techniques for simulating creep damage evolution at...

TECHNIQUES FOR SIMULATING CREEP DAMAGE EVOLUTION AT WELDS WITHEMPHASIS ON EVALUATING LONGITUDINAL SEAM PEAKING IN

HIGH-TEMPERATURE PIPING SYSTEMS

Phillip E. Prueter, P.E.The Equity Engineering Group, Inc.

Shaker Heights, Ohio [email protected]

ABSTRACT

Realistically simulating the creep response of welded com-ponents can help quantify the risk associated with operating in-service, high-temperature equipment and can validate new com-ponent designs in the power generation and petrochemical indus-tries. Detailed finite element analysis (FEA) is employed in thisstudy and is coupled with generalized, non-linear creep simula-tion techniques to investigate the elevated temperature responseof welds. Depending on original heat treatment, creep damageprogression is known to be accelerated by the mismatch in prop-erties of the base metal, weld deposit, and heat affected zone(HAZ). This mismatch results in stress intensification that canaccelerate creep damage near a weldment (typically in or adja-cent to the HAZ). In this paper, the effect of implementing anelastic damage parameter that adjusts the stiffness of the materialas a function of creep damage is examined. This type of dam-age mechanics model has a significant impact on the predicteddamage evolution near weld deposits and can realistically mimicobserved in-service failures. Additionally, commentary on dif-ferent creep damage failure criteria is provided. The simulationspresented utilize the Materials Properties Council (MPC) Omegacreep methodology, with particular emphasis on the behavior ofhigh-temperature, low chrome (1-1/4 Cr 1/2 Mo) piping withlongitudinal weld seam peaking.

Application of these techniques to high-temperature, lowchrome piping is relevant as there have been numerous relatedcatastrophic failures in the power generation and petrochemicalindustries attributed to weld seam peaking. Commonly, weldpeaking occurs during fabrication due to angular misalignment

of rolled plate. Furthermore, for many fusion-welded piping fab-rication standards, no tolerance for peaking is specified. Lo-cal peaking can induce significant local bending stresses, andfor components that operate in the creep regime, the presenceof peaking can lead to an increased risk for creep crack initia-tion, propagation, and eventual rupture. An overview of somewell-known historical low chrome piping failures is providedin this paper, and a literature review on existing creep analysisand peaking measurement methodologies is offered. Addition-ally, the remaining life of low chrome piping systems is esti-mated and the sensitivity in results to variations in key param-eters is highlighted; these parameters include operating temper-ature, magnitude of peaking, and the effect of heat treatment.The simulation techniques discussed in this paper are not onlyvaluable in estimating remaining life of in-service components,but detailed analysis can help establish recommended weld seampeaking fabrication tolerances, appropriate manufacturing prac-tices, and practical inspection intervals for high-temperature pip-ing systems.

INTRODUCTIONDepending on original heat treatment of a welded compo-

nent, creep damage evolution is known to be accelerated by themismatch in properties of the base metal, weld deposit, and heataffected zone (HAZ). As discussed in [1], this mismatch in ma-terial properties can result in stress intensification and increasedtriaxial stress that can accelerate creep damage near a weldment;this damage typically occurs in or directly adjacent to the HAZ.

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Proceedings of the ASME 2018 Symposium on Elevated Temperature Application of Materials for Fossil, Nuclear, and Petrochemical Industries

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Local weld seam misalignment or peaking can induce substantiallocal bending stresses in the pressure boundary of any componentfabricated from rolled plate (including pressure vessels and largediameter piping). As described herein, for welded componentsthat operate in the creep regime, the presence of local peakingcan lead to an increased proclivity for creep crack initiation orpropagation and eventual rupture of the pressure boundary. Fur-thermore, as outlined in this study, computational finite elementanalysis (FEA) can be coupled with creep test data and repre-sents a useful tool in realistically simulating the creep damageevolution at welds, with or without longitudinal seam peaking.

Generally, local weld seam peaking occurs during the pipeor cylindrical vessel manufacturing process due to angular mis-alignment of the rolled plate at the weld seam location (wherethe pipe locally deviates from a proper circular cross-section). Asketch highlighting the concept of local longitudinal weld seampeaking is shown in Figure 1.

Figure 1. Sketch of Angular Misalignment in a Cylindrical ShellLongitudinal Weld Seam [2].

Additionally, for many fusion-welded piping fabricationstandards, no specific tolerance for longitudinal weld seam peak-ing is specified. Furthermore, for many in-service piping compo-nents with longitudinal weld seams, the amount of angular mis-alignment at the welds is unknown and no documented peakingmeasurements are available. Trying to measure peaking in thefield can be challenging due to access limitations, removal of in-sulation, and for large piping systems, the sheer length of pipingcan make inspection of weld seams costly and impractical. Thismakes quantifying the risk associated with elevated temperatureoperation of low chrome piping systems with longitudinal weld

seams problematic. Given these limitations, computational anal-ysis can offer perspective on remaining life sensitivity to peak-ing and the overall risk for creep damage accumulation. Thesegeometric unknowns, along with variations in original heat treat-ment and material properties also makes performing FEA-basedstudies challenging. Numerous documented failures in the powergeneration and refining industries of low-chrome piping systemswith weld seam peaking that operated in the creep regime (asdiscussed in more detail below) highlight the potential for catas-trophic failure, especially in the absence of known peaking mag-nitudes and inspection data.

