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Kota Sawada

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[Material Degradation of Heat-Resistant Steels during Long-Term Creep Exposure](https://mdr.nims.go.jp/datasets/cf9d15e6-6884-4fcd-bcd0-7508d1fbe944)

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Material Degradation of Heat-Resistant Steels during Long-Term Creep ExposureMaterial Degradation of Heat-Resistant Steels during Long-Term Creep ExposureKota Sawada+National Institute for Materials Science, Tsukuba 305-0047, JapanCreep strength and material degradation of heat-resistant steel were discussed, focusing on microstructural changes during long-term creepexposure. The materials and creep rupture data were obtained from Creep Data Sheet Project by National Institute for Materials Science. Fordetermining the allowable stress, the 100,000 h creep rupture strength should be evaluated based on creep rupture data. The selection of TTP anddividing the creep rupture data set are important to evaluate long-term creep strength when creep strength degradation occurs in the long term. Inorder to establish the allowable stress, minimum creep strength must also be evaluated. The minimum creep strength depends on the standarderror of estimate (SEE) for creep rupture data. If the heat-to-heat variation of creep strength is large even for the same material specification, thevalue of SEE is estimated to be large. Therefore, it is important to clarify the reason for the heat-to-heat variation of creep strength in terms ofmicrostructural changes and chemical compositions. Microstructural degradations were reported for carbon steels, low alloy steels, 9–12%Crferritic steels and austenitic heat-resistant steels, focusing on precipitates and dislocation structures. The reason for the heat-to-heat variation ofcreep rupture strength was discussed in term of the difference of solute atoms, segregation and grain size.[doi:10.2320/matertrans.MT-M2026047](Received April 24, 2026; Accepted May 30, 2026; Published August 25, 2026)Keywords: heat-resistant steels, long-term creep, microstructural changes, heat-to-heat variation1. IntroductionWith renewable energy increasingly being introducedworldwide to reduce CO2 emissions from the energy sector,thermal power plants are needed to cover the fluctuations inoutput from renewable energy. Under such circumstances,thermal power plants are operated not as base load but underpartial load, leading to cyclic loading on the components ofthe plants [1]. Accordingly, not only the creep propertiesbut also the creep-fatigue properties should be evaluated toimprove the reliability of such components [2, 3]. On theother hand, the fuel used in coal-fired thermal power plantswill be switched from coal to hydrogen and/or ammonia,thus reducing CO2 emissions [4]. Therefore, it is stillimportant to evaluate the creep strength of heat-resistantsteels to ensure the safe operation of thermal power plants.In the nuclear power sector, next-generation nuclearreactors such as fast reactors and high-temperature gas-cooled reactors are actively being developed to meet theincreasing global demand for power [5, 6]. The design lifeof these reactors is several decades [7], and so the ultra-long-term creep strength of heat-resistant steels and alloys shouldbe evaluated to establish the allowable stress. For example,60-year creep rupture strength was estimated for ASMEGrade 91 steels to establish the allowable stress for theASME Boiler and Pressure Vessel code [8].When introducing new heat-resistant steels and alloys, theallowable stress should be established based on long-termcreep strength evaluations because the wall thickness ofboiler tubes and pipes is designed based on the allowablestress. Recently, ASME Grade 93 [9] and HR6W [10]developed in Japan were standardized based on long-termcreep data and the allowable stress for thermal power plantswas established. For existing heat-resistant steels, allowablestress and chemical compositions were reviewed based onlong-term creep data. For creep strength enhanced ferriticsteels, Grade 91, 92 and 122, the allowable