doi: 10.1111/j.1460-2695.2010.01496.x

Assessment of creep rupture properties for dissimilar steels weldedjoints between T92 and HR3C

YI GONG, 1 J IAN CAO, 1 L I -NA J I , 1 CHAO YANG, 1 CHENG YAO, 1 ZHEN-GUO YANG, 1 JUN WANG, 1

XIAO-MING LUO, 2 FU-MING GU, 2 AN-FANG QI , 3 SHANG-YUN YE 3 and ZHENG-FE I HU 4

1Department of Materials Science, Fudan University, Shanghai 200433, PR China, 2Shanghai Institute of Special Equipment Inspection & TechnicalResearch, Shanghai 200062, PR China, 3Shanghai Boiler Works Ltd., Shanghai 200245, PR China, 4School of Materials Science and Engineering,Tongji University, Shanghai 200092, PR China

Received in final form 4 May 2010

A B S T R A C T Dissimilar steels welded joints, between ferritic steel and austenitic stainless steel, arealways encountered in high-temperature components in power plants. As two newgrade ferritic steel and austenitic stainless steel, T92 (9Cr0.5Mo2WVNb) and HR3C(TP310HCbN), exhibit superior heat strength at elevated temperatures and are increas-ingly applied in ultra-supercritical (USC) plants around the world, a complete assessmentof the properties for T92/HR3C dissimilar steels welded joints is urgently required. Inthis paper, metallographic microstructures across the joint were inspected by optical mi-croscope. Particularly, the creep rupture test was conducted on joints under differentload stresses at 625 ◦C to analyse creep strength and predict their service lives, whiletheir fractograph were observed under scanning electron microscope. Additionally, finiteelement method was employed to investigate residual stress distribution of joints. Resultsshowed that the joints were qualified under USC conditions, and T92 base material wascommonly the weakest part of them.

Keywords creep rupture; dissimilar steels welded joints; finite element method; HR3C;T92.

I N T R O D U C T I O N

With the worsening of global energy crisis and envi-ronmental pollution, higher energy utilization and lowerCO2 emission are presently the two prior criteria for de-signing fossil power plants,1 which will still serve as thedominant energy form for at least 20 years.2–4 As for fossilpower boilers, it is well known that the enhancement ofsteam parameters can facilitate improvement in thermalefficiency, which can result in reduction in not only thefossil costs but also the CO2, SO2 and NOx emission.5

Concretely, compared with supercritical (SC) boilers (24MPa, 565 ◦C), modern ultra-supercritical (USC) boilersoperating around 30 MPa and 600 ◦C can lead to bothan increase of 8–10% in thermal efficiency (from 37%to 45–47%) and a reduction of 20–25% in CO2 emis-sion.1,6–10

Development of USC boilers with increasingly highersteam parameters is an added incentive for boiler material

Correspondence: Z.-G. Yang. E-mail: zgyang@fudan.edu.cn

research. Consequently, heat-resistant steels, mainly re-ferring to ferritic steels and austenitic stainless steels, havebeen pursued simultaneously since the emergence of USCboilers for applications in boiler components, includingsuperheater, reheater, header, turbine, steam piping andso on. Subsequently, for the sake of reliable services inUSC boilers, a wealth of research has been carried outon these heat-resistant steels as F12 (X20CrMoV12.1),T91 (9Cr1MoVNb), TP347H (18Cr10NiNb), etc. to in-vestigate characteristics of their base materials and perfor-mance deterioration after long-term services.11–29 Amongthem, the T92 (9Cr0.5Mo2WVNb), approximately simi-lar to T91 but with a little modification in chemical com-positions for preferable high temperature properties thanT91, and the HR3C (TP310HCbN) are the two typicalrepresentatives of new grade ferritic steels and austeniticstainless steels for their superior heat strength above620 ◦C, and will certainly have a broad applicationprospect in forthcoming USC boilers. However, appli-cations of T92 and HR3C are currently a bit constraineddue to the lack of sufficient literature and experience on

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Table 1 Chemical compositions and heat treatment conditions of T92 and HR3C samples (wt%)

