The Weldability of Nitrogen-Containing
Austenitic Stainless Steel:
Part II—Porosity, Cracking and Creep Properties
Enhanced nitrogen enrichment is conducive to improved cracking resistance
and improved strength characteristics in austenitic stainless steel weld metals
BY T. OGAWA, K. SUZUKI, AND T. ZAIZEN
ABSTRACT. Porosity in an austenitic stainless steel weld, due to its nitrogen content being near the solubility limit, can be
curbed by adding elements that enhance
the solubility of nitrogen in the weld
metal; an alternative is to weld in an
atmosphere containing some amount of
oxidizing gas. There is no deterioration in
the resistance to hot cracking in austenitic
stainless steel weld metal enriched with
nitrogen as long as it is maintained in the
austenite-delta ferrite dual phase. Nitrogen exhibits a substantially favorable
effect on the resistance to hot cracking in
fully austenitic stainless steel weld metal.
In tests at relatively high temperatures,
weld metal enriched by nitrogen proved
to be stronger than the base metal,
presumably due to precipitation hardening in addition to solid solution strengthening. The creep properties, especially
the ductility, of Type 308 stainless steel
weld metal at 550 and 600°C (1022 and
1112°F), could be improved significantly
by nitrogen coexisting with some amount
of vanadium. At very high temperatures
(about 900 to 1,000°C, i.e., 1652 to
1832°F), the strength of a nitrogenenriched austenitic stainless steel weld
was approximately double that of conventional Type 308 stainless steel welds,
and its creep strength was also excellent.
Introduction
It has been well known that nitrogen
can be employed to increase the high
temperature strength of austenitic stainless steel (Ref. 1). Nitrogen, which is an
austenite stabilizer like nickel, has the
advantage of being much less expensive
than nickel. When reducing the amount
of nickel used is the main objective, a
fairly large amount of nitrogen (0.03% to
about 0.50%) is used.
Intergranular corrosion, intergranular
stress corrosion cracking, pitting and
crevice corrosion can occur when austenitic stainless steel is used for welded
structures in the corrosive environments
encountered in the chemical and nuclear
industries. In general, one of the most
important causes for these is attributed to
the sensitization arising from the formation of a chromium-depleted zone by the
precipitation of chromium carbides
Cr23C6 at the grain boundaries (Ref. 2).
In order to prevent sensitization in the
welds, the carbon content in the material
must be decreased, and stabilization by
niobium and titanium can be beneficial
(Ref. 3). Lowering the carbon content,
however, results in a decrease in material
strength, especially in the yield strength
employed in design stress. Therefore,
low carbon material called for in the
design and fabrication codes for pressure
vessels has to be enriched with nitrogen.
Based on a paper presented at the 63rd
In the use of cryogenics such as for
Annual A WS Convention in Kansas City, Misfusion power, the demand for austenitic
souri, during April 25-30, 1982.
stainless steel as a structural material has
T. OGAWA and K. SUZUKI are Senior
been increasing in recent years. As the
Researchers and T. ZAIZEN is General Managmaterial used in fusion power equipment
er, Steel Materials Laboratory III, Products
Research & Development Laboratories, Nip- is in rather heavy sections, improving the
pon Steel Corporation, Fuchinobe, Sagamiha- strength of the material is a major concern. The improvement of the material
ra, Kanagawa-Ken, Japan.
strength through the addition of nitrogen
can be far greater at a cryogenic temperature than at an elevated one (Ref. 4).
The creep rupture strength of austenitic stainless steel weld metal is often a little
lower than that of the base metal. This
tendency is more conspicuous when the
applied stress is lower and rupture time is
longer. Weld metal rupture elongation
also, when compared to that of the base
metal, is generally lower.
With Type 304 stainless steel which is
to be used in liquid-metal fast breeder
reactors, it is considered that the elongation of 10" hour (h) rupture at 550°C
(1022°F) should be a value in excess of
15%. This has been a troublesome problem with standard Type 308 stainless steel
weld metal. For this reason, CRE Type
308 stainless steel weld metal has been
developed that meets the end-use
requirements by controlling the residual
elements such as phosphorus, titanium
and boron (Ref. 5). In this paper a procedure that assures appropriate creep rupture strength and sufficient rupture ductility is presented. These are accomplished
through the strict control of nitrogen
content together with the addition of
nominally 0.10% vanadium.
Nitrogen-containing austenitic stainless
steel has been employed in high-temperature environments (about 900 to
1,000°C or 1652 to 1832°F) such as the
radiant tubes of heating furnaces and
thermal reactors for automobile exhaust
gas control.
Part I to this paper was published in
1982 (Ref. 6); it described the effect of
nitrogen on pitting corrosion resistance
of the welds in the weldability of nitrogen-containing austenitic stainless steel.
Part II, which is presented here, describes
the effect of nitrogen on porosity occur-
WELDING RESEARCH SUPPLEMENT 1213-s
Table 1—Chemical Compositions of Base and Filler Metals, Wt-%
Type
material
Base metal
Filler Metal
(a)
<b)
Stainless
steel(a>
C
Si
Mn
P
S
Ni
Cr
Type 304
304LN
316LN
Type 310 S
Steel A
Steel B
0.055
0.013
0.010
0.060
0.025
0.095
0.49
0.57
0.55
0.87
0.89
1.95
0.89
0.89
0.95
0.92
1.46
1.59
0.030
0.027
0.028
0.025
0.023
0.025
0.005
0.009
0.006
0.003
0.005
0.004
8.66
8.95
11.14
19.65
12.90
13.60
18.23
18.50
16.76
24.90
24.50
24.20
Type 308
308V
316LN
Steel A
Steel B
0.050
0.052
0.013
0.030
0.108
0.33
0.18
0.42
0.77
1.99
1.89
1.78
2.19
1.43
1.63
0.020
0.030
0.014
0.021
0.031
0.004
0.003
0.006
0.003
0.002
9.81
10.23
12.50
12.20
13.31
19.92
20.11
19.01
24.15
24.13
Mo
Nb
V
—
—
—
—
-
—
-
—
-
2.10
—
0.83
0.75
-
2.40
0.94
0.94
Note' b)
N
0.026
0.075
0.066
0.027
0.345
0.250
I
j
|
12, and 20 mm thick
J
0.032
0.016
0.054
0.346
0.241
0.120
-
12, and 20 mm thick
30, and 150 mm thick
f
\
1.6
2.0
2.4
3.2
4.0
mm <t>
mm 0
mm 4>
mm a,
mm 0
Types 304LN and 316N base metals, and 308V and 316LN filler metals are classified "Tentative.'
tj) designates electrode diameter.
