Performance of P91 Thick Section Welds Under Steady and ... · PDF filePerformance of P91 Thick Section Welds ... expertise in materials development and defect assessment ... Information

Performance of P91 Thick Section Welds OMMI (Vol. 1, Issue 3) Dec. 2002

Performance of P91 Thick Section Welds Under Steady and Cyclic Loading Conditions: Power Plant and Research Experience I A Shibli, European Technology Development, UK [email protected]

Dr Ahmed Shibli has over twenty five years of experience in power plant R&D and life assessment/extension issues, with particular expertise in materials development and defect assessment under creep and creep-fatigue conditions. His previous experience includes working for British Steel, Mitsui Babcock Energy, ERA Technology and as a consultant/expert for the UK and European power plant companies. At present he is leading a number of large projects on behalf of the European Commission and international industry including a new planned initiative on the assessment of P91 welded components.

Note: This Paper is based on a recent review carried out by ETD on the use of 9Cr martensitic steel in power plant both as thick section components (headers, steam pipes etc.) and thin section tubing in superheater plant sections. In this paper only the behaviour of thick section welded components will be discussed. ABSTRACT Limited plant experience indicates that although the relatively new ASTM P91 martensitic steel is highly regarded and its use as thick section components (such as headers, steam pipes etc.) has been very beneficial to the power plant industry in raising plant temperature and efficiency, this steel, like its predecessor low alloy ferritic steels, may be equally prone to Type IV cracking. This can result in failures relatively early in life. Information from limited R&D work on cross-weld specimens and the testing of feature specimens of P91 appears to support this view. The R & D work also indicates that creep fatigue interaction under plant cycling conditions may adversely affect the performance of this type of steels. The Paper suggests that now that the 9Cr steel use has become somewhat wide spread both in conventional steam raising and CCGT plant there is a risk that P91 steel may start showing mid-life Type IV cracking problems and thus the behaviour of such steels needs to be better understood and procedures and tools developed for repair. 1. BACKGROUND AND INTRODUCTION By 1987 it was apparent that P91 did not suffer from reheat cracking and had enough creep ductility to resist weld creep cracking. However, it was recognised by this time that P91, like all ferritic steels, is potentially susceptible to Type IV cracking, e.g. [1].


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The basis for this review was that the P91 steel has been in use, in a limited number of plants, since the late eighties/early nineties, and that although the steam temperatures in most of these older units were between 540 - 565°C (below the maximum design temperature of 600°C for this steel), there have been some disturbing reports, admittedly few in number, about P91 failures. All of these failures were of the Type IV variety in which cracking took place in the fine grain section of the HAZ. These failures are a matter of concern especially because they occurred at lower temperatures than the maximum design temperature of 600°C for such steels and the present operating temperature of up to 590°C in some of the new higher efficiency European and Japanese plant where 9Cr martensitic steels have been introduced more recently. A prime aim of this review was to forecast whether this could presage a series of mid-life problems in units using this material. Furthermore, assuming that operators may experience a number of failures in future the review outlined potential repair weld procedures, although the weld repair aspect is not covered in this paper. The review also covered the performance of dissimilar metal welds and of thin section tubing in both steam and flue gas environments but again these aspects are not discussed here. Market forces and competition are now dictating the power plants to be cycled on daily basis, to follow load demands and maximise profits. This “cycling” is driving a redesign of existing and new plants. One redesign solution has been to use higher strength P91 steel for pressure vessel construction which allows pressure-containing components to be made with thinner sections. Thinner components require less time to reach thermal equilibrium and have fewer tendencies towards thermo-mechanical damage mechanisms. In addition, the reduced section thickness increases pipework flexibility. Because of this potential benefit a number of operators/owners of the existing plants are substituting these new higher strength steels for the older materials, especially when a plant is moved from base load to cyclic operation. In fact in the case of new HRSGs it is understood that some manufacturers offer P91 as the only material for thick section components. However, recent limited R&D work throws some doubt on the validity of the claim that the relatively thinner section components of P91 may be less susceptible to fatigue cracking and this aspect has been discussed in this paper. 2. P91 THICK SECTION COMPONENT FAILURES 2.1. Recent Published Failures Five/six known incidents of cracking and failures occurred in the UK which has the longest experience of service (installtion of P91 thick section section headers in the UK started in 1989). All of these failures occurred in the West Burton power plant. As usual, after a few years of confidentiality, information on these failures has now been published [2, 3, 4, 5]. Four of these failures occourred in ‘bottle’ type joints, and one in a header end cap with another header end cap showing cracking during later investigations/ inspections. The first bottle failure occurred after a service of only 20,000 hours while the end cap failure occurred after a service of mere 36,000 hours. In all cases the nominal operating temperature was 565oC. It should be emphasised that these were the longest serving thick wall components in the UK and perhaps