Elements C Cr Mo V Nb Ni Mn P S Si N Al W B

T92 Sample 0.11 8.76 0.36 0.21 0.059 0.25 0.46 0.016 0.002 0.39 0.044 0.01 1.63 0.0033ASME 0.07–0.13 8.50–9.50 0.30–0.60 0.15–0.25 0.04–0.09 ≤0.40 0.30–0.60 ≤0.020 ≤0.010 ≤0.50 0.03–0.07 ≤0.04 1.50–2.00 0.001–0.006

SA-213T92

HR3C 0.06 24.63 / / 0.49 20.29 1.24 0.012 0.001 0.39 0.24 / / /Sample

ASME ≤0.10 23.00–27.00 / / 0.20–0.60 17.00–23.00 ≤2.00 ≤0.030 ≤0.030 ≤1.50 0.15–0.35 / / /SA-213TP310HCbN

Heat treatment conditions:T92: 1050 ◦C × 20 min (normalizing) + 760 ◦C × 60 min (tempering).HR3C: solution-treated at 1110 ◦C minimum.

creep rupture performances of them and their weldedjoints, let alone the dissimilar steels welded joints betweenthem. Therefore, a thorough assessment of the compre-hensive properties, particularly the creep properties ofT92/HR3C dissimilar steels welded joints, seems prettyurgent.

In this paper, besides various conventional mechanicaltests including tensile test, bending test and hardness sur-vey, optical microscope (OM) was also applied to inspectthe metallographic microstructures across the dissimilarsteels welded joint between T92 and HR3C. Moreover,creep rupture test was particularly employed under dif-ferent load stresses at 625 ◦C to investigate the creepfeatures of the joints, whose fractograph was then ob-served by using scanning electron microscope (SEM) aswell. Furthermore, the residual stress distribution of thewelded joint after welding was calculated by using finiteelement method (FEM), which was a tentative approachto evaluate residual stress of the dissimilar steels weldedjoint between these two novel materials through the com-putational simulation method. Finally, based on the anal-ysis results, not only the creep rupture performances andthe degradation curves of T92/HR3C dissimilar steelswelded joints were reported, but also the mechanism ofvoids initiating creep rupture was concretely discussed,which may have critical significance in both service-lifeprediction and future heat-resistant steels preparation forboiler components.

E X P E R I M E N T A L

Tested materials were nominal T92 and HR3C heat-resistant steels with scales of 48O D × 8.4 mm thickand 48.26O D × 10.16 mm thick, respectively. Chemicalcompositions as well as heat treatment conditions of theirbase materials are listed in Table 1, which are in accor-dance with the requirements of ASME SA-213 T92 andTP310HCbN specifications.30 Etched in agent of picricacid (2, 4, 6-trinitrophenol) 1.25 g, HCl 20 ml, ethanol10 ml and H2O 10 ml for 40 s, the metallographic mi-

Fig. 1 Metallographic microstructures of tested base materials (a)T92, 1500× (b) HR3C, 200×.

crostructure of T92 sample is presented in Fig. 1a, whichdisplays a typical tempered lath martensitic microstruc-ture. Similarly, metallographic microstructure of HR3Csample was also obtained after being etched in the agentof CuSO4 4 g, HCl 20 ml and ethanol 20 ml for 20 s.As is shown in Fig. 1b, HR3C presents a fine-grained

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Table 2 Chemical compositions of welding wire ERNiCr-3 (wt%)

Elements C Mn Fe P S Si Cu Ni Ti Cr Nb+Ta

ERNiCr-3 welding wire 0.030 2.90 1.30 0.004 0.001 0.04 0.01 72.5 0.31 20.0 Nb 2.40ASME SFA-5.14 (AWS) ERNiCr-3 ≤0.10 2.5–3.5 ≤3.0 ≤0.030 ≤0.015 ≤0.50 ≤0.50 ≥67.0 ≤0.75 18.00–22.00 2.0–3.0

Fig. 2 Dimension of creep rupture test specimen.

austenitic microstructure with average grain size ofabout 7, which conforms with the requirement that thegrain size of TP310HCbN must be coarser than 7 (in-cluding 7) in ASME specification.30

The T92/HR3C dissimilar steels welded joints werewelded by means of gas tungsten arc welding (GTAW)with pure argon gas (Ar) as the shielding gas and AWSERNiCr-3 (corresponding to INCONEL 82/182) as thewelding wire, whose chemical compositions are listed inTable 2.31 Subsequently, the welded joints were subjectedto the post weld heat treatment (PWHT) at 760–770 ◦Cfor 2 h to eliminate the residual stress.

