Weldability of Invar and Its Large-Diameter Pipe

Hot-cracking resistance of reheated Invar weld metal is improved by reducing sulfur and impaired by adding

titanium and boron

BY T. OGAWA

ABSTRACT. Sulfur remarkably increases the hot-cracking susceptibility of the reheated Invar weld metal in the double-bead Varestraint test. The detrimental effect of sulfur is so great that its content should be reduced to less than 0.002% in order to attain the high cracking resis­tance as compared to Type 300 series stainless steels. The morphology of impact fracture surfaces of the reheated weld metal at -196°C ( -321 °F) is closely related to the sulfur content. The materials, the sulfur content of which is decreased to less than 0.002%, result in completely ductile fracture, while the ones bearing more than 0.002% sulfur clearly exhibit a typical habit of columnar grain boundary with brittle-like fracture surfaces.

Cracking in the reheated Invar weld metal occurs in two ways, each working independently or reciprocally. One is the liquation of some low-melting constitu­ents in the grain boundary of the weld metal, caused by the reheating of the weld metal by a subsequent pass. Anoth­er is the grain boundary embrittlement caused by the formation and growth of various compounds on the succeeding weld runs.

Some titanium and boron additions to Invar increased markedly the hot-crack­ing susceptibility of the reheated weld metal by the subsequent pass. Chromi­um, which essentially gives rise to an increase of the intrinsic thermal expan­sion coefficient of Invar, has a beneficial effect for suppressing the crack occur­rence in the reheated weld metals.

Paper presented at the 66th Annual AWS Meeting, held April 29 to May 3, 1985, in Las Vegas, Nev.

T. OCAWA is a Senior Researcher, Welding Technology Lab, R&D Laboratories 11, Nippon Steel Corp., Fuchinobe Sagamihara Kanagawa, lapan.

Based on these results, a 60-ton heat was made in a mill run, and it was wrought and cold-rolled to 0.5-, 0.7- and 1.5-mm (0.02-, 0.03- and 0.06-in.)-gauge sheets. These sheets were welded in butt, fillet and edge joints, all of which resulted in satisfactory weldment proper­ties. From another mill heat, the 11.0-mm (0.4-in.)-thick, 66-cm (26-in.)-diameter and 5-m (16.4-ft)-long welded pipes were made in a production run using the GTAW and plasma arc welding process­es. No cracks or discontinuities could be found in seam and girth welds, and the fracture surfaces of tested material at -196CC were completely ductile with superb toughness.

Introduction

Invar, 36% Ni-Fe steel (Refs. 1-3), has a very small thermal expansion coefficient, good formability, and excellent tough­ness, especially at low temperature. This is due to its austenitic structure. There have long been active moves toward the use of Invar for cryogenic service storage vessels, utilizing its small thermal expan­sion coefficient feature. Examples are seen in its use for LNG carriers in France and Sweden, and for liquid fuel transport pipe by NASA (Refs. 4-6). Technological developments are being pursued in the application of Invar for large under­ground LNG membrane tanks, as demon­strated in the Gas Transport-McDonnell Douglas method (Refs. 7, 8).

Compared with Type 304 austenitic stainless steel, which is widely used for corrosion-resistant storage vessels, Invar is rather expensive and somewhat differ­ent in weldability and corrosion resis­tance. For these reasons, Invar was used mostly in 0.5- to 2.0-mm (0.02- to 0.08-in.)-thick gauge sheet.

Since Invar is a fully austenitic steel, it has been presumed that hot cracking is likely to occur on welding (Refs. 9-14). In

fact, several papers have reported that the weld bead reheated on multi-pass welding or repair welding has a high tendency to crack (Refs. 4, 15, 16). Partic­ularly, the earlier extensive research report by C. E. Witherell (Ref. 13) has been quite suggestive, and pointed out a conspicuously preferential effect of an addition of titanium and manganese to improve the hot-cracking susceptibility of Invar. On the other hand, in the mill process of making high nickel-iron alloys such as Invar, some troublesome prob­lems have been experienced, namely, a greatly accelerated oxidation of the slab ingot surfaces in the heating furnace, leading to heavy cracks in the hot-rolling process, which seem to be involved in the common cracking mechanism on multi-run welds. Therefore, more than 50 experimental heats were made to exten­sively investigate and clarify the effects of the chemical compositions on hot-crack­ing susceptibility on welding in order to expand the range of Invar's application to large welded steel structures. Based on these results, 11.0-mm-thick Invar plates have been developed and, on a produc­tion basis, the 66-cm-diameter, 5-m-long large welded pipe of this alloy was made for LNG applications. Evaluation was made of the various performance prop­erties and the reliability and integrity of welded joints of Invar to determine its useability in large welded steel struc­tures.

Experimental Procedure

The chemical compositions of the experimental Invar heats studied are listed in Tables 1 and 2. Those of the two 60-ton heats made on a mill production basis, commercially available Invar sheet, and austenitic stainless steels commonly used are also listed in Table 2. Experimen­tal 25- and 50-kg (11- and 22-lb) ingots were forged, rolled, and finally solution

WELDING RESEARCH SUPPLEMENT | 213-s

Table 1—Chemical Compositions of Laboratory Heats (25 kg)

