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Recent Welding-Related Pipeline Technology Advances at TWI


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.


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 in optimising 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 of pipelines. 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 as TIG 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 the development 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 transmission pipelines.

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 another area 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 of assessing 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 level and 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 renewed interest 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 the start/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 to the 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, but the 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 recent times 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 and beyond. 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 a power 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 height dimensions 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, the speeds 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 impact toughness. 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 in commercially 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 good low 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 a 0.6mm 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 wall thickness).

Fig.1. Macrosection of Yb fibre laser-MAG arc hybrid weld in X80 line pipe steel made at 1.8m/min

Fig.1. Macrosection of Yb fibre laser-MAG arc hybrid weld in X80 line pipe steel made at 1.8m/min

The welds were inspected both visually and by radiography and were found to be of good quality with no internal defects. Samples were also taken from these welds and used for hardness cross-weld tensile and Charpy impact testing. The samples produced acceptable toughness and tensile failure occurred in the parent material in all cases. However, the hardness of these welds is relatively high and the use of this technology for corrosion-critical applications would need further consideration.

The reliability of the process in industrial applications requires thorough evaluation. Although the laser at TWI appears to be a relatively robust design and has delivered very high availability for processing, this is the first commercial system of this size under evaluation and more data will be needed before industry is fully convinced of its capabilities. Similarly, work needs to be carried out to fully develop an integrated pipe welding system capable of being used in a production environment.

From the development carried out, the following conclusions can be drawn:

  1. Over the past ten years laser technology has developed to a stage where reliable, high-power and efficient Yb fibre laser power sources are now available capable of offering productivity benefits for pipelay.
  2. The hybrid Yb fibre laser-MAG arc process can be used at high processing speed (1.8m/min) to weld a 9mm ligament in a single welding pass with low and acceptable flaw levels.
  3. Welded deposits have been tested for strength and the impact toughness and have been shown to be acceptable to current pipeline application standards.
  4. The developed procedure has been demonstrated in the overhead, vertical up and flat welding positions capable of being translated to a 5G vertical up girth welding procedure in linepipe.

2.2 EB welding of thick tubular components and pipelines without a vacuum chamber

Electron beam (EB) welding offers many advantages for thick-section fabrication, particularly when applied to large structures, where significant savings are anticipated because of the rapid joining rate achievable. Examples of this include the use of EB welding for the future fabrication of structures such as monopile foundations for offshore wind-turbines fabricated from rolled steel with thicknesses in excess of 80mm. Appropriate application of EB welding in a single pass is anticipated to lead to cost and time saving in excess of 50% compared with more conventional fabrication practice. To date, however, the full potential of the EB process has not been realised commercially for thick-section welding and large structures because of restrictions associated with working at a high-vacuum, with the entire structure to be welded enclosed in a vacuum envelope.

TWI has demonstrated that operating the EB process in a reduced pressure range of 0.1-10mbar, rather than in high-vacuum (~10-3 mbar), offers the possibility of eliminating the need for a vacuum chamber by permitting the practical use of local sealing and pumping on a large structure. In adopting the Reduced Pressure Electron Beam (RPEB) process variant, problems of achieving adequate sealing on the component are much reduced and the effect of weld pool emissions and out-gassing of the component on the gun performance are eliminated. So far, RPEB welding has only been applied industrially in a few specific cases. It is envisaged that many more applications of the process can be promoted by further development of practical local sealing devices.

In addition to the RPEB process, another system has been developed at TWI which allows transmission of high-power beams into air (i.e. without any vacuum requirements). This system is capable of welding steel and copper of thicknesses greater than 25mm at speeds approaching 2m/min in a single pass and has recently been configured to permit pulsed operation.

The following further describes the development of local vacuum systems for field deployment of RPEB welding and the optimisation of the Non-Vacuum (NV) EB process, illustrating the potential for using both methods in cost-effective fabrication of large-diameter, thick-section pipelines. The section is based on work by Punshon and Sanderson (2007).

TWI first developed its RPEB technology in the 1990s which allowed operation of a high-power electron gun with the work piece at a reduced pressure in the range 0.1-10mbar. In this system, the electron gun electrode geometry was carefully designed to maintain a vacuum pressure in the gun electrode enclosure of ~10-6mbar whilst the beam was delivered into a reduced pressure of ~1mbar. Where the beam exits the column, an overpressure helium gas feed can be used as an option which reduces scattering of the beam and provides a background welding atmosphere of helium which assists in prevention of weld pool oxidation. With this development came the possibility of working either with big chambers pumped to a coarse vacuum pressure, thus minimising the pump-down time, system cost and operating sensitivity, or with local seals and pumping applied to weld joints on components too large to be contained entirely in a vacuum envelope.

This concept was tested and demonstrated successfully in the laboratory at TWI for application to offshore pipelay of large-diameter, thick-walled pipeline (Figure 2) and illustrated that welds could be made in this pressure regime with consistent high quality and significantly improved process tolerance compared with conventional high-vacuum EB welding. In particular, for a working distance range of 50-500mm, welding performance was shown to be independent of working distance (Figure 3). A target pressure of 1mbar was selected as the best compromise for reduced pressure operation in terms of simplification of vacuum engineering and reliable welding performance. An important consideration of one-shot welding process development is toughness which will be covered in Section 5.1.

