocedure has been used successfully in pipeline installation projects and in-service assessment of pipelines subjected to plastic straining, e.g. due to lateral bucking and ground movement.

The purpose of this section, based on the work by Pisarski and Cheaitani, (2007), is to establish margins against possible failure when using the procedure as currently formulated inDNV-RP-F108. This is done by comparing the J driving force curves predicted this procedure. with those obtained by numerical finite element analyses (FE).

The pipe analysed has an outside diameter of 12¾in and wall thickness of 20.6mm and is nominally to API 5L Grade X65 strength. Yield and tensile strength of the parent pipe is 485MPa and 594MPa, respectively. The pipe containsa girth weld which has a yield and tensile strengths which overmatch the corresponding parent pipe properties by approximately 20%. The work hardening rates of both materials are assumed to be the same. Pre-existing internal surfaceflaws were considered in the finite element analyses, located at the weld root region of the weld fusion boundary. They were semi-elliptical in shape with the following depths and lengths: 3x50, 6x25, 6x50 and 9x90mm. The nominal weldwidth for this location was 8mm, representative of the narrow weld preparations typically employed for these applications. The weld profile was treated as parallel-sided. Two series of analyses were undertaken: the first assumed ahomogeneous material characterized by the parent pipe tensile properties, the second included the girth weld with a higher strength so that the effects of strength overmatching were modelled. The straining of the pipe when it is bentonto the reel was modelled by applying a bending moment to the pipe model which contains the girth weld and flaw.

Assessment of these conditions were also carried out according to DNV-RP-F108 procedures to generate crack driving force curves as a function of applied strain. For each of the flaws considered, two crack driving force curves wereobtained (the J-integral is used as the crack driving force in this section). In order to make a direct comparison with the results from the numerical analyses, the first curve was estimated with no allowance for welding residualstresses. The second curve was obtained assuming that initial welding residual stresses of yield magnitude exist in the region containing the flaw. The initial residual stress was allowed to relax, as the applied stress increased,according to the rule recommended in BS 7910.

Figure 20 shows one example of the relationship between the J driving force and remote applied strain predicted by numerical analysis (labelled as 'J (FE)') and according to the reeling procedure for the 6 x 50mm flaw.Results are shown from numerical analyses obtained on homogenous material and on models where the tensile properties of the weld metal overmatch those of the parent pipe by approximately 20%. Both sets of numerical analyses wereconducted without including welding residual stresses. Two sets of results based on the reeling procedure, obtained with and without including welding residual stresses as specified above, are also shown (both assume homogeneousmaterial).

sphgpjuly07f4.gif
Fig.20. 6 x 50mm surface flaw at weld root fusion boundary. Homogeneous parent pipe properties were assumed in the reeling procedure (DNV-RP-F108) and both homogeneous and weld strength mismatched materials wereemployed in the numerical (FE) analyses

For homogeneous material without including welding residual stresses, it can be seen that the reeling procedure underestimates the J driving force for applied strains greater than about 0.75%. For strains less than around 0.75%,there is good agreement in predicted J values for the two analysis methods (except for the 9 x 90mm flaw size, not shown, where the driving force calculated using the reeling procedure does exceed the FE results at all strain levels.Generally, the reeling procedure is used to justify acceptance of relatively small flaws, certainly less than half way through the wall thickness, therefore, the non-conservative nature of the reeling analyses is of concern. However,there are mitigating factors that reduce the possibility of obtaining non-conservative assessments.

One of the requirements of the reeling procedure is that the weld metal strength must overmatch that of the parent pipe. As discussed in section 5.2, overmatching 'protects' flaws located in both the weld metal and at the fusionboundary. This 'protection' can be observed in Figure 20, which shows that the predicted driving force for plastic straining conditions is reduced significantly and below that predicted by the reeling procedure. If the weldwidth were increased, the J driving force would further decrease. As already seen in Section 5.2, weld metal strength under matching will significantly increase crack driving force and is therefore not currently permitted by thereeling guidelines.

Another requirement of the reeling procedure is to include effects of welding residual stresses. For a conventional assessment, an initial welding residual stress of yield magnitude is assumed to exist in the region containing theflaw. As the applied stress is increased, the residual stress is allowed to relax down to 40% of the yield strength of the parent metal. However, a lower level of residual stress is likely to be present in the actual weld at appliedstrains above yield. The residual stress will increase the driving force for the reeling analysis as illustrated in Figure 20. The reeling analysis curve including the effects of welding residual stress lies well above allother curves including the numerical analysis curves obtained assuming homogeneous material and the reeling analysis curves.

