HIGH STRENGTH PIPELINE STEELS – ISSUES IN THECONTROL OF FRACTURE R. M. Andrews Advantica Ltd United Kingdom ABSTRACT As an economic high pressure pipeline design can not completely remove the risk of fracture, a fracture control plan is required. This documents the measures which will control the likelihood of a fracture occurring and how the consequences are limited if a fracture should occur. When using high strength steels there are additional considerations beyond those involved in conventional designs as a result of the higher stresses, the different material behaviour and more severe service conditions. This paper reviews the issues involved in the fracture control of high strength steel pipelines, including crack initiation and propagation. It is considered that the existing models for toughness independent prediction of crack initiation are likely to be applicable to high strength steels, but there may be concerns for mechanical damage resistance in ultra high strength grades such as X100. Ductile crack propagation is the other main area requiring resolution. 1 INTRODUCTION It has to be accepted that an economic high pressure pipeline design will not completely eliminate the risk of fracture. The designer must therefore take measures to control the likelihood of a fracture occurring, and then to limit the consequences if a fracture should occur. Some design codes, for example the Australian code AS 2885 [1], now require the production of a formal fracture control plan. Even if a plan is not formally required by the design code or statutory regulators, a formal plan will be beneficial to the operator and the investors financing the project. The plan sets out how the pipeline design attempts to minimize the occurrence of fractures and reduce the consequences of a fracture if one should occur. Fracture control should consider a range of factors, such as crack initiation and propagation in the parent pipe material and seam welds, fracture of girth welds and fittings. Some of the general issues involved in fracture control have been discussed and reviewed elsewhere [2]; this present paper concentrates on the issues involved in fracture control in high strength steel pipelines. High strength steel pipelines present particular challenges as the stresses are inevitably higher, which increases the driving force for fracture. When operating under adverse conditions, such as low temperatures, the problems will be magnified, as it will be necessary to achieve the required properties at these lower temperatures. It is also likely that an adverse environment will be relatively sparsely populated, which will encourage the use of higher design factors than would be considered acceptable in a populated area. This will further increase the stresses. Adverse environments will often introduce other problems such as extensive ground movement due to landslides, seismic activity and discontinuous permafrost. These will place additional loads or strains on the pipeline which need to be taken into account in the control of fractures. As a simple example of the effect of moving to a higher grade of steel, consider a pipe of constant diameter and thickness where an increase in pressure is accommodated by increasing the grade and so the working stress. The fracture driving force in terms of Crack Tip Opening Displacement, δ, is related to the stress and defect size by an equation of the form: E K Y I σ δ 2 = (1) Here σ y is the yield strength and E the elastic modulus. Ignoring residual stresses, the stress intensity factor K I is related to applied stress σ and defect size a by: a Y K I π σ = (2) 2 where Y is a dimensionless factor dependent on the loading and the component and defect geometry. Substituting (2) into (1) and noting that most codes use a design factor which is related to (σ/σ Y ), the toughness required to tolerate a constant size of defect increases linearly with applied stress, or effectively with the pipe grade: E a Y Y π σ σ σ δ 2 = (3) Thus increasing in grade from X60 to X80 at a constant design factor requires an increase in CTOD toughness of 33%; moving from X60 to X100 would require a 66% increase. Although simplistic, this analysis illustrates the effect that increasing pipe strength has on other property requirements. These issues will be discussed further below. FRACTURE INITIATION IN LINEPIPE Fracture Initiation in Parent Material To obtain the best resistance to fracture initiation in the parent pipe, the material must have sufficient toughness to give a failure dominated by plastic collapse. Once this level is reached, the critical defect size is controlled by the material strength properties and is no longer affected by toughness. The most widely used models for plastic collapse of axial pipe defects were developed in the early 1970s by workers at Battelle [3] and Shannon [4] at the former British Gas ERS. For a through–wall defect this model takes the general form: flow T M σ σ = (4a) − + = 2 2 4 2 0135 . 0 255 . 