Use of Plasma Arc Welding Process to Combat Hydrogen Metallic Disbonding of Austenitic

Stainless Steel Claddings

Cracking resistance increased when the plasma arc process with a hot wire filler metal was used to clad Cr-Mo base metal

BY O. A. ALEXANDROV, O . I. STEKLOV A N D A. V. ALEXEEV

ABSTRACT. A separation type crack, metallic disbonding, occurred between austenitic stainless steel weld metal cladding and 2V<iCr-1Mo base metal in the hydrodesulfurizing reactor of an oil refining plant. For stainless steel cladding, the submerged arc welding (SAW) process with a strip electrode is usually applied, but the authors experi­mented with the plasma arc welding (PAW) process with hot wire electrode for the cladding. The metallic disbond­ing is considered to be attributed to hy­drogen accumulation at the transition zone and has been generally studied on a laboratory scale using an autoclave . The authors used a electrolytic hydro­gen charging technique for the sake of experimental simplicity and made a comparison with the results for gaseous hydrogen charging. The main conclu­sions obtained were as follows:

The PAW stainless steel weld metal cladding is more resistant to metallic dis­bonding than the identical weld metal deposited by the SAW process.

Increased resistance to the disbond­ing wi th the PAW process is explained by the desirable microstructure and properties of the first layer of weld metal at the transition zone, i.e., fine austenite grains close to the interface with a min­imum of austenite coarse-grain bound­aries paralleling the fusion line, and the

O. /. STEKLOV is Director of the Welding En­gineering Department and A. V. ALEXEEV is Chief of the Plasma Arc Welding Research Group, State Academy of Oil and Gas, Moscow, Russia. O. A. ALEXANDROV is a Welding Engineer at NACAP Nederland B. V., The Netherlands.

small width and hardness of the transi­tion zone, which included the marten­sitic layer as-welded and the carbide layer after postweld heat treatment (PWHT).

Electrolytic hydrogen charging pretty well reproduces the results of autoclave gas phase charging.

Introduction

The metal l ic d isbonding phe­nomenon is a typ ical hydrogen em­br i t t lement problem occurr ing at the interface between austenitic stainless steel we ld metal c ladd ing and low-alloy ferritic base metal. The disbond­ing investigated occurred in the shut­down period of a hydrodesul fur iz ing reactor operating at high temperature and high hydrogen pressure (Refs. 1-12). It typical ly results from hydro­gen saturation of the interface region when the pressure vessel exposed to

KEY WORDS

Metallic Disbonding Austenitic Stainless Steel Stainless Steel Clad Weld Metal Plasma Arc Welding Transition Zone Austenite Grain Carbide Layer Pressure Vessels Cathode Charging

hydrogen is cooled d o w n . Figure 1 shows the distr ibution of hydrogen in the wal l of a vessel dur ing operat ion and after cool ing (Refs. 8, 9).

This part icular problem depends both on operating procedures (hydro­gen pressure, temperature, exposure cycles, cool ing and heating rates) and on interface character i t ics (type of c ladding and base metal , carbide layer, coarse austenite grains).

In spite of many investigations (Refs. 4, 7, 10-12), the mechanism of metal l ic d isbonding is sti l l not clear. It is supposed (Refs. 4 , 7) that the cracks mainly occur at the front edge of the hardened carbide layer, wh ich is the result of a carbon migration from the base metal to the weld metal dur­ing PWHT and intense carbide precip­itation at the interface wi th in the weld metal. In this report, the authors look at the inf luence of the we ld ing pro­cess on the cladding so as to give rec­ommendations regarding the safe op­eration of pressure vessels containing a high-temperature, high-pressure hy­drogen environment. For the stainless steel c ladding, the PAW process was used. This process has the fo l l ow ing advantages (Refs. 13, 14): shal low penetrat ion into the base metal , low di lut ion values (as little as 5%), a weld metal w i th very low carbon content, good transfer of a l loy ing elements, and high productivity. The PAW spec­imens were compared wi th test speci­mens obtained with the SAW process, a commonly used process for stainless steel c ladd ing of pressure vessels (Refs. 2, 3).

506-s I NOVEMBER 1993

Materials and Experimental Procedures

Clad Sample Production

The PAW process using a hot wire filler metal was performed with experi­mental equipment, which was designed in the research laboratory at the State Academy of Oi l and Gas in Moscow. This equipment is shown in Fig. 2. A schematic is shown in Fig. 3.

The base metal was identical for all test specimens. It is 2'/.iCr-1Mo steel, which is usually chosen for high-tem­perature, high-pressure hydrogen envi­ronments to avoid hydrogen attack prob­lems (Refs. 1-3). Stainless weld metal cladding either of Cr25-Ni1 3 or Cr1 9-Ni9 was used for the first layer and ei­ther Cr20-Ni9-Nb-V or Cr1 9-N i l O-Nb was used for a second layer. The chem­ical composit ion of the second layer does not influence the disbonding (Refs. 1, 2). Welding processes were of two types: SAW wi th a strip electrode and PAW wi th a hot wire electrode. The welding conditions of these processes are given in Table 1. Chemical compo­sitions of the base metal and consum­ables are given in Table 2. Chemical analyses of the first layer of deposited claddings are given in Table 3.

Metallic Disbonding Test

Two types of cracking tests were per­formed, and electrolytic charging was done by a special technique (Ref. 4) at various current densities and times. Specimens for this test were machined to the configuration shown in Fig. 4. Fig­ure 5 shows a setting situation for a spec­imen in the electrolyte. The electrolyte used was an aqueous solution of 5% sul­furic acid, to which Na2S203 x 5 H 2 0 was added. After the hydrogen charging specimens were left in the air for more than 24 h, they were cut normal to the fusion boundary at the middle section. Crack examination was carried out by optical microscopy.The crack percent­age was fixed on a crack's length with reference to a specimen's width.

