The Effects of Minor Elements on the
Welding Characteristics of Stainless Steel
Sulfur, oxygen and deoxidant contents determine
penetration, while the silicon-to-aluminum ratio controls
slagging characteristics
BY B. POLLARD
ABSTRACT. The effects of systematic
variations in the aluminum, sulfur, silicon,
manganese, phosphorus, molybdenum
and copper contents on the penetration
and slagging characteristics of laboratory
heats of Type 304 stainless steel have
been investigated. The effects of aluminum and sulfur on penetration in tungsten inert gas welding can be satisfactorily
explained in terms of their effects on the
temperature coefficient of surface tension and the theory of surface tension
driven fluid flow. However, the results
for silicon and manganese are more complex and, while not inconsistent with the
surface tension driven fluid flow model,
indicate that other factors besides the
temperature coefficient of surface tension must be taken into account. In the
case of silicon, its influence on the viscosity of the molten metal appears to have a
significant effect on penetration. As
expected, phosphorus, molybdenum and
copper had no effect on penetration
since they are not surface active elements, or likely to interact with surface
active elements and, therefore, have no
effect on the temperature coefficient of
surface tension.
inclusions in the steel because it tended
to approach an equilibrium corresponding to temperatures near the surface of
the weld pool, which were much higher
than those encountered in steelmaking.
Provided that sufficient amounts of the
appropriate deoxidant was present,
these high temperatures favored the formation of slags that were richer in the
high-stability oxides, such as those of
calcium, aluminum, titanium and silicon,
and poorer in the low-stability oxides,
such as those of manganese and chromium, than the inclusions. In general, the
slagging characteristics of stainless steel
were found to be determined primarily
by the silicon-to-aluminum ratio of the
steel. Patch type slags, which caused an
uneven weld bead and irregular penetration, were formed when the silicon-toaluminum ratio was less than 50 and the
less troublesome globular slags formed
when it was greater than 50.
Compositional limits for stainless steel
with welding characteristics suitable for
the production of welded tubing are
given.
Part I—Penetration Characteristics
Penetration values for production
heats of Types 304 and 316 stainless
steels showed a strong dependence on
sulfur content, but were lower for the
laboratory heats, which was attributed to
their lower oxygen contents.
Slag formation during the GTA welding
of stainless steels was found to be caused
primarily by the oxide inclusions in the
steel. However, EDS analyses indicated
that the composition of the slag was not
always equivalent to a mixture of all the
B. POLLARD is with the Research Center,
Armco, Inc., Middletown, Ohio. The paper
was prepared while the author was with I TV
Steel Co., Independence, Ohio.
Paper presented at a conference on "Progress
in Resolving Undesirable Trace Element Effects
on the Weldability of Steels," held November
21, 1985, in Charlotte, N.C.
202-s | SEPTEMBER 1988
Introduction
The effects of minor elements on the
welding characteristics of stainless steel
are of particular interest to steelmakers
KEY W O R D S
GTA Welding
Stainless Steels
Minor Element Effect
Weld Characteristics
Weld Penetration
Weld Slagging
Welded Tubing
Si-to-AI Ratio
Sulfur-02 Deoxidants
Weld Bead Geometry
who supply stainless steel strip to welded
tubing manufacturers. Welded stainless
steel tubing is made by continuously rollforming a strip of steel into a tube and
welding the butt joint of the strip together. For wall thicknesses up to 0.125 in.
(3.18 mm), the welding process most
commonly used is gas tungsten arc
(GTA). Since no filler wire is used, welding
characteristics of the steel are determined completely by its composition.
The t w o welding characteristics of particular concern to tubemakers are penetration and slagging. Poor penetration characteristics limit welding speed and, therefore, have an adverse effect on the
economics of the tubemaking operation.
In extreme cases, they can result in material that is unweldable because there is
essentially no difference between the
current required for full penetration and
that which causes excessive melt
through. The second problem, slag formation, is covered in Part II.
The problem of heat-to-heat variations
in the penetration characteristics of stainless steel was first reported in the literature by Oyler, et al. (Ref. 1), in 1967.
However, it was not until 1975 that it
became a major problem for steelmakers. At that time, steelmakers noticed a
significant increase in the number of complaints from tube-making customers
regarding the poor penetration characteristics of 300 Series stainless steels.