1 1 t R c Rt c M T (4b) where the quantities are defined as: c half length of a through wall crack, mm R pipe radius, mm t pipe wall thickness, mm σ hoop stress, N/mm 2 σ flow flow stress, defined below M T is the Folias factor, which accounts for the stress concentrating effect of bulging at the ends of the crack. The form used here is the original three term approximation used in [3], although other approximations have been widely used. The models, and subsequent variations, differ in their definition of the flow stress, and this can have a small effect on the tolerable crack length as the pipe grade increases. Shannon defined the flow stress as 1.15 * SMYS, and using this definition it can be seen from (4) that the tolerable crack size is independent of material grade when operating at a constant design factor. A problem with this definition for high strength steels with relatively limited work hardening capability is that it can result in a flow stress that exceeds the tensile strength. For example, with X80 (551 N/mm 2 ) yield strength material the flow stress would become 92 ksi (634 N/mm 2 ), which is above 3 the specified minimum tensile strength of 90 ksi. This can be avoided by adopting the definition used in BS 7910 [5], which defines the flow stress as the average of the yield and ultimate strengths. Battelle [3] originally defined the flow stress as SMYS + 10 ksi, which gives a flow stress which is not a constant proportion of yield. Figure 1 shows the tolerable flaw size calculated using (4) for a 36” nominal pipe size, 12.7 mm wall pipe operating at 80% SMYS for a range of pipe grades. The flow stress is defined as SMYS + 10 ksi, and it is seen that there is a small decrease in tolerable flaw size as the grade increases. Increasing the pipe grade has a bigger effect on the toughness required to ensure that this flow stress dependent behaviour is obtained. Once the flow stress dependent condition is attained, there is no benefit in terms of fracture initiation resistance from further toughness increases. A model relating flaw length to Charpy toughness was developed [3] using a strip yield model and calibrated against burst test results: = flow T flow c M c K σ σ π σ π 2 sec ln 8 2 2 (5a) c v c A E C K 1000 2 = (5b) Here the following additional quantities are defined: K c 2 fracture toughness defined as energy/(unit area of crack surface) C v Charpy toughness, Joules E Elastic modulus of steel, N/mm 2 A c fracture area of the Charpy specimen, mm 2 The fracture toughness is related to the Charpy energy by the empirical relation (5b); this has been modified from the form presented in [3] to convert to SI units – when using US customary units the factor of 1000 becomes 12. Figure 1 also shows the Charpy toughness required for the same pipeline design as considered previously to tolerate a through-wall crack with a length of 90% of that obtained with fully ductile material. The values increase significantly as the grade increases. For X80 this analysis gives a requirement of 59 J in a full thickness specimen; API 5L (42 nd edition) requires 68 J for X80 material at 0 °C for PSL 2 material. However, for X70 the API 5L requirement is only 27 J at 0 °C, whilst this analysis suggests 47 J is needed. In practice, for gas transmission pipelines these values will be replaced by the requirements for ductile crack arrest, which are likely to be higher. This model is semi-empirical, and was calibrated against relatively low toughness and low strength (X60 and below) materials. Although it appears that this approach was used in the design of the Alliance Pipeline [2], there must be some concern that the model is appropriate for higher strength grades such as X100. 4 Work sponsored by the European Coal and Steel Community [6, 7] on trial production of 36” and 56” X100 material showed that the failure pressure of part – wall flaws was conservatively predicted by (4a), modified for the part wall thickness flaw geometry. The flow stress was defined using the BS 7910 definition, the average of the yield and ultimate strengths, as this was the only definition giving a flow stress below the ultimate strength. Although the predictions were conservative, the degree of conservatism was not as great as for lower strength materials. Results from ring tension and vessel tests [8] on first generation prototype X100 material are shown in Figure 2. These show that the model developed by the former British Gas and included in Annex G of BS 7910 [5], for predicting the failure pressure of metal loss defects can be used for both volumetric metal loss (corrosion) and crack like defects in X100 grade steels. The original model had already been shown to be appropriate for grades up to X80. This model is similar in form to that in (4), but uses a different definition of the Folias factor and uses the tensile strength rather than a flow stress as the material strength parameter. Overall the limited data on the crack initiation resistance of high strength steels are inconclusive. Data for X100 appear to show that existing industry models are usable, albeit with small margins. Recent work in Japan [9] has, however, suggested that the toughness dependent model, equation (5), may become non-conservative at Charpy toughness levels over 130 J, as the data in [9] shows non-conservatism for X80 material. This is a cause for concern and further work is required to resolve this