Failure Mechanism of High Temperature Components in Power Plants

8/12/2019 Failure Mechanism of High Temperature Components in Power Plants

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R. Viswanathan

J. Stringer

Electric Power Research Institute,

Palo Alto, CA 95070

Failure Mechanisms of HighTemperature Components inPower PlantsThe principal mechanisms of failure of high temperature components include creep, fa-

tigue, creep-fatigue, and thermal fatigue. In heavy section components, although cracksmay initiate and grow by these mechanisms, ultimate failure may occur at low tempera-tures during startup-shutdown transients. Hence, fracture toughness is also a key consid-

eration. Considerable advances have been made both with respect to crack initiation andcrack growth by the above mechanisms. Applying laboratory data to predict component

life has often been thwarted by inability to simulate actual stresses, strain cycles, sectionsize effects, environmental effects, and long term degradation effects. This paper will

provide a broad perspective on the failure mechanisms and life prediction methods andtheir significance in the context utility deregulation. S0094-42890000103-1

1 Introduction

Many utility power companies in the U.S. are aggressively pre-paring themselves to face a possible transition from being regu-lated monopoly companies to deregulated free market competi-tors. Reducing the cost of power production is paramount forstaying competitive in the new scenario. Reducing capital costs bydeferring replacement of expensive components and reducing op-erating and maintenance O&M costs by optimizing operation,maintenance and inspection procedures will both be key strategicobjectives for utilities. This poses a significant challenge to thetechnical community since two apparently opposing needs willneed to be reconciled. On the one hand, the need for improvedplant efficiency and availability will dictate more severe and cy-clic duty schedules which result in more severe creep-fatiguedamage and warrant increased attention to the components. On theother hand, the need to reduce O&M costs may result in fewer,shorter, and lower quality maintenance and inspection outages;thus, placing the components at greater risk of failure. The chal-

lenge to the technical community, therefore, is to develop toolsand techniques that will permit more rapid, cost-effective, andaccurateassessment of condition of critical components, both off-line and on-line. In addition to assessing the current condition,these tools must also be capable of evaluating the impact of alter-native strategies for operation, inspection and maintenance. It iscrucial therefore that the high temperature research community bemore intimately familiar with the specific needs of the industry.This paper will bring out some of the industry perspectives re-garding high temperature failures and illustrate them with somefailure examples pertaining to creep, thermal fatigue and em-brittlement. A detailed review of the failure mechanisms affectingthe integrity of utility and chemical plants can be found in 1.

2 Some Industry Perspectives on Failure of Compo-

nents

2.1 Failure Definition. The industrial definition of failure isoften quite different from the textbook definition. A component, inpractice, is deemed to have failed when it can no longer serve itsintended function safely, reliably and economically. Any one ofthese criteria can constitute failure. For example, a steam turbineblade whose tip has eroded, affects turbine efficiency and hence

affects the economics of operation adversely. The blade wouldtherefore be replaced even though it can continue to operate. An-other example is the case of defective piping which could con-

tinue to be operated by decreasing the temperature or pressure butthe resulting loss of output warrants replacement of the piping.Component failures are thus defined in terms of functionalrather than structural failures. Most of the time, replacement ofparts is based on economic considerations, rather than technical.

2.2 The Cost of Failure. Catastrophic failures of compo-nents occur rather infrequently, but when they do, they take aheavy toll on human lives in addition to the costs of repairs,replacement power, and litigation costs. The total cost of the hotreheat pipe failure at the Mojave Power Station in 1986, includinglitigation costs and downtime costs cleanup, pipe replacement,cost of replacement power is estimated to be in excess of 400million dollars. This was in addition to loss of several lives. Fail-ure of most stationary components are generally less costly thanthat of rotating components. For instance, a typical tube leak re-

sults in two days of forced outage, while a header or pipe leakmay result in four days of outage. Major ruptures in piping mayinvolve several months of outage. In the case of major failures,e.g., steam turbine rotor damage, the consequential damage in-volving the wreckage of the entire turbine can be severe. Missilegeneration also needs to be considered. In combustion turbines,breakage of a single first row blade can damage all other down-stream components. In such cases, the unit may be unavailable fornearly six months.

Considering that one day of outage can cost $500,000 in lostrevenue in a 500 MW plant, six months of downtime costs nearly100 million dollars. Unavailability costs can skyrocket even morein the deregulated market during peak demand periods in the sum-mer months when electricity prices may be as high as $5000 perMWh in the spot market for short periods. It is important to re-

member, therefore, that the major cost of failure is invariably thecost of unavailability of the unit, compared to which componentreplacement costs and repair costs pale into insignificance .

