Jinyang Zheng
Cunjian Miao
Yaxian Li
Institute of Process Equipment, Zhejiang University, Hangzhou, 310027 P. R. China
Ping Xu
1Institute 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 Influence
Factors of Mechanical Properties
of Austenitic Stainless Steels for
Cold Stretched Pressure Vessels
Cold stretched pressure vessels from austenitic stainless steels (ASS) have been widely used all over the world for storage and transportation of cryogenic liquefied gases. Cold stretching (CS) is performed by pressurizing the finished vessels to a specific pressure to produce 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 in
welding procedure qualification for cold stretched pressure vessels, are considered to be important factors to mechanical behavior of ASS. During the qualification, welded joint will 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. There are two kinds of preloading method. One is to apply required tensile stress rkon specimen
and 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. In order to understand the effects of nickel equivalent, preloading and preloading rate on the mechanical 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 tensile strength (UTS) decreased as the nickel equivalent increased, while the elongation to frac-ture increased. It was more difficult to meet the available mechanical requirements with strain-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 effect on the mechanical properties of preloaded welded joint. These results are helpful for choosing appropriate material and determining a proper preloading method for welding procedure 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 widely used all over the world for storage and transportation of cryogenic liquefied gases, and guidances have been implemented in several standards 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 appendix of ASME BPVC VIII-I: 2011). Cold stretched pressure vessels are manufactured from finished vessels through CS, which is per-formed by pressurizing the finished vessels to a specific pressure to produce the required stress rk. After CS, an amount of plastic
deformation is given to withstand the pressure load. Such vessels will 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 manufacturing and transportation.
Nieqis an important factor to ASS mechanical behavior for cold
stretched pressure vessels. It could be used to describe the austen-ite stability, which has a strong effect on the deformation-induced martensite (DIM) transformation. Because of DIM’s main influ-ence on the mechanical behavior, Nieqmay indirectly affect ASS
mechanical properties. Preloading is another important factor. During welding procedure qualification for cold stretched pressure vessels, the finished welded plate will be preloaded and unloaded, leaving a permanent plastic deformation. Then the preloaded welded joint made from the plate will go through tensile test to obtain its mechanical properties, which will be evaluated by the standard requirements. Preloading makes the qualification differ-ent from the normal one. It introduces an effect similar to the CS effect on vessels into the qualification to help know whether this qualification is proper for cold stretched pressure vessels. Two kinds of preloading are employed now. One is to apply the required stress rk on the specimen and hold this stress until the
strain rate gets lower than 0.1%/h [2–4]. The other is to stretch the specimen to a specific strain of 9% with a quasi-static strain rate [5]. These two preloading methods may lead to different results in mechanical properties of preloaded welded joint. In this paper, these preloading methods were noted as rk-stress-preloading and
9%-strain-preloading, respectively. Preloading rate is also consid-ered to be an important parameter. The effects of strain rate on DIM and mechanical properties of some ASS grades were studied at strain rate between 104 and 103/s [6–8], and the rate was proved to influence the mechanical behavior to a certain extent. Little researches were reported about the effect of preloading rate on 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 stretched pressure vessels, a series of tests were performed with a commercial 1
Corresponding author.
Contributed by the Pressure Vessel and Piping Division of ASME for publication in the JOURNAL OFPRESSUREVESSELTECHNOLOGY. Manuscript received November 3, 2011; final manuscript received April 25, 2012; published online November 21, 2012. Assoc. Editor: David L. Rudland.
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 preloaded welded joints were made. Third, tensile tests were carried out on base metal, welded joint, and preloaded welded joint. Furthermore, some DIM measurements were carried out. Based on these test results, analysis was done and some suggestions were put forward. These findings will be helpful for choosing the appropriate material and determining a proper preloading method in welding procedure qualification.