Gaseous charging was produced in an autoclave containing hydrogen under the fol lowing conditions: 1 5 MPa pres­sure at 430°C for 48 h, fol lowed by air cooling. Dimensions of a specimen for this type of test are illustrated in Fig. 6. All surfaces of the specimen, except the deposited cladding, were surfaced with shielded metal arc welding using Cr25-Ni13 electrodes. The test specimens were left in the air for 14 days and then investigated for metallic disbonding. The extent of the disbonding was evaluated by ultrasonic testing and microscopic

^2i6

Fig. 1 — The distribu­tion of hydrogen in a pressure vessel wall

during processing and after cooling (15 MPa,

400°C, 20"C/h). 1) Dur­ing processing, 2) after 7.5 of cooling, 3) after

15 of cooling.

Table 1 — Welding Conditions of Cladding by SAW and PAW Processes

Frequency Protective of and

Size Travelling Plasmatron Oscillation Plasma-Welding Electrode, Current, Voltage, Speed, Oscillation, Amplitude, formation Process V

SAW 1st 65 X 0.7 800-850 32-34 layer

2nd 65 X 0.7 800-850 32-34 layer

PAW 1st <t> 3 290-310 22-26 layer

1st <j> 4 310-330 26-28

layer

2nd <fi 3 290-310 22-26 layer

m/h

8-10

8-10

5-7

6-8

5-7

Hz

0.5

0.5

0.5

m

0.05

0.05

0.05

Cas

Ar

Ar

Ar

Heat lnput,(a)

k j /mm

10.2-11.8

9.7-11.2

4.0-4.8 2-2.4

4.0-4.8 2-2.4

(a) Numerator — min and max value without heat efficiency coefficient. Denominator —min and max value using heat efficiency coefficient: SAW, 0.95 and PAW, 0.5.

Table 2 — Chemical Compositions of Materials Used

Material

Cr25Ni13 strip Cr19Ni10Nb

strip Cr25Ni13 wire Cr19Ni9 wire Cr20Ni9NbV

wire 2.25Cr-1Mo

plate

Size Electrode,

mm

65 X 0.7 65 X 0.7

0 4 « 3 <t>i

200 X 150 X

Chemical Composition (wt. %)

Mn Cr Other

Ni Elements

0.076 0.69 1.43 - 0.008 22.46 12.95 0.066 - 1.90 0.022 0.007 17.78 10.38 Nb 0.81

0.070 0.82 1.45 0.020 0.015 24.20 13.83 0.030 0.85 1.57 0.020 0.015 19.15 9.80 0.060 1.2 1 1.66 0.030 0.020 20.05 9.25 Nb 1.18;

V 1.15 2 0 0 X 1 5 0 X 5 0 0.100 0.28 1.11 0.020 0.020 2.19 0.54 Mo 0.5

Table 3 — Chemical Compositions of Clad Metal (First Layer)

Type of Electrode

Cr25-Ni13-str ip (SAW)

Cr25-Ni13—wire (PAW)

Cr19-Ni9-wi re (PAW)

C

0.094

0.072

0.047

Si

0.48

0.65

0.72

Che Mn

1.65

1.50

1.27

mical Corr P

0.030

0.026

0.020

position S

0.009

0.008

0.009

(wt -%) Cr

18.50

22.20

19.36

Ni

11.30

13.70

9.41

Mo

0.12

0.10

0.08

W E L D I N G R E S E A R C H S U P P L E M E N T I 5 0 7 - s

Power source of main plasmatron

Additional Power sourse of power sourrse additional plasmatron

Main plasma arc

Feed mechanism )T of wire electrode

Additional plasma arc

Fig. 2 — Equipment for plasma arc welding process. Fig.3-trode.

Schematic of plasma arc welding process with hot wire elec-

e x a m i n a t i o n . T h e c r a c k p e r c e n t a g e w a s c a l c u l a t e d o n the d i s b o n d e d area in re fe rence to the to ta l P A W or S A W c l a d sur face a rea .

A l l s p e c i m e n s w e r e inves t iga ted fo r m e t a l l i c d i s b o n d i n g a f ter P W H T at 6 5 0 ° C fo r 1 2 h ( fu rnace c o o l i n g ) .

Resul ts

Electrolyt ic Hydrogen Charging

T a b l e 4 s h o w s t h e e f f e c t s o f c a ­t h o d i c c h a r g i n g o n t h e d i s b o n d i n g f o r d i f f e r e n t v a l u e s o f c u r r e n t d e n ­s i t y a n d c h a r g i n g t i m e s . It is s e e n t h a t P A W c l a d w e l d m e t a l is m o r e r e s i s t a n t t o t h e d i s b o n d i n g t h a n S A W c l a d m e t a l at a l l c h a r g i n g c o n ­d i t i o n s ( c u r r e n t d e n s i t y 0 . 0 5 - 0 . 2 m A / m 2 , c h a r g i n g t i m e 1 2 - 3 0 h) . T h e S A W s p e c i m e n s h a d c r a c k s w i t h a l l

test c o n d i t i o n s . T h e P A W s p e c i m e n s w i t h C r 2 5 - N i 1 3 a re m o r e r e s i s t a n t t o m e t a l l i c d i s b o n d i n g . A t s o m e test c o n d i t i o n s , c r a c k s w e r e n o t f o u n d . T h e P A W c l a d m e t a l o f t h e t y p e C r 1 9 - N i 9 h a d t h e bes t r e s i s t a n c e to d i s b o n d i n g . Less t h a n 5 % c r a c k s w e r e f o u n d at a l l c h a r g i n g c o n d i ­t i o n s .

