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. A L E X A N D R O V , 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 c o n c l u sions 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 w i t h 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.
506-s I N O V E M B E R 1993
small width and hardness of the transition zone, w h i c h 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 m e t a l l i c d i s b o n d i n g phen o m e n o n is a t y p i c a l hydrogen e m brittlement p r o b l e m o c c u r r i n g at the
interface between austenitic stainless
steel w e l d metal c l a d d i n g and l o w alloy ferritic base metal. The disbonding investigated occurred in the shutd o w n period of a h y d r o d e s u l f u r i z i n g
reactor operating at high temperature
and high hydrogen pressure (Refs.
1-12). It typically results from hydrogen saturation of the interface region
w h e n 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 c o o l e d d o w n . Figure 1
shows the distribution of hydrogen in
the w a l l of a vessel d u r i n g operation
and after cooling (Refs. 8, 9).
This particular p r o b l e m depends
both on operating procedures (hydrogen pressure, temperature, exposure
cycles, cooling and heating rates) and
on interface characteritics (type of
c l a d d i n g and base m e t a l , carbide
layer, coarse austenite grains).
In spite of many investigations
(Refs. 4, 7, 10-12), the mechanism of
metallic d i s b o n d i n g is still not clear.
It is supposed (Refs. 4 , 7) that the
cracks mainly occur at the front edge
of the hardened carbide layer, w h i c h
is the result of a carbon migration from
the base metal to the w e l d metal during PWHT and intense carbide precipitation at the interface within the weld
metal. In this report, the authors look
at the influence of the w e l d i n g process on the cladding so as to give recommendations regarding the safe o p eration of pressure vessels containing
a high-temperature, high-pressure hydrogen environment. For the stainless
steel c l a d d i n g , the PAW process was
used. This process has the f o l l o w i n g
advantages (Refs. 13, 14): s h a l l o w
penetration into the base metal, l o w
dilution values (as little as 5%), a weld
metal w i t h very l o w carbon content,
good transfer of a l l o y i n g elements,
and high productivity. The PAW specimens were compared w i t h test specimens obtained with the SAW process,
a c o m m o n l y used process for stainless
steel c l a d d i n g of pressure vessels
(Refs. 2, 3).
^2i6
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 O i 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,
w h i c h is usually chosen for high-temperature, high-pressure hydrogen environments to avoid hydrogen attack problems (Refs. 1-3). Stainless w e l d metal
cladding either of Cr25-Ni1 3 or Cr1 9Ni9 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 composition of the second layer
does not influence the disbonding (Refs.
1, 2). W e l d i n g processes were of t w o
types: SAW w i t h a strip electrode and
PAW w i t h 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 w h i c h N a 2 S 2 0 3 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 following conditions: 1 5 MPa pressure at 430°C for 48 h, followed 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 Cr25Ni13 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
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
Welding
Process
Frequency
Protective
of
and
Size
Travelling Plasmatron Oscillation PlasmaHeat
Electrode, Current, Voltage, Speed, Oscillation, Amplitude, formation lnput, (a)
kj/mm
V
m/h
Hz
m
Cas
SAW 1st
65 X 0.7 800-850
layer
32-34
8-10
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
10.2-11.8
9.7-11.2
8-10
5-7
0.5
0.05
Ar
4.0-4.8
6-8
0.5
0.05
Ar
2-2.4
4.0-4.8
2-2.4
5-7
0.5
0.05
Ar
layer
2nd
layer
<fi 3
290-310
22-26
(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
Chemical Composition (wt. %)
Size
Electrode,
mm
Mn
Cr
Ni
Other
Elements
Nb 0.81
65 X 0.7
65 X 0.7
0.076
0.066
0.69
-
1.43
1.90
0.022
0.008
0.007
22.46
17.78
12.95
10.38
04
«3
<t>i
0.070
0.030
0.060
0.82
0.85
1.2 1
1.45
1.57
1.66
0.020
0.020
0.030
0.015
0.015
0.020
24.20
19.15
20.05
13.83
9.80
9.25
200
2
0 0X
X 1150
50X50
0.100
0.28
1.11
0.020
0.020
2.19
0.54
Nb 1.18;
V 1.15
M o 0.5
Table 3 — Chemical Compositions of Clad Metal (First Layer)
Type of
Electrode
Cr25-Ni13-strip
(SAW)
Cr25-Ni13—wire
(PAW)
Cr19-Ni9-wire
(PAW)
Che mical Corr position (wt -%)
P
S
Cr
Mn
C
Si
Ni
Mo
0.094
0.48
1.65
0.030
0.009
18.50
11.30
0.12
0.072
0.65
1.50
0.026
0.008
22.20
13.70
0.10
0.047
0.72
1.27
0.020
0.009
19.36
9.41
0.08
W E L D I N G RESEARCH S U P P L E M E N T I 507-s
Power source of
main plasmatron
Additional
power sourrse
Power sourse of
additional plasmatron
Feed mechanism
)T of wire electrode
Main plasma arc
Fig. 2 — Equipment for plasma arc welding
process.