Heat Ni Mn O Al Cr Note

D

Q K R O

L M E F G

911 912 913

921 922 923

931 932 933

941 942 943

951 952 953

971 972 973

T U

35.23

35.82 35.62 35.82 35.49

35.5 35.5 35.6 35.6 35.5

34.1 35.2 36.5

35.6 35.5 35.4

35.3 35.4 35.3

35.8 35.6 35.8

35.8 35.9 35.6

35.8 35.5 35.5

35.52 35.52

0.003

0.018 0.016 0.018 0.018

0.020 0.020 0.017 0.017 0.020

0.002 0.001 0.001

0.014 0.027 0.049

0.022 0.020 0.018

0.023 0.019 0.015

0.021 0.022 0.019

0.020 0.020 0.020

0.014 0.014

0.02

0.15 0.14 0.15 0.16

0.15 0.15 0.15 0.15 0.15

0.07 0.01 0. () I

0.15 0.15 0.15

0.09 0.18 0.27

0.10 0.10 0.10

0.15 0.15 0.15

0.16 0.16 0.17

0.15 0.15

<0.01

0.20 0.23 0.20 0.21

0.20 0.20 0.21 0.21 0.20

0.01 0.01 0.01

0.30 0.30 0.30

0.33 0.33 0.33

0.10 0.33 0.44

0.31 0.31 0.31

0.30 0.30 0.30

0.20 0.20

0.0009

0.0007 0.0004 0.0007 0.0002

0.0008 0.0070 0.0120 0.0160 0.0040

0.003 0.003 0.003

0.003 0.003 0.003

0.003 0.003 0.003

0.002 0.002 0.002

0.002 0.002 0.002

0.002 0.002 0.002

0.0002 0.0002

0.0003

0.0003 0.003 0.005 0.014

0.0006 <0.001 0.0003 0.0003 0.005

0.002 0.002 0.003

0.0016 0.0018 0.00 16

0.0016 0.0016 0.00 16

0.0018 0.0019 0.0017

0.0023 0.0019 0.0021

0.0022 0.0021 0.0024

0.0003 0.0003

0.0053

0.0027 0.0030 0.0024 0.0017

--

0.0030 0.0015

-0.044 0.045 0.043

0.0024 0.0025 0.0025

0.0015 0.0015 0.0013

0.0117 0.0087 0.0083

0.0055 0.0067 0.0058

0.0065 0.0056 0.0040

0.0024 0.0025

--— -------

0.002 0.002 0.002

0.007 0.007 0.006

0.016 0.016 0.016

0.003 0.002 0.002

0.002 0.002 0.002

0.001 0.008 0.019

— —

0.17

0.17 0.17 0.17 0.17

0.14 0.14 0.13 0.13 0.14

0.01 0.01 0.01

0.18 0.18 0.18

0.15 0.16 0.17

0.13 0.14 0.16

0.13 0.30 0.51

0.15 0.15 0.14

0.13 0.13

0.0006

0.0005 0.0010 0.0006 0.0005

— --

0.0005 0.0005

0.0015 0.0013 0.0012

0.0017 0.0012 0.0015

0.0014 0.0012 0.0013

0.0018 0.0017 0.0017

0.0013 0.0014 0.0015

0.0021 0.0020 0.0018

0.0014 0.0035

low

CO

CU

z

U

lo

c 5

( )

<

Z

c, Si, Mn

<* od co

(N

DC u_ 0)

% I

,.

,| Ol

u

1

J

c, u 0)

3 *Z J O 3 c 1 .Q

DO u a

03 J >

co

<r 1 r

Table 2—Chemical Compositions of Laboratory (25- and 50-kg) and Mill (60-ton) Heats, Commercial Invar, and Austenitic Stainless Steels

Ni C Si Mn P S O Cr Ti N B Note

Ti B TiB

CR3 CR2 CR1 M MT1 MT2 MT3

Heat I Heat II

Commercial Invar

Type 304 Type 31OS

35.7 35.7 35.7

36.15 36.56 36.30 35.91 36.22 35.85 36.23

35.5 35.7

36.1

9.4 25.1

0.028 0.021 0.030

0.018 0.016 0.016 0.015 0.021 0.024 0.018

0.025 0.028

0.036

0.05 0.05

0.16 0.17 0.16

0.15 0.15 0.15 0.15 0.15 0.15 0.15

0.17 0.15

0.21

0.75 0.66

0.28 0.20 0.29

0.29 0.30 0.29 3.09 3.12 2.12 4.21

0.33 0.27

0.35

0.94 1.01

0.003 0.003 0.003

0.004 0.004 0.005 0.003 0.003 0.004 0.003

0.003 0.003

0.005

0.026 0.026

0.0010 0.0004 0.0009

0.005 0.004 0.004

<0.001 <0.001 <0.001 <0.001

0.0011 0.0010

0.002

0.007 0.004

-

0.0051 0.0075 0.0033 0.0021 0.0018 0.0019 0.0018

0.0033

0.0066

-

0.16

0.16

7.97 4.15 2.14 0.06 0.06 0.06 0.06

0.17 0.23

-

18.3 19.9

0.028

0.011

1.06 1.07 1.06

0.0015 0.0035 0.0013

0.0030 0.0014 0.0021 0.0010 0.0010 0.0009 0.0008

0.0009 0.0017

0.0022

0.016 0.020

0.0020 0.0021

i i

C CQ OD O a LT

Add

it Ti

See

O C JN

J ? " o ._- .SP •z= \- u-"O . - 0) < ^ l/l

Taj "*Tt

Fig.

14

fille

r m

t 1.

6- &

5

Commercial 0.5-, 0.7-mm g au

l\

ro QJ

JZ

.6 SS

c

E

Q. L U

' I

t OJ

Si

$e

Commercial 5.0-mm gauge

Note: 1.6- and 2.4-mm-diameter matching filler metal was made f rom Heat I.

214-s I AUGUST 1986

impact-fractured surface observed

sl i t -cut by saw

Fig. 7 - Double-bead Varestraint test. A-Schematic sketch of crack occurrence; B-Cross-section of double beads

8

7

E

. 6 Ul (0 "* r E 5 -o 3)

c 4 -c ao

4) 0

U

o 2

1

0 X s p Si

-

SINGLE-BEAD GTAW, Ar shielding | 70A —7.5cm/min / STRAIN2.4% /

rS I I D

0.0003 0.0009 0.02

Q | R | 0 0.0003 0.0050 0.0140 0.0007 0.15

0.0007 0.15

0.0002 0.16

Fig. 2 test

2 w 3 4

Augmented s t ra in (%)

( A ) {%) (B)

-Effect of sulfur and phosphorus on hot-cracking susceptibility in single-bead Varestraint

heat-treated at 1050°C (1922°F) in the laboratory.

Evaluation of the hot-cracking suscepti­bility was performed mainly with two variations of the Varestraint test (Ref. 17): 1) the conventional longitudinal single-bead method for investigating weld cracking in the heat-affected zone of the base metal and the solidification cracking of the weld metal, and 2) the double-bead method, which simulates the reheated weld beads in repair welding or multi-layer welding, as illustrated sche­matically in Fig. 1A. In the double-bead Varestraint test, cracks actually occur on the first bead (Bead I) of the reheated weld metal by the time of the next weld run, on the second bead (Bead II) of the next weld run itself, and in the heat-affected zone of the base metal along Bead II. The welding variables employed for GTAW in the Varestraint test were 70 A and 16 V, with a travel speed of 75 mm/min (3 ipm) on 5.0-mm (0.2-in.)-thick plates. Cracks of the Varestraint speci­mens were examined on the as-tested surfaces with a 30X binocular micro­scope. They were measured, using this information.

The crack and fracture surfaces in the welds were examined by optical micros­copy, electron probe microanalysis (EPMA), scanning electron microscopy (SEM) with accompanying energy disper­sive spectroscopy (EDS), and Auger elec­tron spectroscopy (AES).

Since cracking is associated with the extent and behavior of grain boundary embrittlement, particularly careful exami­nations were conducted of the fracture

surfaces in Bead I of the reheated weld metal. The specimens were impact frac­tured at the liquid nitrogen temperature of -196°C, as depicted in Fig. 1B, to ensure good fracture surface condition.

The large 60-ton heats were wrought and rolled to sheets and plates and solu­tion heat-treated in the mill. These sheets and plates were welded by the GTAW, resistance seam welding and plasma arc welding processes, using 1.6- and 2,4-mm (>i6- and ;&Hn.)-diameter matching filler metal. The filler metal was specially made from a part of the same heat of ingots that were used for the large welded pipes. Dye penetrant inspection, bending, Charpy impact, tensile tests, and extensive examinations of bead cross-sections were conducted on each weld­ment. For the welds of the 11.0-mm-thick, 66-cm-diameter pipes, the three-point bending crack tip opening displace­ment (CTOD) test was performed in accordance with BSI • DD-5762. The specimens were machined perpendicular to the welding direction, with the length of the notch being parallel to the surface. In addition, the fissure bend test was employed for multi-layered weld metals made by GTAW with the matching filler metal.

Test Results

Effects of S, P, and Other Chemical Elements on Hot-Cracking Susceptibility

Figure 2A depicts the effect of sulfur content on the total crack length in the single-bead test. Note that the crack length increases rapidly with sulfur con­

tent above 0.005%. Figure 2B shows the effects of sulfur and phosphorus when the augumented strain was varied in the test. Like sulfur, the effect of phosphorus is very great. The test results on commer­cial austenitic stainless steels are also shown for comparison (Ref. 30). Obvi­ously, it is extremely important to reduce the contents of sulfur and phosphorus to an extra-low level in order to attain high hot-cracking resistance compared to a commercial Type 304 stainless steel.