Table 2—Chemical Compositions of Weld Metals Made Using GTAW and Experimental Filler Metals, Wt-%<a) (b) < c)
metal
C
Si
Mn
P
S
Ni
Cr
Nb
N
a
b
c
d
e
f
0.016
0.016
0.016
0.017
0.015
0.015
0.33
0.34
0.39
0.33
0.30
0.37
1.50
1.48
1.54
1.21
1.25
1.03
0.012
0.011
0.014
0.013
0.014
0.012
0.004
0.004
0.005
0.003
0.003
0.002
9.75
9.90
9.40
9.34
9.14
8.94
17.50
18.61
18.66
18.90
18.77
19.03
0.25
0.24
0.25
0.23
0.26
0.24
0.0210
0.0415
0.0620
0.0970
0.1001
0.1118
(a)
Joint design is as shown schematically below:
' G T A W conditions as follows: 250 A; 10 c m / m i n { = 1.67 m m / s or 3.9 ipm); argon shielding.
' Tensile specimen: 10 m m (0.39 in.) diameter; 50 m m (1.97 in.) gage length.
are shown in Table 1, and those of
experimental filler metals and their weld
metals are listed in Tables 2 and 3.
Evaluation of welding hot cracking susceptibility was performed mainly by the
Varestraint test (Ref. 7). Specimen sections were 5.0 mm (0.04 in.) thick and
were machined from steel plates that
were 12, 30, 40 and 150 mm (0.47, 1.18,
1.57 and 5.91 in.) thick. Welding was
conducted using the gas tungsten arc
welding process (CTAW), and the welding conditions were set at 70 A, 16 V, and
7.5 cm/min (i.e., 1.25 mm/s or 2.95 ipm).
The shielding gas was mainly pure argon
gas, but argon and nitrogen gas mixtures
at specified ratios were also employed.
CMS
3
rence and hot cracking susceptibility, as
related to welding techniques. Part II also
describes the effect of nitrogen on the
tensile strength and creep properties of
the welds that are related to the abovementioned uses and areas of applicabilityExperimental Procedures
The chemical compositions of commercial steel plate and filler metals used
The CTAW welds for the tensile test
were made in 45 deg groove angles with
12 mm (0.47 in.) root openings between
12 or 20 mm (0.47 or 0.79 in.) thick base
metal plates. The completed welds were
sectioned to prepare all-weld metal tensile specimens parallel to, and completed
welds were sectioned to prepare weld
joint tensile specimens transverse to the
welding direction. The tensile specimens
had primarily 50 mm (1.97 in.) gage
lengths by 10 mm (0.39 in.) diameter
sections, but some testing was done with
30 mm (1.18 in.) gage lengths by 6 mm
(0.24 in.) diameter sections.
For the creep tests of Type 308 weld
metals, all-weld-metal specimens were
Table 3—Chemical Composition of Experimental Filler Metals and Weld Metals Deposited by SAW Wt-%<a> <b>
Type
material
Stainless steel
t8)
lb)
Si
Mn
P
s
Ni
Cr
V
N
308
No. 1
No. 2
0.051
0.030
0.52
0.36
1.59
1.90
0.011
0.024
0.005
0.003
10.06
9.56
20.54
20.00
0.006
0.053
0.0100
0.0472
308 V
No. 3
No. 4
0.054
0.061
0.12
0.10
1.63
1.87
0.013
0.052
0.005
0.004
10.24
10.85
20.26
20.31
0.094
0.107
0.00950
0.0219
308 V
308
308 V
308
308 V
308
308
A
B
C
D
E
F
G
0.049
0.048
0.057
0.045
0.052
0.035
0.050
0.37
0.80
0.49
0.42
0.36
0.56
0.70
1.44
1.76
1.69
1.72
1.42
1.16
1.05
0.020
0.017
0.017
0.016
0.031
0.018
0.030
0.003
0.005
0.005
0.005
0.004
0.004
0.006
10.26
10.36
10.16
10.39
10.20
9.47
9.00
20.00
20.09
19.71
19.92
19.54
20.21
18.37
0.088
0.006
0.088
0.006
0.079
0.044
0.048
0.0094
0.0111
0.0115
0.0136
0.0252
0.0374
0.0784
Weld metal
Filler metal
C
^-30°V
loint design is as shown schematically at right:
Welding carrif d out with CaO-SiOj-MgO flux at 200-600 A, 26-30 V and 50-220 cm/min (= 8.3 36.7 mm/s. 19.7-86.6 ipm).
f I
i 40
/
-IOQ/I
1
V/~j
^-v
LL
\
cut from the center parts of the welded
butt joints made by GTAW and the
shielded metal arc welding process
(SMAW) in 25 mm (0.98 in.) thick base
metal and from butt joints in 40 mm (1.57
in.) thick base metal welded by the submerged arc welding process (SAW) in
both longitudinal and transverse directions. The creep test specimens had 30
mm (1,18 in.) gage lengths by 6 mm (0.24
in.) diameter sections, except for 15 mm
(0.59 in.) gage lengths in the welds made
using the experimental heat filler metals.
The creep test specimens of steel B with
30 mm (1.18 in.) gage lengths by 6 mm
(0.24 in.) diameter sections were transversely cut from the welded butt joints
made with CTAW in 12 mm (0.47 in.)
thick base metal.
The cracks encountered during Varestraint testing were measured on the
surface of each weld under X30 magnification, and dye penetrant inspection was
performed on the side bend test of the
welds. The microstructures and fracture
surfaces in the welds were observed by
optical microscope, electron probe
microanalysis (EPMA), scanning electron
microscopy (SEM) and transmission electron microscope (TEM). The delta-ferrite
measurements were made using a Ferritescope.
Results and Discussion
Welding Technique
In the welding of nitrogen-containing
austenitic stainless steel, t w o main points
should be kept in mind when considering
the increase in the nitrogen content in the
welds:
0.4
STEEL-A
-0.3
E
1. There is the possibility of porosity
occurring.