problem. As an alternative to using Charpy energy, parent material initiation toughness requirements could be derived using fracture parameters such as CTOD or the J integral and assessment methods such as BS 7910 [5]. This would represent a major change in approach from the use of Charpy energy, and would have implications for plate and pipe manufacture, as at present these parameters are not normally specified as production tests. A compromise solution might be to derive a toughness requirement using CTOD or J and then convert this to a Charpy energy requirement using correlations such as those proposed by Wallin [10]. However, such correlations tend to be conservative and generally appear to have been developed and validated using the low alloy steels typically used for pressure vessel applications rather than high strength TMCP linepipe steels. Seam weld toughness It is generally accepted that it is neither practical nor necessary to require the seam weld toughness to match that of the parent material for gas transmission. This assumes that for UOE pipe the pipe will be laid with the seam welds offset, so that a defect propagating in the seam weld will reach parent metal in the next joint, giving at worst an increase in rupture length of one joint. However, some level of toughness is required in the seam weld to give resistance to crack initiation. This can be set to be equivalent to the level required for flow stress dependent behaviour in the parent pipe, or alternatively levels derived from the pressure vessel codes can be used. If the approach of PD 5500 Annex D [11] is used, this would require the seam weld in X70 and above to achieve 40 J Charpy energy at a test temperature related to the design temperature and the thickness. For some applications this would require testing below 5 the design temperature; for example a 25 mm thick pipe of grade X70 or higher with a design temperature of –20 °C, as might be required for Arctic service, would have to achieve 40 J at -40 °C. Depending on the shape of the transition curve, this may be a more onerous requirement than reaching the levels for toughness independent behaviour, as discussed above and shown in Figure 1. The requirements for seam weld HAZ toughness levels are less clear cut. Low Charpy and CTOD toughness levels have been measured in the seam weld HAZ of X100 materials [12] but the failure pressures of ring specimens of the same material with artificial defects in the HAZ were predicted by a ductile collapse analysis [8]. Analysis using a conventional fracture mechanics assessment to BS 7910 [5] procedures had predicted much lower failure pressures. This dichotomy was resolved [13] by applying constraint based fracture mechanics procedures [14] which showed that the actual loading conditions on a seam weld HAZ defect represented a much lower constraint condition than the standard test pieces. This conclusion is consistent with experience in lower grade pipelines where low toughness heat affected zones have occurred, but there is no evidence that this is a major cause of failure. Similarly, the offshore structural industry has carried out extensive investigations into local brittle zones in the HAZ, [13] but experience has shown that there have not been extensive cases of failures of structures due to low toughness areas. Mechanical Damage Resistance An area where there is little data for high strength steels is that of resistance to mechanical damage. This can take the form of dents, gouges and combined dent- gouges; the most severe form is the dent-gouge where material at the base of the dent has been removed and cold worked. Operators and regulators of pipelines constructed from high strength linepipe will require assurance that the material can resist external damage, and methods for assessing damage detected in service will be required. Models for assessing dent-gouges have been developed by Battelle [15] and the former British Gas (now Advantica) [16-18]. The Battelle model is empirical, using a damage parameter Q: c t d R D C Q v 2 2 = (6) where D is the dent depth and d the gouge depth. The British Gas model is more complex, being based on a fracture mechanics model which includes the stress raising effect of the dent. As in the Battelle model, the material toughness is represented in the British Gas model by the 2/3 Charpy toughness through an empirical calibration. Although the toughness range used in the calibration of the British Gas model was up to 140 J (equivalent to 200 J in a full size specimen), the material grade was limited to X70. A similar restriction applies to the Battelle model, but the toughness was limited to a lower level, 80 J in a 2/3 specimen. As the calibrations against Charpy toughness are empirical, tests are required to determine if the calibrations remain valid for high stress levels. 