There were t w o reasons for these complaints: first, many tube mills had installed
new high-speed welding lines, which
required steel with improved penetration
characteristics in order to exploit the
higher welding speeds which the equipment was capable of; and second, the
argon-oxygen decarburization (AOD)
process was rapidly replacing the electric
furnace (EF) for refining stainless steel.
The most obvious difference between
A O D and EF steel is the oxygen content,
which typically averages about 60 ppm
for A O D steel versus 120 ppm for EF
steel. Therefore, poor penetration was
originally attributed to the low oxygen
3 (CaO)s 4- 2 [AI] M + 3 [5]*,—
3 (CaS)s + (Al 2 0 3 )s
(1)
Low penetration was found to be
associated with high-aluminum residuals,
i.e., aluminum contents up to 0.030%. To
prevent high-aluminum levels in steel for
tubing applications, aluminum additions
were prohibited and penetration was
generally found to be satisfactory when
the aluminum content did not exceed
low-sulfur heats, which were solved by
increasing the sulfur content to 0.015%.
The adverse effect of aluminum and the
beneficial effect of sulfur on weld penetration have been confirmed by more
recent studies (Refs. 7-9). Unfortunately,
although high-sulfur contents result in
good penetration characteristics, they
have an adverse effect on the hot workability, cold formability and pitting corrosion resistance of stainless steel. Therefore, the present practice at LTV Steel
Specialty Products Co. 1 for common
grades of stainless steels, such as Types
304 and 316 and their low-carbon equivalents, is to produce both a regular quality grade, which may be low-sulfur if
desired, and a tubing quality grade in
which maximum aluminum and minimum
sulfur contents are specified so as to
ensure good welding characteristics.
Recently, questions about the effect of
composition on the welding characteristics of stainless steel have again been
raised. The reasons for this resurgence in
interest are primarily economic. First, a
significant reduction in inventory cost
could be achieved by combining regular
and tubing grades; and second, an
improvement in the hot workability of
difficult-to-roll grades such as Type 316,
which would reduce surface defects and
increase yields, could be obtained by
going to lower sulfur contents if this
could be done without sacrificing weldability.
In order to determine if it were possible to produce low-sulfur steel with good
penetration characteristics, a better
understanding of the interaction between the surface active elements oxygen and sulfur, and the deoxidants aluminum, silicon and manganese was
required. Optimum manganese and silicon contents were suggested by Oyler,
et al. (Ref. 1), but no data were provided
to support their recommendations. A
systematic study was, therefore, made of
the effects of aluminum, sulfur, silicon
and manganese on the welding characteristics of Type 304 stainless steel using
laboratory heats in which one element at
a time was varied. The effects of phosphorus, molybdenum and copper additions in amounts typical of those found in
production heats were also investigated.
The results of this study were then compared with penetration data for production heats of Types 304 and 316 stainless
steel.
tion problems were still e x p e r i e n c e d w i t h
7. Now l&L Specialty Products Co
ingots
content of A O D steel (Ref. 2). Evidence
for the beneficial effect of oxygen was
provided by the results of other investigators who showed that oxygen present
as an oxide film improved penetration
(Refs. 3, 4). However, laboratory welding
tests and customer experience revealed
that, although on average EF steel exhibited slightly better penetration than A O D
steel, there were wide variations in penetration for steel produced by either process. Low penetration was found to be
associated with the presence of strong
deoxidants, specifically aluminum, magnesium, cerium and calcium (Refs. 5, 6). If
oxygen increases penetration, then it is to
be expected that strong deoxidants will
reduce penetration. Of the forementioned deoxidants, only aluminum has
been widely used in stainless steel. It is
sometimes added to the desulfurization
mix to facilitate the removal of sulfur by
the reaction
Experimental Procedure
Material
were made with
the aluminum,
Table 1 --Compositions and Penetration Ratios for Laboratory Heats
Heat .
No.