An example of a creep-related crack like flaw occurring ad-jacent to the long-seam weld deposit in a 36-inch catalytic re-former pipe (1-1/4 Cr 1/2 Mo that was originally normalizedand tempered after welding) is shown in Figure 2. The micro-graphs shown in this figure show a crack morphology indica-tive of creep voiding at the crack tip (meaning creep is the likelycause of the crack propagation through-thickness). The effectsof original heat treatment on overall damage tolerance and theevolution of creep-related crack-like flaws near the weld depositare discussed herein.

Figure 2. Micro-Graphs Showing a Creep-Related Crack-Like Flawin the Long-Seam of 1-1/4 Cr 1/2 Mo Catalytic Reformer Piping.

DOCUMENTED INDUSTRY FAILURESReference [3] indicates that early failures of 1 1/4 Cr - 1/2

Mo components such as super-heater outlet headers and pip-ing components operating in the creep regime were attributed tocracking. In 1968, due to several in-service failures, the ASMECode reduced the time-dependent allowable stresses for 1 1/4Cr - 1/2 Mo materials. The allowable stresses at 1,000F and



1,050F were reduced by 16 and 26 percent, respectively [3].This implies that headers and piping components operating inthe creep regime fabricated during the 1950s and 1960s are po-tentially under-designed. A second decrease in allowable stressestook place in the 1989 addenda of the ASME Code [4], where theallowable stresses for 1 1/4 Cr - 1/2 Mo decreased from 6.9 ksito 6.3 ksi at 1,000F and from 4.6 ksi to 4.2 ksi at 1,050F. Thischange was prompted by some of the industry failures duringthe 1980s (discussed below). Additionally, ASME B31.3 intro-duced a weld joint strength reduction factor (W ) in the 2004 Edi-tion [5]. The intent of this factor is to account for long-term be-havior of welds at elevated temperatures in the absence of creeptests (above 950F). Furthermore, ASME B31.1 introduced thisparameter in the 2008 Edition [6].

There have been several catastrophic failures of low chrome,long-seam welded piping systems in the petrochemical and elec-tric utility industries due to long term creep. Major contributorsto these failures included local stress increase from long-seamweld peaking geometry from fabrication and the effects of creepstrength (rate) mismatch of base metal, weld heat affected zone(HAZ), and/or weld deposits at the welded joint. Long-seamfailures have been observed in chrome-moly welds given sub-critical PWHT as well as those that are normalized followed bytempering. The Mohave steam pipe failure (1 1/4 Cr1/2 Mo al-loy) in 1985 and Monroe (2 1/4 Cr - 1 Mo alloy) in 1986 showeddirect evidence of a progression of cracking from sub-surfaceinitiation points in regions of high stress multiaxility due to lo-cal creep property mismatch and joint geometry. The crackingwas preceded by extensive cavitation in the weld HAZ or at theweld fusion line facilitated by the local triaxial stress state as-sociated with the creep rate property mismatches in the weldzone. Essentially, the effects of the different material propertyzones at the weldment lead to stress intensification and triaxialtension that accelerates the rate of cavity growth near the weld-ment (occurs in or adjacent to the HAZ). These systems operatedat roughly 1,000F. It has been confirmed in postmortem inves-tigations (documented in WRC Bulletin 475 [7]) that the creepdamage initiated subsurface due to high triaxial tension and inlong-seam welded piping. That is, the creep damage progressedto an advanced state without any apparent evidence on the out-side surface.

Additional failures in the electric utility industry led to thedevelopment of Electric Power Research Institute (EPRI) inspec-tion guidelines first published in 1987 [8], now in its fourth edi-tion [9]. The original edition provided inspection recommenda-tions for long-seam piping using ultrasonic and metallographicreplication. The current edition now includes detailed summariesof multiple industry failures that have ensued over the years sincethe document was first published and outlines acoustic emissionand advanced ultrasonic inspection techniques.

Low chrome piping failures have also occurred in the re-fining industry. In a 1995 failure investigation, by Dobis etal. [10–12], of a refinery (catalytic reforming service) long-seam

piping failure, the metallurgical investigation and creep testingprogram indicated that the 1-1/4-Cr 1/2-Mo (normalized andtempered) pipe material exhibited greater than average creepstrength and creep ductility. Nevertheless, the pipe failed afteronly 100,000 hours at a nominal hoop stress of roughly 6 ksiwith an operating temperature range of 970F to 1,000F. Re-sults from subsequent detailed finite element stress analyses ofthe failed pipe indicated very high bending stresses were presentin the pipe due to peaking at the long-seam weld. Furthermore,this accelerated the damage at the HAZ fusion line until an 18-inch long through-wall crack developed. This study highlightedthat peaking on the order of 1/8 inch can result in a stress increaseof about 2.5 times the nominal hoop stress from pressure; thiswas subsequently determined to be a significant contributor tothe premature failure. Other documented failures of long-seam,low chrome piping in catalytic reforming service have been doc-umented over the years and as discussed in [10–12], all of thesefailures in the refining industry were due to cracks that initiatedat the inside surface, with essentially no visible external damage.