stresses werereduced [11, 12] because it was clear that the 100,000 h creeprupture strength was an overestimation due to long-termdegradation of creep strength [13, 14].For establishing the allowable stress for thermal powerplants, the 100,000 h creep rupture strength is determinedbased on regression analysis of creep rupture data and time-temperature parameters (TTP) such as the Larson–Millerparameter [15]. TTP suitable for material properties shouldbe selected to accurately estimate the 100,000 h creep rupturestrength. The minimum value of creep rupture strength is alsoneeded to establish the allowable stress [16]. This minimumvalue depends on the statistical dispersion of creep rupturestrength [17]. However, not experimental error but metal-lurgical factors can cause dispersion in some cases even forthe same material specification. It is predicted that thechemical compositions and heat treatment conditions of thematerial specification will be reviewed when metallurgicalfactors strongly affect the dispersion of creep rupturestrength. Therefore, when revising chemical compositionsand heat treatment conditions it is important to clarify themetallurgical factors.The National Institute for Materials Science (NIMS)started the Creep Data Sheet project in 1966 to obtain the100,000 h creep rupture strength of heat-resistant steels andalloys [18]. In this project, creep testing of multi-heat forthe same material specification was performed. Thedegradation of materials and heat-to-heat variation of creepstrength have been investigated for heat-resistant steels andalloys. Furthermore, the effects of chemical composition andmicrostructural factors on the degradation of materials andheat-to-heat variation of creep strength have been examined.In the present overview, creep strength evaluation, allowablestress, material degradation during long-term creep andheat-to-heat variation of creep strength are discussed basedon long-term creep data obtained from NIMS Creep DataSheets.+Corresponding author, E-mail: sawada.kota@nims.go.jpMaterials Transactions, Vol. 67, No. 9 (2026) pp. 1451 to 1461©2026 The Japan Institute of Metals and Materials OVERVIEWhttps://doi.org/10.2320/matertrans.MT-M20260472. Method of Evaluating Creep Strength and AllowableStressFor determining the allowable stress, the 100,000 h creeprupture strength is evaluated based on creep rupture data.Regression analysis of the creep rupture data is performedusing a regression equation of logarithmic stress with TTP toestimate the 100,000 h creep rupture strength. For example,Larson–Miller (LM) [15], Orr–Sherby–Dorn (OSD) [19] andManson–Haferd (MH) [20] parameters are used as typicalTTP:LM TTP ¼ ðT þ 273:15ÞðCþ log tRÞ ð1ÞOSD TTP ¼ log tR �Q=½2:3RðT þ 273:15Þ� ð2ÞMH TTP ¼ ðlog tR � log taÞ=ðT þ 273:15� TaÞ ð3Þwhere tR = time to rupture (h), T = temperature (°C), C, Q,ta, and Ta = optimized constants, and R = the gas constant.The master creep rupture curve for the fitting is expressedas follows:TTP ¼ b0 þ b1 log S þ b2ðlog SÞ2 þ . . .þ bkðlog SÞk ð4ÞTTP ¼ bþ b0S þ b1 log S þ b2ðlog SÞ2 þ . . .þ bkðlog SÞkð5Þwhere S = stress (MPa), b, b0, b1, b2, b3,+ , bk = regressioncoefficients estimated by the least squares method, andk = degree of regression equation.Normally, all the creep rupture data are used for theregression analysis. However, in some cases, the fitting curve(dashed line) overestimates the long-term creep strengthwhen the creep strength abruptly degrades in the long termas shown in Fig. 1. It is recommended that creep rupture dataof less than 500 h should not be used for the regressionanalysis in the ASME code [16]. Region-splitting analysiswas proposed [14] for the regression analysis in the event ofoccurrence of creep strength degradation in the long term. Inthis method, creep rupture data are divided into two regionsby the half value of 0.2% proof stress at each temperature,which corresponds to the proportional limit [21]. Plasticdeformation can easily occur during creep exposure whenstress exceeds the proportional limit. On the other hand, creepdeformation can occur in relatively weak areas such as grainboundaries