F i g u r e 7 s h o w s t h e e f f e c t o f c h a r g i n g t i m e o n t h e d i s b o n d i n g at t h e c u r r e n t d e n s i t y o f 0 .2 m A / m 2 . It is seen t h a t i n c r e a s i n g c h a r g i n g t i m e p r o m o t e s t h e d i s b o n d i n g . H o w e v e r , t h e t o t a l c r a c k l e n g t h is l o n g e s t w i t h t h e S A W s p e c i m e n s at a l l c h a r g i n g t i m e v a r i a t i o n s . T h e d i s b o n d i n g f o r P A W s p e c i m e n s w i t h t y p e C r 1 9 - N i 9 a f t e r 3 0 h w a s less t h a n 5 % , t h e P A W s p e c i m e n s o f t y p e C r 2 5 - N i 1 3 h a d 1 5 t o 2 0 % c r a c k s a n d t h e S A W s p e c i m e n s h a d 6 0 t o 7 0 % c r a c k s .

Autoclave Gas Charging

F i g u r e 8 s h o w s t h e r esu l t s o f a u ­t o c l a v e gas c h a r g i n g . T h e c o r r e l a ­t i o n b e t w e e n a u t o c l a v e gas c h a r g ­i n g a n d c a t h o d i c c h a r g i n g is g o o d . It c a n be s e e n t h a t S A W c l a d w e l d m e t a l g e n e r a l l y c r a c k s m o r e e x t e n ­s i v e l y t h a n P A W c l a d m e t a l . T h e S A W s p e c i m e n s h a d 18 to 2 0 % c r a c k s , w h e r e a s t h e P A W s p e c i m e n s t y p e C r 2 5 - N i 1 3 h a d o n l y 8 % c r a c k s a n d t y p e C r 1 9 N i 9 h a d 2 t o 4 % c r a c k s . T h e c l a d m e t a l t y p e C r 2 5 -N i 1 3 r e s i s t e d d i s b o n d i n g 3 t o 4 t i m e s b e t t e r w i t h t h e P A W p r o c e s s c o m p a r e d t o t h e S A W p r o c e s s . T h e P A W c l a d m e t a l t y p e C r 1 9 - N i 9 , w h i c h has a Cr a n d N i c o n t e n t a p ­p r o x i m a t e l y m a t c h i n g t h e S A W c l a d m e t a l C r 2 5 - N i 1 3 , is m o r e r e s i s t a n t

Side view

overlaid metal

we ld bond base metal

10 mm

Bottom view

7 mm

1 mm

10 mm

2 mm

Fig. 4 — Shape of the specimen for cathodic charging test.

Electrolite

Fig. 5 — Setting of the specimen in the electrolyte.

5 0 8 - s ! N O V E M B E R 1 9 9 3

manual cladded weld metal

second cladded layer first cladded layer

" 7 m m , :

Fig. 6 — Shape of the specimen for autoclave gas charging test.

t o m e t a l l i c d i s b o n d i n g .

Characteristics of Metallic Disbonding

Typ i ca l me ta l l i c d i s b o n d i n g charac­teristics are shown in Fig. 9 for ca thod ic and a u t o c l a v e gas c h a r g i n g . It is no te ­w o r t h y tha t c racks have a t e n d e n c y to o c c u r at t he cen te r o f a s p e c i m e n a n d are l oca ted in t he t r ans i t i on z o n e be ­t w e e n the stain less steel w e l d meta l c l add ing and the base me ta l . A c c o r d i n g to many invest igat ions (Refs. 5, 8, 9), the m e t a l l i c d i s b o n d i n g is a resul t o f t w o types o f c r a c k s : 1) c racks o c c u r r i n g in the carb ide layer at the inter face; and 2) cracks o c c u r r i n g a long coarse austeni te g ra in b o u n d a r i e s near t he w e l d in ter­face . In t he present s tudy (F ig.9) , the cracks m i c r o s c o p i c a l l y w e r e located in the carb ide layer at the interface, as we l l as a l o n g the austeni te gra in boundar ies w i t h i n the c l a d meta l a d j a c e n t to the base meta l . There was a c o m b i n a t i o n o f these t w o types o f c racks , a n d a t rans­granular type of crack was also found in the invest igat ion.

Figure 10 shows examples of the fracture surface for PAW and SAW samples (cathodic

charging). As can be seen, the transgranular fracture surface is generally dominant on both these specimens. There are in some places an intergranular brittle fracture surface. The PAW specimen has more of these places than the SAW specimen. The content of Cr and Ni on the fracture surface is - 1 2 % Cr and - 5 % Ni with the PAW process and ~11 % Cr and - 6 % N i wi th the SAW process. The distribution of these al loying elements (Fig. 11) and data of Table 5 show that these cracks occurred in the transition zone. In the case of the PAW process, the crack is located inside the car­bide layer (-15 pm from the fusion line) and in the case of the SAW process, it is located outside the carbide layer ( - 5 5 - 6 0 pm f rom the fusion line). The transgranular type of cracking is, presumably, connected with the peculiarities of cathodic charging.