2
Test Procedure
2.1 Test Material, Chemical Composition, and Nickel Equivalent. This investigation was carried out on industrially manufactured ASS grade EN 1.4301 in as-received condition. These material plates were available with the thicknesses from 6 mm to 20 mm, and in a hot-rolled and solution heat-treated con-dition. The chemical composition is listed in Table1, as well as Nieq, which was calculated by the following equation with the
consideration of chemical composition, temperature and deforma-tion [9]:
Nieqð Þ ¼ Ni þ 0:65Cr þ 0:98Mo þ 1:05Mn þ 0:35Si þ 12:6C%
þ 0:03 T 300ð Þ 2:3log 100= 100 R½ ð Þ 2:9 (1) whereT is the temperature (K) and R is the deformation (%). Room temperature was considered. The effect of the deformationR caused by 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 the
Nieqvalues of solution annealed (SA) and CS materials are nearly
the same, and the effect of deformation is ignored for convenience. 2.2 Preloading and Tensile Test. Preloading was performed on welded plates using quasi-static strain rates. Both 405 MPa-stress-preloading (the stress rkwas 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 preloading methods are shown in Fig.1. For the 405 MPa-stress-preloading case, the curve starts from O, and goes elastically until yielding
part A. Then, it goes up to B before which the rate is about 1 104/s. The stress of B is r
k (405 MPa), and this stress is held
for 1–2 h until C, while the strain rate keeps decreasing slowly. For 9%-strain-preloading case, the curve goes directly to D with the strain rate of about 1 103/s. The yielding is also located in part
A. The ending D of 9%-strain-preloading has a higher stress and strain than the ending C of 405 MPa-stress-preloading. After being unloaded, the specimens of both preloading case were reloaded in the subsequent tensile test (STT) to the failure with a strain rate of about 1 103/s. A test number summary for Ni
eqinvestigation is
listed in Table2, and the total data numbers are 12, 5, 13 for base metal, 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 done through 9%-strain-preloading. Strain rates of 1 105 and
1 103/s were employed during this preloading, while strain
rate of 2.5 103/s was used in all the STT processes
correspond-ing to the typical rate in tensile tests of engineercorrespond-ing application. The tests with 103/s in preloading and 2.5 103/s in STT were
marked as type A, and those with 105/s in preloading and 2.5 103/s in STT were recognized as type B. Each type of tests
had two specimens.
All the tests were performed in air by using a servo hydraulic MTS 810 tensile testing machine. Specimens with rectangular cross section were used in the tests for investigating the effect of Nieq, and specimens with a 5-mm-diameter and a gauge length of
25 mm were used in the study for the influence of preloading and its rate. All specimens were made according to GB/T 228 [10] (equivalent to ISO 6892) and cut parallel to the rolling direction of the plates. Strain data were measured with an MTS 634.12 F-25 extensometer.
2.3 Measurement of DIM. The a0-martensite contents of specimen were measured during the experiments for the effect of preloading 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 18.75 2 0.051 0.53 1.34 0.030 0.001 18.14 8.22 0.060 19.47 3 0.058 0.49 1.64 0.028 0.001 18.12 8.23 0.051 19.83 4 0.034 0.45 1.02 0.023 0.003 18.01 8.05 0.067 18.61 5 0.055 0.51 1.07 0.027 0.001 18.28 8.15 0.061 19.22 6 0.055 0.64 1.35 0.029 0.013 17.36 8.03 0.041 18.75 7 0.041 0.59 0.80 0.022 0.001 17.71 8.00 0.039 18.17 8 0.034 0.47 0.83 0.027 0.002 17.43 8.02 0.041 17.91 9 0.039 0.49 0.99 0.029 0.003 17.52 8.07 0.044 18.26 10 0.055 0.64 0.94 0.026 0.002 17.42 8.10 0.042 18.43 11 0.043 0.54 0.88 0.027 0.002 17.25 8.09 0.039 18.06 12 0.022 0.44 1.70 0.027 0.008 17.28 8.31 0.040 18.86 13 0.041 0.34 0.88 0.026 0.002 18.35 8.04 0.045 18.63 14 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 of preloading methods
Table 2 A test number summary for Nieqinvestigation
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 12
405 MPa-stress-preloading 1 1 1 1 1 0 0 0 0 0 0 0 0 0 5
(model FMP30, Helmut Fischer GmbHþCo.), which is normally used to measure d-ferrite content in austenitic and duplex steels and to determine the fraction of DIM in austenitic materials. It was found that ferritescope measurement was an efficient way to measure the content of the ferromagnetic a0-martensite phase. The
measurement should be performed on an electropolished surface of 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 results will be converted to actual martensite contents with a calibration curve [11].
3
Influence of Ni
eqon Mechanical Properties
3.1 Requirements of Mechanical Properties. The effect of Nieqon the mechanical properties of ASS and preloaded ASS are
shown from Figs.2–4, while the mechanical properties included yield strength (YS), UTS and elongation to fracture (A, which is always used to describe the elongation to fracture in the relevant standard [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 Table3to help reveal the effects of Nieq. It is clear
in Table 3 that YS increased to about 405 MPa from about 205 MPa after preloading, and A decreased to 25% from about 40%. The UTS for preloaded ASS was not specified, while 520–720 MPa was employed in the Chinese company standard of Zhangjiagang CIMC Sanctum Cryogenic Equipment Co., Ltd [5].