6 The toughness of gouged and dent-gouged material is likely to depend on the amount of strain induced during denting and gouging. Experimental [19] and analytic studies [20] have shown that prior strain reduces the toughness of linepipe steel. The reduction in toughness of pre-strained material is approximately linearly related to the amount of prior strain [20]. The strain induced during damage is likely to be a function of dent geometry and the extent of the damage rather than material grade. As high strength materials tend to have relatively low uniform strains at failure [12], the strain induced in damaging the pipe will be a large proportion of the failure strain. Hence the toughness of externally damaged high strength line pipe may be relatively low when compared with that of a low grade material subject to the same amount of damage. Overall, the resistance of high strength linepipe to external damage is an area where further work is required to demonstrate that current models are applicable, or to develop alternatives. FITTINGS A pipeline will require various fittings such as bends, tees and valves. For bends and tees the toughness levels may be based on those required for control of crack initiation in the corresponding linepipe, or alternatively the requirements from pressure vessel codes could be used. There is no requirement to control crack propagation in fittings, as generally the stresses are lower than in the linepipe to compensate for the lower material grades which are usually used as it is not possible to obtain the high strength levels. Some component standards, such as API 6D [21] do contain toughness requirements, although these are not onerous by linepipe or some pressure vessel standards. The requirements are for Charpy energies based on the tensile strength; for UTS < 85ksi – 20J, 86-100 ksi - 27J, > 100 ksi – 34 J when the minimum design temperature is below -20 °F (-29 °C). Unlike in the pressure vessel codes there is no requirement for increased toughness if the thickness increases; this may be a concern where high pressures require wall thicknesses beyond the normal range. There will also be a requirement for more complex components such as pig traps and filters, although these are generally designed to pressure vessel design codes. The materials allowed by the pressure vessel codes are likely to be much lower in strength than high strength linepipe; for example the UK design code PD 5500 [11] does not include any Carbon-Manganese materials with yield strengths approaching even X70 strength levels. This problem is compounded by the fact that the pressure vessel codes generally use lower design allowables than are accepted for line pipe. In PD 5500 the design allowable stress for Carbon Manganese steels is the lower of (yield/1.5) or (tensile/2.35), but most pipeline codes allow a hoop stress of (yield/1.4) or in some cases (yield/1.25). Obviously the lower strengths and design factors can be offset by an increase in thickness, but this will increase the weight and fabrication costs, and will also make it more difficult to achieve adequate toughness. An alternative would be to move to alternative steels which are not normally used for pressure vessels, such as the HY series alloys used for submarine hull construction. 7 GIRTH WELD DEFECT TOLERANCE The fracture control of a pipeline should consider girth weld failure. This requires control of the defect levels in the weld, particularly where the welds are likely to be subject to large axial loads. It should be noted that the most adverse condition may occur during installation rather than in service; for example offshore pipelaying by reeling will generate strains well above yield. This control of defects must be achieved by inspection after welding combined with defect acceptance criteria, as the post construction hydrotest will not stress the girth welds to a high level. Most girth weld defect acceptance levels have been based on workmanship requirements, but there has been a move towards using acceptance criteria based on a fitness for purpose approach. This allows the defect acceptance levels to be related to the service conditions, and so reduce repair rates without reducing the integrity. Such an approach is suited to pipelines operating in adverse environments, as the defect acceptance levels are related to the actual service conditions. With workmanship criteria, the integrity of the weld is arbitrary and may not be sufficient to tolerate adverse conditions such as strains beyond yield. A good example of a fitness for purpose approach is the set of guidelines produced by the European Pipeline Research Group (EPRG) [22]. These have a three tier approach, with the higher tiers allowing larger defects but with more restrictions on the properties such as toughness. The published guidelines are currently limited to material grades up to X70, but recent work has shown that Tier 2 can be extended to include grade X80 [23], although the margins are becoming smaller than at lower strength levels. Extension beyond X80 to include grade X100 has not yet been considered by EPRG and for this grade application specific testing will be required until sufficient data have accumulated to support an extension. For adverse conditions where the axial strains exceed yield the EPRG guidelines can not be used as they assume an axial load equal to the yield; testing or more advanced analysis is required. An example of this, using a combination of experiments and analysis using a fracture mechanics approach is given in [24], where