Cracking susceptibility of stainless steel subjected to plasma disruption

Fusion Engineering and Design 27 (1995) 499-506

ELSEVIER

Fusion Engineer!ng and Design

Cracking susceptibility of stainless steel subjected to plasma disruption Haruki Madarame a, Toshio Sukegawa a, Hiroshige Inoue

b

a Nuclear Engineering Research Laboratory, University of Tokyo, Tokai-mura, Ibaraki 319-11, Japan b Joining Steel Research Laboratories, Nippon Steel Corporation, Japan

Abstract

The similarities and differences in the cracking susceptibility between welding and resolidification after plasma disruption were examined experimentally using a number of primary candidate alloy samples with. different chemical compositions. The product of the number density and the average depth of the cracks was measured after simulated disruption, employing a hydrogen ion beam as the heat source, and was compared with the Varestraint test result. An adequate correlation was observed between them, which indicates that the cracking susceptibility during plasma disruption can be well estimated from the welding cracking susceptibility.

1. Introduction

An integrated-type first wall made of bare stainless steel is considered in the Fusion Experimental Reactor (FER), the machine which succeeds JT-60 [1]. The very high heat flux during plasma disruption may cause melting of the surface layer of the first wall. If the layer resolidifies without any defects, the same layer repeatedly melts and protects the rest of the wall, and thus the wall lifetime is little affected by the melting. However, a very small solidification crack in the layer may grow during reactor operation to a size larger than the layer thickness and may affect the wall integrity. Although the melting layer thickness has been examined both analytically [2] and experimentally [3-6] by many researchers, very few have reported on the cracks. In a previous paper [7], we have examined quantitatively the effect of heat flux on the size and density of the cracks. In this paper, the effect of the impurity content of the steel on the cracking susceptibility after disrup-

tion is examined and compared with that of welding. Well-established methods have been described [8] to define the cracking susceptibility in the welding process, which have clarified, to a certain extent, the effect of impurities on the weldability. However, the thermal condition during disruption is different from welding. The aims of this study are to reveal the similarities and differences in the cracking susceptibility between welding and resolidification after disruption, and to evaluate quantitatively the effect of impurities on the cracking susceptibility.

2. Testing methods 2.1. Simulated plasma disruption

The experimental method of simulated plasma disruption is basically the same as that described previously [7,9]. The specimen (length, 50 mm; Width,

0920-3796/95/$09.50 © 1995 Elsevier Science S.A. All rights reserved SSDI 0920-3796(94) 00091-3

500

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506

20 mm; thickness, 4 mm) was framed and fixed on a thick copper plate so as not to bend during the thermal transient caused by exposure. After a cooling off period, the specimen was cut in the direction of the width for microscopic observation of the cracks. Most of the cracks were open and grew towards the bottom of the melt layer, although they seldom reached the heataffected zone which did not liquefy. Since the most important factor for the fatigue lifetime is not the number density but the size of the crack [10], the crack length in the direction 0f the thickness (crack depth) was measured in four cross-sections with a total length of 80 mm. The crack density is defined as the total crack depth divided by the length of the cross-section. The Particle Beam Engineering Facility (PBEF) of the Japan Atomic Energy Research Institute (JAERI) was employed as the heat source [11]. The hydrogen beam energy was 63 keV, and the heat flux was uniformly around 90 M W m 2 on the specimen surface for a duration of about 80 ms. Since it fluctuated for each shot because of difficulty in adjusting the ion beam, the melt layer thickness of each specimen was not the same. The melt layer thickness was used as an index of the energy density loaded on the specimen in this experiment. The effect of energy density fluctuation on cracking was checked using several samples as shown in Fig. 1. Although the depth of each crack increased with increasing melt layer thickness, the number decreased [7], resulting in a crack density reduction in thicker melt layers. However, no compensation for the difference in the melt layer thickness was made, because the effect of fluctuation was not large. 2.2. Varestraint test [12]

E E

0.2

I

l

l

l

l

l

l

l

l

l

~

E ~ ~ . .

No.9

"~ t"- 0.1 No.12 "O

~

O

N

o

.