Shown in Fig. 3 are the effects of sulfur and phosphorus on the total crack length of Beads I and II at the augmented strain of 5.0% in the double-bead test. No difference could be found in the hot-cracking tendency on Bead ll in the dou­ble-bead test, compared with that in the single-bead test in Fig. 2A. However, there was quite a difference between the effects of sulfur and phosphorus on the cracking tendency in Bead I of the dou­ble-bead test. With regard to sulfur, when its content exceeds 0.0030%, the increase in the crack length becomes quite remarkable. On the other hand, note that the crack length increases con­spicuously only for phosphorus content above 0.0160%. Figure 4 shows the effects of sulfur and phosphorus on the hot-cracking susceptibility of Bead I in the double-bead test when the augmented strain was varied. From Figs. 3 and 4, it should be noted that extremely high hot-cracking resistance in Bead I can be obtained by restricting the sulfur content to the very low level of less than 0.002%. Note that only a small crack could occur on Bead I of Type 304 and Type 310S in

WELDING RESEARCH SUPPLEMENT 1215-s

(ppm) P S

To ta l c r a c k l e n g t h ( m m ) 3 4 5 6 7 8 9

10

70

120

160

140

type310S 2801 50 type304 300| 50

23.6 34.3

V \ \ V \ V V \ V \ \ \ \ \ \ \ \ V \ \ V \ \ \ \ \ \ \ \ \ \ \ \ V \ \ \ V ^ T ^ V ^ - H _ m -26.0- 59.7

;w=r Effect of P

W F T

YW=

v W W W W W V v V M -

w <¥~

3 - austenitic stainless steel

Fig. 3 — Effect of sulfur and phosphorus on hot-cracking susceptibility in double-bead Varestraint test

the doub le -bead test. Shown in Fig. 5 is the effect of n i t rogen on the cracking tendency of Bead I. If the n i t rogen c o n ­tent is increased f r o m 0.0014% to 0.0035% in the material, the total crack length o f Bead I increases slightly at the higher augmented strain.

Figure 6 indicates the effects o f nickel, ca rbon , silicon, manganese and chromi ­u m in the ext ra- low P and S materials o n the cracking tendency of Bead I in the doub le -bead test. Here, the variat ion of the contents o f each e lement was main­ta ined w i th in the range in wh i ch the increase of the thermal expansion coef f i ­cient cou ld be ignored. N o significant effects cou ld be f o u n d on the w h o l e , though highly pur i f ied materials o f l o w C, l o w Si, and l o w M n exhib i ted excellent resistance to the cracking propensi ty of Bead I, even at the high oxygen level wh ich resulted f rom a relatively large number o f inclusions in the material. Inci­dentally, some a luminum, wh i ch is added as a deoxidizer on the melt ing process, occasionally impairs the resistance to the cracking tendency of Bead I.

Figure 7 depicts the effects of t i tanium and b o r o n on hot-cracking susceptibil ity of Bead I in the test. Both elements, a l though able t o stabilize carbon , n i t ro­gen and oxygen to some extent , have a significantly detr imental ef fect on hot -cracking. Howeve r , it is con f i rmed simul­taneously that the total crack length o f Bead II is hardly a f fected by the a b o v e -ment ioned t i tanium addi t ion, though a

small increase is exhib i ted w i t h the b o r o n addi t ion.

Microscopic Observation of Cracks and Fracture Surfaces

Figures 8 and 9 present examples o f cracks in the surfaces and cross-sections o f the specimens subjected to the aug­men ted strain o f 5.0% in the doub le -bead Varestraint test, respectively. As illustrat­ed schematically in Fig. 8, cracks occur red markedly, even on Bead I of the speci­m e n having a sulfur con ten t o f 0.005%. Moreove r , these cracks w e r e widely dis­t r ibuted f r o m the apex of w h e r e the bend was appl ied, up to 30 m m (1.2 in.) away and in the lower tempera ture regions. Accord ing to we ld ing thermal cycle measurement, the lowest tempera­ture at wh i ch cracking can occur is approximate ly 6 5 0 ° C (1202°F). W i t h specimens having the suff iciently l o w sul­fur and phosphorus conten t of 3 to 7 p p m , such cracks w e r e rarely observed . Figure 9 shows the internal cracks of the Varestraint test specimens wh ich w e r e sect ioned in sequence starting at 4 m m (0.15 in.) f r o m w h e r e the bending was appl ied and wo rk ing towards the point w h e r e the w e l d was started. As seen o n the surfaces o f the specimens, many cracks w e r e observed in Bead I. It is ver i f ied f r o m the cracks in specimen No . 8 that cracks can occur even at the por t i on approx imate ly 28 m m (1.1 in.) away f r o m the apex of the bend in the

3 ̂ r D

Augmented strain {%) Eig. hot

4 —Effect of sulfur and phosphorus on cracking propensity in repair welding

double-bead method GTAW. Ar shielding

: 70A-7.5cm/min 5.0 x 40 x 320mm

2 -»~ 3

Augmented strain [%)

Fig. 5 — Effect of nitrogen on hot-cracking pro­pensity in repair welding

Varestraint test. A lmost all cracks w e r e f ound along the grain boundar ies o f the austenites in the reheated w e l d metals.

Figure 10 illustrates the f racture sur­faces of Bead I and Bead II o f the speci­men w i th 0.014% sulfur content , wh ich w e r e impact- f ractured at the tempera­ture of —196°C. The fracture surfaces primarily exhibit typical co lumnar grain boundaries. It should be no ted that the grain boundar ies in the w e l d metal have embr i t t led greatly, even w h e r e cracks d o not appear on the surface o f the speci­men. Fur thermore, it is significant that the morpho logy o f such f racture surfaces at a very l o w tempera tu re is closely related t o the sulfur content . Specifically, Fig. 11 clearly shows the effect o f sulfur on grain boundary embr i t t lement . W h e n the sul­fur content exceeds 0.002%, the f racture morpho logy definitely changes f r o m a complete ly ducti le f racture t o an embr i t t ­led columnar grain boundary one , being in g o o d agreement w i t h the surface crackings o n the reheated w e l d metal in the Varestraint test. O n the other hand, it was not until the phosphorus content exceeded 0.016% that the transit ion of the m o r p h o l o g y t o an embr i t t led f racture behavior appeared , p rov ing it less o f a

216-s I A U G U S T 1 9 8 6

eleme' nts

Ni

34.1

35.2

36.5

C 0.014

0.027

0.049

Si 0.09

0.18

0.27

Mn 0.10

0.33

0.44

Cr 0.13

0.30

0.51

30

30

30

30

30

30

20

20

20

20

20

20

Total crack length on weld metal (mm)

0 1 2 3 4 5 6

3^

16

18

16

5 3 1

.gr

16

16

16

W N N W Y T W . V . M

18

19

17

23

19

21

Bead I I . 0 3 — I Bead I ^7F=I—l

variation of Ni 'C, 0.001 ,Mn,0.01 vSi,0.01 N, 0.001

^

->=*

variation of C

variation of Si (Al added, 0.16)

v, s S v ' s~T^

\ N \ \ i^J> i _ ^ . a — < variation of Mn

>g=r

S F

SX3?"