defect acceptance levels for welds intended to be reeled were obtained. FRACTURE PROPAGATION The fracture control plan should consider fracture propagation for both safety and economic reasons. For conventional grades and thicknesses relatively little attention has been given in recent years to brittle fracture propagation, as modern high quality linepipe producers can achieve the DWTT shear area requirements usually specified. However, this may not be the case for high strength levels combined with thick walls or low temperatures. Under these conditions both brittle and ductile propagation should be considered. Brittle Crack Propagation Research carried out in Europe has shown that the standard DWTT shear area requirement can be used to avoid brittle fracture propagation in high strength steels at least up to X80 and probably to X100. Results from tests carried out on thick wall X80 material [25] are shown in Figure 3. These show that the transition curve 8 obtained from the DWTT using full thickness specimens is conservative with respect to the curve from West Jefferson tests. The margins are not necessarily this great; data for X70 presented in [25] showed a smaller temperature shift between the two curves. Results on X100 material [7] also show that the DWTT appears to give a conservative prediction of the brittle fracture propagation behaviour. A problem when testing thick walled high strength pipe is that the test machine may not have enough energy to completely fracture the specimen. Possible solutions to this are to test a reduced thickness DWTT specimen and apply a temperature shift to the results or to use the Charpy transition behaviour instead. The data in Figure 3 suggest that the reduced thickness DWTT is only marginally conservative. The comparison between Charpy DWTT and the West Jefferson test is shown by results obtained in an EPRG study on thick walled X65 pipe [26]. This data (reproduced as Figure 4) showed that Charpy transition curves are unconservative compared to DWTT and West Jefferson tests. In addition, this work also showed that large variations may exist between the Charpy transition curves obtained at the centre and at the surface of the plate. Overall, the data presently available suggest that the full thickness DWTT is a satisfactory predictor of the brittle fracture propagation behaviour of high grade linepipe, at least up to X80 and probably for X100, although only limited data are available at this strength level. Ductile Crack Propagation Extensive research is in progress on the control of ductile fractures in high strength steels, and it is not possible to cover all of this in a section of a general review paper. Rothwell [27] gives an excellent review of the history of research into ductile crack propagation and considers many of the general issues. This section will therefore only highlight a few issues of particular concern. The most widely used predictive formula is the Battelle equation which predicts the 2/3 Charpy energy required to arrest a crack as a function of hoop stress and geometry: ( ) 3 / 1 2 5 3 / 2 10 . 382 . 2 Rt C h V σ − = (6) This equation shows the strong dependence of the predicted toughness on the hoop stress, so increases in pipe grade will increase the required toughness by a large amount. In fact this formula underpredicts the required toughness; Figure 5 shows collected results for X80 and the available results for X100 [7]. It is apparent that a correction factor of 1.33 – 1.4 is required to predict the arrest/propagate boundary. Equation 6 was derived by fitting to results from the Battelle Two Curve model, and so is limited to strength grades up to about X70, lean natural gas and pressures below about 100 bar. The full two curve analysis can remove those limitations, and take account of non-ideal gas behaviour, and so is the preferred approach for high strength material used for high pressures. The test results represented in Figure 5 are shown when analysed using the two curve model in Figure 6, where it is seen that the multiplier required for safe predictions increases to about 1.5. This increase is because 9 for the conditions tested the full two curve model tends to predict slightly lower toughness requirements than the simple equation. This is not always the case, particularly for gas exhibiting two phase decompression behaviour. Although Figures 5 and 6 suggest that the fracture propagation behaviour of X100 may be similar to that of X80, the data are limited. Further results are required to confirm this conclusion. Two major projects have recently been completed to provide some data [7, 28], but the results are not yet available in the public domain. The foregoing suggests that empirical adjustments to existing models can be used, but as noted by Rothwell [27] such adjustments may no longer be viable. Leis has shown that improvements can be obtained by extracting that part of the Charpy energy which is associated with crack propagation from the total energy [29]. Improvements may also be obtained by ensuring that the absorbed Charpy energy is well below the test machine capacity, effectively ensuring that the strain rates do not vary significantly during the