The results of cracking tests show that ca­thodic charging pretty wel l reproduces char­acteristics o f metal l ic d isbonding obtained from autoclave gas phase charging, with more pronounced cracking along austenite grain boundaries inside the clad metal close to base metal in one case and the development of transgranular type of cracking in another case. It can be seen that by using cathodic charg­ing it is possible to get even more severe dis-

Table 4 — Effect of Charging Conditions on Metallic Disbonding

Current Testing Time (h)

mA/m 2

0.2 0.1 0.05

12

on * o • • •

18

• a * •n * • D *

24

• a * • a • • a *

30

• 3 * ca* ca •

O D * — no disbonding QLT* — slight disbo.iding (<5% cracks) • • • - s e r i o u s disbonding (>5% cracks) * - SAW process (Cr25-Ni 13) O - PAW process (Cr25-Ni 13) • - P A W process (Cr19-Ni'l)

Table 5 — Dimension of Austenite Grains, Carbide Layer and Transition Zone

Width lam)

Type of Electrode

Cr25Ni13-strip (SAW)

Cr25Ni13-wire (PAW)

Cr19Ni9-wire (PAW)

Grain(a)

(cm)

Carbide Transition Layer(b) Zone

100-500 2 0 - 4 0 8 0 - 1 0 0

200-250

70-750 150-200 50-400 120-160

1 0 - 2 0 2 0 - 3 0

15— 25 25—35

(a) Numerator - min and max dimention o f grains near the weld bond in the overlaid melal. Denominator - mean value. (b) Mean value.

bonding than by using autoclave gas charg­ing.

D i s c u s s i o n

Grain Morphology

Grains in the y phase c lose to the in ­te r face , w h e r e the d i s b o n d i n g occu rs , w e r e f o u n d to be p lanar a n d coarse in

Disbonding, %

A

I

1 / j

<

I —•

^ r A

• . I

i

\

1,

Cr25Ni13 SAW

^ ^ ^ ^ - ^ C r t S N M S PAW

1

| flCM9Ni9 PAW

Testing time, hours

Fig. 7 — Effect of charging time on the metallic disbonding (current density i = 0.2 mA/m2).

Fig.8. — Resistance to the disbonding after autoclave gas charging test.

20

16

12,

8

4

0

Cr25Ni13 SAW

Cr25Ni13 PAW

CM9Ni9 PAW

Type of cladded weld metal

W E L D I N G R E S E A R C H S U P P L E M E N T I 5 0 9 - s

t v O V E R L A I D .

/ METAL I

•" . ? ' N x : - i kf ' i • Cw ' ,<> v v **~-

r f i ^ W % • . • • ; : BASE M E T A L

O V E R L A I D

M E T A L

Fig. 9 — Typical ex­amples of the metal­lic disbonding. 500X. A — SAW specimen, cathodic charging; B — PAW specimen, cathodic charging; C — SAW specimen, autoclave gas charging.

Fig. 10 — The fracture surface of clad weld metal type Cr25-Nil3. A — PAW; B — SAW. 300X.

510-s I NOVEMBER 1993

Cr,%

-20 0 20 40

DISTANCE y<m)

Fig. 71 — The distribution of Cr and Ni in the transition zone. 780X. A — PAW specimen, type Cr25-NH3.

MEIAL \'M

..

• • . .

icrij

• OVERLAID! METAL

|7T|

. . .... - .

ifeAWW 1 1

Hy^Vw** * *^

•H i

HAZ

0 20 40 60

DISTANCE (Mm)

B — PAW specimen, type Cr19-Ni9.

BASE METAL

BjQHSj >j f^SSKk\

^^^^^^^f^^^

^

.

• OVERLAID j METAL

y : . . . - . . • • • •

• : . • : • } 5 ' : * " ' ; • • • •

| C r j H |

,

A

• ' • • '• "i-fjj t \

40 60

DISTANCE (um)

C — SAW specimen, type Cr25-Nil3.

WELDING RESEARCH SUPPLEMENT I 511-s

OVERLAID

^ / ^ i K ^ v * •/•;-: • '4:;":.; #.K

vt ryy:iy>-;ymiym.^:f•

OVERLAID METAL

1 BASE H H

Fig. 12 — The microstructure of the transition zone. 200X. A — PAW specimen, type Cr25-NH3; B — PAW specimen, type Cr19-Ni9;

C — SAW specimen, type Cr25-Ni 13.

OVERLAID •;•a > METAL

• t ~ mm A

yyf-

BASE MFTAL S > V - . » • j .: * & $ & & . .? ^

the case of SAW compared to PAW — Table 5. The worst disbonding occurred in SAW specimens with y coarse grain boundaries parallel to the fusion line — Fig. 1 2C. Specimens wi th a finer grain structure and smaller length of grain boundaries parallel to the weld fusion line (Fig. 1 2A) were more resistant to metallic disbonding. The best crack re­sistance was wi th the clad metal type Cr19-Ni9, which had the smallest austenitic grains and no specific grain boundary near the interface — Fig. 1 2B.