A third factor, which is more difficult to quantify, concerns the effect of the material's resistance to fracture. The reeling procedure recommends that single edge notch tension (SENT) specimens are employed to determine fracturetoughness. These specimens are designed to be more representative of the crack tip constraint conditions of flaws in the pipe girth weld than conventional specimens. However, analyses which have compared crack tip constraint between aSENT specimen and a pipe girth weld flaw (Nyhus et al, 2005; Pisarski and Wignall, 2002) indicate that crack tip constraint in the specimenremains more severe than in the pipe for the range of conditions examined. Consequently, the design of the SENT specimens can provide a degree of conservatism by providing a pessimistic value of fracture toughness.

These three factors mean that the apparent non-conservatism in the reeling analyses at strains above 0.75% is reduced. Similar conclusions with regard to the effects of transferability of fracture toughness parameters for theassessment of mismatched welds were reached by Toyoda, (2002). The requirements for weld strength overmatching and the inclusion of welding residual stresses as a secondary stress in theanalysis ensured that the J driving force estimated according to the reeling procedure was above the driving force estimated by numerical analysis. This conclusion would appear to be supported by satisfactory field experience whenthese methods have been used to set flaw acceptance criteria for girth welds.

Despite the comments made above, the assessment procedure should provide a more accurate modelling of the appropriate driving force curve derived from numerical analysis. There are a number of possibilities including: (i) modifyingthe reference stress solution and (ii) the use of more refined and potentially more accurate models to allow for the effects of welding residual stresses. For (i), a promising approach currently under development is to use J-basedreference stress solutions.

Other current research includes numerical analyses leading to the derivation of crack driving force curves vs. applied strain for both surface and embedded flaws. These analyses include the use of large-strain large-displacementformulation. Findings from these activities will enable the production of improved guidance on assessing the significance of girth weld flaws subjected to significant plastic straining, which will address the weaknesses and limitationsof existing guidance including the reeling procedure considered above.

A further area of current work is the effect of internal pressure. If straining occurs during service, the overall loading condition can be significantly more severe and the tolerances to flaws lower than during installation becausethe pipeline will be pressurised. Current procedures for assessing these conditions are either inadequate or not properly validated. Ongoing work on the effects of the combination of axial straining plus internal pressure on crackdriving force indicates that the crack driving force under biaxial loading conditions can be significantly higher than under axial straining alone.

In conclusion, analyses have been undertaken of girth weld flaws subjected to plastic straining simulating installation by reeling and J driving force curve have been derived. It can be concluded that for 'even-matched' strengthconditions and when welding residual stress is ignored, the J driving force predicted is higher than that predicted by the reeling procedure for strains exceeding 0.75% for most flaws. When a weld metal with a yield strength 20% higherthan the parent pipe is included in the numerical analysis model, the driving force is reduced relative to the even matched strength condition and below that predicted by the reeling procedure for strains up to at least 2%. Byincluding welding residual stress, in accordance with BS7910, in the assessments based on the reeling procedure, the resulting driving force curves lie well above all other curves including the numerical analysis curves obtainedassuming homogeneous material.

6 Overall conclusions

This paper has summarised recent work related to welding-related pipeline technologies.

In the area of pipeline welding and joining process development, recent advances in the use of power beam processes such as hybrid laser-arc and electron beam welding have been described and their potential for improved productivityhas been demonstrated.

With respect to inspection and non-destructive testing, a significant innovation related to pipeline technology is the development of long-range ultrasonic testing. Much progress has been made in developing this technique into apowerful flaw detection and screening tool. Recent advances have concentrated on novel solution to further enhance the capability of the technology to permit flaw sizing.

In relation to corrosion management, much work has been carried out to better understand the effect of corrosive components in the pipeline media. This paper has summarised the finding, particularly their effects on fatigueperformance of pipelines and risers. Mitigation measures such as the use of thermal-sprayed aluminium coatings have also been studied.

Finally, examples of structural integrity assessment work related to pipelines been the support of evaluating the fracture performance of novel high-productivity pipe girth welding processes; the evaluation of mismatch effects toassist in the design of weld metal for high-strength (X100-type) steel developments, and the improvements of flaw assessment procedures for pipelines subjected to plastic straining.

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