impact [30]. This has the effect of giving lower measured values. The Drop Weight Tear Test has also been used to estimate the propagation resistance in terms of energy/unit area of crack extension. The energy in this test associated with crack propagation has been estimated by varying the crack initiator geometry. This has replaced the standard pressed notch with a pre-crack produced by methods such as static loading, using an embrittled weld bead or a machined chevron notch. However, none of these approaches has gained widespread acceptance. The fracture energy can also be estimated by varying the propagation distance by modifying the ligament, as in work supported by PRCI [31]. This approach has used a slot in the back face of the specimen which is then blocked by a hardened shim so that it can take load in compression. Crack propagation then stops when the advancing crack reaches the slot. It is assumed that the crack initiation processes are not affected by this modification, so that the varying ligament produces differing amounts of propagation. This can then by analysed to determine the propagation resistance as a function of energy/unit crack advance. This work initially intended to use the Charpy geometry, but was then re-directed to use the DWTT specimen. An alternative fracture parameter used in pipeline fracture studies is the Crack Tip Opening Angle or CTOA. This parameter can be related directly to the kinematics of the propagating crack [32], although it has had limited use as a practical fracture parameter until the development of finite element modelling techniques. The CTOA parameter lends itself to use as a crack advance condition, as the angle can be related directly to the separation of nodes behind the crack tip. Extensive work on the development of a finite element model of ductile crack propagation was supported by PRCI [33]. This work also provided a method of measuring CTOA using a “two specimen” technique. Other methods of measuring CTOA based on the DWTT geometry have been proposed, for example high speed video recording of the fracture [34]. However, these still retain the disadvantages of this geometry, and indeed it has recently been suggested [35] that the determination of CTOA using a bend specimen is inherently geometry dependent. Recently a technique for measuring CTOA directly in a specimen with a much lower degree of bending has been proposed [36, 37]. This specimen, shown in Figure 7, permits a much longer amount of crack growth than is possible in a Charpy 10 specimen. In addition, it allows a transition to the fully slant growth which is not often seen in the DWTT or Charpy geometries. Ductile failures in Charpy and DWTT tests typically show initial growth by flat tearing followed by a transition to a slant shear on 45° planes. Long propagating ductile fractures in high strength steels typically show only slant fracture. The total absorbed energy will contain contributions from both mechanisms in addition to other losses due to overall bending and indentation. Advanced non-linear numerical damage mechanics models have been calibrated for both flat and slant fracture and used to simulate Charpy tests [38]. The proportions of slant and flat fracture in the simulations were varied from all flat to all shear failure. The predicted energy absorbtion was greatest for all slant fracture and least for all flat fracture. Figure 8 shows the results, where the 100% shear simulation appears to give a good representation of the experimental results. Correction factors derived from this analysis were similar to the empirical values as shown in Figure 6. Further development is required to improve the measurement and modelling techniques and then to transfer the results to full scale tests, but this method appears to be more fundamentally based and could be used for fracture control if it can be developed to a practical production test method. Control and Prediction of Rupture Lengths The final output from the fracture control plan should be the predicted rupture length, or a curve showing the probability of different rupture lengths occurring. Figure 9 shows such a curve, predicted for a typical X80 application using the approach developed on behalf of EPRG by Dawson and Pistone [39]. This analysis implicitly takes account of the scatter in the results (as exemplified by Figures 5 and 6) by deriving distributions for the arrest and propagate toughnesses. These are combined with the toughness distribution of the pipe supply to estimate the proportion of pipe joints which can arrest a crack. This approach has the benefit of being compatible with the trend towards reliability based design and operation of pipelines, but it does require a knowledge of the distribution of pipe toughness. This may not be available for designs which are at the boundaries of pipe manufacturing practice where there is limited data on the variability of production. This analysis, as with others, assumes that the crack arrests within one joint. Researchers at the Japanese High Pressure Line Pipe Committee have developed a dynamic variant of the Battelle two curve approach [40] which can predict when arrest may take more than one joint. At