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elsewhere, as the West Burton Power Station was the first to receive replacement components of P91 when this programme started in the UK in the 1980s. As [2, 3] show the failure was attributed to the overtempering of the base material. This was inferred from the low hardness of the failed casts compared with the most of the other P91 material but was not traced back to the heat treatment records presumably due to their non-availability. High aluminium to nitrogen ratio, albeit within the Standards specfications, was also reported to be the potential culprit as this is thought to result in the formation of coarse aluminium nitrides leaving less nitrogen for other creep strngthening fine nitrides precipitates. However, no direct proof in terms of finding higher amount of aluminium nitride precipitates is known to have been found. In the case of the end cap failures/cracking design involving sharp change of section near the weld section was also blamed although such design has been successfully used in P22 vessels for a long time and was in accordance with the British Standards design specifications.


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Table 1: Summary of West Burton Transition Bottle Failures (UK) Plant West Burton Power Plant

Section Superheater Component Superheater T91 Transition Bottle Service Hours 20,000 hrs Steam Temperature

565ºC

Steam Pressure 15.0 MPa Description of Component

A 75 mm diameter F91 forged rod was bored through to an ID of 35.5 mm and turned to an OD of 60.5 mm. A portion of the length was then further machined to an OD of 48.3 mm with a smooth taper between OD’s creating a transition piece dubbed a superheater “bottle.” The large diameter end of the fitting was butt-welded to 60.5 mm diameter x 12.5 mm wall T91 tube, while the smaller diameter end was butt-welded to 48.3 mm diameter by 6.35mm wall T91 tubing.

Description of Failure

Four such transition pieces failed, the earliest at 20,000 hrs, by cracking through approximately half the circumference of the smaller diameter end of the transition piece in the Type IV band of the heat affected zone.

Failure Analysis Results

There was no evidence of poor welding or installation procedure, but there was evidence that there may have been a deficiency in the heat treatment of the forged stock material.

Fig. 1: Transition bottle failures

P91 60.5 mm Ø 12.5 mm wall

T91 48.3 mm Ø 6.4 mm wall

F91 "Bottle"

Crack


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Table 2: Summary of West Burton Header End Plate Failure (UK) Plant West Burton, UK

Section Superheater Component Superheater Header Service Hours 36,520 hrs with 469 hot starts and 72 cold starts Steam Temperature 565ºC Steam Pressure 15.0 MPa Component The superheater header is sketched below. Description of Failure

Type IV cracking through the HAZ of the F91 forged endplate. Corresponding location on other headers showed severe creep damage.

Failure Analysis Results

Zone IV of the HAZ in the F91 material was found to be approximately 20 HV units softer than Zone IV of the P91, and stress analysis showed that the stress at the origin of the crack was 212 MPa and was as high as 110 MPa at a depth of 5-10 mm into the end-plate forging. It was concluded that the cause of failure was Type IV Creep.

(a)

b) Figs. 2 (a), 3 (b). Header end-cap cracking

242 mm

33 mm

80 mm

56 mm

Crack

P91

F91 P91

F91

Crack in an unfailed header endplate, as detected by NDE (Courtesy: West Burton PS)


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2.2. Unpublished Failures Reportedly more recently there have been at least three other P91 Type IV failures in power plant in Europe but these, as happened in the case of the West Burton plant, are not expected to become public for a few more years due to the plant confidentiality. 3. CREEP RUPTURE AND TYPE IV FAILURES As is well known, the introduction of P91 to the power plant industry was based on the rupture strength of the parent metal. Only short-term weld metal data were available at the start of the use of this steel. However, this situation is now changing. Some of the work carried out as early as late 1980s showed that the cross weld rupture strength of this steel fell much below that of the base metal [1]. Recently a number of other publications on 9Cr Martensitic steels have confirmed this view, e.g. [6,7]. This is shown in Figure 3 for the European 9Cr martensitic steel, E911. Stress, MPa