A variety of mechanical tests for the welded joints werethen successively carried out. Tensile test, bending testand hardness survey were performed at room temperatureaccording to the ASTM E8-04, E290-97a(2004) and E92-82(2003)e2 standards, respectively. Also, metallographicmicrostructure across the welded joint, especially in thetwo heat-affected zones (HAZ) and the weld seam, was in-spected under LEICA DMLM OM. In accordance withASTM E139-06 standard, the creep rupture test was con-ducted at 625 ◦C under load stresses of 110, 120, 130,140, 150, 160 and 180 MPa, respectively. Figure 2 showsthe round bar configuration of the creep rupture speci-men with the size of 10 mm diameter for base materialsand 50 mm gage length for welded joint. After the creeprupture test, micro morphologies of the cross-sections ofthe ruptured samples were observed by using PHILIPSXL30FEG SEM (Eindhoven, The Netherlands).

Residual stress analysis of the welded joint after weld-ing was carried out by means of finite element analysissoftware ANSYS 10.0 (Canonsburg Pennsylvania, USA).The FEM analysis result could clearly reflect the resid-ual stress distribution of the T92/HR3C dissimilar steelswelded joint in a convenient way, which may also theo-retically supplement the results of mechanical and creeptests.

Table 3 Tensile test results of T92/HR3C dissimilar steels weldedjoints

Tensile strengthSample No. (σ s, MPa) Rupture position

1 707 T92 base material2 699 T92 base materialT92 specification ≥620 /HR3C specification ≥655 /

Table 4 Bending test results of T92/HR3C dissimilar steelswelded joints

Bending style Sample No. Test condition Result

Face bending 1 D = 4T, α = 180◦ Qualified2 D = 3T, α = 50◦

Back bending 1 D = 4T, α = 180◦ Qualified2 D = 3T, α = 50◦

‘D’ denotes the bending diameter; ‘T’ denotes the material thickness;‘α’ denotes the bending angle.

R E S U L T S A N D D I S C U S S I O N

Mechanical tests results

As is clear from Table 3, the T92/HR3C dissimilar steelswelded joints exhibit qualified tensile strength, and theT92 base material part is their weakest region under loadstresses. Table 4 reveals the sign that the welded joints alsopresent eligible toughness. In addition, no cracks werefounded on the bended surfaces. Hardness survey resultsare displayed in Fig. 3, which indicates that the T92 basematerial and weld seam are the two low-hardness parts.

Metallographic microstructure inspection

As is well known, a dissimilar steels welded joint is oftenapproximately divided into five regions, i.e. base materialA, HAZ of A, weld seam, HAZ of B and base materialB. In this paper, the metallographic microstructures ofthe five regions across the T92/HR3C dissimilar steelswelded joint were inspected under OM.

Figure 4a presents the metallographic microstructure ofT92 base material after welding, in which no signs of dif-ferential are detected when compared with that of orig-inal material in Fig. 1a. However, in the HAZ of T92,

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Fig. 3 Hardness distribution across theT92/HR3C dissimilar steels welded joints.