2. An increase in hot cracking susceptibility can occur.
Occurrence of Porosity
There is generally an increase in the
nitrogen content in the weld metal when
nitrogen gas is added to argon shielding
gas in GTAW. The test results on the
welds of steel A, steel B, and Type 304
stainless steel can be seen in Fig. 1. The
amount of nitrogen that can be added to
the welds depends upon the alloy compositions of the weld metals due to
differences in the nitrogen solubility in the
alloys on welding solidification. Porosity
can occur when the nitrogen content
involved is in the area of the solubility
limit of the weld metal.
Figure 2 shows the weld surface
appearances of bead-on-plate welding
without filler metals on Types 304 and
310S stainless steel, using an Ar + 25% N 2
shielding gas mixture. Extensive concavity
and porosity occurred in the weld bead
on Type 304 stainless steel, but there was
absolutely no sign of such an occurrence
with Type 310S stainless steel which contains a relatively large amount of chromium (Ref. 6).
The solubility limit of nitrogen in molten metal in the case of arc melting is
reported to be considerably higher than
the level of equilibrium (Refs. 8, 9, 10).
There are, however, different types of
arc welding such as GTAW using an inert
gas and SAW relying mainly on slag-metal
reaction. The research related to nitrogen
solubility is not yet sufficiently well established to be generally applied to welds
made with these processes. The nitrogen
contents in weld metals made using
2 0.2
A STEEL-A
matching composition wire 1.6?
0.1
(0.347%N)
O STEEL-B
matching composition wire 1.6?
(0.269 %N)
type 304 xx
X type 304
type 308 wire, 2.0 /
XX
XX
0 2
5
10
15
N 2 added in shielding gas A r ( V o l . % ) Fig. 1 — Increases of nitrogen in weld metal deposited by GTAW using argon-nitrogen shielding gas.
Welding carried out at 250 A, 15 V, 15 cmAmin (5.9 ipm), 15 Lymin (31.8 cfh). 12 mm (0.47 in.) thick
base metal, using a V-groove weld
Fig. 2 — Porosity in bead-on-plate welds made
with GTA W: A — Type 303 stainless steel
welded with Ar + 25% N2 shielding gas at 70 A
and 7.5 cmymin (3 ipm); B — Type 310 stainless
steel welded with Ar + 25% N2 shielding gas
WELDING RESEARCH SUPPLEMENT 1215-s
E
E
LU
a
O
decrease it (Ref. 11). Consequently, the
porosity occurring through the use of a
large amount of nitrogen near the solubility limit can be remedied by an appropriate increase in, or by the addition of, such
solubility-increasing elements.
On the other hand, the solubility limit
of nitrogen in the weld metal can be
raised by adding a small amount of oxygen gas to the argon shielding gas. As for
welding with an oxidizing atmosphere, it
is reported in the literature that an
increase in the nitrogen solubility occurs
in the presence of nitric oxide (Ref. 10) in
the atmosphere and due to surface activation (Ref. 12) by oxygen on the surface
of molten metal. These principles should
be applicable in SMAW and SAW, where
welding solidification occurs under the
slag which is comprised of various
oxides.
high Si steel
STEEL-B
c\J
-t-j
CU
E
CD
x
o
<
uu
M-
CO
c
J»
Ln
o
z
CtJ
o
a.
O
"TO
co
o
>
LU
Q
CNI
c
X
O
QC
<
nj
cu
LU
CA)
UJ
QC
1_
Ct3
>
0
>
2
a.
o
5
^°
>
LU
L.
0} • Q
o
tr
<
15
1
«o
>
A STEEL-A
OSTEEL-B
±type304
• t y p e 310S
Eco
a
10
N added in shielding gas A r ( v o l . % )
o ^-- a
CD
<u + J CNI
Oi
—i
LU
LU
2
©**
5
fig. J - Influence of nitrogen in shielding gas on the hot cracking susceptibility of austenitic stainless
steel weld metal. GTAW carried out at 70 A, 7.5 cm7min (3 ipm) with a 1.2% augmented strain and
5.0 mm (0.196 in.) thickness
iZ
LU
a.
o
GTAW with pure argon shielding gas in
steels A and B, as shown in Fig. 1, were
approximately 0.27% and 0.17%, respectively, with fluctuations and scattered
values lower than that of the base metal.
In arc welding, there is considerable
variation in the nitrogen solubility,
depending on the alloy compositions.
Elements such as Ti, Zr, Nb, V, Cr, M o
and Mn increase the solubility of nitrogen
in the molten metal, while C, Si and Ni,
Table 4 shows the oxygen and nitrogen contents in weld metals made in
steels A and B when a small amount of
oxygen gas has been added to the argon
shielding gas for GTAW and gas metal arc
welding (GMAW). Compared to the use
of pure argon shielding gas, a fairly large
amount of nitrogen can be enriched
without any occurrence of porosity or
concavity. However, from the viewpoint
of welding techniques, adding oxygen
gas, nitrogen gas or carbon dioxide to the
shielding gas while maintaining strict control of the nitrogen content in the weld
metal, is quite difficult. With GTAW,
electrode erosion resulted in excessive
sputtering and arc instability.
It is also reported that, in connection
with the nitrogen solubility in the liquid
(determined by the molten metal temperature), porosity in welds can be
decreased by increasing the arc lengths
on GTA welding (Ref. 13). It can be said,
in any case, that the occurrence of porosity because of nitrogen can be kept
under control through careful consideration of the above-mentioned points and
through diligent research on how to
enhance the nitrogen solubility.
Table 4—Nitrogen and Oxygen Contents in Weld Metals Made in Steels A and B Using GTAW and G M A W with Different Shielding Gases and
Filler Metals of Matching Compositions
H
Z
Process(a)
Steel
designation
Shielding gas(b)
Nitrogen
content, wt-%
Oxygen
content, wt-%
LU
GTAW
A
A
A
B
B
Ar at 15L/min
A r l - 2% 0 2 at 15L/min
Ar I- 5% C 0 2 at 15L/min
Ar at 15 L/min
Ar I- 2% 0 2
0.27
0.406
0.410
0.172
0.278
0.0041
0.0373
0.0495
0.0030
0.0319
GMAW
A
A
Ar + 2°<, 0 2 at 25L/min
Ar I- 2"o N 2 + 1°o 0 2 at
25L/min
0.393
0.329
0.0487
0.0352
Base metal
A
B
0.351
0.261
0.0047
0.0040
D.
o
—I
LU
>
UJ
a
x
o
cr
<
LU
(/>
LU
cr
1
1
-
Welding conditions as follows: G T A W - 2 5 0 A, 12 V. 10 c m / m i n ( = 1.67 m m / s . 3.94 ipm); G M A W - 3 0 0 A, 30 V. 40 c m / m i n ( = 6.67 m m / s , 15.7 ipm).