under low stress that is less than the half value of0.2% proof stress. For modified 9Cr-1Mo steels, temperedmartensite homogeneously recovered during creep exposureunder high stress. However, preferential recovery occurredaround prior austenite grain boundaries during creepexposure under low stress in modified 9Cr-1Mo steels [13].Figure 2 shows the result of region-splitting analysis forcreep rupture data of ASME Grade 91 steels (modified 9Cr-1Mo steel). The LM parameter was used for the two regions[22]. Creep strength can be accurately evaluated in the longterm even if creep strength degradation occurs in the longterm.Multiregion analysis was also proposed for evaluatingcreep strength [23]. In this method, creep rupture data aredivided into multiple regions based on the activation energyfor creep rupture time. For example, the activation energydecreases in the long term for ASME Grade 91 steel [24].Creep strength can be accurately estimated in the long termby multiregion analysis based on the change of activationenergy for creep rupture time. Region-splitting analysis isuseful when there is not enough creep rupture data for thevery long-term region. On the other hand, multiregionanalysis can accurately evaluate long-term creep strengthwhen the activation energy can be estimated using enoughlong-term creep rupture data.Consequently, both the selection of TTP and dividing thecreep rupture data set are important to evaluate long-termcreep strength when creep strength degradation occurs inthe long term. Therefore, it is important to identify themetallurgical factors causing creep strength degradation inthe long term.The analysis method mentioned above evaluates averagecreep strength. In order to establish the allowable stress,minimum creep strength must also be evaluated [16]. Forexample, the minimum creep strength [17] is expressed asfollows from eq. (1) and eq. (4) in case of the LM parameter:log tR ¼ ðT þ 273:15Þ�1½b0 þ b1 log S þ b2ðlog SÞ2þ . . .þ bkðlog SÞk� � C� 1:65SEE ð6Þwhere SEE = the standard error of estimate. If the heat-to-heat variation of creep strength is large even for the samematerial specification, the value of SEE is estimated to belarge. This means the minimum creep strength is very low. Toimprove the allowable stress, not only average creep strengthFig. 1 Creep rupture strength of ASME Grade 91 steel.Fig. 2 Result of region splitting analysis for creep rupture data ofGrade T91.K. Sawada1452but also minimum creep strength should be increased,indicating that the heat-to-heat variation of creep strengthshould be small. Therefore, it is important to clarify thereason for the heat-to-heat variation of creep strength in termsof microstructural changes and chemical compositions.Equation (6) is also used to estimate the life of operatingplants. The coefficient of SEE in eq. (6) depends on theconfidence interval. The heat-to-heat variation of creepstrength can also affect the estimation of life of operatingplants. For conventional heat-resistant steels, product form,chemical composition and heat treatment condition can affectthe heat-to-heat variation for the same material specification.Recently, creep strength was evaluated for heat-resistantmaterials by additive manufacturing [25, 26]. In this case, notonly the factors mentioned above but also build directionmay be a reason for the heat-to-heat variation of creepstrength [25].3. Material Degradation during Long-Term CreepExposure3.1 Carbon steels and low alloy steelsNormally, carbon steels and low alloy steels consist offerrite and pearlite. Figure 3 shows SEM images of creepruptured specimens of 0.2C silicon-killed steel, JIS STB 410[27]. The spheroidization of pearlite structure occurred aftercreep rupture and was remarkable at higher temperatures inthe long term. Stress promotes spheroidization during creepexposure [28]. However, no remarkable drop in creep rupturestrength was confirmed in the long term for carbon steels[27]. It was reported that a large drop in creep strengthoccurred in the long term for carbon steel with high nitrogen[29]. In this case, the solid solution nitrogen contributed tothe creep