One preventive measure against metallic disbonding is to promote a finer structure without a y coarse grain bound­ary parallel to the fusion line. The metal­lic disbonding usually locates micro­scopically along these grain boundaries (Refs. 2, 5 , 1 5 ) . This is fundamental re­garding resistance to hydrogen embrit­tlement (Refs. 6, 1 3). Finer grains w i l l mean that hydrogen, carbon and harm­ful impurities, such as sulfur and phos­phorus, wi l l be less concentrated at the boundaries. Impurity segregation has a negative influence on the granular ad­hesion and decreases the surface energy value of a crack. Investigations (Refs. 7, 12, 16) show that sulfur, phosphorus, sil icon and carbon influence cracking along the grain boundary. Also, the grain orientation toward stresses at the inter­face when cooling down wil l not be the

512-s I NOVEMBER 1993

same. The formation mechanism of ycoarse

grain boundaries can be stated as fo l ­lows (Ref. 5): the austenite grains at the fusion boundary in the heat-affected zone (HAZ) formed in the 8 —> y trans­formation during cool ing are going to grow into the clad metal. Before that, however, other y grains have already nu­cleated and have been growing in the transition zone near the composite re­gion from the reaction of l iquid —> l iq­uid + 8 —> l iquid + 8+ y during solidif i­cation. Therefore, when the y grains from the HAZ grow only a little into the clad metal, the y grains from the HAZ and the clad metal collide with each other in the transition zone at about 1350°C, and this collision makes the y grain bound­ary parallel to the fusion boundary. This y grain boundary shifts a little accompa­nying the disappearance of 8 during cooling from 1 350° to 1 300°C, and the zone between the grain boundary and the carbide layer formed after PWHT is regarded as the y coarse grain.

Zhang, ef al. (Ref.5), show also that if the y coarse grain boundary is located inside the carbide layer (the intersection of 8 + y—>y boundary line with 1300°C, it is a stopping point for the y grain boundary), and it is effective in prevent­ing cracking.

The authors have used the transition zone transformation (TZT) diagram (Ref. 5) and Fe-Cr-Ni phase diagram (Ref. 1 7) for the design of a new TZT diagram for the investigated specimens on the basis of an imaginary Cr and Ni distribution in the transition zone. Figure 11 shows a gradient in alloy level from the ferritic substrate into the weld metal extending over a distance of 20 to 35 pm in the PAW samples and 80 to 1 00 pm in the SAW samples. The distribution of l iq ­uidus, solidus 8 —> 8 + y a n d S + y —>y transformation temperatures in the tran­sition zone is roughly shown in Fig. 1 3. The abscissa is the distance from the fu­sion boundary to the inside of the clad metal, and the origin is set to the fusion boundary. The right border, namely the composite region (terminology of Zhang, ef ai), is the solidif ication as it proceeds from liquid (L)L + 8—>L+S + y —> 8 + y in type Cr25-Ni13 clad metal — Fig. 1 3A and C. The transition zone in this case, except the part near the composite region, solidifies as a single 8 phase. The l iquidus, the solidus and the 8 + y —> y boundary lines fall nearly monotonously together wi th the dis­tance, but the 8 —> 8 + y boundary line falls a little then rises to a maximum, and again falls near the composite region. Figure 13B shows that the clad metal type Cr1 9-Ni9, including the transition zone, solidifies as primary 8 phase, and y is formed after the completion of soli-

S_ 1400

Q-1300

TZT diagram in PAW cladded metal type Cr26Ni1 3

-Transition zone in overlaid metal

L

I'lOO

20

Distance (yu m]

B TZT diagram in PAW cladded metal type Cr1 9Ni9

-Transition zone in overlaid metal -

L

20 30

Distance [ u m ]

ISOO

IZOO

O 1S00 -

CX £ 1300

TZT diagram in SAW cl

-*—-Transition zone

======-

cT

-•- Carbide layer-*-!

dd

n overla

id metal type Cr25Ni13

L

cT+

X

, I

-

- Ihoo

20 JI0 B0 no Distance [u m]

IZOO "00

Fig. 13—An example of TZT diagram. A — SAW specimen, type

Cr25-Nil3; B — PAW specimen, type Cr 19-Ni9;C —SAW speci­men, type Cr25-Nil3.

dification. It is noteworthy that the tem­perature of the 8 + y —> y boundary line drops below 1 300°C in the vicinity of the fusion boundary.

As can be seen, the intersection of the 8 + y —> y boundary line at 1 300°C in the deposit type Cr19-Ni9 locates in­side the carbide layer formed after PWHT. But this line is always above 1 300°C in type Cr25-Ni1 3 deposits (for both PAW and SAW). This means that no ycoarse grain boundary paralleling the fusion line is formed in the transi­tion zone of the former clad metal. Usu­ally the ygrain boundary paralleling the fusion boundary forms outside the car­bide layer. In both of the other cases, on

the contrary, this specific boundary is hardly formed inside the transition zone. Similar grain morphology was observed in the real specimens. The clad metal of type Cr1 9-Ni9 showed no ycoarse grain boundary paralleling the fusion bound­ary, but in deposites of the Cr25-Ni1 3 type, this specific grain boundary was found.

The difference of the austenite grain morphology in the transition zone be­tween the SAW and PAW clad metal can be explained by the effect of welding process parameters. The PAW process has a low heat energy that leads to a high cooling rate and a short contaction time for the solid and liquid phase during so-

WELDING RESEARCH SUPPLEMENT I 513-s

Fig. 14 — Schaeffler diagram. 1) PAW, Cr25-Ni13, first layer deposit; 2) PAW, Cr19-Ni9, first layer deposit; 3) SAW, Cr25-NH3, first layer deposit.

20

16

12

8

4

0

\

A + M

Martensite \ .

^^

\ ^^^ \ 2,2SC*iMoplate

F \ • \

M \

Austenite

M + F

^

l / / s A + M + F

% Ferrite -.