Rupture Time, Hours

Fig. 3. Rupture strength of E911 cross weld specimens showing Type IV cracking compared with the base metal rupture strength for E911 [7]


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More recently Allen et. al. through their work on E911 steel have shown that the weld strength reduction factor for this steel could be as high as 40% when the failure is in Type IV position [6]. They extrapolated it to 50% for failure to 100,000 hours [6]. This compares unfavourably with 20% strength reduction factor for P22 steel at 550oC – its usual temperature of operation [8]. The criticism of the work on E911 [6] can be that tests carried out on cross weld specimens were conducted at 625oC, higher than the service temperatures. Some work on P91 at 600oC (maximum design temperature for this type of steel) carried out in the project SOTA [9] also showed that P91 specimens tested in creep crack growth mode were more vulnerable to Type IV cracking than P22 specimens tested at 550oC. The UK analysis of experience with low alloy ferritic steels used in power plants is shown in Figure 4. This shows that although the Type IV failures started appearing after ~ 20000 hours most of the failures in these steels appeared after 50,000 hours of service i.e. the Type IV failures were a medium to long term service phenomenon. In the UK where experience with the use of thick section P91 has been the longest (about 45,000 hours to date), even ignoring unpublished failures, at least five failures and one incidence of cracking has manifested itself. This in spite of the fact that the service temperature was only modest (568oC) for this steel and the frequency of use is also very low compared with the conventional ferritic steels.

Fig. 4. Type IV weld failures in conventional ferritic steels [10]


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4. CREEP-FATIGUE/ HIGH TEMPERATURE FATIGUE ISSUES 4.1. Creep and Thermal Fatigue Cracking Phenomenon Many of the power plant components go through some form of fatigue cycling due to normal shutdown and re-start, load following, etc. However, because of privatisation and competition in the electricity industry power plants are often started up to meet peak demand and shut down or partially cooled down during the less demanding times. This has enhanced the risk of cracking due to fatigue. In principle, since creep is both time and temperature dependent, plant cycling (also known as ‘two shifting’) and low load operation would be expected to reduce damage due to long term creep. During a unit start, or at periods of low load running, there may be some circumstances when temperature overshooting or localised overheating may occur. Whilst these problems are well known to plant operators, it should be recognised that the cumulative effect of repeated overheating during thermal and load cycling can give rise to extended periods of operation above the design temperature which may result in accumulation of creep damage or its acceleration. A creep related phenomena in ‘two shifting’, when using conventional low alloy ferritic steels has been degradation in the microstructure and the concomitant reduction in material properties which has already occurred during the previous term of base load operation. These microstructural changes will have occurred simply as a result of exposure to temperature. The most obvious signs of such degradation are the onset of spherodisation in carbon- manganese and 2.25Cr1Mo steels. The implications of this are usually reduced ductility, compared to virgin material, and reduced resistance to creep and/or fatigue cracking. Similar phenomena can occur in 9Cr martensitic steels where the precipitates can coarsen or reprecipitation can occur and the matrix can loose creep strengthening elements. However, by far the most common problem experienced as a result of plant cycling is thermal fatigue damage. This can manifest either in the form of cracking of an individual component or by the mechanical failure of structures. Cracking of a component is attributed to severe thermal gradients arising from excessive steam to metal and through wall temperature differences associated with rapid rates of change of steam temperatures as generally observed during start up, shut down and load changes. The principal components at risk typically comprise any thick walled sections such as boiler superheater headers, steam pipework, valves, HP and IP steam chests and turbine inlet belts. HP heaters and economiser inlet headers are also frequently exposed to similar effects due to rapid cooling by cold feed water. On a wider scale, structures such as boiler framework and tube attachments, boiler supports and pipework support systems are also vulnerable to thermal cycling. Thin walled sections such as boiler tubes, reheater headers etc. are less prone to the problem. A well known example of thermal fatigue cracking in a header ligament is shown in Figure 5.