Fig.2. Local vacuum Reduced Pressure Electron Beam welding system manufactured for offshore pipelay

Fig.2. Local vacuum Reduced Pressure Electron Beam welding system manufactured for offshore pipelay

Fig.3. Transverse sections of RPEB welds in X65 steel pipe of 10.75in diameter by 25mm and 28in diameter by 41mm wall thickness made with gun to work distance of 270mm and 50mm, respectively

Fig.3. Transverse sections of RPEB welds in X65 steel pipe of 10.75in diameter by 25mm and 28in diameter by 41mm wall thickness made with gun to work distance of 270mm and 50mm, respectively

Previous work has illustrated that almost since the first industrial use of electron beam welding there has been a desire to apply the process to large components using local pumping and sealing. With the exception of the TWI/Saipem work (Punshon et al, 1998), all other attempts to apply EB welding using local vacuum have involved systems designed to operate at vacuum pressures lower than 5x10-2 mbar. The attempts to work at high vacuum, although reasonably successful in the short term, were eventually derailed by inconsistent sealing and pumping performance.

The ability to work at reduced pressure greatly improves the potential reliability of local seals and pumping as the need for high levels of cleanliness and sophisticated pumping and sealing technology are eliminated. The system developed employs steel brushes as the primary seal (Figure 4). With two differential pumping stages, a pressure level of less than 1mbar can be achieved reliably on plate with a typical hot rolled surface finish. With this arrangement, it was established that a pressure of less than 1mbar could be achieved reliably in less than 10s pumping time and could be maintained whilst traversing a steel plate with a weld bead on the surface.

Fig.4. Reduced pressure EB local vacuum head:

Fig.4. Reduced pressure EB local vacuum head:

a) concentric brush seals;

b) operating on 45mm thick plate

b) operating on 45mm thick plate

The use of high-power EB with a local, mobile vacuum head, opens a number of industrial applications which hitherto would have required the construction of very large vacuum chambers. In addition, the field welding of large pipeline structures is made possible. This can be achieved either by means of local mobile seals or by use of a locally pumped vacuum chamber and sliding seals see Figure 5. In both cases operation is made simpler, cost effective and reliable by operation at reduced pressure.

Fig.5. Schematic of potential applications for RPEB welding of pipeline structures:

Fig.5. Schematic of potential applications for RPEB welding of pipeline structures:

a) local, mobile vacuum seal;

b) locally sealed vacuum chamber

b) locally sealed vacuum chamber

Early attempts to use NVEB welding, for thick materials were held back by the impractical requirement for very short stand-off distances (~10mm) and reduced weld quality due to the greater degree of beam scattering that occurs at atmospheric pressure. However, recent developments have significantly improved this situation (Sanderson, 2007).

The degree of scattering can be reduced and the working distance range extended by employing higher accelerating voltages. However, this increases the size, weight and bulk of the equipment, making it less favourable for mounting the gun on a robot. One promising line of research is to pulse the electron beam, so that the plasma and gas level in the vicinity of the weld are allowed to decay when the beam is off, but with a sufficiently high duty cycle and power level that the weld pool does not collapse or solidify. This approach increases substantially the penetration performance. It has been found that the penetration range can be increased for a given average power level by using high peak power levels. Sanderson, (2007) has shown that penetration depth in steel can be increased by 50% combined with a corresponding reduction in weld width, particularly in the vicinity of the top bead. Pulsing also appears to alter the solidification mechanism that often leads to solidification cracks in deep NVEB welds.

Figure 6 shows a melt run made at an average power of 36kV and a welding speed of 480mm/min in the flat position in low alloy steel using a pulsed non-vacuum electron beam. It will be noted that the fusion zone is almost parallel sided with a well-rounded tip. Apart from minor pores, the fusion zone was sound.

Fig.6. Flat position, 22mm deep melt run made in low alloy steel with a pulsed NVEB beam

Fig.6. Flat position, 22mm deep melt run made in low alloy steel with a pulsed NVEB beam

In conclusion, the ability to weld thick section material (25-150mm) in a single pass has been one of the driving forces behind the development of high-power EB welding systems with potential for high-productivity and high-accuracy fabrication. The requirement to operate in high vacuum has precluded the use of the process in the fabrication of very large structures as the construction and operation of large vacuum chambers can be difficult to justify economically. The possibility of having portable local vacuum equipment which can be delivered to site and operated on a lease/hire basis improves practicality and economics of thick section EB welding. Operation at reduced pressure where process reliability and performance is better than at high vacuum make this now a practical proposition and work is currently underway to manufacture an industrial system which will realise this opportunity.

Similarly, the possibility to operate at atmospheric pressure with an non-vacuum system is equally attractive particularly where thinner materials are concerned (i.e. <50mm). With recent developments in beam pulsing and gun configuration, it is likely that high-speed welding of even thicker materials will be realised at atmospheric pressure in the near future.