8

O

E E

0

I

0.5

I

I

l

[

l

l

l

r

l

0.30 l

l

l

l

l

0.35

l

t

t

t

l

l

l

No.8

o.4

B_ -O

_.___.____-----~ No.9

E 0.3 E "~ 0.2 0.1

. ~ - - ~ r No.2

I

I

1

[

I

I

I

I

I

I

I

0.30 0.35 Melt layer thickness (mm)

Fig. 1. Effect of energy density on cracking.

u

320

,[

I I

A

I WELDING DIRECTION m

Fig. 2 illustrates the Varestraint testing device. Welding was performed autogenously on the specimen from left to right at 80 A and 10 V with a travel speed of 8 cm rain-1 in Ar gas shielding using the G T A W process. As the arc passed point A in the figure, a yoke bent the specimen downward at a speed of 500 mm s - 1 to conform to the radius of curvature of the top surface of the die block, B. The nominal value of the applied strain in the outer fibres of the specimen was 3.2%. Meanwhile, the arc travelled steadily onward and was subsequently interrupted in the run-off area at C. Hot cracking results when a susceptible microstructure is subjected to a tensile stress. There are several hot cracking tests, and the Varestraint test is the most popular. The length of the crack on the specimen surface running radially in the annular solid-liquid interface zone at the instant of strain application is

l

SPECIMEN

II

/~t, ARC

DIE BLOCK GTAW(TIG) 80A- 8 cm/min Arshield

R~L,~

t / 2 R = 3.2%

Fig. 2. Varestraint test device. proportional to the cracking temperature range for the welding procedure and the level of augmented strain [12]. Two specimens were tested for each sample, and

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506 the length o f every crack was m e a s u r e d to provide quantitative indices of the cracking sensitivity.

~'15 E

i

l

l

l

0

O A Varestraint test

~10 •---~

r-

It is well k n o w n t h a t a suitable a m o u n t o f 6-ferrite distributed in the austenitic matrix is helpful in preventing h o t cracking [13]. As a preliminary experiment, the effect o f &-ferrite was e x a m i n e d using 28 types o f stainless steel (base materials were types 304, 316 a n d prim a r y candidate alloy ( P C A ) ) with different impurity contents. T h e chemical compositions of some o f the P C A s are s h o w n in Table 1 (samples 1 - 7 ) . The experim e n t a l conditions were n o t necessarily the same as the m a i n experiment described in the previous a n d following sections; the heat source o f the simulated disr u p t i o n was the N B I at N a g o y a University, which generated a b o u t twice the crack density as the P B E F at J A E R I . T h e a m o u n t o f 5-ferrite increases with increasing Crcq/Nieq ratio, where Creq = % C r + % M o + 1.5%Si + 0.5%Nb a n d Nieq = % N i + 30%C + 0 . 5 % M n [14]. In Fig. 3 the results are a r r a n g e d using the

0.15~

. . . . . .

l

v

3. Test samples

501

•

• • NBI test

PCA 316 304

[:: 1:: 0.10 E C

%

0 . 0 5 -a 0

~--~

0

0

I

1.0

0.8

,

,

,

,

0 --~

1.2 1.4 1.6 1.8 Creq / Nieq (-)

'

0.00 2.0

Fig. 3. Effect of 6- ferrite on cracking. Creq/Nieq ratio as the abscissa. B o t h the crack density after the simulated disruption a n d the total crack length of the Varestraint test decreased with increasing ratio, a l t h o u g h there was considerable scatter w h e n t h e ratio was low. T h e m a x i m u m crack d e p t h after the simulated disruption a n d the m a x i m u m crack length o f the Varestraint test showed a similar tendency. 6-Ferrite ( 2 . 1 % -

Table 1 Chemical compositions of the test samples Sample

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

Chemical composition (wt.%) C

Si

Mn

P

S

Ni

Cr

Mo

A1

Ti

B

0.049 0.050 0.048 0.050 0.048 0.048 0.049 0.049 0.150 0.280 0.051 0.052 0.049 0.053 0.051 0.050 0.049 0.059 0.052 0.051 0.051 0.049

<0.1 0.56 0.52 0.52 0.54 0.54 0.52 0.51 0.53 0.52 0.12 0.71 1.04 0.51 0.50 0.51 0.51 0.51 0.50 0.50 0.51 0.50

1.50 1.47 1.49 1.48 1.47 1.47 1.48 1.50 1.50 1.49 1.50 1.49 1.49 1.50 1.50 1.49 1.51 1.51 1.51 1.51 1.51 1.50