rupture strength in the short term. However, thecreep rupture strength degraded in the long term becausethe amount of solid solution nitrogen decreased during creepexposure due to precipitation of nitride [29]. In addition tothe spheroidization of pearlite, graphitization can occurduring long-term creep exposure in carbon steels. It wasreported that graphite was formed after creep exposure for232,983.5 h at 400°C in 0.3%C steel [30] although JapaneseIndustrial Standards (JIS) [31] and Japan Petroleum Institute(JPI) Standard [32] note the occurrence of graphitizationabove 425°C. Figure 4 shows SEM micrograph of a creepruptured specimen for 0.3%C steel. Elongated graphite wasobserved after creep rupture even at 400°C [30]. Sphericalgraphite can also be formed after creep rupture [30].Elongated graphite tends to form during creep exposureunder higher stresses [30]. Graphite may serve as the originof fracture of components after long-term operation in powerFig. 3 SEM images of creep ruptured specimens of 0.2C silicon-killed steel.Fig. 4 Graphite observed in creep ruptured specimen of 0.3C steel. 400°C,tr = 232,983.5 h.Material Degradation of Heat-Resistant Steels during Long-Term Creep Exposure 1453plants [33]. Therefore, it is important to identify the time andtemperature of occurrence of graphitization. The time-temperature-precipitation diagrams of graphitization werealso reported for 0.2C, 0.3C and 0.5Mo steels [34].For low alloy steels, ferrite-pearlite, bainite and martensitestructures can be used depending on heat treatment. Thecreep rupture strength of the martensite structure is comparedwith that of the ferrite-pearlite structure for 2.25Cr-1Mo steelin Fig. 5 [35]. The creep strength of martensite is far superiorto that of ferrite-pearlite in the short term while no differencein creep rupture strength was observed between martensiteand ferrite-pearlite structures in the long term. It wasconfirmed that the martensite and pearlite structuresdisappeared after long-term creep exposure (Fig. 6), meaningthat the ferrite structure remained for both steels [35]. That iswhy the creep strength of the martensite structure was almostthe same as that of the ferrite-pearlite structure in the longterm. Therefore, microstructural changes during creepexposure should be retarded to improve long-term creepstrength. It was reported that in 0.5Cr-0.5Mo steels, the long-term creep strength of the ferrite-pearlite structure was higherthan those of bainite, tempered martensite and martensitestructures [36]. In addition to recovery of pearlite andmartensite structures during creep exposure, a precipitationfree zone (PFZ) forms around grain boundaries during creepexposure in low alloy steels [37]. The PFZ spreads withincreasing creep time [37]. Creep deformation can becomeconcentrated in the PFZ because precipitation strengtheningis reduced in the PFZ. It is predicted that creep voids canform on grain boundaries due to the concentration of creepdeformation in the PFZ. Therefore, not only the recovery ofpearlite and martensite structures during creep exposure butalso the formation of PFZ contribute to the creep strengthdegradation of low alloy steels.3.2 9–12%Cr ferritic steelsIn this section, 9–12%Cr ferritic steels with temperedmartensitic structure are discussed. 9–12%Cr ferritic steelsare widely used for ultra-supercritical power plants becauseof their good creep strength [38]. For example, 9–12%Crferritic steels such as ASME Grade 91 [39], Grade 92 [39]and Grade 122 [39] steels have the nominal composition of9Cr-1Mo-V-Nb-N, 9Cr-0.5Mo-1.8W-V-Nb-N and 11Cr-2W-0.4Mo-1Cu-Nb-V, respectively. 9–12%Cr ferritic steels arestrengthened by a tempered martensitic structure, M23C6(M: Cr, Mo, W) carbide and MX (M: V, Nb, X: C, N)carbonitride as shown in Fig. 7 [40]. Mo and W can alsocontribute to solid solution strengthening [41]. Themartensitic structure of creep interrupted specimens is shownin Fig. 8 together with hardness value [42]. The martensiticstructure recovered during creep exposure and this recoverycaused a decrease in hardness during creep exposure. Themartensitic structure disappeared and changed to equiaxedsubgrains after creep rupture as shown in Fig. 8(d). ForFig. 