3 Development of novel non-destructive testing methods for, and reliability of, pipeline inspection

3.1 Long-range ultrasonic testing of pipelines and piping

Very substantial investments were made in the 1970s in the development of both pipeline pigs for electromagnetic and ultrasonic detection of wall thinning, and X-ray crawlers for weld inspection in long runs of pipe. The first X-ray crawler was used for weld inspection during pipelay. Pigs and crawlers were at that time the only means of inspecting for erosion, corrosion and other defects in pipes buried inaccessibly below ground or on the sea bed, and/or encased in concrete or other protective coatings to protect outer pipe walls from corrosion. Both approaches have the capability to inspect long lengths of pipes and pipelines in a short period of time. However, not all pipelines can be inspected in this way. Pipe bends or steep gradients occur, which prevent the 'pig' passing through. There are therefore very large proportions of pipelines that cannot be inspected using these methods.

In cases where the pipes are accessible, inner wall corrosion and other defects can be inspected by double-wall through-transmission radiography. A variety of external pipe crawlers and magnetically adhering robot vehicles have been used in an attempt to deploy the inspection sensors more rapidly on long runs of pipe. However, ultimately, the time taken to achieve total coverage is still of serious economic concern, even with the present state-of-the-art of external robotic deployment. Furthermore, most pipelines in process and manufacturing plant are inaccessible for external inspection by the above techniques, along most of their length, because of their proximity to other pipes and structures.

One technique that has been demonstrated as a viable solution to the problem of inspecting non-piggable or inaccessible pipelines is long range ultrasonic testing (LRUT), also referred to as guided waves, and recent progress in the development of this technology is the subject of this section, based on work by Gan et al, (2007). TWI has pioneered this technique for corrosion and erosion detection and monitoring, which, in principle, allows inspection of long runs of pipe from just one access point. In this technique, a pulsed, guided ultrasonic wave is propagated in a pipe wall from a family of equally spaced ultrasound probes supported by a collar wrapped round the pipe. The wave is reflected from the pipe end and from circumferential welds and defects in the wall, and the reflected echoes are received by the transmitting probes. Therefore, all defects in the entire run of pipe are detected simultaneously, provided they are large enough to produce an echo amplitude above the random noise level.

The promise of the technique as a large-scale monitoring tool stems from the fact that low-frequency guided waves have a very long range in pipes because: (i) absorption in the pipe material is low; (ii) for pipes in air, leakage out of the pipe is very low because of the high acoustic impedance mismatch at the solid-air boundaries (therefore, all the energy propagates down the pipe with little attenuation of the energy density); and (iii) a wave mode with low dispersion can be selected so that the rate at which the wave pulse spreads out in time is small. The wave amplitude therefore decreases only slowly with propagation distance and, correspondingly, the minimum detectable flaw size does not reduce markedly with increasing distance. The technique is very amenable to numerical modelling, which has enabled optimised process parameters to be obtained efficiently and quickly. The test range can thus be defined as the distance at which a flaw which requires detection gives a detectable echo, and the probes need only be repositioned along the pipe at intervals equal to twice the test range, with a small allowance for overlap to achieve total inspection coverage of an indefinite length of pipe. Depending on many factors such as pipe diameter, wall thickness and bend radius, as well as the considerations (i) to (iii) above, the test range can be as much as 100 metres for an uncoated pipe in air.

TWI exploits its LRUT technique through Pi Ltd and Figure 7 illustrates the latest generation of Teletest® long range ultrasonic equipment. Pi Ltd have provided service inspection in the pipeline sector with their LRUT system since the late 1990s and during this time have gained much experience of the growing potential of the long range ultrasound technique. A particularly interesting case was the inspection of a 10 inch diameter buried pipeline. A LRUT survey was carried out on a network of water injection lines at an oil processing facility. The concerns were external corrosion around the soil to air interface and where the coating had been damaged in the underground sections. Several locations were identified as being moderate-to-severe defects and were excavated. At one location, the line was found to be so heavily corroded that repairs were immediately required, see actual damage shown in Figure 8. It is unlikely that this defect would have been detected by other means before it had caused a failure of the line.

Fig.7. Third generation Teletest® system

Fig.7. Third generation Teletest® system

Fig.8. Corroded section of pipeline following detection by LRUT

Fig.8. Corroded section of pipeline following detection by LRUT

Work is now underway to move this technique towards higher sensitivity and global screening by using focussed waves. The focussing facility is based on a long-range phased array approach (Catton and Kleiner, 2005). The principle of focussing long range ultrasonic waves at an arbitrary distance along the pipe and circumferential position is as follows: delays are applied to a LRUT transducer tool, facilitating a greater concentration of energy at a given point, and hence a greater reflection from changes in cross sectional area. More sensitivity in the results can thus be obtained (see Figure 9). In doing so, more quantitative information about defects can be drawn from test data. The technique has been proven both in the laboratory and in the field, using simulated real defects. The field test results corresponded with those of radiography tests conducted on the pipe. Investigations using finite element analyses have also been carried out, thus validating the experimental results.