3.2 Influence of Nieq. It was clear in Fig.2that preloading
increased YS obviously. The 9%-strain-preloading gave a higher value of YS than that of 405 MPa-stress-preloading. Both pre-loaded ASS met the yield strength requirement of 405 MPa. As Nieqincreased, YS of 9%-strain-preloading ASS decreased, while
the curves of both base metal and 405 MPa-stress-preloading ASS varied slightly. The YS data of 405 MPa-stress-preloading were relatively well behaved, the reason of which is that the specimens of 405 MPa preload case were all preloaded to 405 MPa and held under the stress for about 1–2 h.
The UTS variation as a function of Nieqwas showed in Fig.3.
The curves of base metal and 9%-strain-preloading ASS decreased with increasing Nieq. They decreased to the UTS upper
limit of 720 MPa at about 18.45%, and then continued decreasing. Furthermore, the curve of preloaded ASS was higher below the line of 720 MPa. The lower limit of 520 MPa seemed not to be reached. The 405 MPa-stress-preloading curve showed a slightly decreasing trend and varied gradually.
All the curves of A in Fig. 4increased with increasing Nieq.
The values of preloaded ASS were obviously lower than that of base metal, while the curve of 405 MPa-stress-preloading was a little higher than that of 9%-strain-preloading. Additionally, the A of 9%-strain-preloading curve increased to 25% as was required at 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 is determined from the strengthening stress rkin the relevant
stand-ard [2,3] and used in USA. The 9%-strain-preloading is always used in China [5]. Tests with these two preloading methods were carried out and results were showed above. Main differences between these methods are described here.
(1) It is more difficult to meet the available mechanical behav-ior requirements with the 9%-strain-preloading case than with 405 MPa-stress-preloading case. Therefore, when the former case meets the requirements, the latter one can make it too. However, the former one is not sure to meet the requirements even when the latter does.
(2) Time spent on the 405 MPa-stress-preloading is about 1–2 h, which is much longer than that of strain-preloading (about several minutes). Therefore, using 9%-strain-preloading could save much time for engineering applications.
It is concluded that the 9%-strain-preloading is more strict than 405 MPa-stress-preloading about the requirements of mechanical properties, and could spare much more time in preloading.
Fig. 2 The effect of Nieqon YS Fig. 4 The effect of Nieqon A
4.2 Preloading Rate.
4.2.1 Mechanical Properties. After 9%-strain-preloading, the true 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 were listed in Table4, as well as the time spent on these tests. The true stress and strain were calculated from the engineering values through the following formulas
rtrue¼ rengð1 þ eengÞ (2)
etrue¼ lnð1 þ eengÞ (3)
The shapes of preloading curves showed parabolic behavior at both the strain rates of 103/s and 105/s, while in STT the curves between YS point and UTS point were almost linear. YS increased with the increasing strain rate in preloading, and the strength at 103/s was about 30–50 MPa higher than that at 105/s. Some
similar phenomenon has been experimentally observed by others [7]. To considering the data scatter influence, another group test was conducted with the same specimen size and test parameters. The comparison in Fig. 5(b) showed that the true stress–strain curves 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 the stress–strain relationship, which means a specific strain rate could lead to only one stress–strain relationship. Thus, the difference was considered to be caused mainly by the strain rate difference. When loaded in STT, the curves at the same rate 2.5 103/s did not
fol-low their former trends, which may be caused by the difference between strain rates of preloading and STT. However, they looked like each other in STT, and seemed not to be affected by the differ-ence between preloading rates, which could also be found from the data in Table4. In addition, the time spent on the preloading of type A was much shorter.
4.2.2 DIM Transformation. DIM mass fraction in preloading and STT was depicted in Fig.6(a), and its distribution along the gauge after fracture was showed in Fig.6(b). The DIM of types A and B showed the same trend. It was observed that the DIM of type 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% strain and became to be linear after about 12.5% strain. The difference in preloading rates seemed to have a little effect on preloading and nearly no effect on the STT. The final DIM distributions along the specimens’ gauge length after fracture were also nearly the same between types A and B. The contents of DIM increased when the distance from the fracture location decreased.
4.2.3 Work-Hardening Rate. The interaction between the work-hardening rate (dr/de) and true strain was illustrated in Fig. 7. During preloading, both work-hardening rates decreased rapidly, which may be mainly due to the appearance of e phase [16]. Meanwhile, there were only a little DIM content showed in Fig.6(a). When it was loaded in STT, DIM became to grow fast in Fig.6(b), while the work-hardening rate became to increase. It is generally accepted that a0-martensite has a strong effect on work-hardening of ASS, and during the STT the transformation was considered to follow the routes of c!a0or 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 the other hand, the rapid decrease ofdr/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 40
EN10028-7:2008 210 520–720 45
GB 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 — 25
Preloaded welded joint Q/320582SDY7—2008 410 520–720 25
Fig. 5 True stress–strain curves. (a) True stress–strain curves of type A and B; (b) the comparison on true stress–strain curves between different group tests with the same parameters.