Cr25M13 strip

' C£l9NBwire

> 0

/ / / / £r25Ni13wire

A + F

^

Ferrite

Chromium Equivalent = % Cr + % Mo + 1,5 x % Si + 0.5 x % Cb

l idif ication. As can seen from Table 5, the heat input during PAW (2 to 2.4 kj/mm) was four times less than when using SAW (9.7 to 11.2 kj/mm). As a re­sult, the HAZ should be above the A3

transformation temperature for a shorter time with PAW. The welding pool is more overcooled, and more new solidi­fication centers in front of the growing grains are formed from the HAZ. All these factors confirm a smaller size for austenitic grains. As can be seen from Table 5, austenitic grain size close to the interface is on the average 120 to 200 pm in the PAW specimens and 200 to 250 pm in the SAW samples. A smaller width for the transition zone in the PAW samples is a result of smaller depth of penetration into the base metal when using the plasma arc process and lower weld pool mixing (Refs. 14, 18).

The width of the transition zone af­fects the ratio of the y coarse grain be­cause the coarse grain is formed inside the transition zone. It was shown (Ref. 5) that the ratio of the y coarse grain has an increasing linear correlation with the

'C

7100

900

no <

500 t

Fig. L

_

•-

15

o

o o V x

o

1 1

7 5

- Influence of PWHT on

o

C

o

X

disbondinCL.

X X X X

O ' '

1 Q 1

10 so f]t

the metallic disbonding.

JK/?S

contraction of melted clad metal wi th the base metal irrespective of the cool­ing rate. It is thought that the width of the transition zone formed in the melt­ing state, i.e., the distribution of alloy­ing elements is one of the major factors for the ycoarse grain. This can be clearly seen in the case of PAW clad metal of the type Cr19-Ni9, which has an alloy content approximately matching SAW clad metal type Cr25-Ni1 3.

Nature of the Cladding

According to Schaeffler's diagram (Fig. 14), the chemical composit ion of the first layer of deposits is such that the microstructure consists of austenite and ferrite in PAW samples and austenite in the SAW clad metal. The PAW speci­mens of type Cr25-Ni1 3 have - 5 % fer­rite and type Cr19-Ni9 has 5 to 10% fer­rite. Metallographic analyses (Fig. 12) also show the austenite structure with a small quantity of ferrite in the case of the PAW specimens. The ferrite in the austenitic stainless clad metal also af­

fects the grain size by changing the solidif ication pro­cess (Refs. 18, 19). The 8 ferrite estab­lishes new solidifi­cation centers in front of growing columnar grains. Due to this effect, the austenite grains are going to be smaller. In com­paring the size of austenite grains of PAW clad metal for types Cr25-Ni13 and Cr19-Ni9 (Table 5), it can be seen that austenite grains close to the inter­face in the second case are smaller, which depends,

presumably, on the quantity of 8 ferrite, since the heat input in both cases was the same.

Due to the low dilution characteris­tic of the PAW process, this austenite-ferrite microstructure was obtained in the cladding. The di lut ion in the PAW specimens was 7 to 10% , but it was 20 to 25% in the SAW samples. That might explain the higher content of Cr and Ni in the PAW cladding (22.2% Cr and 13.7% Ni) compared with the SAW specimen (18.5% Cr and 11.3% Ni). The consumables are also similar in this re­gard: type Cr25-Ni13 strip (SAW) hav­ing a content of 22.5% Cr and 1 3% N i ; and the wire (PAW) having 24.2% Cr and 13.8% Ni . Carbon content in the first layer was 0.072% with PAW and 0.094% wi th SAW at the same carbon percentage in the consumables.

Nature of the Transition Zone

As is known, the metallic disbonding occurs only in PWHT material (Ref. 7), so the disbonding depends on the struc­ture and properties of the hardened car­bide layer. The influence of PWHT on the metallic disbonding is a result of car­bon migration from the low-alloy base metal to the deposited high-alloy clad metal wi th carbide precipitation at the interface. Figure 1 5 shows the influence of PWHT on the disbonding (Ref. 12). It is necessary to point out that the real PWHT for the pressure vessels usally is 690°C during 24 to 30 h, but the PWHT in the present study (650°C for 12 h) was enough to promote the metallic dis­bonding.

The conclusions of many investiga­tions (Refs. 4, 7, 20, 21) are that the de­crease of carbon migration during the PWHT and the prevention of carbide precipitation in the transition zone de­crease the metallic disbonding. Metal­lographic examination showed that in the as-welded condition the fusion zone of all specimens consists of HAZ, tran­sition area wi th austenite-martensite structure adjacent to the fusion line, and austenite or austenite-ferrite clad metal. This observation fol lows those of nu­merous other investigations (Refs. 2, 7, 18, 19, 22). The martensite layer mor­phology had an open texture that devel­oped in the direction of sol idif ication. The martensite region at the interface is supposed to be an area with less than 7% Ni (Refs. 1 9, 22). As can seen from Fig. 11, this corresponds to the width of the martensite layer (15 to 20 pm) in the PAW samples, which is 3 to 4 times less than in the SAW specimen (-60 pm), mainly due to the low penetration of the base metal in the case of the plasma arc process (Ref. 1 8).

Microhardness tests (Fig. 16) indi-

514-s I NOVEMBER 1993

450

410

370

330

290

250

210

170

130

As - welded

Overlaid metal :

L

i

'

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I I

liW I

on / ,

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Fusion Ii

!