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Fig. 5. An example of a header ligament cracking

4.2. Damage accumulation due to creep -fatigue interaction Materials behave in a complex way when both creep and fatigue mechanisms are present. They usually act synergistically to cause premature failure. Creep strains can reduce fatigue life and fatigue strains can reduce creep life. The American Society of Mechanical Engineers (ASME) recognised the effects of interaction and provides guidance on the interaction between creep and fatigue and its effect on the life expectancies of materials. (See ASME Cases of ASME Boiler and Pressure Vessel Code Case N-47, Rev 29, 1990). Figure 6 demonstrates the interaction and consequences of creep and fatigue for P22 - a typical power plant steel.

cracking


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Figure 6. Interaction and consequences of creep and fatigue (based on ASME N-47) for a typical power plant steel (2.25Cr1Mo) The creep-fatigue interactions are not currently well defined, and the limit line as shown by the ‘elbow’ line in Figure 6 represents the design life limit, expressed in fraction of material creep life and fatigue life. This limit line is used to establish the effect of combining the two mechanisms and demonstrates how they act together to reduce the effect of the individual processes. Original design criteria assumed that the two processes were entirely independent. The line is a highly conservative representation of the phenomenon. It does however serve to demonstrate the effects of the interaction. By way of an example, consider a component originally designed for say 10,000 cycles, which might have been designed to operate in a unit which two-shifts on a daily basis over 30 years. Assume also that the component operates in the creep range and was designed for 150,000 hours operation. If the unit were to operate on a base load regime, it will of necessity, accrue some thermal cycles, probably of the order of 1000 over the projected life. The thick ‘continuous base load operation’ line indicates the effective operation of the component. The actual component life is given by the point where it intersects the ‘elbow’ line. This shows a reduction of the component life to about 75% of its predicted creep life. Similarly, if the component operates on a two shifting unit with 300 cycles per year whilst operating in the creep range, the actual life may be as low as 40% of the anticipated fatigue life, as shown by the intersection of the ‘continuous two shift operation’ line. Where operational cycling is introduced on a former base load unit, it can be seen that the residual life can be greatly reduced to between 40% and 60% of the original design life due to the combined effects of creep and fatigue.


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The key implication is that older units designed for base-load operation and used in this capacity over many years are very susceptible to component failure when they are eventually forced to cycle regularly. This logic obviously applies to components made from both the traditional ferritic steels and the new 9Cr martensitic steels. Thus, while increases in failure rates due to cycling may not be noted immediately, critical components will eventually start to fail. Shorter component life expectancies will result in higher plant EFOR (equivalent forced outage rate), longer scheduled outages, and/ or higher capital and maintenance costs to replace components at or near the end of their service lives. In addition, it may result in reduced total plant life or more capital to extend the life of the plant. The above example, although only figurative, demonstrates how plant cycling originally designed for base load conditions can significantly reduce the integrity of components to a far greater degree than might otherwise have been anticipated. It should be noted that this methodology is highly conservative and does not take account of the timing or amplitude of the thermal cycles which may further influence the remnant life of components. More refined analysis methods are currently being developed. 4.3. Creep-Fatigue Capabilities of High Temperature Alloys As explained above, creep-fatigue interaction of high temperature alloys can reduce life in a non-linear manner. Figure 7 (due to ASME) shows that interaction for P22 steel is quite severe compared to some of the other alloys. The ASME curve in Figure 7 is for base metal but the same rule is expected to apply to the welded components.

Fig. 7. Creep fatigue interaction of some of the high temperature alloys (ASME)


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As discussed earlier, some of the older plant are now using or planning to use P91 as a replacement material and new high efficiency advanced power plant are using 9Cr martensitic steels for thick section components. This is considered useful from fatigue damage viewpoint in the belief that thinner section components made from high strength P91 will be less prone to fatigue cracking during the cyclic operation due to smaller temperature gradients across the wall thickness. However, recent creep and high temperature fatigue crack growth work in the European Commission funded project HIDA [11] on welded P91 and P22 components has suggested that the creep-fatigue interaction in P91 components containing welds could be even more sever than P22, Figure 8. These tests were conducted at constant temperature but the load was cycled.

It must, however, be emphasised at this stage that, although fairly detailed, this work has been carried out on only one cast and at one test temperature of 625oC. Further work on this material is now planned. However, it should also be stated that tests in HIDA, both on large feature specimens (both circumferential and seam welded pipes) and laboratory specimens of the fracture mechanics type have shown the same i.e., that cycling can be more harmful to P91 than to P22. Full details of this work are described in a number of recent publications [12, 13].