2.3 Relative Importance of Crack Initiation VersusGrowth. The general ingredients of a remaining-life-assessmentprocedure for a commonly encountered failure scenario can beillustrated with the help of Fig. 1. Region I corresponds to incipi-ent, microscopic damage events leading up to the initiation of amacroscopic crack. These events include dislocation rearrange-ments, coarsening of precipitate phases, and formation of creepcavities and microcracks. Region II corresponds to propagation ofthe above-mentioned macrocrack. Conventional NDE techniques

Contributed by the Materials Division for publication in the JOURNAL OF ENGI-NEERING MATERIALS AND TECHNOLOGY. Manuscript received by the MaterialsDivision October 15, 1999; revised manuscript received February 15, 2000. GuestEditors: Raj Mohan and Rishi Raj.

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apply to crack detection and sizing in region II. In region II, the

crack grows until it reaches critical size, defined asa c , at whichpoint rapid fracture occurs.

The critical crack size can be defined in a number of ways

based on fracture toughness, ligament size, crack-growth-ratetransitions, or other considerations as appropriate. A commondefinition for many heavy-section components is based on thefracture toughness of the material. The critical crack size, ac, isoften not a constant value but decreases with service exposure dueto embrittlement phenomena. Similarly, many adverse factors canaccelerate the stage II crack-growth behavior, so that the failurepoint is shifted to the leftto shorter times. To perform aremaining-life analysis, information is needed regarding crack ini-tiation, the rate of crack growth, and the failure point, as specifi-cally applicable to the component of interest. Conventional NDEtechniques in stage II are based on the premise that a detectable

crack will form and grow slowly enough to permit periodic in-spections and retirement of the component prior to final failure.There are many instances in which crack initiation alone consti-tute component failure and conventional NDE techniques andfracture-mechanics analyses serve no useful purpose. This is oftena basis of contention between original equipment manufacturersOEM and owners. The OEMs recommend component replace-ment even when no flaws are found, on the premise that the esti-mated value of a c is below the NDE detection capabilities. Anentire vintage of turbine rotors, disks and retaining rings havebeen replaced en masse on this basis, without rigorous technical

justification.Table 1 presents examples of the various circumstances that

might dictate whether component failure is governed by crackinitiation or crack growth. In the case of very brittle materials,

such as the heavily segregated bore of a 30-to-40-year-old rotor,ac may be so small that it is below the limit of detection byconventional NDE techniques. Severely embrittled pressure ves-sels, bolts, and blades may be other examples of this. Highstresses once again have the effect of reducing ac , sometimesbelow levels of detection. If a component has a thin cross sectione.g., a blade or a tube, the remaining ligament can be so smallthat crack propagation is not of importance. In some instances, a cmay be large but the rate of crack growth may be so high thatonce a crack initiates, it reaches critical size rapidly. Many envi-ronmentally induced failures in highly stressed components ex-hibit this behavior. For instance, in generator retaining rings andin steam turbine blades where crack growth under corrosive con-

ditions is encountered, the presence of a pit or pit-like defect iscause for retirement. Initiation of a crack by rapid propagation byfatigue is another example.

In components where failure is governed by crack initiation, thedetection of any defect during an inspection, or, more conserva-tively, the suspected initiation of a crack based on calculations,can be used to retire the component. Many advanced NDE tech-niques which can detect incipient damage evolution prior to crackinitiation are under development industry wide. On the otherhand, many stationary components such as casings, nozzles andheaders are routinely operated with cracks.

Techniques that use crack initiation as a failure criterion in-clude calculations based on history, extrapolations of failure sta-tistics, strain measurements, accelerated mechanical testing, mi-crostructural evaluations, oxide scale growth, hardnessmeasurements, and advanced NDE techniques. For crack-growth-based analysis, the NDE information, results from stress analysis,and crack-growth data are integrated and evaluated with referenceto a failure criterion. The various techniques and their limitationsare described in detail in reference 1.

Analytical models combine operating conditions, materialsproperties and damage rules to estimate the total life consumptionof a component. This approach inherently is inaccurate since op-erating conditions and material properties specific to the compo-nent are usually conservative. The damage rules are also notstrictly obeyed. A greater accuracy in the calculated results isneeded only when crack initiation is the governing mode of fail-ure. On the other hand, most instances of creep fatigue failureshave extensive crack propagation lives and lend themselves toperiodic monitoring by conventional NDE techniques. In suchcases modeling studies to improve the accuracy ad infinitum arenot needed2.