15L/min = 31.8 cfh; 25L/min = 53 cfh.
216-s | JULY 1984
Hot Cracking Susceptibility
argon shielding gas during welding with
Type 310S stainless steel, the result is
rather the opposite, showing a tendency
to decrease total crack length. This desirable effect of nitrogen in reducing hot
cracking is more clearly recognized in the
weld metal of steel B when 10% or 15%
nitrogen gas is added.
Figure 4 illustrates the differences in the
total crack lengths on Type 310S stainless
steel occurring for both pure argon
shielding gas and of Ar + 5% N2 shielding
gas mixture, when the bending strain rate
in the Varestraint test has been changed.
Although the crack lengths when using a
specific shielding gas do not vary appreciably with the strain rate, the crack
lengths at all strain rates are shorter by a
substantially constant value for Ar 4- 5%
N2 shielding gas than for argon shielding
gas.
It is expedient to add nitrogen to the
shielding gas during welding (Refs. 14,15)
in order to examine the effect of nitrogen
on hot cracking. Figure 3 shows the
changes in total crack length that
appeared in the weld metal when the
amount of nitrogen gas added to argon
shielding gas was increased from 2 to 5%
and then to 15% by volume during Varestraint tests on commercial Types 304
and 310S stainless steel as well as steels A
and B. Figure 3 also includes (at its bottom) variations in measured values of the
amounts of delta ferrite (hereafter
referred to as cSF) in each weld.
When the amount of nitrogen added
to Type 304 stainless steel and steel A
and B is increased, the amount of 5F in
welds decreases or disappears, resulting
in an increase in total crack length,
depending heavily on the alloy compositions of the steel tested. In steel B that
contains 2.0% silicon, when <SF in the weld
disappears by the addition of 5% by
volume nitrogen gas, there is a maximum
increase in the total crack length; the
increase is almost three times that which
occurs in Type 310S stainless steel. There
are already several excellent papers
reporting on both the beneficial effects
of <5F and the detrimental effects of silicon
in fully austenitic weld metal on hot
cracking (Refs. 16, 17).
Considerable effort has been expended in the development of nitrogen-added, low-carbon austenitic stainless steels
that are resistant to stress corrosion
cracking; these include stainless steels
such as Types 304LN, 316LN and 347LN
for use in the core spray and recirculation
bypass piping system of boiling water
nuclear reactors, and some of these
steels are already in use.
Figure 5 depicts the results of Varestraint tests on specimens from 30, 40,
and 150 mm (1.18,1.57 and 5.91 in.) thick
plates of these steels. They are all almost
at the same level and demonstrate an
excellent resistance to hot cracking that is
as high as that of commercial Type 304
stainless steel. This is due to the fact that
chromium equivalents and nickel equivalents in the alloy compositions of all of
these autogenous weld metals are regulated to help retain a small amount of
SF.
One of the reasons for the shorter
crack lengths is probably believed to be
that nitrogen can form some stable, highmelting nitride with silicon segregating at
grain boundaries, which would otherwise
increase weld cracking. These nitrides
can also have the desirable effect of
reducing the grain size on solidification
through the stable-nitride-formation, increasing total length of grain boundaries
and thus curbing the detrimental segregation. According to diligent examination of
solidification microstructure, the austenite
grains surely become somewhat fine by
the addition of nitrogen to the weld
metals, but the assumed high-melting
constituents
have
not
yet
been
detected.
In Type 310S stainless steel, the weld
metal is generally composed of austenite
(7) single phase, even without the addition of nitrogen. If an increased amount
of nitrogen gas is, in turn, added to the
—
detrimental effect on weld cracking of
the (7 4- 5)-phase weld metal, 5-phase
being unstable, due to the transformation
to fully 7-phase by the strong ability of
nitrogen as an austenite stabilizer. However, if nitrogen is added to the weld
metal while still maintaining the (7 4- 8)
dual phase, the nitrogen does not
degrade the resistance to weld cracking,
but works to improve it.
Figure 6 shows the submerged arc
welding (SAW) of Type 316LN of 150 mm
(5.91 in.) thick Type 316N stainless steel; it
also shows a typical weld macrostructure. Neither solidification cracking nor
micro-fissures were found, and the
results of the side bend tests in these
The addition of nitrogen could have a
fully a u s t e n i t i c - *
Commercial type 31 OS
(25Cr-20Ni)
f5.0tX40wX350* mm
70A-7.5cm/min
I GTA welding
GTAW shielding gas
i o Ar
• (Ar-t-5%N)
Type310S
-Sb
«Ja3 '
Type316
.
•
„--a--<-\-V~"
g^gS^SSSSSS
-'nr'.IL- ~-r- • •
j
t y p e 304
A55=
L<5b-
o304LN(150mm thick)
• 316LN( 30mm
« )•
...-*-"
-1
aug. s t r a i n 1.2%
70A-7.5cm/min
thick. 5.0mm
2
3
4
5
augmented strain rate
6
(%/sec)
7
8
Fig. 4 — Effect of nitrogen on hot cracking susceptibility in fully austenitic
Type 310S stainless steel weld metal
// \
'
-'-y
Fig. 5 (right) — Results of
austenitic stainless steels
Varetraint
testing on
nitrogen-containing
1
>
'
2
=6P
'
^
<
'
L
3
4
^
A
-
augmented strain, %
WELDING RESEARCH SUPPLEMENT 1217-s
0.015C—0.35Si-1.2Mn—9.0~9.5Ni-17.5~19.0Cr—0.25Nb
500
(72.5)
10
a
[ d e p o s i t e d m e t a l by G T A W
2 5 0 A - 1 2 V - 1 5 c m / m i n , A r shi Iding gas
1.3.2mm d i .
450
(65.3)
a
-c
ti
t e s t t e m p e r a t u r e ; 300°C
400(58.0H-
d e l t a - f e r r i t e of deposited metal ; 3 ~ 4 %
10
'10mm
•=
700
| (101.5)
specimen
diameter
50mm gage l e n g t h
las welded
o
CD
weld metal
y
> *
CB
E
650
(94.3)
base metal
t e s t t e m p e r a t u r e ; R.T.