5 Creep rupture strength of 2.25Cr-1Mo steels at 550°C.Fig. 6 TEM images of creep ruptured specimens of 2.25Cr-1Mo steel. (a)STBA24 (initial: ferrite-pearlite): 550°C, tr = 46,816.8 h. (b) ASTMA542 (initial: martensite): 550°C, tr = 85,906.2 h.Fig. 7 (a) TEM micrograph of as received material of ASME Grade 91steel, (b) Result of elemental map by STEM-EDS for as received materialof ASME Gr.91 steel. Red particle: M23C6, Blue particle: V-rich MX,Green particle: Nb-rich MX. (online color)K. Sawada1454ASME Grade 91 steel, the creep strength abruptly degradesin the long term as shown in Fig. 2. The martensitic structurehomogenously recovered during creep exposure under highstress while preferential recovery of martensite occurredaround prior austenite grain boundaries under low stress asshown in Fig. 9. Therefore, the preferential recovery cancontribute to the creep strength degradation in the long term[13], as was also reported for Grade 92 and Grade 122 steels[43]. The martensitic structure can statically recover duringaging [23]. It was reported that creep strength degradationcan occur when the static recovery of the martensitic structureis remarkable during aging [23].Coarsening of M23C6 and MX precipitates occurs duringlong-term creep exposure, causing a decrease in precipitationstrengthening. This process is basically explained by Ostwaldripening [44, 45]. It was reported that the addition of B canretard the coarsening of M23C6 during creep exposure in9%Cr steel [46]. Creep stress and/or strain promotes thecoarsening of precipitates during creep exposure [47]. Inaddition to the coarsening of precipitates, the formation of aharmful phase can occur during creep exposure and decreaselong-term creep strength. Modified Z phase [Cr(V,Nb)N] isharmful to creep strength [48] because its formationconsumes fine MX precipitates during long-term creepexposure as shown in Fig. 10. It was reported in 12%Crsteels that Cr in the matrix diffuses to MX precipitates duringcreep exposure and MX precipitates gradually transform tomodified Z phase [49]. Figure 11 shows the time-temper-ature-precipitation diagram of modified Z phase in 9–12%Crsteels [50]. An increase in Cr content promoted the formationof modified Z phase during creep exposure as shown inFig. 11. An increase in Ni content can also promote theformation of modified Z phase during creep exposure [51,52]. Accordingly, the Cr and Ni contents should be reducedto avoid the formation of modified Z phase during creepexposure in 9–12%Cr steels. Danielsen et al. reported theeffect of elements on the driving force of the formation ofmodified Z phase, using Thermo-calc [53]. The addition ofN can increase the driving force. Laves phase such as Fe2Moand Fe2W is formed during creep exposure [54, 55]. Thisphase can provide nucleation sites for creep voids becausethe size of the Laves phase is larger than those of M23C6 andMX [56]. The Laves phase can contribute to precipitationstrengthening, but the effect of Laves phase formation oncreep strength is unclear because the formation consumesFig. 8 TEM micrographs of as received and crept materials of ASMEGrade 91 steel. (a) Before creep, (b) 600°C/50,000 h, (c) 600°C/70,000 h, (d) 600°C/80736.8 h.Fig. 9 Preferential recovery around prior austenite grain boundary aftercreep rupture in Grade 91 steel. 600°C, 110MPa, tr = 24.189.1 h. Dashedline: prior austenite grain boundary.Fig. 10 Formation of modified Z phase during creep exposure togetherwith disappearance of MX.Fig. 11 Effect of Cr content on time-temperature-precipitation diagram ofmodified Z phase. (online color)Material Degradation of Heat-Resistant Steels during Long-Term Creep Exposure 1455solute Mo and W which contribute to solid solutionstrengthening [41].3.3 Austenitic heat-resistant steelsFigure 12 shows the creep rupture strength ofSUS304HTB (18Cr-8Ni) [57] and KA-SUS304J1HTB(18Cr-9Ni-3Cu-Nb-N) [58]. The creep strength is higher inKA-SUS304J1HTB than in SUS304HTB. The creep strengthof KA-SUS304J1HTB was improved by the addition of Cuand Nb to SUS304HTB [59]. However, the creep strength ofboth steels degrades in the long term. The difference in creepstrength between the two steels becomes small in the longterm because the creep strength degradation is remarkable forKA-SUS304J1HTB as compared with SUS304HTB. Nor-mally, no precipitates except for inclusions are observedbefore creep exposure for both steels because solution heattreatment is performed. M23C6 and sigma phase are formedduring creep exposure for both steels.For KA-SUS304J1HTB, the time-temperature-precipita-tion diagram of M23C6 and sigma phase during aging areshown in Fig. 13. The sigma phase is formed after M23C6formation during aging. Figure 14 shows SEM micrographsof the grip portion of creep ruptured samples for KA-SUS304J1HTB. The precipitates shown in black are M23C6particles. Fine M23C6 particles are formed on grainboundaries after short-term aging. These particles contributeto grain boundary strengthening [60]. On the other hand,coarse sigma phase is formed on grain boundaries andconsumes M23C6 particles after long-term aging as shownin Fig. 14. Therefore, the formation of sigma phase can causecreep strength degradation in the long term. Fine NbXparticles and Cu phase are also formed after short-term creepin KA-SUS304J1HTB [61] in contrast to SUS304HTB. TheNbX particles change to modified Z phase after long-termcreep [61]. The modified Z phase can contribute toprecipitation strengthening because the size of the modifiedZ phase is small in contrast to 9–12%Cr steels. Cu phaseis coherently formed in the matrix after short-term creepand contributes to precipitation strengthening [62]. However,the coherency disappears due to coarsening after long-termcreep, contributing to creep strength degradation [62]. Thecreep fracture mode changes depending on stress andtemperature for both steels [63, 64]. Transgranular or wedgetype fractures can occur under a high stress condition andcracking at the sigma phase/matrix interface is seen undera low stress condition [63, 64]. Figure 15 shows a TEMmicrograph after long-term creep exposure for KA-SUS304J1HTB. It is clearly seen that a PFZ is formedaround the sigma phase on grain boundaries [65]. It ispredicted that creep deformation is concentrated on the PFZand creep voids can form at the sigma phase/matrix interface.Schematic illustrations of microstructural changes after creepexposure are compared between SUS304HTB and KA-SUS304J1HTB as shown in Fig. 16. M23C6 precipitates areobserved in the grain interior for SUS304HTB while fineCu phase, NbX and modified Z phase are distributed in thegrain interior for KA-SUS304J1HTB after short-term creepexposure. Therefore, the precipitation strengthening in thegrain interior is higher in KA-SUS304J1HTB than inSUS304HTB. Sigma phases are formed on grain boundariesFig. 12 Creep rupture strengths of KA-SUS304J1HTB and SUS304H.Fig. 13 Time-temperature-precipitation diagram of M23C6 and σ phase inKA-SUS304J1HTB.Fig. 14 SEM micrographs of grip portion of creep ruptured specimens ofKA-SUS304J1HTB. (a) 650°C, 7636.9 h. (b) 650°C, 44602.7 h. Blackparticles: M23C6. Arrows indicate σ phase.K. Sawada1456after long-term creep for both steels together with theformation of PFZ. This indicates that deformation resistanceis higher in the grain interior than in the PFZ after long-termcreep exposure. The difference in the deformation resistancebetween the grain interior and the PFZ is higher in KA-SUS304J1HTB than in SUS304HTB because the precip-itation strengthening in the grain interior is very high inKA-SUS304J1HTB. This leads to the remarkable drop increep strength for KA-SUS304J1HTB as compared withSUS304HTB as shown in Fig. 12. For example, the effectof PFZ formation on creep strength can be explained by a“core–mantle” model [66] to explain the grain-size de-pendence of the creep rate.4. Heat-to-Heat Variation of Creep StrengthFigure 17 shows the creep rupture strength of carbon steel,JIS STB 410 [27]. The data were obtained from 9 heats thatwere in the range of the material specification. Three steel-makers supplied the materials for creep testing [27]. Theheat-to-heat variation of creep rupture strength