Table 4 Mechanical properties during preloading and STT
Type
Test step and
strain rate YS/MPa UTS/MPa A/% RA/% Time
A Preloading at 103/s 310 — — — 90 s (1.5 min)
STT at 2.5 103/s 500 750 61.5 78.0 3–4 min
B Preloading at 105/s 280 — — — 9000 s (2.5 h)
also be explained by the stress–strain behavior. In elastic loading, the change of strain was small for large changes of stress, which led to an initial highdr/de. The dr/de decreased rapidly when the material was yielding, and leveled off since the stress–strain response 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 between work-hardening rates occurred at the beginning of yielding (when the value of work-hardening rate was high) could still make a big effect on the strength.
4.2.4 Flow Stress. Figure 8 showed the dependence of the flow stress on the square root of a0-martensite fraction of types A and B. The flow stress is defined as the true stress after yielding, and it began from about 300 MPa and 500 MPa during preloading and STT, respectively.
It was found that the flow stress of both type A and B seemed to have a linear relation with the square root of the a0-martansite fraction during STT and some part of preloading, and similar phe-nomenon was found in tensile tests by Fang and Dahl [18] and Juho [7] in tensile tests. When in preloading, the flow stress decreased as the square root of the a0-martansite fraction decreased and finally behaved vertically, which could be explained by the little variation of DIM content closed to zero dur-ing preloaddur-ing from Fig.6(a). It is accepted that the flow stress of the 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 accumulatransforma-tion of dislocatransforma-tions in the austenite phase, thus indirectly increase the flow stress. Therefore, it has been suggested that the linearity is not necessary to indicate the relationship between flow stress and DIM content, it is only an indirect effect of DIM on the flow stress through the dislocation density. These curves of type A and B showed that different pre-loading rates may affect the curves of prepre-loading a little but had no 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 of these influence factors on the mechanical behavior of ASS for cold stretched pressure vessels.
(1) Increasing Nieqlowered the YS of 9%-strain-preloaded ASS
while it increased that of base metal. A weak effect caused by varying Nieqwas found on the YS of 405
MPa-stress-pre-loaded ASS. All the YS of preMPa-stress-pre-loaded ASS were higher than 405 MPa, which is the minimum requirement of the YS. Increasing Nieqalso lowered the UTS of base metal and
9%-strain-preloaded ASS. These UTS would be higher than 720 MPa when Nieqwas lower than about 18.45%. As UTS
is limited to 520–720 MPa, it was better for Nieqto be higher
than 18.45%. There was also a small variation on the UTS of 405 MPa-stress-preloaded ASS due to the changing Nieq.
Higher Nieqcaused higher value of A. For preloaded ASS, Fig. 6 DIM transformation during the tests. (a) DIM mass
frac-tion as a funcfrac-tion of true strain in preloading and STT; (b) distri-bution of DIM mass fraction along the specimen gage length after fracture.
Fig. 7 The correlation between work-hardening rate dr/de and true strain in preloading and STT
Fig. 8 Flow stress as a function of the square root of the a0 -martansite fraction
the lowest value of 25% was required. Results showed that the A of 9%-strain-preloaded ASS became higher than 25% when Nieqwas higher than 19.15%.
(2) In the welding procedure qualification for cold stretched pressure vessels, 9%-strain-preloading is stricter than 405 MPa-stress-preloading about the requirements of me-chanical properties. When the meme-chanical properties results of 9%-strain-preloaded ASS meet the requirements, those of 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 rate increased the strength of true stress–strain curve and led to a little higher DIM. The results of two preloading rates (103/s and 105/s) slightly differed on the work-hardening rate and the relationship between flow stress and square root of a0-martansite fraction. However, with different rates in preloading and the same rate of 2.5 103/s in STT, effect was hardly found on STT mechanical behavior of preloaded ASS.
Acknowledgment
The authors gratefully acknowledge the financial support from the International Science and Technology cooperation project (2010DFB42960) and National Key Technology R&D Program (2011BAK06B02-05).
Nomenclature
rk¼ the specific stress for cold stretching (MPa)
Nieq¼ nickel equivalent (%)
T¼ the temperature in the test (K)
R¼ the deformation caused by cold stretching (%) YS¼ yield strength in tensile test (MPa)
UTS¼ ultimate tensile strength in tensile test (MPa) A¼ elongation to fracture (%)
RA¼ reduction in area (%) reng¼ engineering stress
rtrue¼ true stress
eeng¼ engineering strain
etrue¼ true strain
dr/de¼ the work-hardening rate of stress–strain curve c¼ austenite phase
a0, e¼ a0martensite phase and e martensite phase
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