600
(87.0)
0.04
0.02
0.06
nitrogen content,
0.12
0.10
0.08
%
Fig. 7 —Effect of nitrogen on Types 347L and 347LN stainless steel sheets and weld metals
Table 5—Room Temperature Mechanical Properties of Welds in Steels A and B Made Using
GTAW and Filler Metals of Matching Composition**' (b>(c>
Ultimate tensile
strength, MPa
(1000 psi)
Elongation,
%
Vickers
hardness no.
(10 kg load)
Steel A
814
(118)
34
257
Steel B
794
(115)
32
240
Steel A
804
(114)
Steel B
784
(116)
Steel A
784
(114)
52
185
Steel B
775
(112)
55
210
Material designations
Fig. 6 — Submerged arc welding of 150 mm
(5.9 in.) thick Type 316LN stainless steel: A welding machine; B — weld macrostructure
w e l d s w e r e f a v o r a b l e . Similar f a v o r a b l e
results h a v e b e e n attained w i t h Types
304LN a n d 347LN stainless steel.
Deposited metal
Welded joint
The Strength of Welds
T w o mechanisms o f t h e strengthening
effect of n i t r o g e n are w e l l k n o w n . O n e is
interstitial solid solution strengthening;
t h e o t h e r is p r e c i p i t a t i o n hardening. T h e
relative
effectiveness
of
the
two
strengthening mechanisms d e p e n d s o n
the t e m p e r a t u r e at w h i c h t h e material is
employed.
Effect of Nitrogen on Tensile
Strength
of Welds. Figure 7 illustrates t h e effect of
n i t r o g e n o n the tensile strengths of l o w c a r b o n stainless steel w e l d metals m a d e
b y G T A W at r o o m t e m p e r a t u r e a n d
3 0 0 ° C (572°F). In a d d i t i o n , T a b l e 2 s h o w s
t h e chemical c o m p o s i t i o n s of the w e l d
metals in Fig. 7 m a d e b y C T A W using
e x p e r i m e n t a l filler metals. These filler
218-s|JULY 1984
Base metal
ta)
<w
<cl
loint design is as shown schematically at right:
C T A W conditions as follows: 250 A; 10 c m / m i n ( = 1.67 m m / s , 3.93 ipm).
Tensile specimen: 10 m m (0.39 in.) diameter; 50 m m (1.97 in.) gage length
metals w e r e s w a g e d f r o m the same heats
as the 25 m m (0.98 in.) thick base metals
tested, the strength o f w h i c h is illustrated
as a line o f data in Fig. 7.
N i t r o g e n has b e e n f o u n d t o be m o r e
e f f e c t i v e f o r increasing t h e material
strength of the w e l d metal than o f t h e
base m e t a l . This can p r e s u m a b l y b e
explained b y t h e fairly large d e g r e e o f
H
h-12
p r e c i p i t a t i o n hardening o c c u r r i n g w i t h
multipass w e l d i n g t h e r m a l cycles in t h e
w e l d m e t a l ; this is in a d d i t i o n t o solid
solution
strengthening
which
mainly
w o r k s o n t h e base m e t a l . Some a m o u n t
o f SF, w h i c h is usually i n t r o d u c e d in t h e
austenitic stainless steel w e l d s , w a s also
found to contribute somewhat to increasing their strengths. Table 5 shows
the ultimate tensile strength of welds
made in steels A and B at ambient temperature; Fig. 8 illustrates the tensile
strengths of these steels and Type 304
stainless steel welds up to 1,000°C
(1832°F). According to Fig. 8, the strength
of nitrogen-containing steel welds is
about 1 % MPa (28.4 ksi) at 900°C
(1652°F), almost two times that of the
Type 304 stainless steel weld; it has a
somewhat higher value of approximately
118 MPa (17.1 ksi) at 1,000°C (1832°F).
However, no significant differences
between the nitrogen content of nominally 0.35% and 0.25% in the welds can
be observed. Moreover, even when
more nitrogen was added to the argon
shielding gas during CTAW, no significant
increase in the strength of the t w o stainless steel welds could be found at relatively higher temperatures. Consequently, although it is surely true that nitrogen
exhibits a remarkably great effect on
enhancing the strength, the strength of
the welds (such as those in steels A and B)
containing rather larger amounts of nitrogen is not so sensitive to moderate
changes in nitrogen content.
Creep Properties Improved by Nitrogen and Vanadium Addition to Type 308
Stainless Steel Weld Metal. Figure 9
depicts the creep rupture properties of
the Type 308 stainless steel (deposited by
SAW) weld metal with different nitrogen
contents at 550°C (1022°F). Creep
strength increased significantly with the
nitrogen content.
Figure 10 compares the effects of
nitrogen on the creep rupture elongation
at 31 kg/mm 2 (44,093 psi) at 550°C
(1022°F) between nominally 0.10% vanadium-supplemented and non-supplemented standard Type 308 stainless steel
weld metals by deposited SAW.
Table 3 indicates the chemical compositions of the experimental welding wires
and weld metals of Type 308 stainless
steel together with some welding conditions used in Figs. 9 and 10 where the
letters " A " through " G " correspond to
the weld metals designated by the same
letters in Table 3. The sample for analysis
was cut from near each creep test specimen, where variation of the nitrogen
content was regulated through the flux
employed in SAW.
Rupture elongation in standard Type
308 stainless steel weld metal decreases
detrimentally with the nitrogen content,
but that in vanadium-supplemented weld
metal exhibits a more moderate decrease. The improvement of creep ductility in Type 308 stainless steel weld metal
by an addition of nominally 0.10% vanadium can probably be ascribed to the
precipitation of fine stable vanadiumcarbo-nitrides at the grain boundaries and
STEEL-A,
matching composition wire
l
STEEL-B.
. . . .
...
•
matching composition wire
welded joint
^
G T A
n
W
H
/10mm d i a m e t e r
G.L. ;50mm
- A — type304, type308 w i r e
250A—15cm/min
Ar shielding gas
^
500
o
(72.5)
o
•
%
aa
CD
o\
^
400
| (58.0)
2
\ \
\ \
\\
\\
jf
§
300
S (43.5)
tn
Ic
200
\
A
\\
X
,
a>
^.