W

variation of Cr

fdouble-bead method GTAW, 70A-7.5cm/min 5.0X40X320^ augmented strain;5.0%

Fig. 6 —Effect of each element on hot-cracking susceptibility in double-bead Varestraint test

contributor to cracking than sulfur con­tent.

Figure 12 depicts the optical micro-structure of heavily precipitated and oxi­dized grain boundaries in the reheated weld metal of Bead I and the typical surface features of impact-fractured grain boundaries. These structures are covered with intricate scales and residues bearing either a network of crevices on the surface or many small speckled dimples, presumably resulting from a large num­

ber of fine precipitates, in addition to these, there are some occasional smooth­er, wavy-appearing surface areas repre­senting the last vestiges of liquid on the fracture surfaces, as later partially seen in Fig. 24.

Figure 13A depicts the results of EPMA analysis at a crack in Bead I of 0.014% sulfur material, and indicates the signifi­cant segregation of sulfur, phosphorus and oxygen at the crack tip along the grain boundary. Some enrichment of sili-

s

- 0 0.0011

a 0.0010

n 0.0004

• 0.0009

-

-

T i

-0.028

" 0.011

5.0 >• 4 0 » 3 2 0

i 1 8 Note .

m i l l heat ' ( 60 t o n ) /

~ 0,0020

0.0021

e x p e r i m e n t a l / i

h e a t s ' ( 2 5 k g ) /

/ A mm / /

70A-7.5cm/min / /

G T A W

t

•

Arsh

-"-"

elding / /

r̂" $ Fig. 7—Effects of titanium and boron on hot-cracking propensity

con was also confirmed according to other analyses. Figure 13B shows the result of an AES analysis made of the grain boundary fracture surface of the same 0.014% S material. A strong concentra­tion of oxygen, sulfur and carbon can be found, even after a 240-s cleaning by argon gas sputtering. For comparison, the AES analysis of the same saw-cut material surface is presented in Fig. 13B.

60-Ton Heats on a Mill Production Basis

Based on the above results, a 60-ton heat of material, the chemical composi­tions of which are listed in Table 2, was made in a production run. As shown in Fig. 14, the results of the double-bead Varestraint test on this material exhibits excellent hot-cracking resistance. The fracture surface of Bead I of this material was, at —196°C, completely ductile, as presented in Fig. 15, where one commer­cial Type 304 stainless steel is also given. The latter has, in general, been consid­ered to have excellent toughness in cryo­genic applications. According to the AES analysis of the impact-fractured surface of this material, only slight segregations of sulfur and oxygen could be found, and they were detected through an approxi­mately 120-s argon sputtering, as dis­played in Fig. 16.

The heat material was hot-rolled and finally cold-rolled to 1.5- and 0.5-mm (0.06- and 0.02-in.)-gauge sheet in a nor­mal production process, and then solu­tion heat-treated in the mill furnace. The butt, lap and edge joints were welded by automatic and manual GTAW processes without filler metal, using several heat inputs. All of these weld beads were examined with a dye-penetrant test. The cross-sections of the beads were exam­ined optically at 200X magnification with quite satisfactory results. There were no cracks or microfissures.

In addition to GTAW, double-folded and triple-folded joints of the sheets

WELDING RESEARCH SUPPLEMENT I 217-s

310S(25Cr-20Ni)

(mm) 10

| bending

OJ^'i fal l^H.

W . (Bead II) 0.4 V 0.5 0.2 „ -^ M t O a S L ? ' 0 . 2 ' n ^

welding direction

• = 0

0.2 0.10.1 0.1

(Bead I)

304(18Cr-8Ni)

[ GTAW I 70A—75mm/min 1 augmented strain 5.0%

3

- 2

(Bead I)

(mm) 5

4

3

2

1 fusion a line

2

3

Invar { S 0.0050* R I P 0.0007%

(mm) 10 5 0.4 0.2

0-2 (-0.6 0.3 '0-5 0 3 welding direction

°-60?0.8o.22o!20'!°-3 o",

(Bead I)

0.2 0.10-1

fusion a line

2

3

F/g. 8-Schematic cracking appearance on the weld beads after double-bead Varestraint test

heat O Bead II Bead I (reheated)

No. 1 X~=

: the weld puddle on bending

•

No. 2

No. 8 28mm from the puddle

Fig. 9 —Hot-cracks in the weld bead of double-bead Varestraint test

w e r e resistance seam w e l d e d . This p r o ­cess has generally been used in fabricat­ing cryogenic storage vessels, such as LNG carriers. The we ld ing variables and some propert ies of the we ldments made of these sheet materials are presented in Table 3 (Ref. 31). N o defects or disconti­nuities in the seam we ld nuggets could be found , as indicated in Fig. 17, and the impact- f ractured surfaces of the w e l d nuggets w e r e complete ly ducti le.

The 11.0-mm-thick, 66-cm-diameter , 5-m-long w e l d e d pipes w e r e made f r o m another mill heat. Fol lowing the press-b e n d fo rm ing of 11.0-mm-thick plates, the seam was jo ined using the plasma arc we ld ing process fo r the initial roo t run and the G T A W process for the subse­quent pass, as seen in Fig. 18. The girth we ld ing of these pipes was by multi-pass manual G T A W , and the finished w e l d ­ments w e r e examined by x-ray and dye penetrant inspect ion tests, w i t h satisfac­tory results. Figure 19 shows the macro-structures o f the seam and girth welds, in wh ich no defects or discontinuities cou ld be f o u n d . The results of Charpy impact, side bend , C T O D and tensile tests are g iven in Table 4, and the impact-f ractured surfaces of the Charpy test specimens are indicated in Fig. 20. The absorbed energy value o f 90 ~ 150 ), at - 1 9 6 ° C (122 ~ 203 ft- lb at - 3 2 1 ° F ) , should be considered sufficient fo r w e l d e d structural material. The f ractured surfaces are substantially ducti le, though some port ions exhibit a slight habitual facet in the austenite grain boundary o f the we ld metal . N o microfissures cou ld be de tec ted in the side bend test, and the C T O D value o f 0.25 ~ 0.61 m m (0.01 ~ 0.024 in.) at - 1 9 6 ° C is comparab le t o

the Type 304 austenitic stainless steel welds.

218-s l AUGUST 1986

'I V (f. heat O;0.014S,0.0002P

r/5W¥-*,

Fig. 10 — Fracture morphology at —196°C on the parts of the weld beads where it is presumed to have been approximately 650''C at the point of bending in double-bead Varestraint test. A — Typical habits of columnar grain boundaries of weld Bead I (reheated) fractured at — 796 °C; B — Fracture surfaces of weld Bead II (not reheated)

Discussion

In the double-bead Varestraint test, cracks occur on the first bead (Bead I of the reheated weld metal), on the second bead (Bead II of the subsequent welding), and in the heat-affected zone of the base metal along Bead II. In the case of Invar,

numerous significantly large cracks were likely to occur on Bead I of the reheated weld metal, whereas with the austenitic stainless steel, cracks were observed mostly on Bead II.