obvious sorbitic microstructure with coarsened lathscould be observed in Fig. 4b. The width of the sorbitelath is nearly two times that of martensite lath, which mayresult in an increase of hardness in HAZ of T92 as wellas a decrease of toughness in this region simultaneously.The mechanism can be explained that coarser laths mayblock the growth of cracks under stresses, and eventu-ally lead to brittle rupture in this region. Near the weldseam, a distinct boundary between HAZ of T92 and weldseam can be observed in Fig. 4c. Figure 4d shows thestripe-shaped austenitic microstructure of the weld seam,whose stripe width has already reached around 20 μm.Correspondingly, like Fig. 4c, Fig. 4e gives the boundarybetween weld seam and HAZ of HR3C. Average grainsize of the austenite in HAZ of HR3C is about 6 (Fig.4f), and carbides as M23C6 have precipitated at the grainboundaries. Compared with the average grain size of 7in HR3C base material, the coarsened grains in HAZ ofHR3C may also lead to increase in hardness in this region.Fig. 4g is the metallographic microstructure of HR3Cbase material, which shows no changes with its originalstatus too. Collectively, the schematic diagram as wellas the metallographic microstructures distribution acrossthe T92/HR3C dissimilar steels welded joint is displayedin Fig. 4h.

Creep rupture test results

Table 5 lists the creep rupture test results under loadstresses ranging from 110 to 180 MPa at 625 ◦C. Thisphenomenon that rupture positions vary with change in

load stresses is in accordance with the results in the lit-erature.23 Together with the tensile test results, it canbe concluded that the T92 base material part is actuallythe weakest region of the T92/HR3C dissimilar steelswelded joint under relatively higher stresses. Neverthe-less, compared with the research data of Falat et al.,24 therupture time of our T92/HR3C dissimilar steels weldedjoints under 120 MPa at 625 ◦C actually even exceedsthe counterpart value of P92/P92 (‘P’ denotes the word‘pipe’, which contains the same chemical compositions of‘T’, i.e. ‘tube’, but with a larger diameter and thickness)similar steels welded joint, 1174 h.

According to the classic theory about creep rupture test,a double logarithmic relationship between load stress σ

and rupture time t, i.e. lgσ versus lgt, can be expressedin Eq. (1), in which the letters A and B both denote thematerial parameters of the tested samples. Accordingly,Fig. 5 presents the plot of lgσ versus lgt and the linearfitting result of it.

lg t = lg A − B lg σ. (1)

Thus, the threshold steam stress for service of theT92/HR3C dissimilar steels welded joints exposed at625 ◦C, which can be determined by extrapolation of thefitted line to 105 h, is 61.89 MPa. However, in practi-cal applications, some other unpredicted factors may alsoaffect the creep strength of welded joints. Hence, safetycoefficient n is always adopted to modify the predictedthreshold stress obtained from linear extrapolation. Con-sequently, permitted stress [σ ] can be expressed as Eq. (2),

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Fig. 4 Metallographic microstructures of the five different regions across the welded joint (a) T92 base material, (b) HAZ of T92, (c)boundary between HAZ of T92 and weld seam, (d) weld seam, (e) boundary between weld seam and HAZ of HR3C, (f) HAZ of HR3C, (g)HR3C base material, (h) schematic diagram of welded joint.

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Table 5 Creep rupture test (625 ◦C) results of T92/HR3Cdissimilar steels welded joints

Sample Load stress Rupture RuptureNo. (σ , MPa) Time (h) position

1 180 378 T92 base material2 160 865 T92 base material3 150 930 T92 base material4 140 1640 T92 base material5 130 2787 HAZ of T926 120 3164 Weld seam7 110 4182 Weld seam

Fig. 5 Double logarithmic plot of load stress versus rupture timefor the welded joints.

and the value of n is generally ranging from 1.2 to 1.65.22

[σ ] = σ 625 ◦C105

n. (2)

In this case, we take n = 1.5, and therefore the permittedstress [σ ] of T92/HR3C dissimilar steels welded jointscan be determined from Eq. (3), whose result value 41.26MPa also exceeds the steam stress of USC conditions.

[σ ] = 61.891.5

= 41.26 MPa. (3)

In terms of service life prediction for high-temperaturecomponents in boilers, the Larson–Miller equation, seenin Eq. (4), is always applied to estimate the allowablestresses under specific steam conditions:32

P = T(C + lg tr ), (4)

where T is the absolute temperature in K, and tr is therupture time in hours. Actually, C is a constant depend-ing on different materials. Owing to the fact that T92base material is the weakest part of T92/HR3C dissimi-lar steels welded joint, the value of C can be determinedfrom T92 matrix material. As for T92, the value of C dif-

Fig. 6 Plot of stresses with LMP for welded joints.