0.025 0.025 0.027 0.027 0.026 0.025 0.029 0.027 0.027 0.026 0.028 0.027 0.028 0.026 0.026 0.026 0.027 0.027 0.027 0.026 0.028 0.026

0.0024 0.0025 0.0021 0.0022 0.0023 0.0023 0.0020 0.0014 0.0016 0.0018 0.0017 0.0018 0.0016 0.0014 0.0014 0.0014 0.0015 0.0019 0.0017 0.0014 0.0017 0.0015

16.62 16.54 16.65 16.64 16.64 16.68 16.65 16.80 16.77 16.68 16.76 16.63 16.59 16.65 16.59 16.60 16.68 16.63 16.70 16.61 16.64 16.68

14.90 14.81 14.83 14.79 14.78 14.82 14.80 14.92 14.84 14.84 14.96 14.89 14.89 15.02 14.91 14.86 15.10 14.96 14.95 15.05 15.09 15.10

2.10 2.07 2.11 2.11 2.09 2.10 2.09 2.16 2.15 2.15 2.16 2.14 2.13 2.14 2.11 2.11 2.11 2.10 2.12 2.08 2.12 2.09

0.028 0.030 0.030 0.030 0.007 0.028 0.029 0.029 0.029 0.027 0.033 0.032 0.032 0.034 0.035 0.037 0.014 0.066 0.290 0.035 0.035 0.039

0.25 0.25 0.005 0.06 0.23 0.25 0.51 0.27 0.27 0.26 0.27 0.27 0.27 0.006 0.51 0.80 0.26 0.26 0.27 0.28 0.28 0.28

0.0022 0.0030 0.0024 0.0031 0.0026 0.0001 0.0024 0.0020 0.0022 0.0022 0.0019 0.0019 0.0020 0.0020 0.0020 0.0021 0.0020 0.0020 0.0020 0.0012 0.0063 0.0114

Nitrogen, 0.004-0.007 wt.%; oxygen, 0.001-0.004 wt.%.

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506

502

3.4%) was observed in all specimens of types 304 and 316 stainless steel after the Varestraint test, while no 5-ferrite was observed in the P C A specimens. The result confirmed the beneficial effect of 5-ferrite, although other factors also affected the cracking susceptibility especially of the P C A samples. The first wall material must withstand a high neutron fluence as well as a high heat flux during disruption. Thus it is worthwhile improving the cracking susceptibility by impurity control in P C A s which have good mechanical properties on neutron irradiation. 5-Ferrite, showing ferromagnetic properties, may be undesirable in the first wall, since it disturbs the magnetic fields for plasma confinement. The chemical compositions of the PCAs used in this study are shown in Table 1. Samples 2 and 8 were standard samples, and the effect of each element was investigated by comparison with these standard samples. Samples with high sulphur and oxygen content were not prepared since they exhibited a rough and wavy resolidification surface which affected the crack distribution and made the measurement difficult [9].

4. Experimental results and discussion 4.1. Cracking sensitivity indices The test results are listed in Tables 2 and 3. Fig. 4 shows the relation between the crack density after

0.4 E E E 0.3

l

0

g

O3

g 0.2

0

0

0 0

c30.1

°O O°

0.0 2.0

~ I

No.5 I

4.0 6.0 8.0 1 0 . 0 12.0 Total length per weld (mm)

Fig. 4. Relation between the crack density after disruption and the total crack length of the Varestraint test. disruption and the total crack length of the Varestraint test. These are the clearest indices of the cracking tendency, the product of the number and the size of the cracks. There is an adequate correlation between them; however, some data do not follow the general trend, especially sample 5. The reason will be discussed in detail later. The size is more important than the number of cracks in terms of the fatigue lifetime after disruption [ 10]. Fig. 5 compares the measured m a x i m u m crack depth after simulated disruption with the measured maximum

Table 2 Results of simulated plasma disruption Sample

Tm

n

Dens.

Dmax

Sample

Tm

n

Dens.