Sulfur increases remarkably the hot-cracking susceptibility of the reheated weld metal —Bead I on the double-bead

Varestraint test. From the test results of heats made for the laboratory and from the mill production run, it is evident that superb resistance to hot-cracking in the reheated weld metal of Invar can be attained by lowering the sulfur content to. less than 0.002%. As shown in Fig. 3,'a fairly large crack length was still recog­nized on the reheated weld metal of the specimens which had their sulfur content reduced to 0.003%. Moreover, these cracks on the reheated weld metal can occur at a somewhat higher temperature, even at approximately 650°C. Conse­quently, they are widely distributed to the region about 30 mm (1.2 in.) from where the bending was applied on Bead I in the double-bead test.

The morphology of impact-fractured surfaces of the reheated weld metal

fusion line

Bead II - - —Bead I (reheated)

Fig. 11 - Effects of sulfur on grain boundary embrittlement in weld Bead I (reheated) in an area 20 mm away from the apex of the bend in the Varestraint test

Fig. 12 —Optical microstructures and SEM images of grain boundaries. A — Precipitation and oxidation at grain boundaries in the part of Bead I where it is 30 mm away from the apex of the bend in the double-bead Varestraint test; B —Scales and residues in grain bound­aries in Bead I

WELDING RESEARCH SUPPLEMENT | 219-s

Heat O. EPMA on a crack around the weld puddle on bending.

1, 2 0 -

10,. , A J

—Hi;

•.'•''-• v , . * W7> " •-.•'..v-''-

--'^V""»V'-VaAvl^»Y V*V*V<SV*'W\/M*a>-'WV'

After 240-sec cleaning through argon sputtering ® Fractured columnar

grain boundary surface heat O: 0.0014S.0.00Q2P

D - Electron volt (eV)

100 200 300 400 500 600 700 800

Fig. 13 - EPMA and AES results in Bead I after double-bead Varestraint test. A - Segregations of S, O and P at the tip of a crack in Bead I; B - Strong Auger electron intensity of O, S and C on the fractured columnar grain boundary of Bead I

(Bead I) at -196°C is closely related to the sulfur content. The Invar materials with their sulfur content reduced to less than 0.002% have completely ductile fractures with excellent toughness, while

2 3 4 Augmented strain {%)

5 2 3 4 Augmented strain {%)

Fig. 14 — Hot-cracking susceptibility of extra-low sulfur Invar (Heats I and II made in a production run)

the ones bearing more than 0.002% sul­fur content clearly exhibit a typical habit of columnar grain boundary with brittle-like fracture features.

The change of the fracture surface

• " ; > ' . a

•r*1;.'* v-'V-t*

>~*?.--.i\;-TC?

0.4 mm

Fig. 15-Fracture surfaces of Bead I of the Invar mill heat and Type 304 stainless steel used in the double-bead Varestraint test. A — Mill Invar heat at -196°C; B- Type 304 at -196CC

behavior of Bead I, depending on the sulfur content, is quite consistent with the results of surface cracking of Bead I in the double-bead Varestraint test. Note that there is a significant difference between the detrimental effect of sulfur on crack­ing in the reheated weld metal of Invar and that of stainless steel.

Close examination of the surface crackings and the low-temperature, impact-fracture surfaces of the welds reveals that the columnar grain bound­aries are studded or covered with brittle, thick scales and residues, such as sulfides and oxides, and are also, in most cases, accompanied with some predominantly smooth fracture surfaces. This condition occurs, in part, where a liquid film is suggested to play a role in the cracking mechanism. Therefore, cracking in the reheated weld metal of Invar is consid­ered to occur in two ways, each working independently or reciprocally. One is the liquation on the grain boundaries of the weld metal, caused by the subsequent weld cycle activating such low-melting constituents as nickel-sulfate (N3S2, 635°C71175°F) and nickel-phosphate (Ni3P, 875°C/1607°F) (Ref. 18). Another is the grain boundary embrittlement caused by the formation and growth of various compounds of oxides, sulfides, phosphides and carbonitride in grain boundaries during the succeeding weld runs.

In corrosion and oxidation of metals in high-temperature environments, it has been reported that transportation and chemical reaction of oxygen are acceler­ated and activated by sulfur, which is one of the most crucial elements affecting weld cracking in Fe-Ni alloys (Refs. 19-22). In the grain boundaries of the weld metal which has been rapidly reheated to

220-s I AUGUST 1986

Ar sputtering

0.025C-0.17Si-0.33Mn-0.003P-0.001S-35.5Ni 120 sec

Ar sputtering

30 sec

Ar sputtering

10 sec

Ar sputtering

•^f^^P^^msm"

0 100 200 300 400 500 600 700 800 900 Electron Energy eV— •

Fig. 16 — Slightly segregated sulfur and easily removed oxide films on the grain boundary surfaces in the reheated weld metal of Invar

Fig. 17 — Resistance seam welded joint of the sheets made from mill heats. A — Microstruc­ture; B — Fracture surface at —196 °C

near-solidus temperature by the subse­quent welding runs, the following chemi­cal reactions could be conceptually described, according to the hot corrosion theory associated with accelerated oxida­tion:

4Ni-F2S-F 1 / 2 0 2 ^ Ni3S2 + NiO (1)

Ni3S2 + 3 /20 2 -* 3NiO -F 2S (2)

Here, sulfur stems from the segregation on solidification (Refs. 32, 33) and oxygen comes mainly from the air in the outer

atmosphere, presumably through diffu­sion along the grain boundaries (Refs. 34, 35). If Reaction 1 is triggered, Reaction 2 continues to occur in a self-sustaining fashion. Both reactions leave a porous, nonprotective and fragile nickel oxide (NiO) layer on the grain boundaries. Fig­ure 21 schematically illustrates these oxi­dation reactions. This nonprotective layer of nickel oxide is presumably a major cause of the grain boundary embrittle­ment, which leads to the propagation of the cracks. In the case of nickel sulfate (Ni3S2), which can become a liquid at 635°C (1175°F), as depicted in a Ni-S

phase diagram, hot-cracks are likely to initiate in the presence of the thermal stress of the weld metal as it cools.

When welding stainless steel, which usually contains more than 12% chromi­um, a stable and protective chromium oxide film may form on the weld bead surfaces and at the grain boundary, negating the above-mentioned strong activity of oxygen in the weld metal. This would explain the higher resistance to cracking in the reheated weld metal of Type 300 series stainless steel than in Invar. This mechanism is partly ascer­tained by the results of the double-bead

Fig. 18—Appearance welded Invar pipe

of large-diameter

Table 3—Weldment Properties of 11-mm-Thick, 66-cm-Diameter Large Pipe

Seam weld

Girth weld

Charpy Impact Value ()) Weld S

- 1 9 6 ° C

150 143 135 150

129 122 90

105

tetal 20°C

171 160 142 150

145 136 140 143

HAZ - 1 9 6 ° C 20°C

118 180 118 214 152 285 125 210

Tensile Strength (MPa)

(6-mm Diameter) - 196°C 20°C

777 432 773 452 780 448 775 450

742 733 738

482 476 480

Side Bend Test R:2xt

t = 9.5 mm

good good good good

good good good good

CTOD Value (mm) at - 196°C

0.61 0.49 0.53 0.45

0.251 0.474 0.254 0.252

Note: Base metal Charpy impact value (I)-- 1 % ° C .

tong. 224, Trans. 174 at - 1 9 6 ° C ; Tensile strength ( M P a ) - t o n g . 890. Trans. 893 at