Table 6 LMP values and rupture times under stresses of 30, 35and 40 MPa at 625 ◦C

Service stress (MPa) LMP value (×10−3) Rupture time (h)

30 27.15 16806035 27.01 11919840 26.89 87631

fers from 19.5 to 36 according to a variety of referencesinvolved.33,34 However, thanks to sufficient practical ap-plication experiences of T92 in the past decade, the valueof C has been modified according to lots of firsthanddata. Hence, a value of 25 is now commonly adopted forC to predict the service lives of T92 for better accordancewith the actual conditions. Consequently, in this case, therupture data for the T92/HR3C dissimilar steels weldedjoints can be plotted in terms of lgσ versus Larson–Millerparameters (LMP) in Fig. 6, where the LMP is expressedas given in Eq. (5).

LMP = T(25 + lg tr ). (5)

Also, the lgσ versus LMP curve displayed in Fig. 6 is thenpolynomial fitted for the convenience of service life pre-diction for the welded joints under different stresses. Asfor USC boilers, service steam stresses commonly rangefrom 30 to 40 MPa, whose LMP values can then be easilyread from Fig. 6. Table 6 lists the LMP values and theircorresponding rupture times under stresses of 30, 35 and40 MPa, respectively. It is obvious that rupture times un-der 30 and 35 MPa at 625 ◦C both exceed the recom-mended service life of 105 h, even if the rupture time of40 MPa is long enough in practical applications (nearly10 years). Actually, the rupture time under 40 MPa at625 ◦C is still longer than that counterpart values ofP91/P22 dissimilar ferritic steels welded joint (less than

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Fig. 7 Macroscopic morphologies of ruptured welded joints under different load stresses (a) 180, (b) 130, (c) 110 MPa.

19 000 h) and P91/P91 similar steels welded joint (lessthan 32 000 h).26 Relevant data of T92 welded joints havehardly been found.

Fractograph observation of creep ruptured samples

According to the creep rupture test results, the rupturepositions of the T92/HR3C dissimilar steels welded jointsunder different load stresses, mainly including three dif-ferent parts, seen in Table 5. Hence, fractograph of sam-ples under three representative load stresses: 110, 130 and180 Mpa, whose rupture positions were respectively weldseam, HAZ of T92 and T92 base material, were thenconcretely inspected by means of SEM. Fracture posi-tions of joints under 110, 130 and 180 MPa located inthree different positions. Figure 7 was the macroscopicmorphology of them three. Then, samples were cut fromthe rupture positions to observe the micromorphologiesof their cross-sections under SEM.

Figure 8 presents the micromorphologies of cross-section of the ruptured sample under 180 MPa. It can beseen from Fig. 8a that the diameter of the cross-sectionis about 2.5 mm. Compared with its original value of6.0 mm, the reduction in area ψ can be calculated as about80%. As is shown in Fig. 8b, the cross-section is coveredwith densely distributed spherical creep voids, whose di-ameter ranges from 5 to 50 μm. The presence of creepvoids, whose nucleation mechanisms still need to be fur-ther identified, commonly represents a good toughnessof materials at high temperatures. In some region, threeneighbouring voids with diameters of about 5 μm havecoalesced into a larger wave-like one, whose axial widthhas already reached nearly 20 μm, as can be seen in Fig. 8cand d. This is typical evidence for ductile materials thatcreep rupture is led by the coalescence of microscopiccracks which are also generated by the sub-coalescence ofcreep voids.