LITERATURE REVIEW OF COMPUTATIONAL STUDIESMultiple computational studies have been performed to eval-

uate the creep response of low chrome piping components. Sev-eral recent publications by Prueter et al. [1, 13], that relate tothe numerical analysis results presented herein, outline a FEA-based approach for estimating remaining creep life in high-temperature, longitudinal seam-welded piping systems. A com-putational creep-fatigue study and critical flaw sizing method-ology on a low chrome reformer piping tee is offered in Ref-erence [14]. Brown et al. performed an FEA-based study onpeaked piping welds in Reference [15]. Additionally, Iwamotoet al. [16] examined void evolution behavior using FEA-basedsimulations of chrome-moly welds and Chang and Xu [17] in-vestigated multiaxial creep behavior of materials used in powerplants. An illustration of creep damage analysis on high-pressuresteam pipelines using the Materials Properties Council (MPC)Omega method is provided in Reference [18], and a review oftime-dependent material properties of pipeline steel is renderedin [19]. Reference [20] presents a study that attempted to estab-lish a flaw size acceptance criteria for 1-1/4-Cr 1/2-Mo catalyticreformer piping operating at high temperatures. This paper em-phasizes that little-to-no local peaking could be tolerated in thepresence of a crack-like flaw. Additionally, several papers thataddress the topic of applying acoustic emission testing to mon-itoring in-service, low chrome, elevated temperature piping aredocumented in References [21–23].

CREEP METHODOLOGYThe Materials Properties Council (MPC) Omega creep

model, outlined by Prager in Reference [24], is employed in thisstudy (and discussed in Part 10 of API 579 [2]) and has the gen-



eral form:

log10εco = −

((Ao + ∆sr

Ω ) +

[1

460 +n T

][A1 + ...

...+A2(Sl) +A3(Sl)2 +A4(Sl)

3

]) (1)

where εco is the initial creep strain rate, A0 through A4 are theOmega strain rate coefficients (typical values are defined in An-nex F of API 579 [2] for a given material), ∆sr

Ω is the adjustmentfactor for creep strain rate to account for the material scatter band(a range of -0.5 to +0.5 can be used), nT is the evaluation tem-perature for the nth time increment, and Sl is the log base 10 ofthe effective stress (nσe) as defined below:

Sl = log10(nσe) (2)

Additionally, the effective stress for the nth time increment is afunction of the principal stresses following:

nσe =1√2

[(nσ1−nσ2)2+(nσ1−nσ3)2+(nσ2−nσ3)2

] 12

(3)

Furthermore,

log10Ω =

((Bo + ∆cd

Ω ) +

[1

460 +n T

][B1 + ...

...+B2(Sl) +B3(Sl)2 +B4(Sl)

3

]) (4)

where B0 through B4 are the Omega parameter coefficients(again, typical values are defined in Annex F of API 579 [2] for agiven material), Ω is the uniaxial Omega damage parameter, and∆cd

Ω is the adjustment factor for creep ductility (a range of -0.3to +0.3 can be used). Additionally,

Ωm = ΩδΩ+1n + αΩnBN ,

Ωn = max

[(Ω− nBN ), 3.0

] (5)

where Ωm and Ωn represent multiaxial and uniaxial Omega dam-age parameters, respectively. The term αΩ is a triaxiality param-eter based on the state of stress, and equal to 3.0 for pressurizedspheres or formed heads, 2.0 for pressurized cylinders or cones,and 1.0 for all other components and stress states. The MPCOmega parameter (δΩ) is established as shown below,

δΩ = βΩ

[(nσ1 +n σ2 +n σ3)

nσe− 1.0

](6)

where βΩ is the MPC Project Omega parameter (equal to 0.33)and the Bailey-Norton Coefficient (nBN ) is defined following:

nBN = −

([1

460 +n T

][A2 + 2A3(Sl) + 3A4(Sl)

2

])(7)

From these equations, the remaining life (nL) at stress level nσe,temperature nT , and for the nth time increment, nt, takes theform:

nL =1

εcoΩm(8)

Finally, the accumulated creep damage (Dc) for the nth time in-crement is calculated per the following equation (again, nt is de-fined as the time increment where the component is subjected toa stress level of nσe):

Dc =

N∑n=1

ntnL

(9)

It is noted that the Omega creep model does not directlyinclude primary creep, it conservatively bounds primary behav-ior. According to wording taken directly from Part 10 of API579 [2], Primary creep is not included in the model [Omega]described above. Primary creep at the effective stresses typicallyassociated with construction code designs is extremely small anddoes not contribute to life shortening creep damage in a signifi-cant way. In addition, at high stresses, the Omega model, whilenot treating primary behavior explicitly does predict acceleratedcreep rates and assigns damage to the relaxation associated withthis creep. This treatment of relaxation is more conservative than



letting primary creep relax without damage. It is recognized thathigh principal stresses occur at discontinuities. For those multi-axial situations the effective stresses are reduced and significantprimary creep does not occur.