present this has only been validated against a limited data set, but if extended it would offer the potential for demonstrating arrest in situations where the conventional analysis produces marginal results. This would also analyse longer joints, as the present database of fracture test results has been generated on test pipes typically 10 – 12 m in length. It is not clear how these results can be applied to the 18 m joints which are now available from some mills. The Japanese analysis also suggests at that very high pressures, above around 200 bar, the toughness requirements may actually fall. This is due to the decompression behaviour of the gas. If this prediction is correct, this would remove an obstacle to the use of ultra high strength line pipe at very high pressures. Other problems such as the availability of large diameter fittings of the required class rating would still require 11 resolution. At present there are no full scale ductile crack propagation data for pressures above about 180 bar, and so this prediction is unverified. Where it is not possible to guarantee crack arrest using the pipe body toughness, an alternative is to use crack arrestors [41]. These are more likely to be accepted in remote areas where the hazards to public safety from a long rupture length may be low, although the environmental effects of a long rupture must also be considered. The arrestor spacing is likely to be determined on economic grounds, by balancing the additional costs of installing arrestors against the time to repair a long rupture. Design methods for arrestors are not well established, and there is no validation for large diameter pipelines in very high strength linepipe such as X100. CONCLUDING REMARKS This paper has reviewed the issues involved in fracture control for high strength steel pipelines, considering both crack initiation and propagation. Most of the approaches used for fracture control in conventional materials and environments will be appropriate. The major areas where further work is required include: • The resistance of high strength steels to mechanical damage. • Girth weld defect tolerance • Ductile crack propagation and arrest ACKNOWLEDGEMENTS The author wishes to acknowledge input and discussions with many present and former colleagues in Advantica and elsewhere, particularly EPRG and PRCI. REFERENCES CITED [1]. Standards Australia. Pipelines - Gas and liquid petroleum Part 1: Design and Construction AS 2885.1-1997. Homebush, New South Wales: Standards Australia; 1997. [2]. Eiber RJ, Carlson L and Leis B. Fracture control requirements for gas transmission pipelines. In: Pipeline Technology Proceedings of the third international pipeline technology conference Volume I, Brugge, Belgium, May 21-24 2000. Edited: Denys R. Amsterdam: Elsevier Scientific 2000, pp437-53. [3]. Kiefner JF, Maxey WA, Eiber RJ and Duffy AR. Failure stress levels of flaws in pressurized cylinders. 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Probabilistic evaluation of the safety embodied in the EPRG recommendations for shear fracture arrest toughness. 3R International 1998;37(10/11):728-33. [40]. Makino H, Kubo T, Shiwaku T, Endo S, Inoue T, Kawaguchi Y et al. Prediction for crack propagation and arrest of shear fracture in ultra high pressure natural gas pipelines. ISIJ International 2001;41(4):381-8. [41]. Venton PB and Dietsch AE. Design of crack arrestors. In: Proceedings of the International Seminar on Fracture Control in Gas Pipelines. Edited: Rothwell AB. Lidcombe New South Wales: Welding Technology Institute of Australia 1997, pp91-9. 15 0 20 40 60 80 100 60 70 80 90 100 110 120 Grade, ksi T o u g h n e s s ( J ) o r l e n g t h ( m m ) Charpy toughness Crack length Figure 2 Failure pressures (normalized by plain pipe failure pressure) and predictions for ring tension tests on prototype grade X100 pipe. Data from [8]. Figure 1 Variation of flow stress dependent crack length and Charpy toughness required for flow stress dependence in a 36 inch, 12.7 mm wall pipeline operating at 80% SMYS 120 140 160 16 100 120 0 20 40 60 80 -100 -80 -60 -40 -20 0 20 40 Temperature [°C] S h e a r A r e a [ % ] Charpy V surface Charpy V mid-wall DWTT West Jefferson Test Figure 4 Transition curve results from tests on 27.5 mm wall 36 inch grade L450 (X65) linepipe. Data from [26]. Figure 3 Transition curve behaviour of 26 mm wall thickness X80 linepipe; data from [25]. 17 Figure 5 Ductile fracture propagation test database results for X80 and for X100 (data from [7]) plotted using the Battelle short form equation, eq. (6). 18 Figure 6 Ductile fracture propagation test database results for X80 and for X100 (data from [7]) plotted using the Battelle two curve approach. Figure 7 Fractured CTOA test specimen [36] showing transition to fully slant fracture Figure 8 Numerical simulations of instrumented Charpy impact tests on a X100 linepipe showing varying proportions of simulated slant and flat ductile fracture [38]. 19 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 Rupture Length, pipes 10 -7 10 -6 10 -5 10 -4 10 -3 10 -2 P r o b a b i l i t y Figure 9 Prediction of ductile rupture lengths for a typical X80 pipeline application. 10 -1 10 0 20