Dmax

1 2 2 3 4 5 6 7 8 8 9 9 10 10

0.28 0.30 0.27 0.31 0.28 0.28 0.31 0.28 0.28 0.36 0.28 0.35 0.33 0.33

0.36 0.50 0.93 0.56 0.94 0.48 0.63 1.05 0.58 1.08 1.85 1.14 2.41 1.55

0.023 0.041 0.056 0.027 0.044 0.029 0.032 0.098 0.043 0.062 0.168 0.104 0.354 0.209

0.134 0.191 0.137 0.094 0.087 0.141 0.118 0.232 0.198 0.151 0.236 0.238 0.432 0.385

11 11 12 12 13 14 15 16 17 18 19 20 21 22

0.31 0.30 0.27 0.33 0.31 0.30 0.36 0.37 0.33 0.37 0.39 0.37 0.35 0.34

0.88 0.71 0.93 0.56 0.78 0.45 0.46 0.81 0.90 0.28 0.29 0.81 0.76 1.36

0.055 0.048 0.088 0.052 0.098 0.028 0.046 0.098 0.067 0.021 0.025 0.052 0.102 0.206

0.141 0.148 0.244 0.240 0.361 0.102 0.246 0.316 0.174 0.166 0.220 0.136 0.373 0.394

Tin, melt layer thickness (mm); n, number density (mm-1); Dens., crack density (mm mm-l); D ..... maximum crack depth by extreme value analysis (mm).

503

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506

Table 3 Results of Varestraint test Sample

N

Lt

Lma x

Sample

N

Lt

Lmax

Sample

N

Lt

Lmax

1 2 3 4 5 6 7 8

20.5 17.5 19 19.5 21.5 19.5 27.5 10.5

6.6 7.1 5.8 7.9 11.2 6.6 8.4 5.5

0.92 1.07 0.95 1.13 1.19 0.93 1.03 0.97

9 10 11 12 13 14 15 16

14.5 15.5 13.5 17 12 10.5 8 10.5

7.1 11.1 4.4 5.6 5.2 4.7 4.1 6.6

1.14 2.69 0.78 0.91 1.12 0.86 1.07 1.45

17 18 19 20 21 22

10.5 11 22 14 15.5 14

5.1 4.9 8.1 4.1 8.3 9.7

0.96 1.00 1.04 0.72 1.63 1.94

N, number of cracks per weld (average of two specimens); L t , total crack length per weld (average of two specimens) (mm); L . . . . maximum crack length at weld by extreme value analysis (mm).

0.5

E E0.4 e-

i

0

Q,)

0

•

0.5

i

0 0

0

o~@

o

0

_o~0.3

0

i-

0

0 0

0

0

~'0.3

-O

CI30

E .m

E 0.2

~0.2 X

E go.4

0 0.1

0.1

0.0

I

I

1.0 2.0 3.0 Maximum length at weld (mm)

0.0

I

I

1.0 2.0 3.0 Maximum length at weld (ram)

Fig. 5. Relation between the measured maximum crack depth after disruption and the measured maximum crack length of the Varestraint test.

Fig. 6. Relation between the statistically treated maximum crack depth after disruption and the statistically treated maximum crack length of the Varestraint test.

crack length of the Varestraint test. The maximum crack is often found near the periphery after simulated disruption in which the temperature has a gradient parallel to the surface as well as in the thickness direction. The melt layer sometimes moves towards the periphery; therefore a deeper crack than the melt layer thickness does not necessarily reach the heat-affected zone beneath the layer; the thickness is not defined as the distance from the swollen surface to the bottom of the layer but from the original surface [7]. The size of a few cracks in the peripheral swelling does not seem to represent the crack size of the specimen. The average "crack depth is not a good index either, since it is affected by the omission of small cracks. Therefore statistical treatment was needed.

The maximum crack depth in Table 2 and the maxim u m crack length in Table 3 are not raw measured data; they were obtained by extreme value analysis assuming a Gumbel distribution (the double exponential extreme value distribution). Commercial software for personal computers was used, where the return period of 100 was assumed to give a cumulative probability of 99%. A relation appears, as shown in Fig. 6, between the statistically treated values; the crack size after disruption is large as long as the sample exhibits long cracks at welding. However, samples showing short cracks at welding sometimes have fairly large cracks after disruption. Thus long cracks at welding lead to deep cracks, while short cracks do not always result in shallow cracks after disruption. The crack