\
\\
\
(29.0)
\
*
\
*
100
| (14.5)
*4
\
\
cP
\
D
400
600
800
1000
(752) (1112) (1472) (1832)
temperature, °C(°F)
Fig. 8 —Strength of welded joints in nitrogen-containing austenitic stain/ess steel at elevated
temperatures
in the matrix during multipass welding or,
heterogeneously, on dislocations in the
creep process at elevated temperatures.
The fine precipitates tend to trap and
effectively stop otherwise freely moving
dislocations; this results in the retardation
of climbing and tangling of cell structure
in dislocation distribution (Refs. 18, 19).
Figure 11 shows the microstructures by
TEM of Type 308 stainless steel weld
metal deposited by SAW after creep
rupture at 23 kg/mm 2 (32,714 psi) and
600°C (1112°F). Dislocations are still rather uniformly distributed in the matrix in
vanadium-supplemented weld metal of
Fig. 11B. On the other hand, in the
standard Type 308 stainless steel weld
metal of Fig. 11 A, cell structures of dislocations can be observed to be extensively developed, resulting in a remarkable
decrease of their density in the matrix. In
both microstructures, austenite-ferrite
interface boundaries are heavily decorated with precipitates, which were
determined to be mainly chromium-carbides by electron probe microanalysis
(EPMA).
Several black round compounds found
in Fig. 11 were also confirmed to be
silicon-manganese-aluminum oxide inclusions by EPMA. The embrittlement and
failure of the weld metal during the creep
process is believed to occur due to:
1. The development of uniformly dispersed dislocations to cell structures
through redissolvement of vanadium-carbo-nitrides, otherwise immobilizing dislocations, with creep times.
2. The precipitation and its growth of
chromium-carbo-nitrides
and
sigmaphase at the austenite-ferrite interfaces.
Figure 12 depicts microstructures corresponding presumably to these processes observed by TEM of vanadium-supplemented weld metal after 1,000 and 3,000
hour (h) creep times at 18 kg/mm 2
(25,600 psi) applied stress at 550°C
(1022°F). Dislocations in the matrix
formed a cell structure with some dislocations within solidification cell. At 1,000 h
rupture, the austenite-ferrite boundaries
were quite clearly covered with many
WELDING RESEARCH SUPPLEMENT 1219-s
6.0 ^
6.0 t
9
N
3
5
P
£ (49.8)
h
A
0.0094% N
A
\^
0.0374% N
\
30
(42.7)
70
^
60
\
\
s.
\
A
§ 50
\ v A
550"C(1022T)
,
vanadium added
0.0784% N
V
25
(35.6)
7-weld metal
\
i . 1 . ...1
.
A"
. IM.,1
10 J
10
.
,.!...,
10-
rupture time, hr
Fig. 9—Effect of nitrogen on creep rupture
stainless steel weld metal deposited by SA W
strength
of Type 308
c
_o
« 40
Q.
3
30
(550°C(1022°F)
l31kg/mm2(44093psi)
20
| A ; t y p e 308
I o ; v a n a d i u m - a d d e d t y p e 308
_i
Fig. 10 (right) — Effect of nitrogen on creep ductility of Type 308 stainless
steel weld metal deposited by SA W
i
0.001
i
11 u
I
0.01
i
i
i
i
11
II
I
and sigma-phase precipitates at and adjacent to the austenite-ferrite interfaces in
vanadium-supplemented weld metal was
Fig. 11— TEM microstructures (after creep rupture) of Type 308 stainless steel weld metal
deposited by SAW: A—standard Type 308
weld metal, 23 kgymm2 (32,714 psi), 118 hat
600 °C(1112 °F); B - vanadium added to Type
308 weld metal, 23 kgymm2, 160 h at 600°C
Fig. 12 — Dislocation and precipitation structural changes during creep test of Type 308 weld
metal (with vanadium) deposited by SAW:
A- 1000 h at 550°C (1022°F), 18 kgymm2
(25,602 psi); B-3000
h at 550c'C, 18 kgy
mm2
i
0.1
n i t r o g e n in t h e w e l d m e t a l ,
precipitates with crystal habits, growing
inside of the delta-ferrite area. The
growth rate of chromium-carbo-nitride
220-s|JULY 1984
i
1.0
%
verified by TEM observations to be rather
slower than that in standard Type 308
stainless steel weld metal at longer creep
times.
Based on these results, nominally
0.10% vanadium-supplemented Type
308 filler metal was made on a production scale (1 ton heat), and creep behavior was investigated on the specimens
from the welds made by GTAW, SMAW
and SAW with the filler metal. Figure 13
illustrates the test results, where the number near each datum plotted denotes the
measured rupture elongation value. The
chemical compositions and tensile test
results together with some welding
parameters are indicated in Tables 6 and
7, respectively.
In Fig. 13, the creep strengths of vanadium-supplemented Type 308 stainless
steel weld metals were quite satisfactory
for any welding process when compared
with the limit line of 550°C (1022°F)
minimum time to rupture. This limit was
derived originally from ASME Sec. Ill,
Code Case N-47. The creep elongations
are also sufficiently above the desired
value of 15% at 104 h creep rupture and
are consistent with the results of testing
the above-mentioned experimental weld
metal.
In general, the creep rupture strength
of the weld metal from SMAW and SAW
is rather lower and its creep elongation is
40
(56.9) L
I
^
30
•j» (42.7) S32
30
)39
W.
I
E
20
.£ (28.4)
T
0 34
~^~-~^
•GTAW
10
(14.2)
Fig. 13-Creep
(1022 °F)
c
Si
Mn
P
S
Ni
Cr
V
N
O
" - » * ^
oSMAW(lime-coated)
(%)
SMAW<b»
SAW<a)
50
(MgO-Si02-CaF2)flux\
number '• elongation
I
• 18
3
01
>S_J
weld metal
n40
25
A
' ASAW
Table 6—Chemical Compositions of Type
308 Stainless Steel Weld Metal with
Vanadium Additions as Deposited by SAW,
SMAW and GTAW, Wt-%
6.0 0
test temperature ;550C(1022'F)
\
A S M E - S e c . Ill -Code Case N-47
minimum time t o r u p t u r e
0.060
0.29
2.07
0.033
0.003
10.64
10.13
0.085
0.0323
0.0524
0.055
0.46
1.65
0.035
0.004
9.63
19.07
0.11
0.0152
0.0455
GTAW
0.045
0.17
1.76
0.032
0.002
9.99
20.10
0.12
0.0145
0.0044
<a)
10 3
r u p t u r e time, hr
10'
properties
of Type 308 stainless steel (with vanadium) weld metal at
higher than that from GTAW. The number of welding passes required in the
welding of heavy sections by GTAW is
usually somewhat larger than the number
required with SMAW and SAW. Therefore, the amount of precipitates and
residual strains due to multipass welding
thermal cycles is assumed to become
rather larger with GTAW. Thus, the
material strengthening effect arising from
these causes becomes more remarkable,
resulting in a decrease in the creep ductility. Note the values of ductility around
10 4 h in Fig. 13 (Ref. 24).