Compared with the fractograph of ruptured sample un-der 180 MPa, the fractograph under 130 MPa displays

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Fig. 8 SEM of the cross-section of creep ruptured sample under 180 MPa (a) total morphology, (b) dimples and voids, (c), (d)three-neighbouring voids.

an obvious brittle rupture morphology, seen in Fig. 9a.Lots of slim dissociation steps resulting from narrow sor-bite laths can be found on the cross-section. Meanwhile,creep voids can also be observed in Fig. 9b. However, theamount of voids which have grown in a wide and shal-low form is far less than that on the cross-section under180 MPa. Although a small amount of dimples can beobserved in Fig. 9b as well, they are not the dominant fac-tors of rupture. This phenomenon may be attributed tothe highest hardness and residual stress in HAZ of T92.

Figure 10 displays the fractograph of creep rupturedsample under 110 MPa. An obvious macroscopic disso-ciation step rather than macroscopic creep voids can bedetected in Fig. 10a. Meanwhile, it can be learned that thereduction in area ψ is less than 5%. Actually, magnifiedby 500 times, the zoom is filled with randomly distributedcreep voids, seen in Fig. 10b. Furthermore, neighbouringvoids also have the potential to coalesce into larger ones,as shown in Fig. 10c. However, at least three differentialsincluding void depth, void area and distribution density

can be determined among the voids generated, respec-tively, under 110, 130 and 180 Mpa. On the one hand,as is discussed above, the wide and shallow voids with di-ameter of about 10 μm (Fig. 10d) may be accounted forthe low toughness of the weld seam. On the other hand,the larger distribution density of the creep voids under110 MPa than that generated under 130 MPa may be ac-counted for the lower hardness of the weld seam. Thus,the fractograph of the ruptured sample under 110 MPapresents an intermediate micromorphology between thatunder 180 and 130 Mpa, which indicates its intermediateproperties of weld seam.

Finite element method results

FEM is the most widely used computational simulationmethod for its superiorities as convenience, effectiveness,accuracy, etc., and is always applied in physical field anal-yses including thermal field, force field, magnetic field,etc. and their coupled fields.35–37 In this case, the residual

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Fig. 9 SEM of the cross-section of creep ruptured sample under130 MPa (a) total morphology (b) shallow voids.

stress of the T92/HR3C dissimilar steels welded joint af-ter welding was calculated by the FEM software ANSYS10.0.

As the width of the welded joint is not sufficiently large,their effect on residual stress distribution is generally ne-glected. Thus, the three-dimensional (3D) thermal-stresscoupled field analysis can be simplified as 2D axisymmet-ric problem. The 2D FEM model (after being meshed)is shown in Fig. 11a by using PLANE 13 2D coupledfield solid element. Also, birth–death element was appliedin the weld seam with 21 layers to simulate the weldingprocedure, seen in Fig. 11b, of all the layers that wereinitially dead, the lowest layer was firstly activated whenbeing welded, then the rest ones would be activated layerby layer above the former ones sequentially. Displace-ment in horizontal and vertical directions of the left andthe right borders of the welded joint set zero were theboundary conditions. The physical properties of T92 andHR3C base materials as well as ERNiCr-3 welding wireused in calculation are listed in Fig. 12 and Table 7,12,38

among which the shear modulus is commonly regardedas one-tenth of the elastic modulus.

The results of the computational analysis are presentedin Fig. 13, which shows the Von Mises equivalent stressdistribution across the welded joint. It can be obviouslyobserved from Fig. 13a that residual stress mainly ac-cumulates at the two HAZ regions. The largest resid-ual stress occurs at the HAZ of T92 region, seen in Fig.13b, which has already reached around 330 MPa. Mean-while, the residual stress of the HAZ of HR3C is alsoup to around 250 MPa. This could be ascribed to thepreferable ductility of austenite HR3C, as it is able to off-set stresses through plastic deformation. The FEM resultverified the fact that residual stresses usually accumulateat HAZ of the undermatched part, i.e. the ferrite HAZ, inferrite/austenite dissimilar steels welded joints. A conclu-sion can then be put forward that the selection of weldingwires with similar strength of ferrites could effectively re-lieve the residual stresses of the joints after welding; inother words, it can increase the comprehensive strengthof the joints.