In the FEA simulations described herein, the Omega creepmethodology discussed above is utilized using a specializeduser subroutine that calculates creep strain, damage accumula-tion, and stress relaxation throughout the model (accomplishedthrough a viscoelastic analysis step). Essentially, creep strainrates are calculated as a function of stress and temperature atdifferent points in the FEA model and composite damage is es-tablished based on the user-defined loading histories. The sub-routine also permits the user to account for changes in the elasticmodulus of the material as damage accumulates. This is accom-plished through the introduction of an elastic damage parameter(De) where the elastic modulus (E) of the material is adjusted asa function of damage at any point (D) throughout the model astime elapses following:

E = E0(1−De),

De = 0 if D < Dcrit,

De =D −Dcrit

Dm −Dcritif D ≥ Dcrit

(10)

where E0 is the undamaged elastic modulus and Dcrit (the criti-cal value of damage below which there is assumed to be no elas-tic damage), and Dm (the maximum prescribed creep damagethreshold up to which the relationship between D and De is as-sumed to be linear) are defined by the user. Lastly, creep materialproperties of base metal, HAZ, and the weld deposit can be dif-ferentiated in FEA using this approach. These properties can bebased on creep test data directly, where Omega and strain rate co-efficients can be adjusted or fit to experimental test data for basemetal, HAZ, or weld deposit specimens. As discussed in Refer-ence [25], full-scale HAZ test specimens can be developed forcreep testing using Gleeble simulation techniques (where ther-mal cycles are applied to a specimen to mimic the behavior ofthe HAZ during welding or stress relief).

LONGITUDINAL WELD SEAM PEAKINGThe severity of longitudinal weld seam peaking in pressure

vessels or piping systems is a critical factor when consideringboth creep and fatigue response. From a fatigue perspective,pressure swing adsorption (PSA) vessels are known to be sus-ceptible to cracking failures at peaked long-seams. Long-seampiping components with peaking are also known to be prone tocreep failure as highlighted in the examples above. Peaking in-troduces a bending stress local to the peaked weld seam, withhigh stress on the outside diameter (OD) of the vessel in the case

of inward peaking and high stress on the inside diameter (ID) inthe case of outward peaking. Normally, peaking is caused dur-ing the initial rolling and welding of the plates into cylindricalcomponents. Part 8 of API 579 [2] provides calculation meth-ods for determining the additional elastic bending stress due tolocal peaking. This closed-form method is given below and istypically used to evaluate vessel long seams (pressure loadingonly) for protection against failure from cyclic loading (fatigue).The ratio of induced bending stress from local peaking to appliedmembrane stress from internal pressure (Rcljab ) is defined as:

Rcljab =6δ

tcCf (11)

where the maximum local peaking magnitude is δ (defined be-low), tc is the corroded wall thickness, and Cf is a correctionfactor for angular misalignment that can be determined from Fig-ure 3. An additional parameter that is required to determine Cfis Sp (defined below):

Sp =

√12(1− ν2)PR2

Eyt3c(12)

In this equation, ν is Poisson’s ratio, P is internal pressure, R isthe mean radius of the shell, and Ey is the elastic modulus.

Figure 3. Correction Factor for Angular Weld Misalignment in theLongitudinal Joint of a Cylindrical Shell [2].

For cylindrical or spherical shells, the angular misalignment



at a given joint can be established by using a centered or rockedtemplate as shown in Figure 4. Commonly, the rocked templateis recommended. The arc length of the template should extendbeyond the locally deformed region resulting from the angularmisalignment and should be established using the inside or out-side radius of the cylinder, as applicable. The template shouldbe notched as shown such that no part of it is in contact with thearea of weld reinforcement (weld cap). Using the rocked tem-plate technique, the maximum deviation can be calculated usingthe following:

δ =b1 + b2

4(13)

Figure 4. Method of Measurement to Determine the Extent of LocalPeaking in Longitudinal Weld Seam [2].

Given the extreme sensitivity in remaining life predictionsto peaking severity, understanding the margin of error associatedwith peaking measurements is critical in establishing the risk forpremature failure from induced bending stresses. Furthermore,

for many in-service piping systems, it is often not practical, noris it feasible to acquire peaking measurements along the entirelength of the longitudinal welds due to access limitations (of-ten requiring extensive scaffolding) and the sheer length of manypiping systems. As discussed herein, ranking the relative riskassociated with different levels of peaking using FEA and es-tablishing past and anticipated future operating conditions is oneway to attempt to quantify the propensity for failure of a certainpiping system.

Many owner-users have found that using the conventionalcentered or rocked templates to measure peaking magnitudes of-ten produces unreliable results that are not consistent or repeat-able. Specifically, the measured peaking magnitude depends onthe width of the template. A recent study by Cayard and Geisen-hoff [26] proposed a novel peaking template (see Figure 5).

Figure 5. Suggested Pivot Template Design for Peaking Measure-ments from Reference [26].



This study points out that based on numerous peaked pipegeometries that were evaluated using conventional peaking tem-plates, the estimated magnitude of peaking was sometimes un-derestimated by as much as 50 percent and measured results werevery dependent on the width of the template. Based on these lim-itations, a proposed pivot template design is suggested in Refer-ence [26], as shown in Figure 5. This proposed peaking measure-ment approach involves the use of a hinged template with a dialgage, and reportedly was able to produce accurate and repeatablepeaking measurements.

FEA SIMULATIONSIn this study, 30 and 36 inch (0.375 and 0.430 inch nomi-

nal thicknesses, respectively), 1 1/4 Cr - 1/2 Mo (ASTM A691-85a material specification [27]) catalytic reforming piping (Class22) systems with long-seams is examined. The original materialspecification has no specific limits for longitudinal weld seampeaking, although tolerances for out-of-roundness are included(current specifications do not explicitly limit long-seam peak-ing). The specification also indicates that double-welded, fullpenetration long-seams are to be used. In this case, the designtemperature of the line is 1,000F and the design pressure isroughly 150 psig. The material was subjected to stress relief viasub-critical PWHT (not normalized and tempered), which per thematerial specification [27], should occur at hold temperatures be-tween 1,150F and 1,375F. To realize the effect of local peakingon bending stress (for both pipe geometries), the above closed-form solutions are used to calculate that added bending stress asa function of peaking magnitude, as shown in Figure 6.