Fig.9. Modelling of time delay LRUT focussing technique

Fig.9. Modelling of time delay LRUT focussing technique

More recently, research is also being carried out into a novel focussing method called small reversal focussing (Ennaceur et al, 2007). The principle is that the timed history of signals received from the pipe including a flaw are recorded by the transducer. Then the pulse received is time reversed and sent again down the same pipe. It is found that the time reversed wave focusses on the region that contains the defect. The method leads to an increase in signal amplitude. Time focussing has an advantage over the time delay approach in that it does not suffer from aberrations in inhomogeneous media. It also has the ability to focus automatically because the technique does not need to know the exact position of the defect. The time reversed signals will focus automatically on the defect as it is the source of reflection. From initial experimental results, it is expected that corrosion defects as small as a few percent of the pipe wall cross section with a 95% probability of detection will be detectable with this approach.

In conclusion, TWI has pioneered the LRUT technique for corrosion, erosion and metal loss detection and monitoring which in principle allows inspection of long runs of pipe from just one access point. This provides the ability to inspect complete pipelines, rather than just those regions accessed by pigs and crawlers. Furthermore, the inspections are cost effective and reliable, and are done with the minimum of disruption.

The LRUT system developed is capable of full coverage assessment of pipe lengths of typically 30m in either direction from the test position, although in ideal conditions significantly greater lengths of pipeline can be inspected from a single test point. Recent developments, such as the ability to focus the ultrasonic waves, have added a further dimension to the technique's potential.

3.2 Reliability of inspection for flaws in pipe welds

For flaw detection, different methods are applicable according to whether the outcome of a trial is recorded as a binary variable, ie hit/miss data (typical of enhanced visual techniques such as magnetic particle inspection); or a continuous variable, ie a signal amplitude relative to a given threshold (typical of ultrasonics or eddy current testing). In the hit/miss case, the method of analysis is further divided into methods that group the data, and methods that treat it as a whole to calculate a probability of detection (POD) curve. The continuous method requires the signal amplitude and a threshold for detection. The POD is produced from a set of data that contains more information than the hit/miss method, and this can allow a smaller number of flaws to be used. Wherever possible, POD should be estimated from signal amplitude data rather than hit/miss data. For a given sample size, this will generally yield a more accurate POD estimate, because it will be based on more information. This section, based on work by Schneider and Rudlin, (2004), will focus on signal amplitude data.

For flaw sizing, attention is usually focused on the amount of undersizing that can be allowed, which can then be used to set realistic acceptance criteria. However, it may also be important to quantify the amount of any oversizing, since this can cause unnecessary repairs or plant shutdown. For simplicity, it is assumed that the inspection reliability is being quantified in terms of a single factor (usually the through-wall extent of the flaw) and that all other factors are representative of the site inspection. In practice, this means that these other factors are either carefully controlled or are randomly sampled.

Figure 10 shows the results of automated ultrasonic testing (AUT) of pipeline girth welds. The pipeline operator wished to qualify the inspection to detect planar manufacturing flaws of 2mm through-wall extent (TWE) at 90% POD with 95% confidence. A total of 31 planar flaws were deliberately introduced into four test welds. Figure 10 gives the maximum amplitudes recorded by the system, and shows that all 31 flaws were detected. Data shown as 'saturated' lie at an indeterminate value above the plotted amplitude. These points do not all lie at the same amplitude because different saturation limits applied to different channels of the AUT system.

Fig.10. AUT response versus size data for planar flaws in pipeline welds

Fig.10. AUT response versus size data for planar flaws in pipeline welds

Twenty nine of the flaws are less than 4mm in TWE, and statistical methods can be used to show with 95% confidence, that flaws of 4mm TWE are detected with 90% POD. But this method cannot be used to establish 90% POD with 95% confidence for flaws smaller than 4mm, even though all 31 flaws gave signal amplitudes above the threshold. Moreover, since there are no non-detections, it is impossible to establish a functional dependence of POD on TWE. The usual method of dealing with such data is to use more advance statistical methods, and Figure 11 shows the regression line (the central solid line) fitted through the response versus size data of Figure 10 by the method of Maximum Likelihood. This method properly accounts for the saturated data and provides estimates of the slope and intercept of the regression line (and the spread of data about this line) that maximises the likelihood of obtaining the observed data.

Fig.11. Regression analysis of response versus measured sites

Fig.11. Regression analysis of response versus measured sites

In this sense, the resulting estimates are those that agree most closely with the observed data. Figure 11 plots the 'exact' data only (ie where the signal was not saturated) even though the saturated data is taken into account in estimating the position of the regression line. This is why the regression line does not lie centrally among the data points.

Figure 11 also shows estimates of the 10th and 90th percentiles of the distribution of signal amplitudes about the regression line, which correspond to the threshold levels that are needed to achieve 90% POD and 10% POD, respectively. The centrally positioned straight lines are the best estimates of the percentiles, whereas the curved lines correspond to one-sided 95% confidence limits on these percentiles. The lowest dashed curve indicates that flaws of 2mm TWE are detected well above the inspection threshold with a POD of 90%, with 95% confidence.