2.4 Need to Know Failure Scenario. Although cracks mayinitiate and propagate by creep or creep fatigue the final failuresmay occur by a different mechanism. In thin section components,the critical crack size may correspond to a wall thickness, liga-ment size or a crack size exceeding which crack propagation israpid. In heavy section components a frequently encountered fail-ure scenario involves crack initiation and growth under steady-state operating conditions at high temperature followed by finalbrittle fracture at low temperatures under start-stop transients.

Loading at low temperatures such as hydrotesting of piping andpressure vessels or overspeed testing of rotating components atlow temperatures can also cause failure.

To illustrate this point, a normal operating sequence, based onthe analysis of the Gallatin HP-IP rotor experience is depicted inFig. 23. Region A consists of a warm-up period after which theroll-off commenced Region B. During roll-off, the rotor wasgradually brought up to speed. Once the synchronous speed wasreached then loading began in Region C. The load was thenslowly increased until steady state was reached after a few hours.The estimated values of the temperature and the tangentialstresses at the bore region of the rotor at the failure location 7throw as a function of time are illustrated in Fig. 2. The transient

Fig. 1 Illustration of a remaining-life-assessment procedurefor a common failure scenario involving crack initiation andpropagation. Aembrittlement phenomena. Bunanticipatedfactors excess cycling, temperature excursions, corrosion,metallurgical degradation, improper material, excessivestresses. See text for definitions of regions I and II.

Table 1 Examples of circumstances governing crack-initiationand crack-propagation controlled failure

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stresses reach a peak value of about 520 MPa 74 ksi and leveloff at a value of 330 MPa 47 ksi at longer times. Approximatevalues ofa c corresponding to the temperature and stress transientdepicted in Fig. 2 can be estimated using the lower bound of theKIC data found in the literature. During the warm up and cool-offperiods, the stresses are small so that the ac values are large,despite the low values of K

IC. Near steady-state conditions, the

reduced values of the steady-state stresses compared to the peaktransient stresses, coupled with the high values of KICobtaining atthe higher temperature lead once again to large values ofa c. Theworst combination of stress520 MPaand fracture toughness71MPam, resulting in minimum values ofa c 0.73 cm, occurs atsome intermediate time, four hours after loading 3, but prior toreaching steady state.

3 Examples of High Temperature Failures

Failure mechanisms at high temperatures include creep, thermalfatigue, corrosion, erosion, and hydrogen attack. In addition, em-brittlement phenomena occurring at high temperatures, e.g., car-bide coarsening, sigma phase formation, temper embrittlement,etc. can facilitate rapid brittle fracture at low temperatures during

transient conditions. This section will describe issues associatedwith creep, thermal fatigue and embrittlement. Overviews onother mechanisms may be found in reference1.

3.1 Creep. Creep damage can take several forms. Simplecreep deformation can lead to dimensional changes that result indistortions, loss of clearance, wall thinning etc. Examples aresteam turbine casings, blades, and piping systems. Localized de-formation can cause swelling and eventual leaks in headers, steampipes and superheater reheater SH/RH tubes. Long term creepfailures generally tend to be brittle failures involving cavitationand crack growth at interfaces and at highly stressed regions. Thecavitation form of damage has been found in SH/RH tubes, rotorserrations, occasionally rotor bores, highly stressed areas in pipingsystems and at weldments. The most common weld failures havepertained to dissimilar welds in superheater/reheater tubing, welds

in headers and in hot reheat and mainsteam piping.

3.1.A Dissimilar Metal Weld Failures. Dissimilar metalwelds are used to join ferritic steel to austenitic stainless-steeltubing and piping in many high-temperature applications in en-ergy conversion systems. The most widespread application is inthe superheaters and reheaters of fossil-fired electric power gen-eration boilers. The materials joined in this application are usuallylow alloy ferritic steels such as 1-1/4Cr-1/2Mo or 2-1/4Cr-1Mo toaustenitic stainless steels such as 304H, 316H, 321H, or 347H. Animportant feature of these welds is that the materials joined havesignificantly different metallurgical and physical properties. Forexample, the materials differ greatly in thermal expansion coeffi-

cient which imposes additional stresses. The welds may be madeeither by induction pressure welding the two materials directlytogether or by shielded metal arc welding with a gas tungsten arcroot pass using austenitic stainless steel filler metals or nickel-base filler metal. Figure 3 illustrates a typical DMW failure.Welds of this type have been in use for a great many years. Dur-ing the late 1950s, failures were encountered with a number ofwelds made using austenitic stainless-steel filler. Typically, failureoccurred by the development of low ductility cracking in the low-alloy steel very close to the weld fusion line. These experiencesprompted research to find better methods of joining the dissimilarmaterials. This led to the use of nickel-base filler metals.