examination. The crack percentage
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 r e f e r e n c e t o t h e t o t a l P A W or S A W
c l a d surface area.
A l l specimens were investigated for
m e t a l l i c d i s b o n d i n g a f t e r P W H T at
650°C for 1 2 h (furnace cooling).
Results
Electrolytic Hydrogen Charging
T a b l e 4 s h o w s t h e effects of c a t h o d i c c h a r g i n g on the d i s b o n d i n g
for different values of current dens 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
resistant to the 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 ditions (current density 0.05-0.2
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 specimens had cracks w i t h all
Fig.3trode.
Additional
plasma arc
Schematic of plasma arc welding process with hot wire elec-
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 are m o r e resistant
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
conditions, cracks were not f o u n d .
The P A W clad metal of the type
C r 1 9 - N i 9 had the best resistance 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 tions.
Figure 7 shows the effect of
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 s e e n 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
promotes the disbonding. However,
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
time variations. The d i s b o n d i n g for
PAW specimens with type Cr19-Ni9
a f t e r 3 0 h w a s less t h a n 5 % , t h e
P A W specimens of type C r 2 5 - N i 1 3
had 1 5 to 2 0 % cracks and the S A W
specimens had 60 to 7 0 % cracks.
Autoclave Gas Charging
Figure 8 s h o w s the results of 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 b e s e e n t h a t S A W c l a d w e l d
metal generally cracks more extensively than PAW clad metal. The
S A W s p e c i m e n s h a d 18 to 2 0 %
cracks, whereas the P A W specimens
type C r 2 5 - N i 1 3 had o n l y 8 % cracks
and type C r 1 9 N i 9 had 2 to 4 %
cracks. The clad metal type Cr25N i 1 3 resisted d i s b o n d i n g 3 to 4
times better w i t h the P A W process
c o m p a r e d to the S A W process. The
PAW clad metal type 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 proximately matching the SAW clad
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
overlaid
metal
Side view
7 mm
weld bond
base metal
1 mm
10 mm
2 mm
Bottom view
10 m m
Electrolite
Fig. 4 — Shape of the specimen for cathodic charging test.
508-s ! N O V E M B E R
1993
Fig. 5 — Setting of the specimen in the
electrolyte.
manual cladded weld metal
second cladded layer
first cladded layer
" 7 mm,:
Table 4 — Effect of Charging Conditions on
Metallic Disbonding
Current
Fig. 6 — Shape of the specimen for autoclave gas charging test.
to metallic disbonding.
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 P A W
specimen has more of these places than the
S A W specimen. The content of Cr and Ni on
the fracture surface is - 1 2 % Cr and - 5 % N i
with the P A W process and ~11 % Cr and - 6 %
N i w i t h the S A W process. The distribution of
these alloying elements (Fig. 11) and data of
Table 5 s h o w that these cracks occurred in
the transition z o n e . In the case of the P A W
process, the crack is located inside the carbide layer (-15 p m from the fusion line) and
in the case of the S A W process, it is located
outside the carbide layer ( - 5 5 - 6 0 p m f r o m
the fusion line). The transgranular type of
cracking is, presumably, connected with the
peculiarities of cathodic charging.