WELDING RESEARCH SUPPLEMENT 1221-s

- —

B

• K f - f i

"%

ii«

•* * '%

-%.y

,»'

1mm •

fig. 19—Macrostructures of large-diameter Invar pipe. A —Seam weld made by plasma arc and CTA welding; B— Girth weld made by CTAW

Fig. 20— Weldment impact fracture surfaces of large-diameter Invar pipe at —196 "C. A -Seam weld (2v E-196°C= 135 I); B - Girth weld (2vE-196°C= 122 I)

10

9

8

^ • s

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S~' JZ

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ra > 1

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-

•

/ / /

/ /

-

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- r"

r - /•

/

/

/ / /

/ •

double-bead method GTAW; 70A-7 .5cm/min augmented s t ra i n 5.0%

/

/ ' r / y

Y

/

/ / / / / / / / r / / / /

M T 3 0.001

— 1.06 4.21

i— /

/ /

/

/ / /

/ / .—, / / / / /

/ / / / / / / / /

M T 2 0.001

— 1.07 2.12

rt

/ / / /

/ / /

/

7

' t / /

/ / / / / / ' T 1

M T 1 0.001

— 1.06 3.12

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0.001 — -

3.09

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— 0.30

~

T

i L±J

r1

/ /

(aH

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-0.29

_

£

CR3 0.005 7.97 —

0.29

Porous A/. Si N i 0 _ oxides .

j r sulfides s

phosphide ,

reheated weld -' metal surface

•Carbo-nitride

36%-Fe alloy

Nickel sulfide Ni3Sa NiS

sol idif ication grain boundary

Fig. 22 — Effect of chromium on hot-cracking susceptibility in double-bead Varestraint test

Fig. 21 -Schematic presentation of the assumed oxidation reactions at the grain boundary of the reheated Invar weld metal

Varestraint test on some experimental heats, as shown in Fig. 22. It should be noted in Fig. 22 that the addition of some chromium, which essentially gives rise to an increase of the intrinsic thermal expan­sion coefficient of Invar, very effectively suppresses cracks occurring in the reheated weld metal of Bead I.

Accordingly, there would be almost no susceptibility to hot-cracking in the reheated dissimilar weld metals made with austenitic stainless steel, although the concern for solidification-cracking of the weldments still remains (Refs. 23, 24, 29).

Another example of cracking, due to the embrittlement of the grain bound­aries, which is partly correlated with some liquation-cracking, can be ob­served in the resistance seam welded joint of triple-folded sheets of commer­cial Invar, as shown in Fig. 23. The weld metal of the resistance seam weld epitac-tially solidifies and develops from the base metal toward the center of the weld nugget in an elliptical, step-like way, as schematically depicted in Fig. 23A. Note that cracks do not occur in the center part of the weld nugget where final solidification takes place, but do occur heavily in both convex parts of the ellip­tical region adjacent to the fusion bound­ary, which is considered to have finished solidifying prior to the center part. In the case of fully austenitic stainless steel sheets, such as Types 310S and 316, hot-cracking on resistance seam welding generally occurs at the center part of the weld nuggets (Ref. 25). As presented in Fig. 23B, the impact-fracture surface of the seam weld nugget mainly consists of embrittled grain boundaries, including some areas exhibiting the vestiges of a liquid-film-like feature.

Figure 22 certainly indicates again a detrimental effect of titanium on the susceptibility of the reheated weld met­als, which is quite different from the results reported earlier by C. E. Witherel. However, there are actually some dis­crepancies between Fig. 7 and Fig. 22 on the total crack length of Bead I.

There is no clear explanation for the

222-s I AUGUST 1986

80/

A 100 kg 2900 A 1.7m/min 390V

andard current

0.7 • t

0.5 [ I

0.7

2100A

A a

'f $ V

•u f #,

A

^

I 0.2mm

v. : A B

Fig. 23 — Weld joints of commercial Invar sheet. A — Resistance seam joints welded with a relatively high heat input; B —Crack fracture surfaces in the seam joints mentioned above

profoundly deleterious effect of some titanium or boron additions (Figs. 7 and 22) to Invar. Actually, strongly adhering, thick oxide films were detected by AES analysis of the embrittled grain bound­aries of Ti-bearing Invar (Figs. 24 and 25), and some titanium compounds were detected by EDS, as shown in Fig. 26. Although AES analysis failed to reveal the enrichment of titanium in the grain boundaries, the localized titanium enrich­ment, through the grain boundary sweeping mechanism presented by ). C. Lippold (Refs. 26, 27), might facilitate the embrittlement of grain boundaries by means of accelerated oxidation in the boundary. This can occur in a similar manner as nickel sulfate and through considerable precipitation of carbonitride in it (Ref. 28). Moreover, it is certain, from

Fig. 22, that the addition of some amount of titanium slightly increases the hot-cracking propensity of the heat-affected zone of the base metal along Bead 11.

From the extensive investigation made by the double-bead Varestraint test, the effects of each chemical composition mentioned above, including S, P and Al, are summarized in Table 5. Here, the crack on Bead I is regarded as reheated weld metal cracking, and that of Bead II, just around the weld puddle, as solidifica­tion cracking. In fact, some cracks, appar­ently due to grain boundary embrittling, were occasionally found on the lower temperature region of Bead II.

The detrimental effects of sulfur and phosphorus on the cracking sensitivity of Bead 11 in the double-bead test of Invar (which is thought to be solidification

I'.'jS t ,

W< :, iK> • ' , * ' l v 4I,

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Fig. 24 — Fracture morphology at —196"C on the part of the weld Bead I where it is presumed to have been approximately 700°C at the point of bending in the double-bead Varestraint test

cracking) are nearly equivalent to the ones in the single-bead test of fully aus­tenitic stainless steel (Ref. 12). Hence, both contents should be reduced to an extra-low level of less than 0.002%. As shown in Fig. 22, the addition of 2 to 4% manganese to Invar somewhat increases

WELDING RESEARCH SUPPLEMENT 1223-s

s o. o > U J

o X u oc < Ul cn UJ oc

E OL O - j U J

> Ui Q X U IE < UJ c/> Ui QT

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0 0 2 b C 0 1 6 S i - 0 2 8 M n - 0 . 0 0 3 P - 0.001 S - 3 5 . 7 Ni

; 0 . 0?8T i

"rnVtiin^Tl

vH|l|fi

A r s p u t t e r i n g

900 sec

!i"l I

r s p u t t e r i n g

150 sec

fif&

A r s p u t t e r i n g

300 ^"<~

A r s p u t t e r i n g

0 100 200 300 400 500 600 700 800 900 ELECTRON ENERGY eV

. r rrfr ^ S '<?/

A r s p u t t e r i n g

_ i i j • ' • • • • • • • • -

1 0 0 . "

Fig. 25 —Strong adhering, heavy oxide films on the grain boundary surfaces in the reheated weld metals of titanium-bearing invar

.1.028C-0.16Si - 0 . 2 8 M n - 0 . 0 0 3 P ^ 0 . 0 0 1 S - 3 5 . 7 N i - 0 . 0 2 8 T i