There are many more small spherical
SAW Calmed out using M g O - S i 0 2 - C a F j flux al 500 A, 34 V.
and 40 c m / m i n ( = 6.67 m m / s , 15.7 ipm)
(b)
S M A W carried o u ! with lime-coated electrodes.
10 4
550"C
inclusions of less than about 0.5 ix diameter in the weld metal made by SMAW
and SAW than in that made by GTAW.
The areas around these inclusions are
likely to be sinks for multiplicating and
moving dislocations in the creep process,
preventing their density from increasing
excessively in the matrix. As a result, the
material hardening in these welding processes is believed not to be as conspicuous when compared with that of the
GTAW.
creep behavior of austenitic stainless
steel (illustrated in Fig. 14) has been examined (Ref. 20). For the most part, rupture
elongation decreases with the percentage of nitrogen at each temperature.
Rupture time increases remarkably with
increasing nitrogen content at 650° C
(1202°F) but decreases at around
1,000°C (1832°F) as the nitrogen content
exceeds approximately 0.20%. There
seems to be no definable dependence on
nitrogen over a range from 0.10% to
0.30% nitrogen content at 900°C
(1652°F).
Creep Strength of Nitrogen-Enriched
Austenitic Stainless Steel at Higher Temperatures. The effect of nitrogen on the
Centrifugally cast pipes made from
grades HH and HK 40 steel have been
used as heat resistant material in the
Table 7—Mechanical Properties of Type 308 Stainless Steel Weld Metal with Vanadium Additions as Deposited by SAW, SMAW and GTAW
Proof stress,
MPa
(1000 psi)
Ultimate tensile
strength,
MPa
(1000 psi)
R.T.
372
(53.9)
550
Test
temperature,
Process'31
Elongation, %
Reduction
of area, %
601
(87.2)
44.0
55.3
237
(34.4)
349
(50.6)
23.9
56.0
R.T.
379
(54.9)
562
(81.5)
45.1
52.2
550
272
(39.4)
344
(49.9)
26.6
60.4
R.T.
584
(84.6)
39.4
68.0
550
367
(53.2)
27.8
30.2
oC(b)
SAW
SMAW
GTAW
Joint
designed)
(a)
Welding conditions as follows: S A W - c a r r i e d out with M g O - S i 0 2 - C a F 2 flux at 500 A. 34 V and 40 c m / m i n ( = 6.67 m m / s . 15.7 ipm); S M A W - c a r r i e d out with lime coated electrodes at 135 A, 22 V,
and 15 c m / m i n ( = 2.5 m m / s , 5.9 ipm); C T A W carried out w i t h Ar shielding gas at 200 A, 11 V and 15 c m / m i n ( = 2.5 m m / s , 5.9 ipm).
R.T. = r o o m temperature; ° C = 5 / 9 (°F - 32).
tc)
loint designs A and B are as shown b e l o w schematically;
<b>
y^W-^l
15-
r 45 ^/
i
CM
i
-|12 —
(d)
Tensile specimen: 10 m m (0.39 in.) diameter; 50 m m (1.97 in.) gage length.
WELDING RESEARCH SUPPLEMENT 1221-s
Conclusions
1~3Si-1.6Mn-9~13Ni-18-24Cr-0.08~-0.29N
1. When the nitrogen content in
welds is increased to near the solubility
limit, weld discontinuities such as porosity
and concavity can and do occur. These
can be prevented by adding elements
that enhance the nitrogen solubility of
weld metals and by welding in an argon
shielding atmosphere containing a small
amount of oxidizing gas.
2. The addition of nitrogen can have a
detrimental effect on the weld cracking
of (y + 8) dual phase weld metal, 5-phase
being unstable when its transformation to
fully austenite-phase occurs. However, if
nitrogen is added to the weld metal while
still maintaining the (7 + 5) dual phase,
there is no deterioration in the resistance
to hot cracking. Nitrogen can substantially improve the resistance to weld cracking in fully austenitic weld metals.
2844Psi)
3. Nitrogen has the excellent effect of
increasing weld strength at room and
elevated temperatures. Particularly, the
weld strength surpasses that of the base
metal in the temperature range to about
300°C (572°F) and is presumably related
to the precipitation hardening; this is
additional to solid solution strengthening
which mainly works on base metal.
0.10
0.15
0.25
0.20
4. Nitrogen
increases the creep
strength of welds markedly at elevated
temperatures, and vanadium improves
creep ductility considerably when nominally 0.10% vanadium is present as in
Type 308 stainless steel weld metal.
5. At extremely high temperatures
(about 900 to 1,000°C, i.e., 1652 to
1832°F), the strength of a nitrogenenriched austenitic stainless steel weld is
approximately twice that of conventional
Type 308 stainless steel weld metal. The
welds of nitrogen-enriched austenitic
steel, such as steel B, exhibit excellent
creep strength at 1,000°C (1832°F) and
are as strong as the base metal.
0.30
nitrogen in sheets, %
Fig. 14 - Creep properties
of nitrogen-containing
chemical industry for years; however,
they are liable to deform locally and crack
during service at elevated temperature.
To eliminate such a drawback in these
alloys, steel B, which is a nitrogenenriched stainless steel, has been developed as wrought material by the hotrolling steel-making process. One of the
causes of the strengthening in steel B is
assumed to be in the lowering of the
stacking fault* energy owing to the additions of silicon and nitrogen that are
associated with inhibiting the climbing in
dislocation movement (Refs. 2 1 , 22 23);
another cause is assumed t o be the
retarding chromium nitride precipitation
by silicon (Ref. 20).
Figure 15 depicts the creep test results
of steel B and its welded joints made
using GTAW with matching composition
filler metal, where the creep strength of
the welded joints is almost the same as
for the base metal. No deterioration in
*"Stacking fault" —associated with the movement of discontinuities.
