Jinyang Zheng

Cunjian Miao

Yaxian Li

Institute of Process Equipment,

Zhejiang University,

Hangzhou, 310027 P. R. China

Ping Xu1

Institute of Applied Mechanics,

Zhejiang University,

Hangzhou, 310027 P. R. China

e-mail: pingxu@zju.edu.cn

Li Ma

Abin Guo

Institute of Process Equipment,

Zhejiang University,

Hangzhou, P. R. China

Investigation on InfluenceFactors of Mechanical Propertiesof Austenitic Stainless Steels forCold Stretched Pressure VesselsCold stretched pressure vessels from austenitic stainless steels (ASS) have been widelyused all over the world for storage and transportation of cryogenic liquefied gases. Coldstretching (CS) is performed by pressurizing the finished vessels to a specific pressure toproduce the required stress which in turn gives an amount of plastic deformation to with-stand the pressure load. Nickel equivalent (Nieq) and preloading, which is introduced inwelding procedure qualification for cold stretched pressure vessels, are considered to beimportant factors to mechanical behavior of ASS. During the qualification, welded jointwill be preloaded considering the effect of CS on pressure vessels. After unloading, the pre-loaded welded joint will go through tensile test according to standard requirements. Thereare two kinds of preloading method. One is to apply required tensile stress rk on specimenand maintain it for a long time (stress-controlled preloading). The other is to stretch speci-men to a specific strain of 9% (strain-controlled preloading). Different preloading and pre-loading rates may lead to differences in mechanical behavior of preloaded welded joint. Inorder to understand the effects of nickel equivalent, preloading and preloading rate on themechanical behavior of ASS for cold stretched pressure vessels, a series of tests were con-ducted on base metal, welded joint, and preloaded welded joint of ASS EN1.4301 (equiva-lent to S30408 and AISI 304). As regards to the preloaded welded joint, the ultimate tensilestrength (UTS) decreased as the nickel equivalent increased, while the elongation to frac-ture increased. It was more difficult to meet the available mechanical requirements withstrain-controlled preloading case than with stress-controlled preloading case. Rates of pre-loading had some effect on the mechanical properties of welded joint but nearly no effecton the mechanical properties of preloaded welded joint. These results are helpful forchoosing appropriate material and determining a proper preloading method for weldingprocedure qualification. [DOI: 10.1115/1.4007039]

Keywords: mechanical property, austenitic stainless steel (ASS), cold stretching (CS),chemical composition, nickel equivalent, preloading, preloading rate

1 Introduction

Cold stretched pressure vessels from ASS have been widelyused all over the world for storage and transportation of cryogenicliquefied gases, and guidances have been implemented in severalstandards such as AS 1210 Supplement 2:1999 [1], EN 13458-2:2002 [2], EN 13530-2:2002 [3], and ASME Code Case 2596-2008 [4] (which is being implemented in the mandatory appendixof ASME BPVC VIII-I: 2011). Cold stretched pressure vessels aremanufactured from finished vessels through CS, which is per-formed by pressurizing the finished vessels to a specific pressureto produce the required stress rk. After CS, an amount of plasticdeformation is given to withstand the pressure load. Such vesselswill get a higher proof strength, a lighter weight (about 50–70%of the conventional one with the same load carrying capacity),and thus a lower cost and energy consumption in manufacturingand transportation.

Nieq is an important factor to ASS mechanical behavior for coldstretched pressure vessels. It could be used to describe the austen-ite stability, which has a strong effect on the deformation-inducedmartensite (DIM) transformation. Because of DIM’s main influ-ence on the mechanical behavior, Nieq may indirectly affect ASS