J t - N o v - * ] i < « i « r »••"» • * • « » • » • 4f«CfI TIM *n • •-2MIV mtv/cM r i t i Hi M l * LOMftnr •< rt' I M MIN- * ' * •

l i t **

S I I C MF

Fig. 26 — Titanium oxides on the grain boundary surfaces in the reheated weld metal of titanium-bearing Invar

the total crack length on Bead II of the double-bead test. The cause of this is not clear. It is quite evident from the far fewer cracks in the heat-affected zone of the base metal compared to the weld metal that reducing the segregation or enrichment of chemical elements such as S, P, Ti, B, Al and Si on the grain boundary is indispensable for suppressing crack occurrence. As learned from the cracking

mechanism in the reheated weld metal, another important factor in the welding process is to prevent the molten weld metal from picking up oxygen from the air. According to the fissure bend test, which employed three-layered, multi-run GTA welded specimens on 11-mm Invar plates and a matching filler metal of extra-low-sulfur content, some microfis­sures could be produced in the weld

Table 4—Effect of Each Element on Hot-Cracking

Range (%)

0.0003

0.0140

Crack on Bead l(a»

Crack on Bead ll<b»

(c) (c)

(S0.002) (SQ.003)

0.0002 (e) f (S0.010)

0.0160 l ' (<0.008)

Ni 34.1

s 36.5

no effect no effect

Al

N

C

Si

no addition

J 0.020

£0.0035

0.014

S 0.049

0.09

/ 0.27

(d)

(e» (<0.0035)

somewhat beneficial at higher content

no effect

no effect

(e) (<0.0035)

no effect

somewhat beneficial at lower

content

Mn

O

Cr

0.10

/ 0.44

S 0.045

0.13

/ 0.51

somewhat beneficial at higher content

no effect

no effect

no effect

no effect

no effect

[a)Grain boundary embrittlement. So l id i f i ca t ion cracking. [ c )Maximum detriment. [d)Considerable detriment : e )Minimum detriment.

metal when an insufficient argon shielding gas or a mixture of argon and CO2 gas were used. These facts suggest that inert gas welding processes such as GTAW, plasma arc welding, and laser welding are preferable to SMAW, SAW and GMAW.

A 60-ton heat of material was made in a production run, and wrought and cold-rolled to 0.5- and 1.5-mm-gauge sheets. These sheets were welded in butt, fillet and edge joints using automatic or manu­al GTAW, each resulting in satisfactory weldment properties. Resistance seam welding was also conducted on triple-folded joints of the sheets, presenting defect-free nuggets, as shown in Fig. 17.

From another production heat, large Invar 11-mm-thick, 66-cm-diameter pipes were welded, using the GTAW and plas­ma arc welding processes with a match­ing filler metal. The seam and girth weld­ments of these pipes proved to have excellent weldability, with no cracks or microfissures, and superb toughness at -196°C.

224-s I AUGUST 1986

Conclusion

1. In double-bead Varestraint tests, cracks occur on the first bead (Bead I —the weld metal reheated by the next pass), the second bead (Bead II —the subsequent pass), and the heat-affected zone of base metal along Bead II. In the case of Invar, numerous significantly large cracks were likely to occur on Bead I. This was quite different in the case of the stainless steel. Here the cracks were mostly observed on Bead II.

2. Sulfur and phosphorus remarkably increase the hot-cracking susceptibility of Bead I in the double-bead Varestraint test. In particular, the detrimental effect of sulfur is so great that its content should be reduced to less than 0.002% in order to attain the similar high hot-cracking resistance of a commercial Type 304 stainless steel.

3. The morphology of the impact-fracture surfaces of the reheated weld metal (Bead I) at -196°C (-321°F) is closely related to the sulfur content. The materials, the sulfur content of which is decreased to less than 0.002%, result in completely ductile fracture with excellent toughness, while the ones bearing more than 0.002% sulfur content clearly exhibit a typical habit of columnar grain bound­ary with brittle fracture surfaces.

The change of the fracture surface behavior of Bead I as a result of the sulfur content is quite consistent with the results of surface cracking of Bead I in the double-bead Varestraint test.

4. The fracture surfaces of cracks on Bead I show a typical columnar austenite grain boundary, and are enriched with such elements as S, P, Si, O and N.

5. Cracking in the reheated Invar weld metal is considered to occur in two ways, each of which works independently or reciprocally. One is the liquation by the subsequent welding pass of the grain boundary of the weld metal, due to such low-melting constituents as nickel sulfate (Ni3S2, 635°C/1175°F) and nickel phos­phate (Ni3P, 875°C/1607°F). The other is the grain boundary embrittlement caused by the formation and growth of various compounds of sulfides, phosphides, oxides and carbonitride in the grain boundaries of the succeeding welding passes.

Close examination of the cracking sur­faces and the low-temperature impact-fracture surfaces of the welds reveals that the columnar grain boundary sur­faces studded with brittle, thick residues, such as sulfides and oxides, are, in most cases, accompanied with some predomi­nantly smooth fracture surfaces. A liquid film is suggested to play a role in the cracking mechanism. In addition, there is a significant difference between the de­trimental effect of sulfur on cracking in the reheated weld metal of Invar and that

of stainless steel. In the corrosion and oxidation of metals in high-temperature environments, it has been reported that the transportation and chemical reaction of oxygen are accelerated and activated by sulfur in the grain boundaries of Fe-Ni alloys. On welding of stainless steel, which usually contains more than 12% chromium, a stable and thin chromium oxide film may be formed on the weld bead surfaces and at the grain boundary, resulting in the above-mentioned strong activity of oxygen in the weld metal being curbed. This explains the higher resis­tance to cracking in the reheated weld metal of Type 300 series stainless steel as compared to Invar.

6. Some titanium and boron additions to Invar material increased markedly the hot-cracking susceptibility of reheated weld metals. It was presumed that these additions facilitated the embrittlement of grain boundaries through the precipita­tions of many small carbonitrides.

7. The detrimental effects of sulfur and phosphorus on the cracking sensitivi­ty of Bead II in the double-bead Vares­traint test of Invar (thought to be solidifi­cation cracking) are quite similar to the ones on fully austenitic stainless steel.

8. A 60-ton heat was made in a pro­duction run and was wrought and cold-rolled to 0.5-, 0.7- and 1.5-mm (0.02-, 0.03- and 0.06-in.)-thick gauges. These sheets were welded in butt, fillet and edge joints, resulting in satisfactory weld­ment properties.

9. From another mill heat, large Invar pipes of 11-mm (0.4-in)-thick gauge, 66-cm (26-in.) diameter and 5 m (16 ft) long were made by welding with the GTAW and plasma arc welding processes. No cracks or discontinuities could be found in the seam and girth welds, and the fracture surfaces were completely duc­tile, with superb toughness at —196 QC (-321°F).

References

1. Invar was developed by Guillaume in France and is the trademark of a product of Metalimphy. Brochures by Creusot-Loire and Inco: Invar M63 and 36% Nickel-Iron Alloy for Low Temperature Service.

2. Hunter, M. A. 1971. Low expansion alloys. Metals Handbook, 8th Edition, Vol. 1, pp. 816-819. American Society for Metals, Metals Park, Ohio.

3. Physics and applications of Invar alloys. 1978. Honda Memorial Series on Material Sci­ence (3). Maruzen Company, Ltd., Japan.

4. Loger, )., and Roere, C. 1976. Construc­tion des grands methaniers a curves integrees en invar cryogenique. Soudage et Techniques Connexes, Mai-)uin:183-195.