Sizing errors are usually assessed by comparing the maximum size of a flaw reported by NDT with the true maximum TWE of the flaw (as determined by metallurgical sectioning), since this is generally the key dimension in determining the failure resistance of the weld. It is usually assumed that ultrasonic sizing errors are normally distributed. This means that confidence limits for sizing errors can be estimated by ordinary linear regression.

Figure 12 illustrates AUT sizing with measured data for the same 31 pipe weld flaws. The lower dashed curve, for instance, indicates that flaws of 4mm TWE will, with 95% confidence, have a measured TWE of at least 2.5mm. This information can be used to derive acceptance criteria based on fitness-for-purpose principles. If an engineering critical assessment shows that the pipeline can tolerate flaws larger than 4mm TWE, then an acceptance criterion based on 95% reliability will require rejection of all welds containing flaws having a measured TWE larger than 2.5mm.

Fig.12. AUT flaw sizes versus sizes determined by sectioning

Fig.12. AUT flaw sizes versus sizes determined by sectioning

In summary, the general purpose of NDT is to provide accurate flaw information. This, in turn, allows engineers to make judgements concerning the future life and/or safety of pipelines. These judgements need to take account of the effectiveness of NDT techniques, both in terms of flaw detection and sizing. This section has shown that there are relevant statistical methods available for quantifying the reliability of pipe weld flaw detection and sizing information which can be used to validate NDT approaches.

4 Corrosion assessment of pipe welds

4.1 Corrosion fatigue of welded risers and pipelines made from C-Mn and stainless steels

C-Mn steel risers are widely used for deepwater applications and are subject to fatigue damage associated with wave or tidal motion and vortex-induced vibration. Fatigue loading is particularly severe at the touchdown area where the near-vertical riser meets the flowline. Risers and pipelines are also exposed to potentially aggressive service environments on both the inside and outside surfaces. The corrosive effect of seawater on the external surface is typically controlled by the application of cathodic protection. However, this can lead to hydrogen generation on the steel surface, and degradation in fatigue performance needs to be accounted for in design. CO2-containing (commonly referred to as 'sweet') and H2S-containing (referred to as 'sour') fluids acting on the internal surface also have a significant effect on fatigue behaviour. The presence of salts and water, acidified by the presence of CO2 or H2S, all contribute to the aggressiveness of the environment.

Laboratory testing to quantify the corrosion fatigue performance of welds in aggressive service environments can be performed in two ways. Endurance data are obtained by testing through-thickness sections of weld and plotting the results on an S-N curve. The effect of environment is then expressed in terms of a fatigue life reduction factor, by comparison with endurance data for tests carried out in air. Standard fatigue design curves, such as those given in BS 7608, can then be offset by this factor. In contrast, when fracture mechanics calculations are carried out to determine critical tolerable flaw sizes, fatigue crack growth rate data are required in the environment of interest, and an upper bound curve is derived for design purposes.

Endurance or fatigue crack growth rate data must be obtained under appropriate environmental conditions. Service environments can be complex and it is important to have an appreciation of the environmental variables that can influence the extent to which fatigue performance is degraded. Mechanical variables such as stress range, stress ratio (R) or cyclic loading frequency can also have an effect on the extent to which an environment affects fatigue performance and their influence needs to be considered.

Details of these parameters have recently been reviewed by Baxter et al, (2007), who evaluate the various influencing parameters in seawater, sweet and sour environments. This section is based on their work. The review assessed a number of relevant publication, but the amount of published data is modest, given the industrial importance of the result. TWI has contributed to filling some of the gaps by carrying out fatigue crack growth rates in seawater at various frequencies.

The effect of very low-frequency cycling has been examined by carrying out so-called 'frequency scanning' tests. These are crack growth rate tests carried out under conditions of constant applied stress intensity factor range 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 possible using 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, the measured 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 a plateau was not reached until a much lower frequency, at least as low as 0.001Hz.

Fig.13. Frequency scanning test data for C-Mn steel in seawater

Fig.13. Frequency scanning test data for C-Mn steel in seawater

A further example of the data reviewed by Baxter et al, (2007) relates to the effect of H2S-containing solution on fatigue endurance. This example has been chosen because there are very few published endurance data for girth welded C-Mn steels in sour environments. The only published data relate to tests carried out by Buitrago and Weir, (2002) where strips extracted from girth welded X80 and X65 pipe were tested in H2S partial pressures of 35-70mbar. Data are reproduced in Figure 14 and it can be seen that fatigue life was reduced by a factor of 10-20 compared to both strip tests and full-scale tests carried out in air.

Fig.14. Fatigue endurance data for girth welds tested in air and sour environment

Fig.14. Fatigue endurance data for girth welds tested in air and sour environment

The experimental data available all indicate that environmental and mechanical test conditions can have a significant effect on the observed fatigue endurance or crack growth rate. The need to ensure that design data are determined from tests carried out under conditions that realistically simulate those encountered in service cannot therefore be over-emphasised.