The use of nickel-base fillers to make DMWs effected a con-siderable life improvement. However, by the mid-1970s failureswere also beginning to occur at a significant rate in DMWs madewith nickel-base filler. These failures were generally macroscopi-cally similar in appearance to the failures that occurred in DMWsmade with austenitic stainless steel filler metals. The forced out-age costs resulting from these increased failure rates caused in-creased utility attention, worldwide. In response to this growingconcern, EPRI Research Project 1874 was developed. As a resultof this project our understanding of the failure causes and correc-tive actions is fairly complete. Inspections procedures and life

assessments codes for DMWs have also been developed. A multivolume report describes these results 4, which have also beenwidely disseminated in the public domain 5.

3.1.B Weld Failures in Headers. A schematic illustration ofa header is shown in Fig. 4. A header is essentially a pipe to whichtubes are welded, spaced either axially or circumferentially. Thespacing between the tubes is known as the ligament. In addition tothe tubes, other pipe-to-pipe connections are also present, eitherintegral with the header pipe or welded to it. These branch con-nections can be T-shape connections, as shown in Fig. 4, or of aY-shape configuration. Numerous pipe-to-pipe and pipe-to-tubeweldments are normally present, as shown in the figure.

Initial signs of creep-related distress in headers often appear atweldswelds at stub-tube inlets, long seams, header branch con-

nections or girth butt joints. With the exception of some cases oflong seam welds, and Type IV cracks in girth welds, creep dam-age in welds is invariably manifested on the outside surface ascavities, cracks, or, in extreme cases, steam leaks. Except in re-gard to long seam welds, concern about catastrophic bursts hasbeen minimal. Although weld-related cracking is generally detect-able and repairable, and although it does not have as great animpact on the over-all component life as does header-body base-metal deterioration, it is important from a life-assessment point ofview for the following reasons: Because weld failures are oftenthe forerunners of damage in the body, they can provide an indexof creep damage and remaining life in the base metal. Failure ofwelds at crucial and multiple locations may constitute the end of

Fig. 2 An illustration of a cold start sequence and associatedvariations of stress , temperature T , and critical flow sizeacas a function of time from start

Fig. 3 Typical dissimilar-metal weld locations and failures

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the life of the header, regardless of the condition of the basemetal. The need for frequent weld repair may prove uneconomicaland justify retirement of a header.

Due to the above reasons, creep-damage assessment of weldshas received considerable attention. Damage characteristics atvarious locations are briefly reviewed here on the basis of theextensive information contained in reference 6.

Stub-Tube Welds. Cracking at both the tube and header sidesof stub-tube welds is the most common type of creep damage inhigh-temperature headers. Although such cracking may lead tosteam leakage and forced outages, it is easily detected and re-paired. The cracking may be attributed to any one of severalcauses, including improper seating of the stub tube, inadequatetube flexibility, improper support of the header, bowing, weld-fabrication defects, and locally excessive temperatures. Metallog-raphy in several instances has shown the creep damage to consist

of cavitation and microcracking at the prior austenite grain bound-aries in the heat-affected zone.

Longitudinal Seam Welds. Plate-formed and seam-weldedheaders have been used in some designs. Detailed failure reportsare available for only one incidence 6,7, in which a crack 864mm 34 in. long, in the weld seam had led to a major leak in asecondary superheater outlet header made of 2-1/4Cr-1Mo steelafter 187,000 hours of service. This failure occurred as a result ofcreep-rupture. The damage was confined to the weld metal, withno evidence of damage in the heat-affected zone or base metal.The cracking had apparently initiated just below the outer surface,broke through to the outer surface at an early stage, and thenpropagated to the inner surface. The reason for the subsurfacecrack initiation was believed to be the inferior metal properties atthe location due to a lower carbon content and tempering of the

weld head by subsequent passes. The most severe cracking oc-curred at the centerline of the weld beads, presumably as a resultof impurity segregation during the last stages of solidification.Boat-shaped samples removed at locations away from the crackedarea showed several degrees of cavitation. For each location, a lifefraction consumed could be estimated on the basis of the modeland the plots described later in this chapter. These values wereborne out by subsequent isostress-rupture tests of samples fromthe various locations 6. More recently another failure has beenreported presumably due to locally high temperatures 8.