5. Schick, W. R. 1982. The development of stainless steel sealing washers for LNG tanker repairs. Welding lournal 61(12):15-21.

6. Clautice, W. E. 1975. Vacuum jacketed piping at the Kennedy Space Center. Welding lournal 54(7):500-509.

7. Brochure by Gaz Transport, lnvar-36% Nickel Alloy for Low Temperature Service.

8. Technical documents concerning the GT/MDC liquefied gas containment system by Gaz-Transport/McDonnell Douglas Corpora­tion.

9. Hemsworth, B., Boniszewski, T., and Eator, N. F. 1969. Classification and definition of high temperature welding cracks in alloys. Metal Construction 2:5-s to 16-s.

10. Hull, F. C 1967. Effect of delta ferrite on the hot cracking of austenitic chromium-nickel stainless steel. Welding journal 46(9):399-s to 409-s.

11. Lundin, C. D„ DeLong, W. T, and Spond, D. F. 1975. Ferrite fissuring relationship in austenitic stainless steel weld metals. Weld­ing lournal 54(8):241-s to 246-s.

12. Ogawa, T, and Tsunetomi, E. 1982. Hot cracking susceptibility of austenitic stainless steels. Welding journal 61(3):82-s to 93-s.

13. Witherell, C. E. 1964. Welding nickel-iron alloys of the Invar type. Welding lournal 43(4):161-s to 169-s.

14. Rundell, G. R., and Nehrenberg, A. E. 1966. Weld metal cracking of Invar in circular patch tests. Welding journal 45(4): 156-s to 160-s.

15. Nakagawa, H., Matsuda, F., Nagai, A., and Sakabata, N. 1980. Weldability of Fe-36%Ni alloy —report I. Transactions of japan Welding Research Institute 9(2):55-62.

16. Ogawa, T, Mori, N., and Nagano, K. August 1982. The weldability of Invar. 89th Committee of Welding Metallurgy, Japan Welding Society.

17. Savage, W. F„ and Lundin, C. D. 1965. The Varestraint test. Welding journal 44(10):433-s to 442-s.

18. Hansen, R. P., Elliott, R. P., and Shunk, F. A. 1958, 1965 and 1969. Constitution of Binary Alloys. 1st and 2nd Editions, McGraw-Hill, New York, N. Y.

19. Simons, E. L., Browing, G. V., and Lieb-hafskey, H. A. 1955. Sodium sulfate in gas turbines. Corrosion 11(12):505-t to 514-t.

20. Goebel, J. A., and Pettit, F. S. 1970. The influence of sulfides on the oxidation behavior of nickel-base alloys. Metallurgical Trans­actions 1(12):3421-3429.

21. Bornstein, N. S., and DeCrescente, M. A. 1971. The role of sodium in the accelerated oxidation phenomenon termed sulfidation. Metallurgical Transactions 2(10):2875-2883.

22. Spengler, C, and Viswanathen, R. 1972. Effect of sequential sulfidation and oxidation on the propagation of sulfur in an 85Ni-15Cr alloy. Metallurgical Transactions 3(1):161-166.

23. Bellware, M. D. 1964. How to weld Invar. Welding Engineer (11):41-43.

24. Gottlieb, T., and Shira, C. H. 1965. Fabrication of iron-nickel alloys for cryogenic piping service. Welding journal 44(3): 116-s to 123-s.

25. Ogawa, T., and Nakamura, H. 1974. Unpublished research report of Nippon Steel Corporation.

26. Brooks, ). A. 1974. Effect of alloy modi­fications on HAZ cracking of A-286 stainless steel. Welding lournal 53(11):517-s to 523-s.

27. Lippold, J. C. 1983. An investigation of heat-affected zone hot cracking in an alloy 800. Welding journal 62(1):1-s to 11-s.

28. Shindo, M„ and Kondo, T. 1976. Corro­sion of nickel-base heat resistant alloys in simulated VHTR coolant helium at very high temperature. Tetsu to Hagane 62(12):66-75.

WELDING RESEARCH SUPPLEMENT | 225-s

29. Lundin, C. D., and Chow, C. P. D. 1983. Hot cracking susceptibility of austenitic stain­less steel weld metals. Welding Research Council Bulletin 289(11):22-77.

30. Ogawa, T., Suzuki, K., and Zaizen, T. 1984. The weldability of nitrogen-containing austenitic stainless steel: part II —porosity, cracking and creep properties. Welding jour­nal 63(7):213-s to 223-s.

31. Soya, I., Takashima, H., and Tanaka, Y.

August 1984. -\n application of stress intensity factor to fatigue strength analysis of welded Invar sheet for cryogenic use. Presented at the ASTM National Symposium on Fracture Mechanics, Albany, New York.

32. Fleming, M. C 1974. Solidification Pro­cessing, p. 37. McGraw-Hill Material Science and Engineering Series.

33. Lippold, |. C . and Savage, W. F. 1979. Solidification of austenitic stainless steel weld­

ments: part I —a proposed mechanism. Weld­ing journal 54(12):362-s to 374-s.

34. Peterson, N. L. 1975. Diffusion in Solid, Recent Development, eds. A. S. Nowick and |. ). Burton, p. 167. Material science series, Aca­demic Press.

35. Wolf, ). S., and Evans, E. B. 1962. Effect of oxygen pressure on internal oxidation of Ni-AI alloys. Corrosion 18(4):129-t to 136-t.

Fatigue Design of Aluminum A report entitled "Development of Recommended Specifications for Fatigue Design of Aluminum

Structures" has been prepared by Professors W. W. Sanders, Jr., and J. W. Fisher for the Aluminum Association. The specifications are intended to be included in the next edition of Specifications for Aluminum Structures, published by the Aluminum Association. Those interested in reviewing and critiquing the recommended specifications may obtain a copy of the report by writing to: W. W. Pritsky, Aluminum Association, 818 Connecticut Ave., N. W., Washington, DC 20006.

WRC Bulletin 312 February 1986

Joining of Molybdenum Base Metals and Factors which Influence Ductility By A. J. Bryhan

This report discusses the current status of the joining technology of molybdenum-based metals and alloys. Information of practical significance is included, which will assist in both the design and utilization of molybdenum-based metals.

Publication of this report was sponsored by the AMAX Metals Group—Research Laboratory and the Reactive and Refractory Metals Committee of the Welding Research Council. The price of WRC Bulletin 312 is $14.00 per copy, plus $5.00 for postage and handling. Orders should be sent with payment to the Welding Research Council, Ste. 1301, 345 E. 47th St., New York, NY 10017.

WRC Bulletin 313 April 1986

Computer Programs for Imperfection Sensitivity Analysis of Stiffened Cylindrical Shells By R. L. Citerley

This report contains documentation for four computer programs used in the imperfection sensitivity analysis of cylindrical shells. The four programs are based upon Donnell's equation for cylindrical shells. The formulation of each program is worked in detail so others may make modifications. The input and output instructions are provided.

Publication of this report was sponsored by the Subcommittee on Shells and Ligaments of the Pressure Vessel Committee of the Welding Research Council. The price of WRC Bulletin 313 is $12.00 per copy, plus $5.00 for postage and handling. Orders should be sent with payment to the Welding Research Council, Ste. 1301, 345 E. 47th St., New York, NY 10017.

226-s | AUGUST 1986