As described above there are detrimental effects of sour fluids on the fatigue performance of riser girth welds in C-Mn steel pipe. Results indicate that a reduction in fatigue life may occur in sour fluids by a factor of up to 20, particularly at higher stress ranges. Consequently, there is a need to explore the use of corrosion-resistant alloys (CRAs), either stainless steels, nickel alloys or titanium alloys as alternative riser materials, either as a solid or clad product. The main candidate stainless steels are high strength grades such as 12%Cr supermartensitic, 22%Cr duplex and 25%Cr superduplex grades, or austenitic grades such as 316 stainless steel or Ni-Fe alloy 825, but these have fairly low strength and would have to be used as a clad product. This section, based on work by Woollin et al, (2005), presents an example results of corrosion fatigue crack growth rate behaviour of welded supermartensitic and superduplex stainless steels in seawater with cathodic protection and in a sour environment.

Three high-strength stainless steels, which could be used to construct a riser without the need for clad product, were chosen: (i) UNS S41426 12%Cr supermartensitic stainless steel welded by the pulsed MIG process using a proprietary superduplex wire (Zeron 100X), (ii) UNS S31803 22%Cr duplex stainless steel plate, heat treated to 1350°C and immediately water quenched to simulate a high temperature HAZ microstructure with 62% ferrite, and (iii) UNS S39274 25%Cr superduplex stainless steel welded by the TIG process with a proprietary superduplex wire (Zeron 100X).

Full details of the experiments carried out and the results obtained are give by Woollin et al, (2005) but, as an example, Figure 15 shows the fatigue crack growth rate results for the 25%Cr superduplex stainless steel welds in a sour brine solution. At both ambient temperatures and 80°C, crack growth rates were substantially higher in the sour environment than in air. Crack growth rates were up to ten times higher than in air at higher ΔK values, particularly in the mid-ΔK range. At high and low ?K values, the effect of the environment is apparently reduced, especially at 20°C in the high ΔK range, although it is noted that only a fairly small amount of data is available in the low ΔK regime.

Fig.15. Fatigue crack propagation data for superduplex stainless steel weld metal centreline notched specimens tested in air and a sour brine with 25mbar H2S

Fig.15. Fatigue crack propagation data for superduplex stainless steel weld metal centreline notched specimens tested in air and a sour brine with 25mbar H2S

The effects of the various environments on fatigue crack growth rates observed by Woollin et al, (2005), may provide the basis of preliminary design recommendations. Thus, in the presence of sour brine, for a structure which experiences low-frequency fatigue loading, a fatigue life reduction factor of at least 5-10 may be required for superduplex stainless steel welds. However, since this appears to be particularly relevant at relatively high crack growth rates, it might not be necessary to apply such factors except for low-cycle fatigue applications. There is a need therefore for endurance testing of specimens with geometry representative of real components to determine accurately the fatigue life in such corrosive environments.

The sensitivity to hydrogen of high-strength stainless steels and the consequent reduction of fatigue behaviour under conditions that introduce hydrogen at the crack-tip suggests that a more suitable choice for a sour-service, corrosion-resistant riser may be a material with good intrinsic resistance to hydrogen embrittlement, such as an austenitic stainless steel or a nickel alloy. Due to their low strength, it is most appropriate to use these in the form of internal cladding on a higher strength C-Mn steel pipe. A second alternative would be a material with very highly stable passive film, such as a titanium alloy, such that no appreciable hydrogen pick-up occurs at the crack-tip. Further data are required to explore the most suitable alternative.

4.2 Characterisation and corrosion behaviour of thermal sprayed aluminium coatings

Thermal sprayed aluminium (TSA) coatings are widely specified for the protection of steels from aqueous corrosion in seawater environment, e.g. offshore structures, risers, pipe components and ship structures. Good toughness, low maintenance requirements, and long-term service life are among the desirable properties offered. The basic properties of a TSA coating are its barrier characteristics, combined with good adhesion and an ability to provide cathodic protection to exposed steel. It has been reported that an optimised 200µm thickness TSA coating provides a service life in excess of 30 years in a splash zone environment. Key factors for optimisation include TSA alloy composition; surface preparation; choice and application of a sealer; service application; electrochemical/galvanic exposure conditions; coating application techniques; and process parameters.

The stringent requirement of the offshore industries to have a maintenance-free coating system, with an expectation of up to 50 years life in severe environments, has increased interest in TSA coatings as a viable economic alternative to paint-based systems. Although it has been demonstrated that TSA deposited by conventional spray systems can work well over extended periods, incidents of premature coating failure due to blistering and detachment have been recorded, indicating that coating quality and coating application procedure is extremely important. Porosity, oxide content and non-uniformity in the TSA coatings produced by conventional systems are believed to result in reduced corrosion protection and shorter lifetime. Recent developments in spraying equipment via modification in nozzle and gun design, and process parameters (such as higher gas flow rates and gas pressures) are believed to produce TSA coatings with improved properties. Characteristics of TSA coatings applied onto C-Mn steel, using conventional and newer wire thermal spray systems have therefore been assessed and are summarised in this section, based on the work by Shrestha and Sturgeon, (2005).