mechanical properties. Preloading is another important factor.During welding procedure qualification for cold stretched pressurevessels, the finished welded plate will be preloaded and unloaded,leaving a permanent plastic deformation. Then the preloadedwelded joint made from the plate will go through tensile test toobtain its mechanical properties, which will be evaluated by thestandard requirements. Preloading makes the qualification differ-ent from the normal one. It introduces an effect similar to the CSeffect on vessels into the qualification to help know whether thisqualification is proper for cold stretched pressure vessels. Twokinds of preloading are employed now. One is to apply therequired stress rk on the specimen and hold this stress until thestrain rate gets lower than 0.1%/h [2–4]. The other is to stretch thespecimen to a specific strain of 9% with a quasi-static strain rate[5]. These two preloading methods may lead to different results inmechanical properties of preloaded welded joint. In this paper,these preloading methods were noted as rk-stress-preloading and9%-strain-preloading, respectively. Preloading rate is also consid-ered to be an important parameter. The effects of strain rate onDIM and mechanical properties of some ASS grades were studiedat strain rate between 10�4 and 103/s [6–8], and the rate wasproved to influence the mechanical behavior to a certain extent.Little researches were reported about the effect of preloading rateon the mechanical properties of preloaded welded joint.

In order to understand the effects of Nieq, preloading, and pre-loading rate on the mechanical behavior of ASS for cold stretchedpressure vessels, a series of tests were performed with a commercial

1Corresponding author.Contributed by the Pressure Vessel and Piping Division of ASME for publication

in the JOURNAL OF PRESSURE VESSEL TECHNOLOGY. Manuscript received November 3,2011; final manuscript received April 25, 2012; published online November 21,2012. Assoc. Editor: David L. Rudland.

Journal of Pressure Vessel Technology DECEMBER 2012, Vol. 134 / 061407-1Copyright VC 2012 by ASME

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms

ASS grade EN1.4301 (equivalent to S30408 and AISI 304). First,the chemical composition of ASS was measured to calculate Nieq.Second, preloading was done on the welded plate and preloadedwelded joints were made. Third, tensile tests were carried out onbase metal, welded joint, and preloaded welded joint. Furthermore,some DIM measurements were carried out. Based on these testresults, analysis was done and some suggestions were put forward.These findings will be helpful for choosing the appropriate materialand determining a proper preloading method in welding procedurequalification.

2 Test Procedure

2.1 Test Material, Chemical Composition, and NickelEquivalent. This investigation was carried out on industriallymanufactured ASS grade EN 1.4301 in as-received condition.These material plates were available with the thicknesses from6 mm to 20 mm, and in a hot-rolled and solution heat-treated con-dition. The chemical composition is listed in Table 1, as well asNieq, which was calculated by the following equation with theconsideration of chemical composition, temperature and deforma-tion [9]:

Nieq %ð Þ ¼ Niþ 0:65Crþ 0:98Moþ 1:05Mnþ 0:35Siþ 12:6C

þ 0:03 T � 300ð Þ � 2:3log 100= 100� Rð Þ½ � � 2:9

(1)

where T is the temperature (K) and R is the deformation (%). Roomtemperature was considered. The effect of the deformation R causedby preloading is calculated to be 0.09 (the value of 2.3log[100/(100�R)]), which is less than 1% of the value of Nieq. Thus theNieq values of solution annealed (SA) and CS materials are nearlythe same, and the effect of deformation is ignored for convenience.

2.2 Preloading and Tensile Test. Preloading was performedon welded plates using quasi-static strain rates. Both 405 MPa-stress-preloading (the stress rk was selected to be 405 MPa accord-ing to Ref. [4]) and 9%-strain-preloading were involved. The sche-matic diagram of the stress–strain curves of these preloadingmethods are shown in Fig. 1. For the 405 MPa-stress-preloadingcase, the curve starts from O, and goes elastically until yielding

part A. Then, it goes up to B before which the rate is about1� 10�4/s. The stress of B is rk (405 MPa), and this stress is heldfor 1–2 h until C, while the strain rate keeps decreasing slowly. For9%-strain-preloading case, the curve goes directly to D with thestrain rate of about 1� 10�3/s. The yielding is also located in partA. The ending D of 9%-strain-preloading has a higher stress andstrain than the ending C of 405 MPa-stress-preloading. After beingunloaded, the specimens of both preloading case were reloaded inthe subsequent tensile test (STT) to the failure with a strain rate ofabout 1� 10�3/s. A test number summary for Nieq investigation islisted in Table 2, and the total data numbers are 12, 5, 13 for basemetal, 405 MPa-stress-preloading welded specimen and 9%-strain-preloading welded specimen, respectively.