Girth Welds. Four types of creep damage and cracking asso-ciated with weldments for both headers or piping have been

cataloged by Chan et al. 9. Each of the four creep damage typesare identified below and shown schematically in Fig. 5.

Type IDamage which is longitudinal or transverse in theweld metal and remains entirely within the weld metal.

Type IIDamage that is longitudinal or transverse in the weldmetal, but grows into the surround HAZ.

Type IIIDamage in the coarse-grained region.Type IVDamage initiated or growing in the intercritical zone

of the HAZ the transition region between the fully-transformed,fine-grained HAZ, and the partially-transformed parent basemetal.

Both axial and circumferential cracks have been observed indamaged girth butt welds, with cracking being found in the weldmetal and/or the HAZ. The axial cracking has been attributed tointernal pressure loading and pipe swelling, whereas the circum-ferential cracking has been associated with combined pressure andpiping system loads. Several instances of girth weld cracking hasbeen reviewed6. In one instance, circumferential cracking alongthe coarse-grain HAZ was attributable to stress-relief crackingprior to service. Axial creep cracking across the weld metal hasbeen attributed to a combination of pipe swelling and poor weldductility. Circumferential cracking in the intercritical regions ofthe HAZ has also been observed in both Cr-Mo-V and Cr-Mosteels. This type of cracking, known as Type IV cracking, occursat the end of the HAZ adjacent to the unaffected parent metal.Type IV cracking is generally attributed to localized creep defor-mation in a soft zone in the intercritical region under the actionof bending stresses. Field experience suggests that Cr-Mo-V steelsmay be more susceptible to cracking than Cr-Mo steels and thatoperation at 565C1050Frather than at 540C 1000Fmightfurther exacerbate the problem. Because most of the headers inthe United States are made of Cr-Mo steels and operate at 540C1000F, the problem has not been encountered to any significantdegree.

Fig. 4 Schematic illustration of an elevated-temperatureheader courtesy of B. W. Roberts, Combustion Engineering,Inc.

Fig. 5 Four types of damage in girth welds in relation to mi-crostructure8




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Branch-Connection Welds. Several instances of cracking inbranch-connection welds have been observed. Such cracking hasoccurred on both the header side and the branch side 6, and inboth the HAZ and the weld metal. An example of creep cavitation

in the CGHAZ of a header T section is shown in Fig. 6.

Summary of Creep Cracking in Header Welds. Numerous in-stances of cracking at various locations in header welds, as de-scribed above, have been reviewed by Ellis et al. 6. The salientfacts brought out in this review are as follows. 1 Most weldfailures are creep failures and are clearly evidenced by creep cavi-tation.2Cracks can occur in the weld metal, in the coarse-grainHAZ, or in the intercritical zone of the HAZType IV cracking.3 Cracks in the weld metal are generally attributable to lowerstrength or lower ductility of weld metal. 4 Cracks in the HAZcan arise as a result of hoop stresses, system bending stresses, orresidual stresses due to stress relief. 5 Frequently, the directionof alignment of creep cavities, which is normal to the tensileloading direction, gives a clue to the nature of the system stressesinvolved.

3.1.C Failures in Seam Welded High Energy Piping. Sev-eral categories of pipes carrying high temperature/pressure steamcontain welds that may be of concern. Main steam pipes are pipesthat carry steam at 538565C to the high pressure turbine. Thesepipes are small in diameter and do not contain seam welds. Hence,only girth welds are of concern. The mainstream pipes are how-ever, often connected to the steam header using thick-walled seamwelded piping. In addition, hot reheat pipes which carry steam at538565C but at a lower pressure than the main steam pipetothe reheat IP turbine, and are frequently made of seam weldedpiping. Failure of seam welded pipes used in HRH piping as wellas in header link piping has been of major concern to industry.Failure experience with respect to high energy piping has beenreviewed by Wells and Viswanathan 10.

There have been at least 17 major instances of seam weldedpipe failures including 3 cases of catastrophic rupture, 5 leaks and9 incidents of major cracking. An example of a catastrophic fail-ure is shown in Fig. 7. The failures are generally brittle with a fishmouth appearance.