Back-scattered SEM images of example coatings are shown in Figure 16. The images show layered structures of the impacted molten particles with different amounts of coating porosity and oxide stringers typical for thermal sprayed coatings. The high velocity (HV) coating shows the least variation in coating thickness, ranging between 222-242µm, while the arc sprayed (AS) coating shows variation in coating thickness ranging between 224-303µm. The HV coating also displayed the lowest amount of porosity at about 2.4%, whilst the AS coating had about 6.3% porosity. The HV coating has a very smooth surface finish with about 5.5µm Ra value, while the AS coatings were rougher at about 15.4µm Ra. Measured Vickers hardness values were also higher for the coatings prepared with the newer system.

Fig.16. Back-scattered electron image of a cross section of: a) arc sprayed; and

Fig.16. Back-scattered electron image of a cross section of:

a) arc sprayed; and

b) HV wire flame sprayed aluminium coatings

b) HV wire flame sprayed aluminium coatings

The results of the electrochemical polarisation tests conducted in a nitrogen-purged 3.5wt% NaCl solution at 20°C suggest that there are differences between the coatings produced by various systems. In particular, the AS coating was shown to have a less negative corrosion potential than those exhibited by the coatings sprayed using the HV system. This is believed due to the higher amount of oxide and porosity in the AS coating, giving a mixed potential of the coating and steel system upon immediate exposure to the test solution. The HV coating displayed a large passive region, where the current density was low. This large passive region was similar to that of wrought aluminium and suggests that the coating on its own could act as a highly protective barrier. The AS sprayed coating did not display such a passive region.

Corrosion rates of thermal spray deposit with a 5% uncoated area were used to determine the rate at which the aluminium coating was lost when immersed in a 3.5wt% NaCl solution at ambient temperature. The calculated corrosion rates were assumed to be only due to dissolution of the aluminium coating. The corrosion rates of TSA coatings produced with the new system with a 5% uncoated area were about 11-12µm/y, in contrast to the AS coating which exhibited a rate of about 20µm/y, about two times higher. The corrosion rates for the specimens without uncoated area were substantially lower, as expected. Reports in the literature indicate that longer-term protection can be expected in seawater based on field trials and actual service experience. In seawater, the corrosion rate will be lower due to the formation of a calcareous deposit on the exposed steel (cathode) and corrosion product films on the TSA (anode) surface and hence a longer coating life can be expected.

5 Structural integrity assessment of pipeline welds

5.1 Fracture Integrity of reduced pressure electron beam girth welds for deep water pipelines

The Reduced Pressure Electron Beam (RPEB) process offers potential economic advantages for producing high-integrity girth welds in heavy-wall pipe for deep-water pipe lay applications (Belloni and Punshon, 1997),. It is a single-pass welding process for which the welding time is relatively independent of pipe wall thickness. The most critical factor for weld quality is the beam to joint alignment. This is assured by means of an on-line seam tracking system which compensates in real time for any deviation of the beam from the joint line. Furthermore, unless special steel compositions are chosen which resist grain coarsening, the high effective heat input inherent in the welding process means that the melted region in autogenous welds and the grain coarsened HAZ can have relatively poor fracture toughness, as also considered recently by Prof Toyoda and colleagues (Ohata et al, 2007). Nevertheless, girth weld fracture resistance may be adequate for pipeline applications when standard commercial pipe steels are employed if it can be shown that the girth weld is tolerant to small flaws that are on the limit of reliable sizing by non-destructive testing. This section, based on work by Pisarski and Punshon, (2004), therefore describes the results of full-scale pipe bend tests conducted to demonstrate that RPEB girth welds can be resistant to brittle fracture even if small welding flaws should exist.

Three full-scale pipe tests were carried out to evaluate the fracture performance of RPEB girth welds in pipe to API 5L Grade X65. Two of the pipes tested were 24in OD x 31.8mm WT with a composition designed for sour service (pipe tests 1 and 2). The third pipe was 28in OD x 38.0mm WT and was designed for sweet service (pipe test 3). The girth weld in pipe test 1 was an autogenous RPEB weld and was expected to have relatively poor fracture toughness in both weld and HAZ, when assessed using conventional fracture mechanics tests. An all-weld metal tensile test indicated a 0.2% proof stress of 516MPa. This compared with a 0.2% proof strength of 467MPa measured in the parent pipe. A shallow artificial surface flaw, inclined at approximately 5° to the girth weld was introduced to intersect weld metal, fusion boundary and HAZ. The flaw was 150mm long and 4mm deep. In order to focus attention on the performance of the HAZ, the second pipe test was welded with a pre-placed shim of high-purity Ni. This produced weld metal with 2-3% Ni content and significantly higher fracture toughness than that obtained by autogenous welding. A deliberate welding flaw representing a missed joint, was induced by temporarily disabling the seam tracking system and by using the beam deflection system to direct the beam to miss the root over a distance of 45mm.The third pipe test was on a pipe different from the first two but the girth weld was made in the same way as the second pipe test except that the seam tracking/beam deflection system was not disabled. An artificial flaw was introduced into the fusion boundary on the inside diameter of the pipe by spark erosion (EDM). The flaw was 1.6mm deep and 20mm long. Ultrasonic inspection detected the flaw and sized it to be 1.7mm deep.