Comprehensive analysis

In terms of creep voids, they are commonly generated un-der the interaction among high temperature, load stressand ageing time. Till date, there are several classic con-troversial theories of the emergence of creep voids.39,40

By Greenwood,41 Argon,42,43 Raj and Ashby,44 the initia-tion positions of void nucleation were argued; by Hull andRimmer,45 Needleman46 and Hancock,47 the controllingfactors of void growth were debated; and by Stowell,48

Nicolaou and Semiatin,49,50 and Chokshi,51 the coales-cence procedures of voids were discussed. In this paper,based on their works, a four-stage ‘nucleation to crack’mechanism of creep voids was put forward to clearly un-derstand the mechanisms of creep voids generation forthe T92/HR3C dissimilar steels welded joints.

1 NucleationIn this stage, the creep voids with irregular shapes aregenerated from the effect of grain boundaries slidingand/or grain matrix deformation under specific loadstress and temperature in creep process. Accordingly,the voids commonly nucleate at the grain boundaries ofthe weakest part or the part with largest residual stress oftested material. As for the T92/HR3C dissimilar steelswelded joints, the rupture positions i.e. the voids nucle-ation positions varied under different creep conditions,seen in Table 5. This may be accounted for the differentproperties of the five regions across the joint.

2 GrowthCreep voids continuously grow under constant temper-ature and load stress in this growth stage. However, the

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Fig. 10 SEM of the cross-section of creep ruptured sample under 110 MPa (a) total morphology, (b) dimples and voids, (c)four-neighbouring voids, (d) magnification of void.

Fig. 11 Meshed FEM model of T92/HR3Cdissimilar steels welded joint (a) total modeland (b) welded joint.

ultimate void volumes may vary in a wide range of valuesaccording to different toughness of different materials.As for ductile materials like T92, the preferable tough-ness ensures a good freedom for the voids to grow, whichmay lead to a great amount of deep voids. However, interms of brittle materials like HAZ of T92, voids’ growth

is constrained in a wide and shallow form with only asmall amount owing to the rigidness of matrix mate-rials.52 Meanwhile, in the process of voids growth, theoriginal voids with irregular shapes transform to uni-formly spherical ones. This can be explained that voidswith irregular shapes possess higher surface-free energy,

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Fig. 12 Physical properties versus temperatures of (a) T92, (b) ERNiCr-3 and (c) HR3C.

Table 7 Physical properties of T92, ERNiCr-3 and HR3C

Specific ThermalDensity heat expansion Poisson

Material (kg/m3) (J/kg·K) (10−6/◦C) ratio

T92 7850 Seen in Fig. 12 Seen in Fig. 12 0.3ERNiCr-3 8300 460 12 0.3HR3C 8640 500 Seen in Fig. 12 0.3

which will lead to a strong potential to minimize thesurface area to reduce this surface-free energy. Thus,the spherical-shaped voids, which own minimal surfacearea, are produced from the original irregular-shapedones by means of atoms diffusion. Consequently, thespheroidized voids decrease the extent of stress concen-tration and then increase the toughness of materials athigh temperatures.

3 CoalescenceCoalescence is usually a familiar phenomenon in cav-ities initiating degradations for materials.53 With theenlargement of at least two neighbouring voids, theytend to be combined into a larger one, which may ag-gravate the damages on the materials. In creep process,neighbouring voids commonly coalesce orientate alongthe grains boundaries,54 which will act as the initial formof cracks.

4 Crack formationWhen the voids orientate coalesce to a certain ex-tent, micro cracks emerge. As for ductile materials, thegrowth of cracks is always alleviated by the spherical-shaped creep voids, which will finally result in the oc-currence of dimples. But for brittle materials, crackspropagate quickly and then merge into macroscopicones, which may finally lead to intergranular rupture.Thus, the fractograph of creep-ruptured samples forbrittle materials always presents dissociation fracture

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Fig. 13 Von Mises residual stressdistribution of T92/HR3C dissimilar steelswelded joint (a) total distribution, (b)magnification of HAZ.

morphology with obvious dissociation steps, while forductile materials, it commonly shows a morphology cov-ered with voids.