Characteristics of Metallic Disbonding
Typical metallic disbonding characteristics are s h o w n in Fig. 9 for c a t h o d i c
a n d a u t o c l a v e gas c h a r g i n g . It is n o t e w o r t h y that cracks have a tendency to
o c c u r at t h e c e n t e r o f a s p e c i m e n a n d
a r e l o c a t e d in t h e t r a n s i t i o n z o n e b e t w e e n t h e s t a i n l e s s steel w e l d m e t a l
c l a d d i n g a n d t h e base m e t a l . A c c o r d i n g
to m a n y investigations (Refs. 5, 8, 9), t h e
m e t a l l i c d i s b o n d i n g is a r e s u l t o f t w o
t y p e s o f c r a c k s : 1) c r a c k s o c c u r r i n g i n
the c a r b i d e layer at the i n t e r f a c e ; a n d 2)
cracks o c c u r r i n g a l o n g coarse austenite
grain b o u n d a r i e s near the w e l d interf a c e . In t h e p r e s e n t s t u d y ( F i g . 9 ) , t h e
c r a c k s m i c r o s c o p i c a l l y w e r e l o c a t e d in
the c a r b i d e layer at the interface, as w e l l
as a l o n g t h e a u s t e n i t e g r a i n b o u n d a r i e s
w i t h i n the clad metal a d j a c e n t to the
base m e t a l . T h e r e w a s a c o m b i n a t i o n o f
these t w o types o f cracks, a n d a transg r a n u l a r t y p e of crack w a s also f o u n d in
the i n v e s t i g a t i o n .
The results of cracking tests show that cathodic charging pretty well reproduces characteristics o f metallic d i s b o n d i n g 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-
Figure 10 shows examples of the fracture
surface for P A W and S A W samples (cathodic
Testing Time (h)
mA/m2
12
18
24
30
0.2
0.1
0.05
on *
•a*
•n *
•a*
•3 *
•a •
ca*
ca •
o •
• •
•a*
• D *
O D * — n o 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
Grain(a)
Cr25Ni13strip (SAW)
100-500
200-250
70-750
150-200
50-400
120-160
Cr25Ni13wire (PAW)
Cr19Ni9wire (PAW)
(cm)
Carbide
Layer(b)
Transition
Zone
20-40
80-100
10-20
20-30
15— 25
25—35
(a) Numerator - min and max dimention o f grains near the
weld b o n d in the overlaid melal.
Denominator - mean value.
(b) Mean value.
bonding than by using autoclave gas charging.
Discussion
Grain Morphology
G r a i n s in the y phase c l o s e to the i n terface, w h e r e the d i s b o n d i n g occurs,
w e r e f o u n d t o be p l a n a r a n d c o a r s e in
Disbonding, %
20
i 1,
Cr25Ni13 SAW
16
^ r
A1
/
A\
j
^^^^-^CrtSNMS
I
PAW
8
<
I
12,
1
—••
.
I|
flCM9Ni9
PAW
4
Testing time, hours
Fig. 7 — Effect of charging time on the metallic disbonding
density i = 0.2 mA/m2).
(current
0
Cr25Ni13
SAW
Fig.8. — Resistance to the disbonding after
autoclave gas charging test.
Cr25Ni13
PAW
CM9Ni9
PAW
Type of cladded weld metal
W E L D I N G RESEARCH S U P P L E M E N T I 509-s
tv OVER L A I D .
/ METAL
•"
. ?'
N
x:-i
kf
r f i ^ W
'
i • Cw ' ,<>
I
v
v
**~-
%•.••;: B
A
S
E
METAL
OVERLAID
METAL
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.
510-s I NOVEMBER 1993
A — PAW; B — SAW. 300X.
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
• OVERLAID!
METAL
\'M
HAZ
icrij
..
ifeAWW 1
1
|7T|
••
.
0
Hy^Vw****^
20
40
60
DISTANCE (Mm)
•Hi
.
. . .... - .
B — PAW specimen, type
Cr19-Ni9.
• OVERLAID
j METAL
BASE
METAL
.
:
y
BjQHSj >j f^SSKk\
•:.
•
:
.
•
}
5
.
.
'
:
*
"
-
.
.
•
•
•
•
' ; • • • •
|CrjH| A
40
^^^^^^^f^^^
^
60
DISTANCE ( u m )
,
•
'
•
•
'• "i-fjj t \
C — SAW specimen, type
Cr25-Nil3.
WELDING RESEARCH SUPPLEMENT I 511-s
OVERLAID
METAL
OVERLAID
^ / ^ i K ^ v * •/•;-: • '4:;":.; #.K
vt ryy:iy>-;ymiym.^:f•
1
BASE
HH
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.
a
OVERLAID •;• >
METAL
• t ~
mm A
yyf-
BASE
MFTAL S > V - . » • j .: * & $ & & . .?