After the comparison between the two preloading methods, fur-ther tests focused on the effect of preloading rate were donethrough 9%-strain-preloading. Strain rates of 1� 10�5 and1� 10�3/s were employed during this preloading, while strainrate of 2.5� 10�3/s was used in all the STT processes correspond-ing to the typical rate in tensile tests of engineering application.The tests with 10�3/s in preloading and 2.5� 10�3/s in STT weremarked as type A, and those with 10�5/s in preloading and2.5� 10�3/s in STT were recognized as type B. Each type of testshad two specimens.

All the tests were performed in air by using a servo hydraulicMTS 810 tensile testing machine. Specimens with rectangularcross section were used in the tests for investigating the effect ofNieq, and specimens with a 5-mm-diameter and a gauge length of25 mm were used in the study for the influence of preloading andits rate. All specimens were made according to GB/T 228 [10](equivalent to ISO 6892) and cut parallel to the rolling directionof the plates. Strain data were measured with an MTS 634.12 F-25extensometer.

2.3 Measurement of DIM. The a0-martensite contents ofspecimen were measured during the experiments for the effect ofpreloading and its rate, with an instrument named Ferritescope

Table 1 Chemical composition and Nieq of ASS EN 1.4301(wt %)

No. C Si Mn P S Cr Ni N Nieq

1 0.036 0.45 1.08 0.028 0.003 18.18 8.00 0.064 18.752 0.051 0.53 1.34 0.030 0.001 18.14 8.22 0.060 19.473 0.058 0.49 1.64 0.028 0.001 18.12 8.23 0.051 19.834 0.034 0.45 1.02 0.023 0.003 18.01 8.05 0.067 18.615 0.055 0.51 1.07 0.027 0.001 18.28 8.15 0.061 19.226 0.055 0.64 1.35 0.029 0.013 17.36 8.03 0.041 18.757 0.041 0.59 0.80 0.022 0.001 17.71 8.00 0.039 18.178 0.034 0.47 0.83 0.027 0.002 17.43 8.02 0.041 17.919 0.039 0.49 0.99 0.029 0.003 17.52 8.07 0.044 18.2610 0.055 0.64 0.94 0.026 0.002 17.42 8.10 0.042 18.4311 0.043 0.54 0.88 0.027 0.002 17.25 8.09 0.039 18.0612 0.022 0.44 1.70 0.027 0.008 17.28 8.31 0.040 18.8613 0.041 0.34 0.88 0.026 0.002 18.35 8.04 0.045 18.6314 0.061 0.47 1.11 0.028 0.003 18.41 8.05 0.055 19.22

Fig. 1 The schematic diagram for the stress–strain curves ofpreloading methods

Table 2 A test number summary for Nieq investigation

No. 1 2 3 4 5 6 7 8 9 10 11 12 13 14 Total

Base metal 1 1 1 1 1 1 1 0 1 1 0 1 1 1 12405 MPa-stress-preloading 1 1 1 1 1 0 0 0 0 0 0 0 0 0 59%-strain-preloading 1 1 1 1 1 1 0 1 1 1 1 1 1 1 13

061407-2 / Vol. 134, DECEMBER 2012 Transactions of the ASME

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms

(model FMP30, Helmut Fischer GmbHþCo.), which is normallyused to measure d-ferrite content in austenitic and duplex steelsand to determine the fraction of DIM in austenitic materials. Itwas found that ferritescope measurement was an efficient way tomeasure the content of the ferromagnetic a0-martensite phase. Themeasurement should be performed on an electropolished surfaceof specimen where the effect of cold work introduced by machin-ing was eliminated and it should also be performed where the dis-tribution of magnetic content was uniform. The measuring resultswill be converted to actual martensite contents with a calibrationcurve [11].

3 Influence of Nieq on Mechanical Properties

3.1 Requirements of Mechanical Properties. The effect ofNieq on the mechanical properties of ASS and preloaded ASS areshown from Figs. 2–4, while the mechanical properties includedyield strength (YS), UTS and elongation to fracture (A, which isalways used to describe the elongation to fracture in the relevantstandard [10]). Requirements for EN 1.4301 mechanical proper-ties in relevant standards [2–5,12–14] are also shown in these fig-ures as well as Table 3 to help reveal the effects of Nieq. It is clearin Table 3 that YS increased to about 405 MPa from about205 MPa after preloading, and A decreased to 25% from about40%. The UTS for preloaded ASS was not specified, while520–720 MPa was employed in the Chinese company standard ofZhangjiagang CIMC Sanctum Cryogenic Equipment Co., Ltd [5].