)

Base meta!

d H

P^J

Y.> rO&

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y f > - 0

After PWHT

Overlaid metal Base metal

5" s M l ^ 8 yp—o~ -o

450

410

370

330

290

250

2I0

170

130

H,

0,3 0,1 0 0,1 0,3

DISTANCE, mm

A

0,5 0,3 0,1 0 0,1 0,3

DISTANCE, mm

B

0,5

Fig. 76 — Microhardness distribution (100 g) near weld interface. • PAW specimen, type Cr25-Nil3; 3 PAW specimen, type Cr19-Ni9; O SAW specimen, type Cr25-Nil3.

cated that the HAZ of the 2%Cr-1Mo steel is about 240 to 280 HV. As the fu­sion boundary was approached, a low hardness value was recorded (200-220 HV) as a result of decarburization dur­ing welding. As can be seen, this effect is less for the PAW specimens, possibly as a result of a higher cooling rate and less development of the diffusion pro­cess. Once the fusion line was crossed, microhardness rose rapidly, reaching a peak before fall ing rapidly for the bulk of the first layer (240-275 HV). The hard­ness of the martensite zone was highest for the SAW cladding (450 HV). The PAW cladding had values of 410 HV (Cr19-Nj9) and 345 HV (Cr25-Ni1 3).

After PWHT, carbide precipitation along the fusion line was found (the dark layer at the interface on the stainless steel side — Fig. 12), and a decar-bonization zone developed in the base metal. Precipitation was also seen on the austenite grain boundaries close to the interface. The carbide precipitation had clearly occurred in the region that was martensitic in the as-welded condi­tion. The martensite layer during PWHT had a structural transformation, yet it kept the morphological peculiarities of virgin martensite (acicular structure). The decomposition of the original inter­facial martensite structure during the tempering can seen in the case of de­posited metal for types Cr19-Ni9 and Cr25-Ni13 (SAW). The hardness peak is higher in the as-welded condition com­pared to after the heat treatment. The cladding during PWHT is also struc­

turally changed as a consequence of dis­persion hardening and resolidification.

After PWHT, the hardness in the HAZ is reduced, and that region, which in the as-welded condit ion had a value of about 240 to 280 HV, was found to be 180 to 200 HV after heat treatment. The hardness was lowest near the boundary (1 55-180 HV), which is also lower than in the as-welded condit ion. The weld metal contains a hardness peak just in­side the stainless steel layer. This is the region clearly showing carbide precipi­tation. The width of the carbide layer along the fusion line after PWHT was ir­regular, but it was less in the PAW spec­imens (1 5-25 pm) compared to the SAW specimens (20-40 pm) —Table 5. In the plasma arc deposits, the hardness peak is considerably smaller (360-370 HV) than in the submerged arc deposit (-420 HV). A little farther from the boundary the hardness falls again, but in the SAW sample, the hardened zone is wider and harder. At 100 pm from the fusion line, SAW values were around 375 HV as compared to 305 HV (Cr19-Ni9) and 235 HV (Cr25-Ni1 3) in the PAW sam­ples.

The results of the present study show that the metallic disbonding occurred in the stainless clad metal close to the base metal where the martensitic structure was found. Also, the transgranular type of cracking was observed in this zone. This region wi th the martensitic struc­ture can include the carbide layer after PWHT if using welding processes with a low di lut ion rate, such as PAW, or it

can be wider than the carbide layer with the SAW process. So not only the struc­ture and properties of the hardened car­bide layer influence metallic disbond­ing, but also the properties of the whole region where the martensite structure can be formed during welding. It was pointed out (Ref. 6) that the metallic dis­bonding increases with dilution because with a high dilution value more carbon w i l l be present at the interface along with a wider and more irregular marten­sitic layer.

The results obtained indicate that the smallest width and lowest hardness for the fusion boundary martensite in the as-welded condit ion and the interface hard zone after PWHT on the stainless steel side, including the carbide layer, are displayed in the PAW samples.

The better resistance to metallic dis­bonding (cracks in the carbide layer and the transgranular cracks in the decom­posite martensitic layer), after PAW pro­cess, can be explained in this case by improved properties in the transition zone.

Conclusions

The main conclusions obtained are as follows:

1) The plasma arc process for de­positing stainless steel cladding is more resistant to metallic disbonding than the submerged arc welding process. Type Cr25-Ni13 clad weld metal generally cracks more than type Cr19-Ni9 clad weld metal.

W E L D I N G RESEARCH SUPPLEMENT I 515-s

2) The increasing resistance to meta l ­l ic d i sbond ing in the case o f the plasma arc w e l d i n g process can be e x p l a i n e d by the f a v o r a b l e charac te r i s t i cs o f th is sur fac ing process, w h i c h inc lude a lower heat ene rgy a n d p e n e t r a t i o n , h igher c o o l i n g rates, shorter t i m e fo r the so l id and l i q u i d phase d u r i n g s o l i d i f i c a t i o n , and lower m i x i n g and d i l u t i on . A l l these fac tors c o n t r i b u t e to t he f i ne gra ins o f austeni te st ructure ad jacent to the w e l d in te r face , the least leng th o f y coarse g ra in b o u n d a r y pa ra l l e l to the f us i on l i ne , a n d the sma l les t w i d t h and h a r d ­ness of the t ransi t ion zone , i nc lud ing the mar tens i t i c layer (as -we lded) and the carb ide layer (after PWHT) .

3) The results o f the c a t h o d i c cha rg ­ing test p re t ty w e l l r e p r o d u c e d the re­sults o f the au toc lave gas phase c h a r g ­ing test. The c racks a l o n g the g ra in b o u n d a r i e s w e r e m o r e p r o n o u n c e d in the austen i te stainless c l ad meta l c lose to the base metal in one case, and trans­granular fa i lu re deve loped in the t ransi ­t i on zone w i t h the decompos i te mar ten­sit ic structure in another case.

References

1. Steklov, O. I., Alexeev, A. V., and Alexandrov, O. A. 1988. Disbonding of austenitic stainless clad steel pressure vessels containing hydrogen. Ts lNTlKhlM-NEFTEMASh, Moscow, pp. 1-24.