The three pipe samples were made into test pieces 6000mm long and placed in a four-point bend rig. The notched area was located at the 12 o'clock position, on the maximum tension side. The first pipe test was conducted at a temperature of 13°C. The other two tests were conducted at 0°C. Standard fracture toughness tests were also prepared and tested to generally meet BS 7448:Parts 1 and 2 requirements. The specimens included both deep and shallow notches in the WM and HAZ position.

The first full-scale pipe bend tests survived an overall strain of 2.6% before the pipe was unloaded. Local strains in the weld metal were about 0.5% lower. The test set-up and curvature achieved in the pipe is shown in Figure 17. A metallographic section taken at mid-length of the notch confirms that the tip was located in the fusion boundary and that initiation by ductile tearing had taken place. Fractography and metallography indicated that approximately 0.2mm of crack growth had taken place at an applied CTOD of 0.98mm.

Fig.17. First pipe bend test RPEB girth weld at a strain of 2/6%

Fig.17. First pipe bend test RPEB girth weld at a strain of 2/6%

The second pipe fractured in a brittle manner at 0°C and a stress of 276MPa or 60% of the parent pipe SMYS. Post-test fractography showed a much larger and more irregular shaped flaw than expected. In fact, two fusion boundary flaws were created. The first was an irregular shaped surface flaw on the pipe inside diameter approximately 45mm long with a maximum depth of 15mm. The second was an embedded flaw, ahead of the first flaw, which was about 4.5mm high and 11mm long.

The third pipe test survived an overall strain of 3.01% at 0°C before being unloaded. The local strain reached in the girth weld was lower, at about 2%, probably because of the significantly overmatched weld metal strength compared with the parent pipe. This may explain the relatively low applied CTOD of 0.2mm, although the notch depth was 1.6mm. Post-test metallography confirmed that the notch tip was located on the grain coarsened HAZ side of the weld, very close to the fusion boundary. Indeed, ductile tearing (about 0.1mm) had just commenced.

The HAZ and weld metal fracture toughness values representative of pipes 1 and 2 determined were similar and there was only small increase in toughness when temperature was increased from 0 to 10°C. There was a weak effect of notch depth, higher weld metal fracture toughness can be achieved in shallow surface notched specimen compared with deeply notched through-thickness specimens (maximum values around 750MPa√m for shallow-cracked specimen compared with maximum values of around 500MPa√m for deeply cracked specimens).The HAZ of the girth weld representative 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 and initiation 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 therefore unrepresentative 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 (or stress 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 six shallow-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 is expected 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 the pipe 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 of pipe 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 using deeply 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 can behave 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 loaded predominantly 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 welds contain 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 RPEB process, 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 a particular 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 alloy levels. 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 and tends 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 more care 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 more holistic 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 joint geometries 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 to require 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 girth weld 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 element analysis; 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). An alternative 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 and material/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 strength pipeline 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 generate the 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 by the ratio of the mismatch limit load to the limit load for homogeneous material.

The mismatch ratio M is defined as Mx = σYwYb, where σYw or σYb are flowstresses 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, in general, 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 long surface 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 be representative 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 in different 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 strength mismatch) 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 of allowable 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 undermatched welds (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 the difference 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 effects than weld metal flaws. Nevertheless, there is a clear benefit of using narrow welds if undermatching is likely.

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

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 and illustrate 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 stress cannot 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 occur before 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).

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)

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 that CTOD 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 predicted by 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 specifically analysing 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 to be 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 fracture toughness 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 material subjected 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 are specifically 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 benefits of 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 girth weld 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 there is 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 been used 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 be acceptable 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, not height. 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 nothing inherently 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 fracture mechanics 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 Recommended Practice 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 the procedure 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. The procedure 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 in DNV-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 contains a 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 surface flaws 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 weld width 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 a homogeneous 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 bent onto 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 were obtained (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 residual stresses. 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 were conducted 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 homogeneous material).

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 were employed in the numerical (FE) analyses

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 were employed 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 fusion boundary. 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 weld width 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 the reeling 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 the flaw. 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 applied strains 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 all other 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 fracture toughness. 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 a SENT specimen and a pipe girth weld flaw (Nyhus et al, 2005; Pisarski and Wignall, 2002) indicate that crack tip constraint in the specimen remains 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 the assessment 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 the analysis 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 when these 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) modifying the 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-based reference 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-displacement formulation. 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 limitations of 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 because the 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 crack driving 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' strength conditions 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% higher than 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%. By including 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 obtained assuming 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 productivity has 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 a powerful 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 fatigue performance 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 to assist 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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