is very clear atall testing temperatures even in the range of the materialspecification. At 400°C, the difference in creep rupture timeis one order of magnitude. It is predicted that the SEE valuein eq. (6) will be large, indicating that the allowable stressis estimated to be small because of low minimum creepstrength. The reason for the large difference in creep rupturetime is due to the difference in Mo content [67], althoughthe Mo content is not specified in JIS STB 410 [29]. Therelationship between stress and LM parameter of JIS STB410 is shown in Fig. 18 together with Mo content [67]. Thecreep strength apparently increases with increasing Mocontent even in the range of the material specification. Itwas reported that an increase in Mo-C atomic pairs canincrease the creep strength in carbon steel based on thecalculation of the equilibrium solute elements and atomicpairs by Thermo-calc [68]. Therefore, it is possible that anappropriate allowable stress can be established if the Mocontent is specified in JIS STB 410.Fig. 15 PFZ around σ phase on grain boundary in creep ruptured specimenof KA-SUS304J1HTB. 700°C, 60MPa, tr = 50214.3 h.Fig. 16 Schematic illustrations of microstructural changes during creep exposure for SUS304HTB and KA-SUS304J1HTB. Left handside: after short-term creep, Right hand side: after long-term creep. (online color)Fig. 17 Creep rupture strength of carbon steel (JIS STB 410).Material Degradation of Heat-Resistant Steels during Long-Term Creep Exposure 1457Figure 2 shows the creep rupture strength of ASMEGrade 91 steel tubes [22] including 6 heats from severalsteel-makers. The heat-to-heat variation of creep rupturestrength is clearly seen at 600°C to 650°C. No correlationbetween Ni and Al contents and creep rupture strength wasconfirmed [69] although it was reported that the increasein Ni [70] and Al [71] content decreases the creep rupturestrength. The creep rupture strength at 650°C is shown inFig. 19. It seems that the creep rupture data can be dividedinto two groups by sampling year. The solid and opensymbols mean the heat sampled in the 1990s and 2000s,respectively. However, the sampling year cannot be thereason for the heat-to-heat variation. In addition to thechemical compositions, a difference in initial microstructurecan cause a difference in creep rupture strength. Figure 20shows the results of elemental mapping of Cr afterrenormalizing for as received samples of ASME Grade 91.The Cr segregates along the longitudinal direction of theboiler tube for MGG heat as compared with MGB heat asshown in Fig. 20. Normally, M23C6 and MX precipitates areformed after normalizing and tempering in the steel.However, M23C6 and MX precipitates can dissolve in thematrix after renormalizing. The Cr segregation shown inFig. 20 corresponds to Cr from dissolved M23C6 because Min M23C6 is Cr and Mo for Grade 91 steel. MGB and MGGheats were sampled in the 1990s and 2000s, respectively.Therefore, the creep strength is lower in MGG heat than inMGB heat, meaning that the Cr segregation decreases thecreep strength [69]. It was reported [69] that the decreasein number of M23C6 precipitates during creep exposurewas more remarkable in MGG heat than in MGB heat.Furthermore, recovery of martensitic structure during creepexposure was faster in MGG heat than in MGB heat [69].The precipitates and martensite structure strengtheningsignificantly decrease during creep exposure in the stronglysegregated heat. The difference in these microstructuralchanges can cause the difference in creep strength betweenthe two heats. The effect of segregation along the longitudinaldirection of the boiler tube on creep strength was reportedfor not only Grade 91 but also Grade 92 steels [72]. Ideally,the effect of chemical composition on the creep strengthshould be discussed for a material in which segregation iseliminated. Since it may be difficult to eliminate thesegregation during the manufacturing process of a commer-cial steel, it is recommended to specify the reduction of thesegregation in the material specification.For 700°C, the creep rupture strengths of austenitic heat-resistant steels