V154
1329
1330
1331
1332
1333
V155
1334
V156
1335
1336
1337
1338
1339
1340
1341
1342
1343
1344
1345
1346
1347
1348
1349
1350
1509
1510
1511
1512
1513
'ercent
C
Mn
Si
S
P
Cr
Ni
Mo
Cu
Al
N
O, ppm
D/W
0.040
0.034
0.044
0.038
0.037
0.033
0.038
0.015
0.035
0.026
0.046
0.042
0.035
0.045
0.044
0.040
0.043
0.037
0.034
0.040
0.041
0.037
0.035
0.035
0.045
0.072
0.065
0.060
0.065
0.060
0.050
0.12
0.50
0.96
1.38
2.02
1.41
1.46
1.47
1.45
1.48
1.42
1.54
1.43
1.44
1.46
1.46
1.49
1.44
1.48
1.53
1.47
1.55
1.56
1.52
1.42
1.41
1.45
1.47
1.44
0.50
0.47
0.49
0.49
0.47
0.50
0.01
0.05
0.22
0.26
0.73
0.92
0.49
0.51
0.50
0.49
0.49
0.50
0.50
0.48
0.50
0.49
0.53
0.52
0.52
0.46
0.44
0.46
0.47
0.45
0.014
0.009
0.013
0.013
0.014
0.013
0.014
0.012
0.015
0.014
0.014
0.012
0.004
0.007
0.010
0.017
0.026
0.014
0.014
0.014
0.014
0.014
0.013
0.016
0.015
0.005
0.007
0.014
0.028
0.052
0.016
0.012
0.014
0.014
0.013
0.014
0.015
0.013
0.017
0.014
0.015
0.013
0.014
0.013
0.014
0.014
0.014
0.005
0.005
0.019
0.014
0.014
0.014
0.016
0.013
0.023
0.023
0.024
0.024
0.022
18.79
18.92
18.93
19.09
18.95
17.28
18.78
19.09
18.75
19.04
19.19
19.08
17.66
19.10
18.95
18.88
19.00
18.98
18.80
18.85
18.96
18.90
16.95
18.80
19.00
18.52
18.75
18.51
18.56
18.48
8.50
8.39
8.57
8.60
8.51
8.79
8.37
8.59
8.35
8.82
8.74
8.52
9.26
8.55
8.60
8.53
8.55
8.58
8.40
8.65
8.68
8.60
8.75
8.74
8.43
8.44
8.47
8.49
8.46
8.46
0.01
0.01
0.02
0.03
0.03
0.04
0.01
0.04
0.01
0.04
0.04
0.03
0.03
0.03
0.03
0.03
0.03
0.02
0.03
0.03
0.03
0.03
0.03
0.44
0.03
0.02
0.03
0.03
0.03
0.03
0.05
0.04
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.45
0.05
0.05
0.05
0.05
0.05
0.001
0.005
0.002
0.001
0.001
0.001
0.001
0.001
0.001
0.005
0.001
0.001
0.002
0.004
0.009
0.002
0.010
0.004
0.001
0.002
0.010
0.015
0.031
0.001
0.002
0.019
0.029
0.040
0.033
0.021
0.024
0.018
0.018
0.024
0.024
0.015
0.024
0.034
0.025
0.033
0.019
0.018
0.014
0.024
0.023
0.055
0.052
0.029
0.029
0.014
0.049
0.015
0.033
0.029
0.010
0.065
0.072
0.075
0.73
0.076
194
178
159
195
152
158
197
160
206
183
151
149
156
194
132
166
121
160
125
164
174
156
74
174
197
68
88
73
52
75
0.472
0.493
0.551
0.542
0.549
0.508
0.423
0.418
0.475
0.440
0.525
0.477
0.546
0.528
0.432
0.533
0.491
0.533
0.535
0.537
0.473
0.460
0.455
0.533
0.543
0.383
0.384
0.415
0.528
0.510
WELDING RESEARCH SUPPLEMENT 1203-s
Table 2 - Compositions and Penetration Ratios for Types 304 and 304L Production Heats
Thickness
Heat No
Grade
in.