Figure 6. Influence of Local Peaking on Bending Stresses for the36-inch and 30-inch Piping Systems at Design Pressure.

This figure shows the severe increase in calculated bending stress

as local peaking magnitude increases. Furthermore, for outwardlocal peaking, the elevated bending stress is maximum on theinside surface of the cylinder.

Since the pipe material was not normalized and temperedafter welding in this case, adjusted Omega properties are em-ployed. For the weld deposit, standard 1 1/4 Cr normalized andtempered properties (from API 579 [2]) are employed. Standardannealed 1 1/4 Cr creep properties (from API 579 [2]) are em-ployed for the base metal. HAZ creep properties are adjustedto achieve debited creep properties: the creep rate is increasedby roughly a factor of 4 compared to typical annealed 1 1/4 Crmaterial (accomplished by increasing the magnitude of the firststrain rate coefficient) and Omega is decreased by roughly 25percent (accomplished by increasing the magnitude of the firstOmega coefficient). A half-symmetry, two-dimensional general-ized plane strain FEA model is constructed as shown in Figure 7(an out of plane axial force, proportional to the internal pressure,is applied to simulate axial stress in the pipe cross-section). Theweld deposit, HAZ, and base metal are accounted for by parti-tioning the model to represent a typical double-V weld.

Figure 7. Finite Element Model Material Definitions.

This analysis assumes an idealized cross-section of the long-seam welded piping. Specific weld geometry, including amountof penetration, weld cap contours, actual HAZ contours, etc.,are not known (and would vary for different piping cross sec-tions). Thus, the idealized weld geometry presented in Figure 7is assumed. Furthermore, actual material property details (forthe base metal, HAZ, weld deposit) and original welding andPWHT documentation are not available (as is often the case forin-service components). Reasonable estimates are used, and non-linear viscoplastic analysis is carried out as described below.

There are several parameters in this study that influence



creep response and remaining life predictions, including oper-ating pressure, temperature, magnitude of peaking, and treat-ment of material properties. Parametric studies were carried outto investigate all of the parameters, and some of these compar-isons are discussed herein. The first set of comparisons exam-ines overall sensitivity to peaking magnitude for design condi-tions (roughly 150 psig at 1,000 F) for a duration of 500,000hours. The contour plots below highlight how the extent of creepdamage increases as local peaking magnitude increases. Figure 8compares creep damage for varying levels of local peaking, as-suming the material properties discussed above (lower creep re-sistance assumed for the HAZ).

Figure 8. Influence of Local Peaking Magnitude on Creep DamageAccumulation after 500,000 Hours at Design Conditions (1,000F).

In Figure 8, the consequence of additional bending stressesfrom peaking on creep damage is evident; even 1/16 of an inchof local peaking causes a significant portion of the HAZ to reach100 percent damage (this amount of sensitivity to local peak-ing is consistent with other published literature such as Refer-ence [20]). In this figure, it is noted that the 1/4 inch peaked sim-ulation failed to converge due to excessive damage after roughly350,000 hours. Even the model with no peaking reaches 100 per-cent damage locally in the HAZ. Of course these simulation re-sults reflect design conditions (worst-case operation for 500,000

hours). The effect of operating at lower temperatures is notableas outlined below. Additionally, to evaluate the sensitivity tooperating temperature, several cases are considered for the 1/8inch peaked pipe at design internal pressure: 1,000F, 975F,950F, and 925F. Figure 9 shows creep damage accumulationafter 500,000 hours for these different constant operating tem-peratures. This figure demonstrates the significant sensitivity tooperating temperature.

Figure 9. Effect of Temperature on Creep Damage Accumulationafter 500,000 Hours at Design Pressure (1/8 inch of Local Peaking).

To compare the effects of original heat treatment on creepresponse of the long-seam weld region, a simulation with uni-form normalized and tempered Omega properties is compared tothe model with the above mentioned creep properties (adjustedto approximate the behavior of the HAZ for typical stress re-lief via PWHT) in Figure 10. Normalizing and temperating afterwelding is assumed to remove the property mismatch betweenthe HAZ and the surrounding base metal and weld deposit, es-sentially eliminating the tendency for accelerated damage accu-mulation in the HAZ. As expected, removing the creep propertymismatch between the weld deposit, HAZ, and base metal in theFEA results in a significant improvement in damage accumula-tion for the normalized and tempered model. The local damageaccumulation on the ID near the weld centerline (symmetry plane



in the normalized and tempered FEA model) is very likely an ar-tifact of the assumed weld geometry. That is, an ID weld cap thateliminates the relatively sharp corner or stress concentration as-sociated with the idealized weld geometry would likely result ineven less creep damage accumulation at the center of the weld.

Figure 10. Effect of Original Heat Treatment on Creep Damage Ac-cumulation after 500,000 Hours at Design Pressure (1/8 inch of LocalPeaking at 975F).

In the weld HAZ or at the weld fusion line, damage accu-mulates faster for the simulated PWHT material properties; thisaccelerated damage is facilitated by the local triaxial stress stateassociated with the creep rate property mismatches in the weldzone. These trends relating to peaking magnitude, operating tem-perature, and material heat treatment are observed in the non-linear simulations of both 30 and 36 inch pipe configurations.