512-s I N O V E M B E R 1993
^
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 w i t h 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 w i t h the clad metal type
Cr19-Ni9, w h i c h 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, will 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,
silicon and carbon influence cracking
along the grain boundary. Also, the grain
orientation toward stresses at the interface when cooling down will not be the
same.
The formation mechanism of ycoarse
grain boundaries can be stated as f o l lows (Ref. 5): the austenite grains at the
fusion boundary in the heat-affected
zone (HAZ) formed in the 8 —> y transformation during cooling 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 liquid —> liquid + 8 —> liquid + 8+ y during solidification. 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 i q 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 solidification 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 liquidus, the solidus and
the 8 + y —> y boundary lines fall nearly
monotonously together w i t h 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-
TZT diagram in PAW cladded metal type Cr26Ni1 3
-Transition zone in overlaid metal
L
I'lOO
S_ 1400
Q-1300
20
Distance
(yu m]
B
TZT diagram in PAW cladded metal type Cr1 9Ni9
-Transition zone in overlaid metal -
L
20
Distance
ISOO
IZOO
30
[um]
TZT diagram in SAW cl dd id metal type Cr25Ni13
, I
-*—-Transition zone n overla
O
1S00 -
L
======-
-
cT
- Ihoo
cT+
CX
£ 1300
X
-•- Carbide layer-*-!
20
JI0
B0
Distance
no
IZOO
"00
Fig. 13—An example
of TZT diagram. A —
SAW specimen, type
Cr25-Nil3; B — PAW
specimen, type Cr 19Ni9;C —SAW specimen, type Cr25-Nil3.
[u m]
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
% Ferrite -. >
20
\
Austenite
Cr25M13 strip
16
/
/
£r25Ni13wire
^
A+ M
12
'
Martensite
Fig. 14 — Schaeffler
diagram. 1) PAW,
Cr25-Ni13, first layer
deposit; 2) PAW, Cr19Ni9, first layer deposit;
3) SAW, Cr25-NH3,
first layer deposit.
/
0
/
C£l9NBwire
\ .
A+F
^^
8
^
l /
/
sA +M+F
4
\
F
•
0 M
^^^
\ 2,2SC*iMoplate
M+F
\
\
Ferrite
\
Chromium Equivalent = % Cr + % Mo + 1,5 x % Si + 0.5 x % Cb
lidification. 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 A 3
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).
contraction of melted clad metal with
the base metal irrespective of the cooling rate. It is thought that the w i d t h 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, w h i c h 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 composition 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% ferThe width of the transition zone afrite. Metallographic analyses (Fig. 12)
fects the ratio of the y coarse grain bealso show the austenite structure with a
cause the coarse grain is formed inside
small quantity of ferrite in the case of
the transition zone. It was shown (Ref.
the PAW specimens. The ferrite in the
5) that the ratio of the y coarse grain has
austenitic stainless clad metal also afan increasing linear correlation with the
fects the grain size
by changing the
solidification pro'C
cess (Refs. 18, 19).
The 8 ferrite estabo
7100
lishes new solidification centers in
front of growing
_
C
columnar grains.
Due to this effect,
the austenite grains
900
o
o
are going to be
smaller. In comX
paring the size of
austenite grains of
disbondinCL.
PAW clad metal
o
o V x
X
XX
X
no < •for types Cr25Ni13 and Cr19Ni9 (Table 5), it
o
O
'
'
can be seen that
austenite
grains
close
to
the
inter1
Q
1
1
1
500
10
so
f]t J K / ? S
7
5
face in the second
t 15
case are smaller,
Fig. L
- Influence of PWHT on the metallic
disbonding.
which
depends,
514-s I N O V E M B E R 1993
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 austeniteferrite microstructure was obtained in
the cladding. The dilution 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 2 4 . 2 % Cr
and 13.8% N i . Carbon content in the
first layer was 0.072% with PAW and
0.094% w i t h 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 with 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 w i t h austenite-martensite
structure adjacent to the fusion line, and
austenite or austenite-ferrite clad metal.
This observation follows 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 solidification.
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 ( - 6 0 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-
As - welded
After PWHT
450
450
Overlaid metal
:
410
Overlaid metal
Base meta!
Base metal
L
410
I
370
iW
330
370
H,
I
l
330
I
on !