3.2 Influence of Nieq. It was clear in Fig. 2 that preloadingincreased YS obviously. The 9%-strain-preloading gave a highervalue of YS than that of 405 MPa-stress-preloading. Both pre-loaded ASS met the yield strength requirement of 405 MPa. AsNieq increased, YS of 9%-strain-preloading ASS decreased, whilethe curves of both base metal and 405 MPa-stress-preloading ASSvaried slightly. The YS data of 405 MPa-stress-preloading wererelatively well behaved, the reason of which is that the specimensof 405 MPa preload case were all preloaded to 405 MPa and heldunder the stress for about 1–2 h.

The UTS variation as a function of Nieq was showed in Fig. 3.The curves of base metal and 9%-strain-preloading ASSdecreased with increasing Nieq. They decreased to the UTS upperlimit of 720 MPa at about 18.45%, and then continued decreasing.Furthermore, the curve of preloaded ASS was higher below theline of 720 MPa. The lower limit of 520 MPa seemed not to bereached. The 405 MPa-stress-preloading curve showed a slightlydecreasing trend and varied gradually.

All the curves of A in Fig. 4 increased with increasing Nieq.The values of preloaded ASS were obviously lower than that ofbase metal, while the curve of 405 MPa-stress-preloading was alittle higher than that of 9%-strain-preloading. Additionally, the Aof 9%-strain-preloading curve increased to 25% as was requiredat about 19.15% of Nieq, and then continued growing.

4 Influence of Preloading on Mechanical Behavior

4.1 Different Preloading. The 405 MPa-stress-preloading isdetermined from the strengthening stress rk in the relevant stand-ard [2,3] and used in USA. The 9%-strain-preloading is alwaysused in China [5]. Tests with these two preloading methods werecarried out and results were showed above. Main differencesbetween these methods are described here.

(1) It is more difficult to meet the available mechanical behav-ior requirements with the 9%-strain-preloading case thanwith 405 MPa-stress-preloading case. Therefore, when theformer case meets the requirements, the latter one canmake it too. However, the former one is not sure to meetthe requirements even when the latter does.

(2) Time spent on the 405 MPa-stress-preloading is about1–2 h, which is much longer than that of 9%-strain-preloading (about several minutes). Therefore, using 9%-strain-preloading could save much time for engineeringapplications.

It is concluded that the 9%-strain-preloading is more strict than405 MPa-stress-preloading about the requirements of mechanicalproperties, and could spare much more time in preloading.

Fig. 2 The effect of Nieq on YS Fig. 4 The effect of Nieq on A

Fig. 3 The effect of Nieq on UTS

Journal of Pressure Vessel Technology DECEMBER 2012, Vol. 134 / 061407-3

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms

4.2 Preloading Rate.

4.2.1 Mechanical Properties. After 9%-strain-preloading, thetrue stress–strain curves of type A and B were shown in Fig. 5(a).YS, UTS, A and reduction in area (RA) of the specimens werelisted in Table 4, as well as the time spent on these tests. The truestress and strain were calculated from the engineering valuesthrough the following formulas

rtrue ¼ rengð1þ eengÞ (2)

etrue ¼ lnð1þ eengÞ (3)

The shapes of preloading curves showed parabolic behavior atboth the strain rates of 10�3/s and 10�5/s, while in STT the curvesbetween YS point and UTS point were almost linear. YS increasedwith the increasing strain rate in preloading, and the strength at10�3/s was about 30–50 MPa higher than that at 10�5/s. Some

similar phenomenon has been experimentally observed by others[7]. To considering the data scatter influence, another group testwas conducted with the same specimen size and test parameters.The comparison in Fig. 5(b) showed that the true stress–straincurves were almost the same with the same rate parameters. Fur-thermore, some formula was established with the strain rate as a pa-rameter for face centered cubic material [15] to model thestress–strain relationship, which means a specific strain rate couldlead to only one stress–strain relationship. Thus, the difference wasconsidered to be caused mainly by the strain rate difference. Whenloaded in STT, the curves at the same rate 2.5� 10�3/s did not fol-low their former trends, which may be caused by the differencebetween strain rates of preloading and STT. However, they lookedlike each other in STT, and seemed not to be affected by the differ-ence between preloading rates, which could also be found from thedata in Table 4. In addition, the time spent on the preloading oftype A was much shorter.