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506

504

length of the Varestraint test is a measure of the cracking temperature range since every crack is generated at the instant of bending. The position of crack initiation relative to the liquid-solid interface is not known during simulated disruption, and the intensity of the thermal stress causing cracks is not controlled but determined by certain properties of the material. Other properties such as the mobility of the melt layer may also affect the size. The relation in Fig. 6 indicates that the cracking temperature range plays an important role in the crack size after disruption. The size and number of the cracks after disruption are not correlated [7]; this is also the case for the results of the Varestmint test. However, the product of the size and number of the cracks is correlated fairly well with the size. The coefficient of correlation between the total and maximum crack lengths, excluding sample 5, is as high as 0.80 because the scatter of the crack number is not large. In the following, we focus on the total combined crack length of the Varestraint test and the crack density after simulated disruption in order to avoid duplication.

4.2. Effect of impurities It is known that impurities£ such as sulphur and phosphorus, are detrimental to }cracking prevention [ 13]; samples for checking the effect of these elements were not prepared in the experiment. The effect of carbon content is illustrated in Fig. 7. A decrease in carbon content is beneficial to crocking prevention during plasma disruption and welding, although a much higher carbon content may also have a beneficial effect

~'15 E -o

.

.

.

.

.

.

.

.

0.3

i

o Total length

E

E

• Crack density

~'15 E -o

. . . . . . . .

,10

/o No.5

t-

O

5

O ¢0

~6 F-

E E E 0.2 E

• Crack density

%

I3_

" O

0.3

i

o Total length

v

....

o

o

•

1

I

O c-

"O

o

0.1 -g o

. . . . . . . .

10- 2

I

''

10-1 AI content (wt~)

0.0

Fig. 8. Effect o f a l u m i n i u m o n cracking.

[15]. Fig. 8 shows the effect of aluminium; the Varestraint test indicates that an extremely low content of aluminium is detrimental, whereas the simulated disruption test does not. Aluminium is known to be beneficial to hot working, which suggests that steel with extremely low aluminium content lacks ductility at elevated temperature. A high strain rate in the Varestraint test is considered to be the cause o f the discrepancy between the results in Fig. 8, and the cause of the exceptional data in Fig. 4. Boron segregates at grain boundaries forming chemical compounds of low melting point; therefore a high boron content results in high cracking susceptibility. The boron effect was confirmed in both the simulated disruption and Varestraint tests as shown in Fig. 9. The results in Fig. 10 suggest that a decrease in titanium is beneficial, but the effect is very small and

~'15 E -o

'"1

. . . . . . . .

i

. . . . . . . .

i

0.3

o Total length

E E E

• Crack density

E

0.2 vE

$10 Q. t-

o

O

0~

"O

o

q) C

0

"O

•

o

5

0.1

"

!

o

O

0 10- 2

t

n

n

i

i i i , l

i

i

10-1 C content (wt°,%)

,

i

irn

9.0 10°

Fig. 7. Effect of c a r b o n o n cracking.

09

o

0

C

O

O

0.2 E

$10

o

5

C

•

0,1

0

O

O

"5 b0

,,,I

10-4

,

. . . . . . .

i

. . . . . . .

10- 3 B content (wt°~)

,I

10- 2

Fig. 9. Effect of b o r o n on cracking.

0.0

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506

~'15 E

. . . . . . . .

0.3

I

o Total length • Crack density

E E E 0.2 vE

1"3 I'-

0

t--

" 0

O

5

>,

0

0

t¢.O

0 -O

0

0.1

O

0

O

0

"6

0

.......

I

,

. . . . . .

I

10 - 2

. . . . . . . .

10 -1

I

10 °

0.0

Ti content (wt%) Fig. 10. Effect of titanium on cracking.

~'15

g

. . . . . .

i

. . . . . . . .

I

03

o Total length • Crack density

I11

E

13. r-"

"0

<

0

0

5

C

0

•

0.1

0 0

,

, ,,,,I

We are grateful to Dr. M. Akiba of the Japan Atomic Energy Research Institute for his help in the PBEF experiment. We are also indebted to Professor T. Kuroda of the National Research Institute of Fusion Science for the NBI experiment at Nagoya University.