/
,
290
i J--D--C
250
r^'~\J^*i
210
'
290
j^-0"-!]
)
"K^-ri
r*^
d
P^J y f > - 0
250
Y.>
H rO&
yp—o~ -o
s5"M l ^
170
Fusion Ii
130
0,3
0,1 0
0,1
0,3
0,5
0,3
DISTANCE, mm
0,1
0
0,1
0,3
8
2I0
170
130
0,5
DISTANCE, mm
B
A
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 falling 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 decarbonization 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 condition 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 condition. 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 w i t h the martensitic structure can include the carbide layer after
PWHT if using welding processes with
a low dilution 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 w e l d i n g . 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 condition 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) T h e increasing resistance to m e t a l lic d i s b o n d i n g in the case o f t h e plasma
arc w e l d i n g p r o c e s s c a n b e e x p l a i n e d
by the f a v o r a b l e characteristics of this
s u r f a c i n g process, w h i c h i n c l u d e a l o w e r
heat e n e r g y a n d p e n e t r a t i o n , h i g h e r
c o o l i n g rates, s h o r t e r t i m e f o r t h e s o l i d
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 ,
a n d l o w e r m i x i n g a n d d i l u t i o n . A l l these
factors c o n t r i b u t e to the fine grains of
a u s t e n i t e s t r u c t u r e a d j a c e n t t o the w e l d
i n t e r f a c e , t h e least l e n g t h o f y c o a r s e
grain b o u n d a r y parallel to the fusion
line, and the smallest w i d t h and hardness of the t r a n s i t i o n z o n e , i n c l u d i n g the
martensitic layer (as-welded) and the
c a r b i d e layer (after P W H T ) .
3) T h e results o f t h e c a t h o d i c c h a r g i n g test p r e t t y w e l l r e p r o d u c e d t h e results o f t h e a u t o c l a v e gas p h a s e c h a r g i n g test. T h e c r a c k s a l o n g t h e g r a i n
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
t h e a u s t e n i t e stainless c l a d m e t a l c l o s e
to the base metal in o n e case, a n d transg r a n u l a r f a i l u r e d e v e l o p e d in the t r a n s i t i o n z o n e w i t h the d e c o m p o s i t e m a r t e n sitic structure in a n o t h e r case.
References
1. Steklov, O. I., Alexeev, A. V., and
Alexandrov, O . A. 1988. Disbonding of
austenitic stainless clad steel pressure vessels
containing
hydrogen.
TslNTlKhlMNEFTEMASh, Moscow, pp. 1-24.
2. Technical report of weld overlay disbonding. Symposium on Heavy W a l l Pressure Vessel. ATB, Moscow, 1985. 1(Q): 1-7.
3 . 0 h n i s h i , X., Fuji, A. 1984. Effect of strip
overlay conditions 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 w e l d i n g . 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, N e w York, p. 3 3 1 339.
9. Naito, K., Okada, H., and Watanabe,
J. 1980. Study on hydrogen embrittlement of
pressure vessels overlaid with stainless steel.
Hydrogen embrittlement of transition zone
between weld overlay and base metal. Pressure Engineering, 18(5): 3 9 - 4 6 .
10. O h n i s h i , X., Chiba, R., and W a t a n abe, J. 1985. Hydrogen induced disbonding
of stainless steel overlay w e l d . Symposium
on Heavy W a l 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 w e l d 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 w e l d i n g 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
trhlin 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.
Critical evaluation of constitution of
chromium-iron-nickel system. International
Metals Review 1 : 2 1 - 3 8 .
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 N o : 1558596, December.
Method of cladding.
22. Z e m z i n , V. N. 1966. W e l d e d 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 CONFERENCE PROCEEDINGS
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 stateof-the-art developments in the underwater industry including w e l d i n g equipment and processes,
mechanical and internal w e l d properties, maintenance and inspection procedures, and w e l d i n g
applications in shallow and deep water. (Softbound) Code: CP-391 List: $50.00 AWS Members: $37.50
To order, write or telephone: Order Department, American Welding Society, 550 N.W. LeJeune Road, P.O.
Box 351040, Miami, FL 33135, 1-800-334-9353, or 1-305-443-9353, Ext. 280 (Outside Continental USA).
Non-AWS members must prepay or have company purchase order.
516-s I NOVEMBER 1993