DEFINITION OF FAILUREFundamentally, the damage progression for all of the mod-

els presented herein starts with very local elevated damage in theHAZ (slightly subsurface near the ID due to increased triaxial-ity) and gradually extends to a larger through-wall portion of theHAZ and surrounding weld deposit and base metal. The excep-tion to this is for 1/4 inch local peaking or higher; for this amountof peaking, damage near the middle the weld (through-thicknessdirection) at the bottom of the V-geometry reaches 100 percentcreep damage very close to the same time, or slightly before theHAZ location slightly sub-surface near the ID. To better comparedamage progression and to attempt to quantify the differences inoverall risk associated with varying operating temperatures andmagnitudes of local long-seam peaking for the 36 inch pipe inquestion at design pressure conditions, Tables 1 and Tables 2 areshown below. Table 1 shows the time (in hours) to reach 100 per-cent creep damage in the HAZ slightly sub-surface near the in-side diameter (location A in Figure 7). Similarly, Table 2 shows

the same comparisons only at a location near the middle of theHAZ at the bottom of the V-geometry (location B in Figure 7).

Table 1 – Summary of Time (Hours) to Reach 100 PercentCreep Damage in the HAZ (ID).

Table 2 – Summary of Time (Hours) to Reach 100 PercentCreep Damage in the HAZ (Middle).

The challenge associated with quantifying the risk of fail-ure for a low chrome piping system with long-seams based oncomputational creep simulation results is a function of the defi-nition of failure. Applying the results of numerical simulationsto predicting creep failure is not straight forward. The difficultyis in defining what truly constitutes failure. Failure could be de-fined as damage accumulation reaching 100 percent at a point;however, this is likely an overly conservative assumption. Alter-natively, failure could be defined as creep damage accumulation



reaching 100 percent over a finite length or through-wall damageaccumulation above a given threshold. For several of the docu-mented steam line industry failures previously discussed herein,a progression of cracking from sub-surface originations in re-gions of high stress multiaxility due to local creep property mis-match and joint geometry were cited as the cause of failure. Fur-thermore, this phenomenon is difficult to fully quantify becauseFEA techniques and numerical predictions are by no means ab-solute. Additionally, there are many unknowns that must also beconsidered such as weld geometry (weld cap contour, amount ofpenetration) and any potential lack of fusion following fabrica-tion. Weld filler metal properties and variations in welding tech-niques are also factors. To this end, a point in the HAZ reaching100 percent damage, or close to it, may indicate susceptibilityto creep crack initiation, but may not be indicative of imminentfailure. As discussed below, inspection can be used as a tool toprovide insight and to better understand the risk associated withcreep damage progression that ultimately leads to failure.

DAMAGE EVOLUTION & ELASTIC DAMAGE MODELTo account for changes in the elastic modulus of the mate-

rial as creep damage accumulates, the introduction of the elas-tic damage parameter as described in Equation 10 above is im-plemented in the FEA. Figure 11 illustrates how accounting fordegradation in material stiffness as a function of creep damageaffects the overall damage progression through the HAZ. Thisfigure shows creep damage contours for a model with 3/16 inchlocal peaking after 225,000 hours of operation at design pressureand 975F. The upper plot is for the model with no elastic dam-age parameter and the lower plot corresponds to a model with thecritical damage threshold (Dcrit) set to 0.1.

In this case, as creep damage surpasses 10 percent, the elas-tic modulus begins to linearly decrease to zero between 10 and100 percent damage (see Figure 13). While using a damagethreshold of 10 percent may not be truly representative of thewelded region (that is, in reality elastic modulus may not start tosignificantly decrease until creep damage is higher and closer toend-of-life), this example highlights how implementing an elas-tic damage parameter can change the overall damage progressionthrough the HAZ at a welded joint. Commentary on the chosenelastic damage threshold is provided below. The challenge asso-ciated with selecting the appropriate damage parameter in com-putational simulations is that material testing is often requiredto establish the correct elastic modulus scaling techniques. Fur-thermore, this damage parameter is also material dependent, withlikely dependency on original heat treatment as well.

Figure 12 shows the damage progression at varying timesfor the same model with an elastic damage threshold (Dcrit) of10 percent shown in Figure 11. This simulation reaches non-convergence at roughly 225,000 hours; this is an artifact of thedamage that has progressed through the entire thickness of theHAZ and reached a magnitude of 100 percent.

Figure 11. Effect of Elastic Damage Parameter on Creep DamageAccumulation after 225,000 Hours at Design Pressure (3/16 inch ofLocal Peaking at 975F).

For the portion of the HAZ nearest to the ID of the pipe,the damage progresses along the edge of the HAZ closest to theadjacent base metal. Contrarily, for the portion of the HAZ near-est to the OD of the pipe, creep damage progresses towards theoutside surface along the edge closest to the weld deposit. Thelatter effect is evident in Figure 12 between 150,000 hours and225,000 hours (predicted to be near end of life for these oper-ating conditions). The implementation of an elastic damage pa-rameter is required to achieve this behavior in the FEA, whichis fairly indicative of how a crack would progress through theHAZ (consistent with documented industry failures). That is, forthe FEA simulations without an elastic damage criterion, dam-age does not progress or evolve in this manner along the edge ofthe HAZ; rather, damage progresses once a larger portion of theHAZ (across the width) has reached elevated damage levels.