In the cases of HRH pipes, the welds generally have a double Vconfiguration and the pipes are generally subjected to a normaliz-ing and tempering treatment. The cracking generally initiates sub-surface at the cusp of the double V and then propagates along thefusion line toward the outside and inside, as shown in Fig. 8. Inthe case of the thicker walled header leak pipes, the weld gener-ally has a U geometry and is subjected to subcritical PWHT. Avariety of cracking modes, including fusion line, Type I and Type

IV cracking have been observed. Failures of most of the seamwelded piping have occurred prematurely and could not be pre-dicted based on simple life-fraction rule calculations. Failures oc-cur due to unique combination of operating and metallurgicalvariables. Some of the contributing factors have been identified tobe operating temperature, pressure and cycling; system stresses;weld geometric factors such as configuration, cusp angle and roof

Fig. 6 Creep cavitation in a T-section of a ferritic steel desu-perheater header in a utility boiler

Fig. 7 Rupture in Monroe No. 1 north hot reheat line

Fig. 8 Macrograph of cross-section at location 6LS1, counter-clockwise side of weld sighting along flow; note ID-connected-cracking, located and detected by UT, and extent of cuspdamage




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angle; welding practice employed; inclusion content and creepstrength mismatch, etc. Currently two failure scenarios have beenpostulated. In one scenario, failure is proposed to involve crackinitiation and propagation stages. In the alternative scenario cavi-ties form and grow and eventually link up into a larger crack.Which of these is operative can determine whether NDE basedmonitoring is viable. A comprehensive review of the subject may

be found elsewhere

11,12. A review of literature on Type IVcracking in girth welds and seam welds may be found in reference

13.Since in many of the early instances of girth weld damage, the

damage consisted of evolution of creep cavities into cracks at thecoarse grained heat affected zone CGHAZ, assessment of dam-age consisted of simply classifying the damage and then recom-mending an appropriate action. Damage was classified as Aiso-lated cavities, B oriented cavities, C linked cavities and Dmicrocracking, as per the German practice, Fig. 9 14. Morequantitative correlations between the degree of cavitation and the

creep life expended have been established based on EPR1 re-search. The results shown in Fig. 10 have provided a clearcutbasis for establishing re-inspection intervals. This approach ishowever valid only for Type III cracking in the CGHAZ. Theevolution of damage in the other cases have not been sufficientlyinvestigated.

While replication is very useful for detecting surface damage,many types of failures such as long seam weld and Type IV dam-age in girth welds originate sub-surface. In these cases, replicationalone is not a reliable method to detect damage. In long seamwelds in hot reheat piping and header link piping, high sensitivityconventional or automated UT, focused beam UT or time-of-flightdiffraction UT methods are needed to ensure safety of the piping.In the case of girth welds however, conventional UT seems to beadequate.

Some forms of creep damage are more manageable than others.For example, if Type I, II, or III creep damage is found, the

Fig. 9 Creep-life assessment based on ca?? classification Fig. 10 Correlation between damage classification and ex-pended creep-life fraction for 14Cr-12Mo steels

Table 2 Fossil power plant components involving creep-fatigue as a common failure cause




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subsequent action can range from record and monitor to someform of repair depending on the severity of damage. AdvancedType IV damage is characterized by profuse intergranular cavita-tion in the creep weak area of the HAZ. It has been suggested thatthe evolution of damage from the observation of cavitation byreplication to macro-cracking can be swift and cannot be dealtwith using the German system. In the absence of enough experi-mental evidence regarding damage evolution, the current ap-proach is to replace completely the affected weldment, if anystage of Type IV damage is confirmed 13.

3.2 Creep-Fatigue Failures. Creep-fatigue damage in-

duced by thermal stresses is of major concern with respect to theintegrity of many high temperature components. The concern hasbeen exacerbated in recent years due to cyclic operation of unitsoriginally designed for base load service.

Table 2 is a sample list of fossil plant components in whichcreep-fatigue has been a dominant failure mode. The list is by nomeans complete, since many components not included here mayalso become subject to creep-fatigue if more severe cycling con-ditions were imposed upon them. The purpose of the table is tomake several key points as follows:

1 Creep fatigue damage is generally the result of thermalstresses induced by constraint to thermal expansion during tran-sient conditions. The constraint may be internal such as in the caseof heavy section components e.g., rotors, headers, drums, cas-

ings where thermal gradients arise between the surface and theinterior or vice versa. Internal constraint may also arise from in-ternal cooling of components subject to rapid surface heating suchas in combustion turbineCT blades. The constraint may be ex-ternal such as in the case of joining of thick sections to thinsections or of materials of different coefficients of thermal expan-siondissimilar metal welds. Since stresses are always thermallyinduced, and since crack initiation occurs in less than 103 cycles,this form of creep fatigue damage is also referred variously asthermal fatigue, thermomechanical fatigue and low-cycle fatigue.These terms will be used interchangeably in this paper.

2 This form of creep fatigue damage may involve large plasticstrains achieved locally at stress concentrations such as rotorgrooves, header bore holes, etc. It may also involve primarilyelastic strains combined with stress relaxation, as occurs for com-bustion turbine blades.