4.2.2 DIM Transformation. DIM mass fraction in preloadingand STT was depicted in Fig. 6(a), and its distribution along thegauge after fracture was showed in Fig. 6(b). The DIM of types Aand B showed the same trend. It was observed that the DIM oftype A was a little higher than that of type B during preloading.The DIM of both types began to grow fast after about 9.0% strainand became to be linear after about 12.5% strain. The differencein preloading rates seemed to have a little effect on preloadingand nearly no effect on the STT. The final DIM distributions alongthe specimens’ gauge length after fracture were also nearly thesame between types A and B. The contents of DIM increasedwhen the distance from the fracture location decreased.

4.2.3 Work-Hardening Rate. The interaction between thework-hardening rate (dr/de) and true strain was illustrated inFig. 7. During preloading, both work-hardening rates decreasedrapidly, which may be mainly due to the appearance of e phase[16]. Meanwhile, there were only a little DIM content showed inFig. 6(a). When it was loaded in STT, DIM became to grow fastin Fig. 6(b), while the work-hardening rate became to increase. Itis generally accepted that a0-martensite has a strong effect onwork-hardening of ASS, and during the STT the transformationwas considered to follow the routes of c!a0 or c ! e ! a0 [17].After 25% strain, the work-hardening rate began to fall gradually,which was also found in metastable ASS grade EN 1.4318 [7].The reason might be the slight effect of a little adiabatic heating,which could strongly prevent the DIM transformation. On theother hand, the rapid decrease of dr/de during preloading could

Table 3 Requirements of mechanical properties for EN 1.4301 in the relevant standards

Material Standard YS/MPa UTS/MPa A/%

Base metal with no preloading ASTM A240/A240M-10 b �205 �515 �40EN10028-7:2008 �210 520–720 �45GB 24511:2009 �205 �520 �40

Preloaded base metal ASME Code Case 2596: 2008 �405 — —Preloaded base metal and welded joint EN 13458-2:2002, EN 13530-2:2002 �410 — 25Preloaded welded joint Q/320582SDY7—2008 �410 520–720 25

Fig. 5 True stress–strain curves. (a) True stress–strain curvesof type A and B; (b) the comparison on true stress–straincurves between different group tests with the same parameters.

Table 4 Mechanical properties during preloading and STT

TypeTest step and

strain rate YS/MPa UTS/MPa A/% RA/% Time

A Preloading at 10�3/s 310 — — — 90 s (1.5 min)STT at 2.5� 10�3/s 500 750 61.5 78.0 3–4 min

B Preloading at 10�5/s 280 — — — 9000 s (2.5 h)STT at 2.5� 10�3/s 505 755 63.5 76.0 3–4 min

061407-4 / Vol. 134, DECEMBER 2012 Transactions of the ASME

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms

also be explained by the stress–strain behavior. In elastic loading,the change of strain was small for large changes of stress, whichled to an initial high dr/de. The dr/de decreased rapidly when thematerial was yielding, and leveled off since the stress–strainresponse was relatively linear during plasticity. This figure indi-cated that work-hardening rate was not affected by the difference

between preloading rates. However, a slight difference betweenwork-hardening rates occurred at the beginning of yielding (whenthe value of work-hardening rate was high) could still make a bigeffect on the strength.

4.2.4 Flow Stress. Figure 8 showed the dependence of theflow stress on the square root of a0-martensite fraction of types Aand B. The flow stress is defined as the true stress after yielding,and it began from about 300 MPa and 500 MPa during preloadingand STT, respectively.