-o

0

o I-

an index of the cracking susceptibility after plasma disruption, and showed an adequate correlation with the cracking susceptibility at welding. The maximum size of the crack after disruption was correlated with the maximum crack length of the Varestraint test when the data were statistically treated assuming a Gumbel distribution. However, when the latter showed smali values, the former varied widely because it was affected not only by the cracking temperature range of the material but also by other physical properties. The cracking susceptibility after disruption was largely dependent on the Creq/Nieq ratio owing to the beneficial effect of ~-ferrite. Decreases in carbon and boron content were also beneficial for cracking prevention in the experimental range. The contents of aluminium, titanium and silicon had only a small effect on the cracking susceptibility.

Acknowledgments

E 0.2 E i--

505

,

,

,

i ,,.,I

10 -1 10 ° Si content (wt°~)

0.0

Fig. 11. Effect of silicon on cracking. no reason is known for this tendency. Fig. 11 shows the

effect of silicon. No remarkable trend is observed. The effect of impurity content on hot cracking has not been fully revealed. Some of the present results can be explained by existing knowledge and some cannot. An adequate correlation is observed between the results of the Varestraint test and simulated plasma disruption, which indicates that the cracking susceptibility during plasma disruption can be well estimated from the cracking susceptibility at welding.

5. Conclusions

The crack density, the product of the number density and the average depth of the cracks, was proposed as

References

[1] JAERI, Conceptual design study of fusion experimental reactor (FER) (FY 1984, 85 Report), Rep. JAERI-M86134, Japan Atomic Energy Research Institute, November, 1986 (in Japanese). [2] A.M. Hassanein, G.L. Kulcinski and W.G. Wolfer, Surface melting and evaporation during disruptions in magnetic fusion reactors, Nucl. Eng. Des./Fusion, 1 (1984) 307-324. [3] M. Seki, S. Yamazaki, A. Minato et al., A simulated plasma disruption experiment using an electron beam as a heat source, J. Fusion Energy, 5 (1986) 181-189. [4] G. Rigon, P. Moretto and F. Brossa, Experimental simulation of plasma disruption with an electron beam, Fusion Eng. Des., 5 (1987) 299-315. [5] H. Yanagi, T. Sukegawa, K. Kobayashi et al., High heat load experiments for first wall materials, J. Nucl. Mater., 155-157 (1988) 402-406. [6] M. Ogawa, M. Araki, M. Seki et al., Experimental study on melting and evaporation of metal exposed to intense hydrogen ion beam, Fusion Eng. Des., 19 (1992) 193-202. [7] H. Madarame and T. Sukegawa, Cracks in the resolidification layer of stainless steel formed by simulated plasma disruption, Fusion Eng. Des., 19 (1992) 339-346. [8] G.M. Goodwin, Test methods for evaluating hot cracking: review and perspective, Proc. 1st Japan-US Symp.

506

H. Madarame et al. / Fusion Engineering and Design 27 (1995) 499-506 i

on Advances in Welding Metallurgy, Yokohama, 1990, pp. 59-78. [9] H. Madarame, T. Sukegawa, H. Yanagi et al., Effect of impurity content in stainless steel on resolidified surface condition after disruption load, Fusion Eng. Des., 9 (1989) 213-217. [10] H. Kobayashi, Y. Arai, Y. Kajiyama et al., Evaluation of fracture strength and residual lifetime of materials damaged by simulated plasma disruption, Fusion Eng. Des., 19 (1992) 307 313. [11] M. Seki, M. Araki, K. Yokoyama et al., Thermal shock tests on various materials of plasma facing components for FER/ITER, Fusion Eng. Des., 15 (1991) 59-74.

[12] W.F. Savage and C.D. Lundin, The Varestraint test, Welding J., 44 (1965) 433-s-442-s. [13] V. Kujanppa, N. Suutala, T. Takalo et al., Correlation between solidification cracking and microstructure in austenitic and austenitic-ferritic stainless steel welds, Weld. Res. Int., 9 (1979) 55-75. [14] A.L. Schaeffler, Constitution diagram for stainless steel weld metal, Metal Progress, 56 (1949) 680-682. [15] Y. Arata, F. Matsuda and S. Katayama, Solidification crack susceptibility in weld metals of fully austenitic stainless steels (Report II)--effect of ferrite, P, S, C, Si and Mn on ductility properties of solidification brittleness, Trans. Jpn. Weld. Res. Inst., 6 (1977) 105-116.