are shown in Fig. 21 [58, 73, 74]. The data areobtained from several heats for each steel. The heat-to-heatvariations of creep rupture strength are clearly observed foreach steel even for the same material specification. It wasreported that a correlation between solute nitrogen contentand creep rupture time was observed for SUS304HTB [75].For SUS321HTB (18Cr-10Ni-Ti), the creep rupture timeincreased with increasing grain size [75]. The increase inboron content increased the creep rupture time ofSUS347HTB (18Cr-12Ni-Nb) even in the range of thematerial specification [75]. The heat-to-heat variation ofcreep strength is observed for not only heat-resistant steelsbut also Ni based alloys such as JIS NCF718-B [76].Fig. 18 Relationship between stress and Larson-Miller parameter forcarbon steel (JIS STB 410). CAA, CAB, CAC, CAG, CAH, CAJ,CAL, CAM, CAN: heat name.Fig. 19 Creep rupture strength of Grade 91 steels at 650°C.Fig. 20 Results of Cr mapping by SEM-EDS for as received sample afterrenormalizing in Grade 91. (a) MGB heat, (b) MGG heat. (online color)K. Sawada14585. Creep Strength Evaluation and Allowable Stress inFutureFor a commercial steel it is difficult to avoid creep strengthdegradation in the long term because the precipitation anddislocation strengthening will decrease during creep exposureas shown in section 3. Furthermore, deformation and fracturemechanisms may change depending on the stress conditions.The methods such as region-splitting analysis and multi-region analysis mentioned in section 2 are suitable forevaluating creep strength when creep strength degradationoccurs. Therefore, for a developed material, not only creeprupture data but also information on microstructural changesduring creep exposure should be considered to evaluate long-term creep strength and establish an allowable stress. Fornext-generation nuclear reactors, the 60-year creep rupturestrength should be evaluated. In this case, creep rupture dataand data of microstructural changes in the very long termwill be needed to clarify whether a method such as region-splitting analysis should be applied or not.In order to establish an allowable stress, the minimumcreep strength estimated from eq. (6) is also needed. Theminimum creep strength decreases when the heat-to-heatvariation of creep strength is large as mentioned in section 2.For developing a new material, in addition to improvingaverage creep strength, it is important to minimize the heat-to-heat variation of creep strength. The range of chemicalcomposition and manufacturing process can affect the heat-to-heat variation of creep strength as mentioned in section 4,indicating that robustness is needed for a new heat-resistantmaterial. Recently, additively manufactured (AM) materialsare being developed as heat-resistant materials [26, 77, 78],and these materials are also actively being standardized in theASME code [79]. For AM materials, printing parameters andbuild direction may be new factors that affect the heat-to-heatvariation of creep strength.It is easy to predict long-term creep strength when noremarkable microstructural changes occur during creepexposure. Ideally, solid solution strengthened materialswithout precipitation strengthening are suitable for heat-resistant materials, but these have not yet been achieved.AcknowledgementThe author thanks Dr. Kazuhiro Kimura, Dr. TomotakaHatakeyama, Mr. Yasushi Taniuchi, Dr. Kaoru Sekido, Mr.Takehiro Nojima and other people involved in the NIMSCreep Data Sheet Project.Open AccessThis paper is open access and licensed under a CC-BY-NC-ND license. You are free to share or adapt the materialsas long as you follow the license term: Attribution,NonCommercial, and NoDerivatives. To view a copy of thislicense, visit https://creativecommons.org/licenses/by-nc-nd/4.0/.REFERENCES[1] S. Zhang and Y. Takahashi: Effect of variable loading on creepproperty of Mod.9Cr-1Mo steel and life evaluation, Proceedings of the60th Symposium on Strength of Materials at High Temperatures, (TheSociety of Materials Science, Kyoto, 2022) pp. 1–5.[2] F. Abe, T.U. Kern and R. 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