It is well-known that the strength of ferritic steels iscommonly not as high as that of austenitic stainless steels.Especially their heat strength at elevated temperatures isnot good. Thus, aged at 625 ◦C under relatively higherload stresses ranging from 140 to 180 MPa, the creep rup-

ture positions of the T92/HR3C dissimilar steels weldedjoints concentrated at the T92 base material part. More-over, the rupture mechanism was predominantly ductilerupture according to the densely distributed deep creepvoids seen on the cross-section of the ruptured sample(Fig. 8b). This may be accounted for the better toughnessof T92 base material for its fine grained narrow martensitelaths. The mechanical test results that the rupture posi-tion of tensile test was T92 base material (Table 3) and

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the hardness of T92 base material was the second lowestpart of all the five regions (Fig. 3) also verified this fact.

With decrease of load stresses in creep test, the heatstrength differential between T92 base material and otherregions of the welded joints did not perform as distinct asthat under higher stresses like 180 MPa. Consequently,the residual stress may play an important role for the creeprupture property of the welded joints. According to theFEM analysis result, the largest residual stress accumu-lated in the HAZ of T92 region. Although the residualstress had been relieved a lot during the PWHT period,it still remained at relatively the largest value in HAZof T92 across the welded joint. Therefore, the creep rup-ture position under 130 MPa was the HAZ of T92 region.Moreover, owing to its sorbitic microstructure with coars-ened laths, which could result in the increase of hardness(Fig. 3) and alleviate the growth of creep voids, the rupturemechanism in HAZ of T92 was brittle rupture. Thus, anydimples could be seldom detected on the cross-sectionof the ruptured sample under 130 MPa, as can be seen inFig. 9a. Meanwhile, the shape of the creep voids exhibitedwide and shallow form, as shown in Fig. 9b.

Weld seam melted and recrystallized in the welding pro-cess, hence the grains of which could fully grow and be-came coarsened seen in Fig. 4d. Also, owing to its inter-mediate properties of toughness, the rupture mechanismof the weld seam under 110 MPa was predominantly brit-tle rupture like that under 130 MPa but the amount ofcreep voids was larger than that under 130 MPa. The for-mer phenomenon could be attributed to the coarsenedaustenitic grains in weld seam while the latter one maybe accounted for the low hardness and high toughness inthis region, too.

C O N C L U S I O N S

1 Regular mechanical properties of T92/HR3C dissimilarsteels welded joints were qualified at room temperature,among which the T92 base material part presented thelowest strength and the weld seam part exhibited thelowest hardness.

2 Aged at elevated temperature as 625◦C, creep rupturepositions of the welded joints accumulated at T92 basematerial part with ductile rupture mechanism due toits lowest heat strength in this region under relativelyhigher load stresses; while rupture positions changedto HAZ of T92 and/or weld seam regions with brittlerupture mechanism due to their low toughness and highresidual stress in this two regions under relatively lowerload stresses closer to the actual USC steam conditions.

3 Variation of the metallographic microstructures in thefive regions across the welded joint greatly affected theheat strength of the joint. Coarsened sorbitic structureof HAZ of T92 as well as coarsened austenitic structuresof weld seam and HAZ of HR3C led to the decrease of

toughness in these three regions. However, the marten-sitic structure of T92 and austenitic structure of HR3Cbase materials did not show any sign of change afterwelding.

4 Based on the service life prediction curves obtainedfrom creep rupture test, the T92/HR3C dissimilar steelswelded joints could ensure safe service for duration of105 h under 30 to 40 MPa at 625 ◦C, which is simi-lar to the USC steam conditions, and their allowablestresses and service lives both exceed those values of thecurrently used steels like T/P 91.

5 FEM analysis results showed that the HAZ of T92 wasthe region with largest residual stress after welding,which was in accordance with the creep rupture testresults under relatively lower load stresses.

Acknowledgements

The work was supported by both National Natural Sci-ence Foundation of China (Grant 50871076) and Shang-hai Leading Academic Discipline Project (Project Num-ber: B113).

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