It was found that the flow stress of both type A and B seemedto have a linear relation with the square root of the a0-martansitefraction during STT and some part of preloading, and similar phe-nomenon was found in tensile tests by Fang and Dahl [18] andJuho [7] in tensile tests. When in preloading, the flow stressdecreased as the square root of the a0-martansite fractiondecreased and finally behaved vertically, which could beexplained by the little variation of DIM content closed to zero dur-ing preloading from Fig. 6(a). It is accepted that the flow stress ofthe material is linearly proportional to the square root of the dislo-cation density, and it has been explained that the DIM transforma-tion causes the accumulation of dislocations in the austenitephase, thus indirectly increase the flow stress. Therefore, it hasbeen suggested that the linearity is not necessary to indicate therelationship between flow stress and DIM content, it is only anindirect effect of DIM on the flow stress through the dislocationdensity. These curves of type A and B showed that different pre-loading rates may affect the curves of preloading a little but hadno effect on STT curves. In addition, the data scatter was consid-ered to have slight effect.

5 Conclusions

Based on the results above, the effects were summarized ofthese influence factors on the mechanical behavior of ASS forcold stretched pressure vessels.

(1) Increasing Nieq lowered the YS of 9%-strain-preloaded ASSwhile it increased that of base metal. A weak effect causedby varying Nieq was found on the YS of 405 MPa-stress-pre-loaded ASS. All the YS of preloaded ASS were higher than405 MPa, which is the minimum requirement of the YS.Increasing Nieq also lowered the UTS of base metal and 9%-strain-preloaded ASS. These UTS would be higher than720 MPa when Nieq was lower than about 18.45%. As UTSis limited to 520–720 MPa, it was better for Nieq to be higherthan 18.45%. There was also a small variation on the UTS of405 MPa-stress-preloaded ASS due to the changing Nieq.Higher Nieq caused higher value of A. For preloaded ASS,

Fig. 6 DIM transformation during the tests. (a) DIM mass frac-tion as a function of true strain in preloading and STT; (b) distri-bution of DIM mass fraction along the specimen gage lengthafter fracture.

Fig. 7 The correlation between work-hardening rate dr/de andtrue strain in preloading and STT

Fig. 8 Flow stress as a function of the square root of the a0-martansite fraction

Journal of Pressure Vessel Technology DECEMBER 2012, Vol. 134 / 061407-5

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms

the lowest value of 25% was required. Results showed thatthe A of 9%-strain-preloaded ASS became higher than 25%when Nieq was higher than 19.15%.

(2) In the welding procedure qualification for cold stretchedpressure vessels, 9%-strain-preloading is stricter than405 MPa-stress-preloading about the requirements of me-chanical properties. When the mechanical properties resultsof 9%-strain-preloaded ASS meet the requirements, thoseof 405 MPa-stress-preloaded ASS could meet too. How-ever, the opposite condition might not be true. In addition,9%-strain-preloading could save much more time for engi-neering applications than 405 MPa-stress-preloading.

(3) The rate of 9%-strain-preloading had an effect on the me-chanical behavior of ASS in preloading. Higher rateincreased the strength of true stress–strain curve and led toa little higher DIM. The results of two preloading rates(10�3/s and 10�5/s) slightly differed on the work-hardeningrate and the relationship between flow stress and squareroot of a0-martansite fraction. However, with different ratesin preloading and the same rate of 2.5� 10�3/s in STT,effect was hardly found on STT mechanical behavior ofpreloaded ASS.

Acknowledgment

The authors gratefully acknowledge the financial support fromthe International Science and Technology cooperation project(2010DFB42960) and National Key Technology R&D Program(2011BAK06B02-05).