Fig 8. Creep-fatigue crack growth curves for some of the high-

temperature alloys (including P91) The work involving feature specimen / pipe tests showed that even very low cycling (0.0001 Hz - three cycles a day with about seven hours hold time in each cycle) of large welded pipes adversely affected the crack growth rate in P91 HAZ. The adverse effect of fatigue is not surprising as work on the P91 cross-weld rupture specimens tested in the HIDA programme has shown that when fracture occurs in the Type IV position the ductility of these specimens could be as low as 1 to 2%. Mohrmann et. al. [14] carried out low cycle fatigue (LCF) tests on welded E911 specimens with and without hold times at room temperature and 600oC. These tests showed a decrease in lifetime by a factor of about two compared with the base


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material. The hold times further reduced the lifetime. Similarly Bicego et. al. carried out LCF tests at 600oC on cross weld specimens of P91, P92 and E911 steels [15] – all 9Cr martensitic steel variants - and showed similarly a life reduction of about 2 compared with the base metal. In summary, the above work shows that: a) creep fatigue interaction effect on P91 weldments can be more severe than on the conventional P22 weldments, and b) the creep-fatigue interaction effect can be more severe on the weldments than on parent metal. The creep-fatigue interaction effect is related to the steady creep strength issues (discussed in Section 3) as failure in both cases was observed to be in the Type IV position. There is now independent evidence that the Type IV position in P91 can be vulnerable to cracking/failure. Thus recent work by Tabuchi et. al. [16] has shown that M23C6 precipitates and Lave phases form faster at the fine grain HAZ region in 9Cr martensitic type steels (compared with the other regions of the weldment) and this makes the Type IV position in these steels very vulnerable.

5. DISCUSSION Thick section P91 components have only seen limited service durations, the longest being that in the UK. From this experience it appears that thick wall components may be subject to Type IV failures even at lower service temperatures than those experienced in the new higher performance ultra supercritical power plant where 9Cr martensitic steels were introduced only a few years ago. The thick wall UK failures of header end caps, at 36000 hours duration, were at first attributed to end cap design which put the weld close to the high stress level position of the change in section (Fig. 2a), although P22 steel has happily experienced the same British Standards design. However, four other failures of ‘transition bottles’ even at shorter service durations of 20,000 hours, where the weld was situated away from high stress position, made the plant operators look for other possible reasons. A comparison of the failed and some of the unfailed components finally made them conclude that a hardness difference of about 10 VPN (Vickers pyramid hardness number also designated as VH) could have been responsible for such failures, the failed components showing a hardness of 10 to 20 VPN below 200VPN. This lower hardness was attributed to the possible over-tempering heat treatment of the failed casts. The other possible reason was cited to be the low N/Al ratio, albeit still within the Standards specifications, of the failed casts compared with the unfailed casts. This has been attributed to the observation that aluminium nitride creates large size precipitates often situated on grain boundaries, which impair creep ductility very significantly. The AlN precipitation can reduce the fine VN particle precipitation which contributes to the creep strength of this steel. Thus, for example, the creep rupture strength in 100,000 hours of a high Al cast of 12Cr martensitic steel was observed to be lower by as much as 30% compared with other casts with lower Al contents [17] indicating that coarse AlN particles do not contribute to precipitation strengthening. The lower rupture strength of low N/Al level steels is also supported by other research findings which have shown higher creep strength for high N level 9Cr steels [18], although, this work has pointed out that this effect is only seen in short term tests.