Numerous publications have attempted to address contin-uum damage mechanics and constitutive models for the creep re-



Figure 12. Creep Damage Progression using an Elastic Damage Pa-rameter (Dcrit=0.1) at Design Pressure (3/16 inch of Local Peakingat 975F).

sponse of materials. Notably, Perrin and Hayhurst implementedan elastic damage term in Reference [28] to investigate the be-havior (Type IV failure) of cross-weld specimens and to compareresults to experiments. Additionally, Hyde et al. [29] developedconstitutive equations for continuum creep damage approxima-tions based on experiments. An additional FEA-based study per-formed benchmark analysis to validate a creep continuum dam-age mechanics approach [30]. One particularly relevant paper byYatomi et al. [31], developed FEA predictions for creep ruptureof notched C-Mn steel specimens and evaluated the effect of in-corporating an elastic damage parameter in the simulations. This

paper asserts that the elastic damage term could be considered asa method to account for micro-cracks, and their ensuing effect onstiffness. In this particular paper [31], an elastic damage param-eter was implemented similar to Equation 10 above, but with thedamage threshold assumed to be zero (basically, elastic modu-lus is scaled down to zero directly as a function of creep damagefrom 0-100 percent damage). In [31], the authors point out thatthe predicted time to fracture initiation is considerably increasedwith the elastic damage model and the predicted time to ruptureis significantly reduced, and thus much closer to actual measuredrupture times in the notched specimens.

An additional comparison presented herein relates to thechosen elastic damage threshold (Dcrit). The example consid-ered in Figure 11 and Figure 12 is examined with an elastic dam-age threshold of 0.8 and compared to the simulation with Dcrit

set equal to 0.1. The elastic modulus at 975F is plotted as afunction of damage in Figure 13. While it might seem intuitivethat the higher the elastic damage threshold, the less severe thepredicted damage progression would be; this is not the case forthis example. Figure 14 shows contours of creep damage andelastic damage parameter (De) at 150,000 hours for the modelwith 3/16 inch of local peaking operated at design pressure and975F.

Figure 13. Elastic Modulus at 975F as a Function of Creep Dam-age for Dcrit = 0.1 and Dcrit = 0.8.

Figure 14 demonstrates that the simulation with the higherelastic damage threshold (implying that an adjustment in stiff-ness does not occur until damage at a given location in the FEAmodel reaches 80 percent damage) predicts damage progressionfurther through the wall thickness in the HAZ (along the bound-ary between the HAZ and the adjacent weld deposit/base metal).This behavior is an artifact of the primary stresses not being re-distributed over a wider portion of the HAZ (as is the case when



Figure 14. Contours of Creep Damage and Elastic Damage Param-eter (Dcrit) at 150,000 Hours for the Model with 3/16 inch of LocalPeaking Operated at Design Pressure and 975F (for Dcrit = 0.1and Dcrit = 0.8).

the elastic damage threshold is 0.1). That is, the damage evolu-tion with a higher elastic damage threshold is more indicative offailure from propagation of crack-like flaws because highly lo-calized damaged regions are surrounded by relatively stiff adja-cent material, thus propagating creep damage through-wall alongthe weld fusion line. Contours of predicted elastic damage pa-rameter (De) in Figure 14 (a value of one indicates that stiffnesshas been completely scaled down to zero) highlight this effect.

These comparisons confirm that accounting for elastic dam-age in computational creep simulations is complex and in mostcases, requires supplemental experimental test data for the mate-rial or welded geometry in question. Furthermore, for elevatedtemperature piping components with longitudinal weld seams,trying to predict realistic creep damage progression is challeng-ing because of unknowns associated with the high-temperaturematerial properties of the weld deposit, HAZ, and base metal.While published literature tends to suggest that implementing anelastic damage parameter in creep simulations is likely appropri-ate, the suitable elastic damage threshold (the creep damage limitat which stiffness begins to decrease), particularly for weldedjoints is largely unknown and is very likely material dependent.

THE INFLUENCE OF FABRICATION DEFECTSOne final comparison is presented to examine how an ini-

tial fabrication defect might accelerate creep damage. Figure 15compares a 36 inch diameter, 1/2 inch thick, 1 1/4 Cr - 1/2 Mopipe with the aforementioned base metal, HAZ, and weld de-posit properties (for a double-V weld) representative of a sub-critical PWHT. This figure shows contours of creep damage after2 years in service for models with 1/8 inch peaking operating atdesign internal pressure and 1000F (an elastic damage thresh-old of 0.8 is assumed in this comparison). The contour plot onthe left corresponds to a model with no initial fabrication defectand the contour plot on the right corresponds to the same modelwith a 1/16 inch deep inside surface-breaking crack-like flaw in-cluded in the base metal directly adjacent to the HAZ. This figuredemonstrates the significant effect the initial defect has on creepdamage evolution and highlights the importance of performingproper inspection to rule out the presence of crack-like flaws orfabrication defects.

Figure 15. Contours of Creep Damage for 36-inch Piping Modelswith 1/8 inch Peaking (Sub-Critical PWHT), Operating at 1000F for 2Years (with and without an Initial Fabrication Defect).