Nomenclaturerk ¼ the specific stress for cold stretching (MPa)

Nieq ¼ nickel equivalent (%)T ¼ the temperature in the test (K)R ¼ the deformation caused by cold stretching (%)

YS ¼ yield strength in tensile test (MPa)UTS ¼ ultimate tensile strength in tensile test (MPa)

A ¼ elongation to fracture (%)RA ¼ reduction in area (%)reng ¼ engineering stressrtrue ¼ true stresseeng ¼ engineering strainetrue ¼ true strain

dr/de ¼ the work-hardening rate of stress–strain curvec ¼ austenite phase

a0, e ¼ a0 martensite phase and e martensite phase

References[1] AS 1210-Supp2, 1999, “Pressure Vessels-Cold-stretched Austenitic Stainless

Steel Vessels.”[2] EN13458-2, 2002, “Cryogenic Vessels-Static Vacuum Insulated Vessels Part 2:

Design, Fabrication, Inspection and Testing.”[3] EN13530-2, 2002, “Cryogenic Vessels-Large Transportable Vacuum Insulated

Vessels Part 2: Design, Fabrication, Inspection and Testing.”[4] ASME Code Case 2596, 2008, “Coldstretching of Austenitic Stainless Steel

Pressure Vessels.”[5] Q/320582SDY7, 2008, “Pressure Strengthening of Cryogenic Vessels From

Austenitic Stainless Steels-Static Vessels.”[6] Hecker, S. S., Stout, M. G., Staudhammer, K. P., and Smith, J. L., 1982,

“Effects of Strain State and Strain Rate on Deformation-Induced Transforma-tion in 304 Stainless-Steel.1.Magnetic Measurements and Mechanical-behav-ior,” Metall. Trans. A, 13(4), pp. 619–626.

[7] Talonen, J., Nenonen, P., Pape, G., and Hanninen, H., 2005, “Effect of StrainRate on the Strain-Induced Gamma ->Alpha’-Martensite Transformation andMechanical Properties of Austenitic Stainless Steels,” Metall. Mater. Trans. A,36A(2), pp. 421–432.

[8] Das, A., Sivaprasad, S., Ghosh, M., Chakraborti, P. C., and Tarafder, S., 2008,“Morphologies and Characteristics of Deformation Induced Martensite DuringTensile Deformation of 304 LN Stainless Steel,” Mater. Sci. Eng., A-, 486(1-2),pp. 283–286.

[9] Chun-chun, X., Xin-sheng, Z., and Gang, H., 2002, “Microstructure Change ofAISI304 Stainless Steel in the Course of Cold Working,” J. Beijing Univ.Chem. Technol., 29(6), pp. 27–31.

[10] GB/T 228, 2002, “Metallic Materials-Tensile Testing at Ambient Temperature.”[11] Talonen, J., Aspegren, P., and Hanninen, H., 2004, “Comparison of Different

Methods for Measuring Strain Induced Alpha’-Martensite Content in AusteniticSteels,” Mater. Sci. Technol., 20(12), pp. 1506–1512.

[12] DIN EN 10028-7, 2008, “Flat Products Made of Steels for Pressure Purposes-Part 7: Stainless Steels.”

[13] ASTM A240/A 240M, 2010, “Standard Specification for Chromium andChromium-Nickel Stainless Steel Plate, Sheet, and Strip for Pressure Vesselsand for General Applications.”

[14] GB 24511, 2009, “Stainless Steel Plate, Sheet and Strip for PressureEquipments.”

[15] Follansbee, P. S., and G.T. Gary, I., 1989, “An Analysis of the Low Tempera-ture, Low and High Strain-Rate Deformation of Ti-6Al-4V,” Metall. Trans. A,20A, pp. 863–874.

[16] De, A. K., Speer, J. G., Matlock, D. K., Murdock, D. C., Mataya, M. C., andComstock, R. J., 2006, “Deformation-Induced Phase Transformation and StrainHardening in Type 304 Austenitic Stainless Steel,” Metall. Mater. Trans. A,37A(6), pp. 1875–1886.

[17] Lo, K. H., Shek, C. H., and Lai, J. K. L., 2009, “Recent Developments in Stain-less Steels,” Mater. Sci. Eng., R., 65(4-6), pp. 39–104.

[18] Fang, X. F., and Dahl, W., 1991, “Strain-Hardening and Transformation Mech-anism of Deformation-Induced Martensite-Transformation in Metastable Aus-tenitic Stainless-Steels,” Mater. Sci. Eng. A, 141(2), pp. 189–198.

061407-6 / Vol. 134, DECEMBER 2012 Transactions of the ASME

Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 10/21/2013 Terms of Use: http://asme.org/terms