222-s | JULY 1984
austenitic stainless steels
creep properties in the welds could be
found, even when varying the nitrogen
content from about 0.20% to 0.25%.
Q.
O
O
•ST
£
E
3.0
(4.27)
2.5
(3.56)
2.0
(2.84)
-O—base metal. 1100 C(2012 F) heat treatment
A
GTAW weld joint, fracture at base metal
after post heat treatment. 1250 C(2282 F)
s°45 "V
6.0^
51£&27"\
37,35>
1.5
(2.13)
I
^_^—1
1
(—**
^A
I
1
^~\
I
S-30-1
3
0066 ** x
1.0
63 \
(1.42)
0.8
(1.14)
*
^oo83
82\
\
t e s t t e m p e r a t u r e ; 1000°C (1832 F)
0.6
(0.853)
10 z
300 500
103
5 X 1 0 3 10 4
ruDtur e time, hr
Fig. 15 - Creep rupture life of steel B as a functionof applied stress at 1000°C (1832 °F)
References
1. Hum, |, K., and Grant, N. J. 1953. Austenitic stability and creep-rupture properties of
18-8 stainless steel. Transactions of ASM
45:105-133.
2. Board, ). A. 1973. Stress-corrosion in
power industry, journal of Institute of Metal
101:242-247.
3. Abe, S., and Ogawa, T. 1979. A modified
Type 347 stainless steel for nuclear power
application. Metal Progress 116(9):61-65.
4. Tobler, R. L, and Reed, R. P. 1978.
Interstitial carbon and nitrogen effects on the
tensile and fracture parameters of AISI 304
stainless steel. Proc. of symposium held at the
107th AIME Annual Meeting in Denver, Colorado, March 2, pp. 17-48.
5. King, R. T., Stiegler, ). O., and Goodwin,
G. M . 1974. Relation between mechanical
properties and microstructure in CRE Type 308
weldments. Welding journal 53(7):307-s to
313-s.
6. Ogawa, T., Aoki, A., Sakamoto, T., and
Zaizen, T. 1982. The weldability of nitrogencontaining austenitic stainless steel: part I —
chloride pitting corrosion resistance. Welding
Journal 61(5): 139-s to 148-s.
7. Savage, W . F., and Lundin, C. D. 1965.
The Varestraint test. Welding Journal 44(10)
433-s to 442-s.
8. Hambert, ]. C , and Elliot, I. F. 1960. The
solubility of nitrogen in liquid Fe-Cr-Ni alloys.
Trans. AIME 218(12):1076-1088.
9. Uda, M., Ohno, S., and Wada, T. 1969.
Solubility of nitrogen in arc- and levitationmelted iron and iron alloys. / of Japan Weld.
Society 38:382-392.
10. Kobayashi, T., Kuwana, T., and Kikuchi,
T. 1971. Effect of atmospheres and welding
polarity o n the nitrogen content of weld. /
Japan Weld. Society 40(3):221-231.
11. Kobayashi, T., Kuwana, T., and Kiguchi,
R. 1972. Effect of alloying elements on nitrogen content of steel weld metals. / Japan
Weld. Society 41(3&4):308-321 and 415-424.
12. Uda, M., and Ohno, S. 1972. Effect of
surface active elements on nitrogen content of
iron under arc melting. / japan Weld. Society
41(7)772-780.
13. Brooks, ]. A. 1975. Weldability of high
N, high M n austenitic stainless steel. Welding
Journal 54(6):189-s to 195-s.
14. Pascarel, L , Hubert, M., and Messager,
C. 1971. Aspects particuliers de la fissuration a
chaud par soudage d'aciers austenitiques;
influence de lazote. Revue de Metallurgie
68(12):809-828.
15. Masumoto, I., Tamaki, K., and Kubo, T.
1975. Use of nitrogen in welding atmosphere.
/ Japan Weld. Society 44(6):492-498.
16. Hull, F. C. 1967. Effect of delta ferrite on
the hot cracking of austenitic chromium-nickel
stainless steels. Welding Journal 46(9):399-s to
409-s.
17. Polgary, S. 1969. The influence of silicon
content on cracking in austenitic stainless steel
weld metal with particular reference to 18Cr-
8Ni steel. Metal Const. Brit. Weld. J. Feb(S):
93-97.
18. Borggreen, K., and Tholen, A. R. 1976.
Creep mechanisms in a vanadium alloyed austenitic
stainless
steel,
Metallurgical
Transactions. A 7A(11):1615-1620.
19. Liljestrand, L. G., and Omsen, A. 1975.
The mechanisms of improved creep strength
in a new austenitic stainless steel. Metallurgical
Transactions. A 6A(2):279-286.
20. Nakasawa, T., Sunami, T., and Abo, H.
1979.
Improvement of high temperature
strength of austenitic stainless steel by additions of silicon and nitrogen. Tetsu to Hagane
65(7):227-236.
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WRC Bulletin 285
July, 1983
Stress Indices and Flexibility Factors for Concentric Reducers
by E. C. Rodabaugh and S. E. Moore
This r e p o r t was developed as p a r t of t h e ORNL Piping Program funded by t h e U. S. A t o m i c Energy
C o m m i s s i o n . The r e c o m m e n d e d stress indices in t h e r e p o r t were i n c o r p o r a t e d into t h e ASME Boiler and
Pressure Vessel Code, Section III, in 1 9 7 7 .
Finite Element Analysis of Eccentric Reducers and Comparisons with Concentric Reducers
by R. R. Avent, M. H. Sadd, and E. C. Rodabaugh
This r e p o r t was developed to provide stress indices for e c c e n t r i c reducers, and includes r e c o m m e n dations f o r relatively minor w o r k changes in t h e NB-3680 p o r t i o n of t h e ASME Code w h i c h would extend
t h e coverage t o include e c c e n t r i c reducers.
Publication of t h e s e reports was sponsored by t h e Sub-Committee
on Piping Pumps and Valves of t h e
Pressure Vessel Research Committee
of t h e Welding Research
Council.
The price of WRC Bulletin 285 is $ 1 4 . 2 5 per copy plus $ 5 . 0 0 for postage and handling. Orders should
be sent w i t h p a y m e n t to t h e Welding Research Council, Rm. 1 3 0 1 , 3 4 5 E. 4 7 t h St., New York, NY
10017.
WELDING RESEARCH SUPPLEMENT 1223-s