3 Table 2 also shows that the industry view of what constitutes

failure is different for stationary components such as headers andcasings and rotating components such as blades/rotors. In theformer case, cracks are tolerated and crack initiation is believed tooccur early 1020 percent life in life-component retirement istherefore based on economics of repeated repairs and growth of acrack to a critical allowable size. Hence, excessive concern withrefining the damage rules is unwarranted in such cases. In rotatingcomponents, such as CT blades and rotor grooves, crack initiationdefines failure since upon crack initiation other failure modes suchas high cycle fatigue may intervene and cause rapid failure. Inthese cases, more refined prediction of damage evolution and

crack initiation would be useful.

Detailed review of literature shows that there are divergentopinions regarding which damage approach provides the best ba-sis for life prediction. It is quite clear that a number of variables,such as test temperature, strain range, frequency, time and type ofhold, waveform, ductility of the material, and damage character-istics, affect the fatigue life 2. The conclusion drawn in anyinvestigation may therefore apply only to the envelope of materialand test conditions used in that study. The validity of any damageapproach has to be examined with reference to the material andservice conditions relevant to a specific application. Broad gener-alizations based on laboratory tests, which often may have norelevance to actual component conditions, do not appear to beproductive. Thus, one should use a tailored, case-specific ap-proach for any given situation.

In most instances of fatigue, the temperature varies along withthe strain, giving rise to what is known as thermomechanical fa-tigueTMF. Depending further on when the hold time is super-imposed, various cycle shapes are possible. In the past, thermalfatigue traditionally has been treated as being synonymous withisothermal LCF at the maximum temperature of the thermal cycle.Consequently, life-prediction techniques have evolved from theiso-thermal LCF literature. The assumed equivalence of isother-mal LCF tests and TMF tests has been brought into questions as aresult of a number of studies.

High tensile strains at high temperature IPwould favor creep,whereas high tensile strains at low temperature OPwould favorcracking of oxide and hence accelerated environmentally induceddamage during subsequent high-temperature exposure. Hence,Kuwabara et al. rationalized that in case of materials where dam-

age is driven by creep, IP cycles would be more damaging than

Fig. 11 Ligament cracking at a tube bore hole viewed from the ID of a header




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OP cycles, for a given strain range15. In other materials, wherethe environmental contribution is significant, OP cycles may bemore damaging than IP or isothermal LCF cycles. In addition toenvironmental effects, differences also arise between cycles interms of the relaxed mean stresses. The relative severity or the

different cycles can also change with material ductility, maximumtemperature and hold time. Consequently, a simple classificationof material behavior is not possible.

A case in point is the ligament cracking encountered in CrMosteel header pipes illustrated in Fig. 11. Cracks initiate in the tubebore holes and are oriented parallel to the axis of the tube borehole. Linking up of cracks between holes on the inside surface ofthe header leads to propagation to form cross ligament cracks.Presence of ligament cracking has been observed in a very largenumber of superheater headers in the U.S. The cracking mode hasbeen identified as creep fatigue. A computer code, B oiler L ife

E

valuation and S

imulation S

ystem BLESS developed recently,incorporates two alternate approaches for predicting crack initia-tion; one involving an inelastic linear damage summation method,and a second approach involving repeated cracking of oxide scaleand oxide notching16. For a variety of cycle histories, the Codepredicts crack initiation occurring in about 20,000 h by the oxidecracking mechanism. The creep-fatigue damage summation ap-proach on the other hand, is inconsistent with the early initiationof cracks observed in headers. Metallography of cracked headershas shown numerous oxide spikes, see Fig. 12, indicating oxidecracking to be the crack initiation mechanism. This exampleclearly illustrates the need for using appropriate thermomechani-cal fatigue data simulative of actual component cycles in predict-ing crack initiation life of components.