Lastly, to demonstrate the effect of original heat treatmentcreep damage evolution in the presence of an initial fabricationdefect, an example of a creep simulation for a normalized andtempered (N&T) 1 1/4 Cr - 1/2 Mo pipe cross section with thesame 1/16 inch deep inside surface-breaking crack-like flaw in-cluded in the base metal directly adjacent to the HAZ is shownin Figure 16. This FEA model has a local peaking magnitudeof 1/8 inch and an assumed operating temperature of 950F. InFigure 16, a micro-graph of a creep-related crack-like flaw froma sample of an in-service N&T catalytic reformer pipe is shownfor comparison. The damage progression observed in the sam-



ple is relatively consistent with the FEA model. Furthermore, itcan be seen that the creep damage (manifesting itself similarlyto a crack-like flaw) does not progress directly along the welddeposit or HAZ (following the double-v weld deposit as shownin Figure 15), as would be expected for a weld that received sub-critical PWHT. Rather, the observed damage in the N&T modelprogresses radially outward through-wall.

Figure 16. Contours of Creep Damage for a 36-inch Piping Modelwith 1/8 inch Peaking (N&T), Operating at 950F for 12 Years (withMicro-Graph of Creep Damage Progression in an In-Service Pipe).

These simulations show that normalizing and tempering af-ter welding (minimizing the creep property mismatch betweenthe weld deposit, HAZ, and base metal) is predicted to improvedamage tolerance and slow creep damage progression in the pres-ence of an initial defect, relative to sub-critical PWHT in low-chrome piping. As discussed above, coupling creep test resultswith detailed FEA, as summarized in this paper, can result inrealistic creep damage evolution predictions and remaining lifeestimates for welded components operating in the creep regime.

CONCLUSIONSThis paper offers examples of simulating the creep damage

evolution at welded joints using the MPC Omega Method, wherehigh-temperatures properties of the base metal, weld deposit, andHAZ can be customized in FEA based on creep test results. Also,the creep response of high-temperature 1 1/4 Cr - 1/2 Mo pip-ing with longitudinal weld seams is very sensitive to the mag-nitude of local peaking as well as historical operating pressuresand temperatures. Realistically predicting the remaining life ofthese piping systems is of great interest to owner-users, as there

have been a number of catastrophic failures of low chrome long-seam welded piping in the petrochemical (typically, in catalyticreforming units) and electric utility industries due to long termcreep. The results presented herein highlight the importance ofquantifying the magnitude of long-seam peaking in piping sys-tems, understanding historical operating temperature and pres-sure trends, and accurately documenting future operating condi-tions. A proposed peaking measurement technique that utilizes apivot template (from Reference [26]) is summarized herein, andmay offer a more consistent and repeatable method for estab-lishing peaking values for in-service components, compared toconventional peaking templates.

For new piping designs, seamless piping is obviously pre-ferred; however, if long-seams are unavoidable, normalizing andtempering after welding is an effective way to improve creepbehavior (by minimizing the creep property mismatch betweenthe weld deposit, HAZ, and adjacent base metal). Additionally,strict tolerances for the magnitude of permissible local peakingshould be enforced. For in-service high-temperature piping sys-tems, rigorous inspection (for both peaking and cracking) canbe cumbersome and often times not feasible due to accessibilityand the sheer length of the piping systems. Focusing inspectionon sections of piping that have historically operated under themost severe pressure and temperature combinations is one wayof attempting to manage the risk associated with potential creepfailures. The accuracy of documented historical operating pres-sures and temperatures in addition to material properties, includ-ing original heat treatment, is also critical.

Computational creep simulation results and parametric stud-ies, such as the ones offered in this paper, are valuable and canbe coupled with an inspection plan to estimate the risk associatedwith continuing to operate in-service piping systems. One diffi-culty with trying to correlate simulation results to remaining lifepredictions revolves around the underlying definition of failure.Creep damage at a point certainly does not necessarily constitutefailure. However, in the absence of measured peaking magni-tudes and inspection histories, conservative estimates should beestablished. Additional unknowns such as weld cap geometry(possible stress concentrations), weld deposit properties, amountof weld penetration, and any lack of fusion, etc., can significantlyreduce creep life.

Based on the FEA results investigated in this study, the in-troduction of an elastic damage parameter, where stiffness isscaled down as a function of creep damage, can drastically af-fect the overall creep damage progression through the HAZ andultimately, influence remaining life predictions. It is generallyaccepted that there is some relationship between creep damageand material stiffness, particularly near end of life. Interestingly,the chosen elastic damage threshold (the creep damage value atwhich stiffness begins to decrease) also has a notable effect ondamage evolution through the HAZ, with a higher threshold re-sulting in damage manifesting itself along the edge of the HAZ ina way that resembles crack-like flaw propagation, in some cases.



Establishing appropriate elastic damage parameters for a givenmaterial or welded geometry usually requires supplemental ma-terial creep testing. Finally, the introduction of an initial fabrica-tion defect near the weld in an FEA model with an elastic damagethreshold of 0.8 significantly increases creep damage accelera-tion along the base metal-HAZ interface (the assumed originalheat treatment influences the damage evolution as well). Thishighlights the importance of quality fabrication and inspectionpractices, since the presence of a crack-like flaw could dramat-ically reduce remaining life and lead to eventual rupture of thepressure boundary.

ACKNOWLEDGMENTSThe authors would like to thank Dr. Martin Prager of the

Materials Properties Council for his guidance relating to the de-velopment of the modified Omega parameters utilized for theHAZ in the FEA simulations presented herein and in Refer-ences [1, 13].

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