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Although a large amount of test work has been carried out on P91 base metal and cross weld specimens little work is known to have been carried out on thick wall welded feature specimens. The exception to this appears to be the HIDA work where tests on both butt and seam welded pipes showed similar findings [13]. These were that: (a) P91 appears to be more vulnerable to Type IV cracking than P22 and (b) that even low cycle fatigue results in enhancement of crack growth in P91 fine grain HAZ (i.e. Type IV position) compared with the steady state crack growth [13]. The limitation or criticism of this work can be that it was carried out on one cast and at one slightly accelerated (625oC) test temperature while the maximum envisaged temperature for the use of P91 is 600oC. However, in spite of the limited nature of the work the uniqueness of the work means that its findings cannot be ignored until results at service temperatures from planned future work [19] on large wall feature test specimens become available. There is of course some support of the HIDA work from [14, 15] as discussed earlier, which shows that even at 600oC low cycle fatigue cycling reduces the life of cross weld specimens by a factor of 2 compared with the base metal. In the project HIDA, the low cyclic life was attributed to the very low cross weld rupture ductility of the 9Cr steels. The finding by Tabuchi et. al. [16] is of interest where it has been shown that during the creep process M23C6 precipitates and Lave phases form faster at Type IV position which will make this position more vulnerable. Indeed it is possible that West Burton failures occurred because of the heavy duty cycling that this power plant (see Table 2) has been undergoing, like many other UK plant. The lower creep rupture strength of the weaker casts due to overtempering or low N/AL ratio or both might have accelerated this process. The question therefore that needs to be asked is: Is it possible that stronger casts may also start showing Type IV cracking at longer service durations or at higher service temperatures? It is well known that 9Cr martensitic steels, like the 12Cr martensitic steels with which some European countries (notably Germany and Denmark) have a much longer experience, are very sensitive to tempering and welding heat treatments. According to V&M Tubes the tempering and PWHT temperatures for 9Cr martensitic steels must be in the correct regime (750-760oC) to give the desired rupture strength [20]. At higher tempering temperatures overtempering will occur resulting in lower rupture strength, and at lower temperatures ductility and other aspects will not be satisfactory. In the case of PWHT the lower PWHT temperature will mean that the hardness difference between Type IV position and weld metal/ coarse grain HAZ will be very high and will make Type IV position more vulnerable [21]. The strict observation of the heat treatment regime of this steel during steel production and the welding cycle is so important that Elsam in Denmark, as one of the more successful users of this steel as thick section components, have developed their own strict criteria of material quality check and control. This is outlined in the original ETD P91 Survey Report [21]. In summary, they ensure compliance of all heat treatment of the base and weld metals to European standards and checking of all heat treatment records in detail including the accuracy of the thermocouples used by the material supplier. This is followed by random checks on hardness and dimensions of the supplied material. Where the heat treatment of any material supply appears to be suspect Elsam carry out metallographic checks including transmission electron microscopy to observe precipitates and microstructure in general and compare it with the normal material.


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In conclusion, both the plant experience and research work shows that it is possible that P91 may start showing mid-life crisis at least in the case of the some of the more vulnerable casts being used around the world, especially in the plant subjected to cyclic operation. It is thus important that further work is undertaken to better understand the behaviour of the welded thick section components of P91 and to develop repair and replacement strategies. 6. CONCLUSIONS So far P91 has been considered to be advantageous steel as thinner wall and smaller component size reduce thermal gradients and hence the adverse effect of fatigue cracking - which in the past has been experienced in thick section components made from the traditional low alloy ferritic steels. However, preliminary R & D studies have shown that P91 welded components may be equally, or even more, prone to Type IV cracking compared with the conventional low alloy ferritic steels, both under creep and creep-fatigue conditions, possibly due to the weakness of this zone and low cross-weld creep ductility of the welded material. From the limited plant experience it appears plausible that cracking could start showing in thick wall components at a rather early stage. This experience also indicates that some casts of this steel may be more vulnerable than others due to the differences in their heat treatment or chemical composition. It is thus important that preparations should now be made for appropriate inspection and repair strategies of plant using this steel. 7. REFERENCES 1. Middleton C, Metcalfe E, ‘A review of laboratory Type IV cracking data in high chromium ferritic steels’, Paper C386/027, Published in IMechE Proceedings, London, 1990. 2. Brett S J, Allen D J and Pacey J, “Failure of a modified 9Cr header endplate”, Proc. Conf. on "Case Histories in Failure Investigation", Milan, Sept. 99, pp. 873-884. 3. Allen D J, Brett S J, “Premature failure of a P91 header endcap weld: minimising the risks of additional failures”, Proc. Conf. "Case Histories in Failure Investigation", Milan, Sept. 1999, pp.133-143. 4. Brett S J, ‘Identification of weak thick section modified 9Cr forgings in service’, Published in the CD version of the Proceedings of the Swansea Creep Conference; Organised by the University of Swansea and EPRI, and held in Swansea, UK, April 2001. 5. Brett S J, ‘The creep strength of weak thick section modified 9Cr forgings’, Published in the Proceedings of Baltica V, Vol. 1, June 2001. 6. Allen D J, Fleming A. ‘Creep performance of similar and dissimilar E911 steel weldments for advanced high temperature plant’, Published in the Proceedings of the