Another example of the critical need for TMF data is in the case

of protective coatings. In the case of coated components such ascombustion turbine blades, cracking of the coating leads to loss ofenvironmental protection from the coating, and, eventually, tocracking of the base metal Fig. 13. The integrity of the coatingdepends upon both the ductility of the coating and the strain-timehistory of the coated blade, as shown in Fig. 14. The strain-to-cracking of the coating is a strong function of temperature, oftengiven by a ductile-to-brittle transition temperature DBTTcurve.The strain-temperature cycle in the engine must lie below theDBTT curve. For the example shown in Fig. 14, coating A willnot crack under normal duty, but will crack during an emergencyshutdown. However, coating B will not crack under either condi-tion. In view of the DBTT behavior exhibited by coatings, it be-comes even more critical in evaluating coated components that

Fig. 12 Oxide notching at ligament cracks

Fig. 13 TMF cracks in GT29CoCrAIYcrating penetrating the INCO 739 basemetal




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thermo-mechanical fatigue tests be performed. Furthermore,simple TMF cycles, in which the maximum tensile strain is madeto coincide with either the peak temperaturein-phase, IPor withthe lowest temperature out-of-phase, OP in the cycle, will leadto unrealistic results. If one were to compare the performance oftwo coatings A and B by conducting an isothermal LCF test or anIP type thermomechanical fatigue strength, one would concludethat both coatings perform well, since both coatings have highstrain capability at high temperature. If the same comparison weremade under more relevant TMF conditions, coating B would bechosen over coating A. Hence, TMF cycles simulative of actualblade cycles must be performed to evaluate the effect of the coat-ing. Such test data on coated components is extremely scarce inthe open literature, and is limited even in proprietary data bases.

One of the major problems in evaluating the applicability ofdifferent life-prediction methods is that in many cases it is neces-sary to use all the available data in deriving the life-predictionmethod, and thus it is not possible to examine the accuracy withwhich a given method describes data not used for the developmentof the method, or outside of the range of conditions, or lives,considered. There also is a scarcity of instances in which serviceexperience has been compared with specific life-prediction meth-ods. In general, the available methods are utilized only to predictthe lives of samples tested under laboratory conditions. Validationagainst component test data in the laboratory and inservice moni-toring of actual equipment would lead to more confidence in theuse of the various rules.

Results from most studies show that even the best of the avail-able methods can predict life only to within a factor of 2 to 3.

Some of the cited reasons for these inaccuracies have already beendiscussed. Some additional reasons are: failure of the methods tomodel changing stress-relaxation and creep characteristics causedby strain softening or hardening, use of monotonic creep datainstead of cyclic creep data, and lack of sufficiently extended-duration test data. All of the damage rules available today are atleast partly, if not totally, phenomenological in nature. They allinvolve empirical constants that are material-dependent and diffi-cult to evaluate theoretically. Extrapolation of the rules to materi-als and conditions outside the envelope covered by the specificinvestigation may result in unsuccessful life predictions and usu-ally result in predictions whose accuracy is difficult to evaluate.One form of extrapolation that is especially difficult to evaluate is

the need to use short time data for long time service. For compo-nents whose service time is from 3 to 30 years, and which mayoperate continuously for hundreds to thousands of hours, use oftest datawhich lasts as much as four months 1/3 of a yearandhas hold times as much as 16 hoursrequires extrapolation of oneto two orders of magnitude. The statistical methods chosen tomake these extrapolations significantly affect the estimated livesof the components. Furthermore, in performing these acceleratedtests, either the strain, the temperature, the frequency, or somecombination must be increased over actual service conditions inorder to produce failure in a reasonable amount of time. Thus,there is always the danger that the physical mechanism of failurein the laboratory test is different from that during service, or thatimportant aspects of the service conditions are not considered dur-

ing the laboratory tests. For application to service components, thestress-strain variation for each type of transient and its time de-pendence must be known with accuracy. The importance of usingrelevant TMF data cannot be overemphasized. This realization hasled to several recent studies in life prediction of combustion tur-bine components using TMF based algorithms.

4 Summary and Conclusions

Creep and creep-fatigue are the principal failure mechanismsaffecting the integrity of components operating at elevated tem-peratures. Creep damage in weldments poses major challengesboth in analytically calculating it and in experimentally reproduc-

ing it. Several alternative damage locations and mechanisms havebeen observed which are often difficult to reproduce in laboratorytests. Fusion line cracking and fine grain heat affected zones FG-HAZcracking has led to catastrophic failure of high energy pip-ing. Thermomechanical fatigue TMF or creep fatigue affectsmany heavy section components as well as internally cooled com-ponents such as combustion turbine blades. It is important thatresearchers focus on component specific rather than genericlifeprediction models with a full understanding of the applicable fail-ure definition, failure scenario and relevant duty cycle. Futureresearch needs to address advanced NDE techniques, on-linemonitoring techniques, TMF mechanisms, and evolution of dam-age and growth of cracks in welds.

Fig. 14 Typical thermomechanical cycle for a first-stage blade, showing leading-edge strain and temperature variations for normal start-up and shut-down, and anemergency shutdown




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References

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Failure Mechanism of High Temperature Components in Power Plants

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