2. Technical report of weld overlay dis­bonding. Symposium on Heavy Wal l Pres­sure Vessel. ATB, Moscow, 1985. 1(Q): 1-7.

3 .0hnish i , X., Fuji, A. 1984. Effect of strip overlay condit ions on resistance to hydro­gen-induced disbonding. Trans. JWS, 1 5(2): 49-55.

4. Matsuda, F., Nakagawa, H., Tsuruta, S., and Yoshida, Y. 1984 Disbonding between 27.Cr-1Mo steel and overlaid austenitic stain­less steel by means of electrolytic hydrogen charging technique. Trans, of JWRI, 1 3(2): 263-272.

5. Zhang, Y., Nakagawa, H., and Mat­suda, F. 1987. Proposal of TZT diagram for microstructural analysis of transition zone in dissimilar metal weld ing. Trans, of JWRI, 16(16): 103-113.

6. Pressoure, C , Chaillet, J., and Valette, G. 1 982. Parameters affecting the hydrogen disbonding of austenitic stainless cladded steels. Current Solution to Hydrogen Prob­lems in Steel. ASM, New York, pp. 349-355.

7. Imanaka, T., Shimomura, I., and Nakano, S. 1985. Hydrogen attack in Cr-Mo steels and disbonding of austenitic stainless weld overlay. Kawasaki Steel Technical Re­port, 13(9): 109-119.

8. Okada, H., Naito, K., and Watanabe, J. 1982. Hydrogen-induced disbonding of stainless steel weld overlay in hydrodesulfur­izing reactor. Current Solution to Hydrogen Problems in Steel. ASM, New York, p. 331 -339.

9. Naito, K., Okada, H., and Watanabe, J. 1980. Study on hydrogen embrittlement of pressure vessels overlaid with stainless steel. Hydrogen embritt lement of transition zone between weld overlay and base metal. Pres­sure Engineering, 18(5): 39-46.

10. Ohnishi , X., Chiba, R., and Watan­abe, J. 1985. Hydrogen induced disbonding of stainless steel overlay we ld . Symposium on Heavy Wal l Pressure Vessel. ATB, Moscow, 1(P): 1-35.

11 . Kinoshita, K., Itoh, H., Ebata, A., and Hattori, T. 1985. Mircoscopical critical con­dition for the initiation of disbonding of weld overlaid pressure vessel steel. Trans. Iron and Steel Inst. Jap.), pp. 505-512.

12. Imanaka, T. 1984. Development of austenitic stainless weld overlay having an

excellent resistivity against disbonding. J. Jron and Steel Institute of Japan, 70(5): 669.

13. Vainerman, A. E., Shorshorov, M. Ch., Veselcov, V. D. and Novoselov,V. S. 1969. Plasma arc welding process for cladding of metals. Mashinostroenie, Leningrad.

14. Steklov, O. I., etal. 1989. A high-pro­ductivity process of plasma arc hot wire sur­facing. Welding International, 12: 1058-1059.

15. Libra, O., and Soukup, K. 1985. K problematice tvoreni vodikem indukovanych trhl in u vysokotlakych nadob s navary. Svaranie, 34(10): 297-303.

16. Sakai, T., Asami, K. , and Katsumata, M. 1 982. Hydrogen induced disbonding of weld overlay in pressure vessels and its pre­vention. Current Solutions to Hydrogen Prob­lems in Steels. ASM, New York, pp. 340-348.

17. Rivlin, V. C , and Raynor, C. V. 1980. Crit ical evaluation of constitution of chromium-iron-nickel system. International Metals Review 1 : 21-38.

18. Livshits, L. S. 1979. Science of metals for welders. Mashinostroenie, Moscow.

19. Gotalskij, Y.N. 1980. Welding of het­erogeneous steels. Mashinostroenie, Leningrad.

20. Tadachi, H., Toshiaki, F., and Kazuhisa, K. 1986. Hydrogen induced dis­bonding of stainless steel overlay weld and its preventive measures. Nippon Kokan Tech­nical Report, 47: 17-22.

21 . Steklov, O. I., Alexeev, A. V., Alexan­drov, O. A., Smirnov, V. I., Semenov, J. N., Bublik V. G., and Ovcharenco L. V. 1989. Patent USSR No: 1558596, December. Method of cladding.

22. Zemzin, V. N. 1966. Welded joints of heterogeneous steels. Mashinostroenie, Moscow-Len i ngrad

A M E R I C A N W E L D I N G SOCIETY C O N F E R E N C E P R O C E E D I N G S

International Conference on Computerization of Welding Information IV Thirty-two papers by professionals from major organizations presenting the latest techniques in the field of computer welding information are included in this 394 page proceedings from the conference held November 3-6, 1992 in Orlando, Florida. This conference was sponsored by the American Welding Society, the American Welding Institute, and the National Institute of Standards and Technology. Topics include data formats and searchable standards, welding engineering applications, quality and non­destructive examination, weld sensing for real-time control, weld controllers and control systems, and databases and welding procedures. (Hardbound) Code CP-1192 List: $125.00 AWS Members: $93.75

International Conference on Underwater Welding This 169 page conference proceedings includes thirteen papers by recognized authorities in the underwater welding field presented at the conference held in New Orleans, LA, March 20-21, 1991. Topics cover state-of-the-art developments in the underwater industry including weld ing equipment and processes, mechanical and internal weld properties, maintenance and inspection procedures, and weld ing applications in shallow and deep water. (Softbound) Code: CP-391 List: $50.00 AWS Members: $37.50

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516-s I NOVEMBER 1993