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5th Charles Parsons 2000 Conference on ‘Advance materials for 21st century turbines and power plant’, 3-7 July, Churchill College, Cambridge, UK. Pp 276-290. 7. Orr J, Buchanan L W, Everson H, ‘The commercial development and evaluation of E911, a strong 9% CrM0NbVWN steel for boiler tubes and headers’, Published in the Proceedings of the Conference ‘Advanced heat resistant steel for power generation’, Editors: R Viswanathan, J Nutting, Conf. held in San Sabastian, Spain, Publisher: IOM Publications, London, UK. 8. Etienne C F, Heerings J H, ‘Evaluation of the influence of welding on creep resistance (strength reduction factor and lifetime reduction factor), IIW doc.IX-1725-93, TNO, Appledoorn, The Netherlands. 9. Shibli I A, ‘Creep and fatigue crack growth in P91 weldments’, Published in the Proceedings of the Swansea Creep Conference; Organised by the University of Swansea and EPRI, and held in Swansea, UK, April 2001. 10. Parker J D, ‘The creep and fracture behaviour of thick section, multi-pass weldments’, Conference Proceedings on ‘Integrity of high temperature welds’, Nottingham, UK, 1998, Published by Professional Engineering Publishing, UK. 11. Shibli, I A., ‘Overview of the HIDA project’, Published in the Proceedings of the 2nd International HIDA Conference on ‘Advances in Defect Assessment in High Temperature Plant’, 4-6 Oct. 2000, held at MPA, Stuttgart, Germany. 12. Shibli I A, Le Mat Hamata N, Gampe U, Nikbin K, ‘The effect of low frequency cycling on creep crack growth in welded P22 and P91 pipe tests’, Paper S4-4, Published in the Proceedings of the 2nd International HIDA Conference on ‘Advances in Defect Assessment in High Temperature Plant’, 4-6 Oct. 2000, held at MPA, Stuttgart, Germany. 13. Le Mat Hamata N. and Shibli I. A. Creep Crack Growth of Seam-welded P22 and P91 Pipes with Artificial- Part 2: Published in the Proceedings of the 2nd International HIDA Conference on ‘Advances in Defect Assessment in High Temperature Plant’, 4-6 Oct. 2000, held at MPA, Stuttgart, Germany. 14. Mohrmann R, Hollstein T, Westerheide R, ‘Modelling of low-cycle fatigue behaviour of the steel E911’, Published in ‘Materials for advanced power engineering 1998’, Proceedings of the 6th Liege Conference, Volume 5, Part 1. 15. Bicego V, Bontempi P, Mariani R, Taylor N, ‘Fatigue behaviour of modified 9Cr steels, base and welds’, Published in ‘Materials for advanced power engineering 1998’, Proceedings of the 6th Liege Conference, Volume 5, Part 1. 16. Masaaki Tabuchi, Takashi Watanabe, Kiyoshi Kubo, Masakazu Matsui, Junichi Kinugawa and Fujio Abe, ‘Creep Crack Growth Behaviour in HAZ of Weldments for W containing High Cr Steel’, Published in the Proceedings of the 2nd International HIDA Conference on ‘Advances in Defect Assessment in High Temperature Plant’, 4-6 Oct. 2000, held at MPA, Stuttgart, Germany.


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17. Kubon Z, Foldyna V, Vodarek V, ‘Optimised chemical composition of 9-12% Cr steels with respect to maximum creep resistance’. Published in ‘Materials for advanced power engineering 1998’, Proceedings of the 6th Liege Conference held in 1998, Volume 5, Part 1, p 375. 18. Wachter O, Ennis P J, ‘Investigation of the properties of the 9%Cr steel of the type 9Cr-0.5Mo-1.8W-V-Nb with respect to its application as a pipework and boiler steel operating at elevated temperatures’, Ph.D. Thesis, Research Centre Julich, Germany, March 1995. 19. Shibli, I A , Planned future work for P91 investigations on feature testing, weld repairs and the behaviour of dissimilar metal welds. Three ETD proposals for Groups Sponsorship, [email protected] 20. Vallourec & Mannessmann Tubes, The T91/P91 Book, 1999. 21. Performance Review of P91 and other 9Cr Martensitic Steels: Plant and R&D Issues, ETD Report No. 1009-IIP-1002, November 2001.

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Performance of P91 Thick Section Welds Under Steady and ... · PDF filePerformance of P91 Thick Section Welds ... expertise in materials development and defect assessment ... Information