mm
70024
71261
93504
70002
70030
70037
70069
70097
70099
70919
71028
71030
803175
894104
91776
304L
304L
304L
304
304
304
304
304
304
304
304
304
304
304
304
0.108
0.064
0.061
0.063
0.071
0.032
0.052
0.050
0.108
0.027
0.060
0.060
0.046
0.046
0.058
2.74
1.63
1.55
1.60
1.80
0.81
1.32
1.27
2.74
0.69
1.52
1.52
1.17
1.17
1.47
Percent
C
Mn
0.023 1.66
0.026 1.63
0.022 1.72
0.051 1.60
0.070 1.66
0.063 1.67
0.055 1.67
0.055 1.68
0.054 1.75
0.044 1.64
0.049 1.70
0.036 1.67
0.050 1.87
0.034 1.84
0.036 1.61
Si
S
0.54 0.023
0.56 0.010
0.69 0.021
0.52 0.017
0.55 0.004
0.57 0.003
0.43 0.018
0.49 0.017
0.52 0.014
0.68 0.002
0.66 0.022
0.61 0.015
0.57 0.025
0.57 0.019
0.61 0.019
sulfur, silicon, manganese and phosphorus contents varied within the following
ranges: aluminum 0 to 0.03%, sulfur 0 to
0.05%, silicon 0 to 1.0%, manganese 0 to
2.0% and phosphorus 0 to 0.020%.
Also, one heat was made with an addition of 0.5% molybdenum and one with
an addition of 0.5% copper, which are
typical maximum values for these elements in Type 304 stainless steel. The
compositions of the laboratory heats are
presented in Table 1. All material was
made into 0.060-in. (1.52-mm) thick
annealed and pickled strip using standard
processing procedures for Type 304
stainless steel.
D/W
P
Cr
Ni
Mo
Cu
0.028
0.029
0.032
0.030
0.025
0.026
0.027
0.030
0.027
0.024
0.026
0.025
0.034
0.034
0.035
18.37
18.40
18.19
18.32
18.38
18.35
18.39
18.26
18.31
18.22
18.34
18.32
18.21
18.42
18.13
8.97
8.88
8.97
8.14
8.92
9.08
8.60
8.18
8.15
8.92
8.17
8.14
8.24
8.28
8.28
0.35
0.23
0.25
0.18
0.16
0.16
0.16
0.15
0.20
0.20
0.18
0.15
0.35
0.30
0.25
0.41
0.35
0.38
0.19
0.19
0.25
0.22
0.23
0.21
0.36
0.30
0.19
0.37
0.33
0.32
Al
N
0.090
0.001
ND
ND
0.079
0.001
0.001 0.006 0.091
0.003
ND
0.090
0.003
ND
0.048
0.001
ND
0.039
0.001
ND
0.050
0.003
ND
0.078
0.004
ND
0.084
0.004
ND
0.037
0.003
ND
0.079
0.003
ND
0.078
0.001 0.007 0.061
0.001 0.007 0.072
0.003
ND
0.080
Penetration Measurements
Welding tests were performed on
strips of metal approximately 13 in.
long X 1.5 in. wide (33 mm X 3.8 mm),
which were washed with mineral spirits,
followed by acetone, before testing. The
test specimens were held in a Jetline
longitudinal positioner with a copper
backing bar containing a 0.50-in.
wide X 0.25-in. deep (12.7-mm X 6.4mm) groove, and weld penetration was
measured using a variable current technique and the welding conditions listed in
Table 4. After striking the arc at a current
level well below that required to penetrate the strip, the current was continuously increased at a rate of 6 A / s to a
value above that required for full penetration before the arc was extinguished.
The bead width was then measured at
the point where the weld bead just
penetrated the underside of the material
Production material for weldability
studies was obtained in the annealed and
pickled condition in thicknesses of 0.020
to 0.110 in. (0.51 to 2.79 mm), depending
on availability. The compositions of the
production heats are presented in Tables
2 and 3.
Ti
O, ppm Measured
ND
64
57
ND
ND
ND
ND
ND
ND
ND
ND
ND
36
37
92
Corrected
0.501
0.371
0.481
0.411
0.411
0.354
0.443
0.521
0.459
0.373
0.456
0.433
0.459
0.460
0.494
0.457
0.367
0.480
0.408
0.401
0.380
0.450
0.530
0.415
0.403
0.456
0.433
0.472
0.473
0.496
and the penetration ratio, D / W , was
defined as the ratio of the strip thickness
to the bead width.
The ratio D / W varies with welding
conditions as well as steel composition,
and the choice of welding conditions
under which D / W measurements are
made is somewhat arbitrary. D / W
increases with welding current to a maximum value when the weld just penetrates the underside of the strip. The
D / W measurements were made at this
current level, because measurements
made under these conditions should
relate more closely to the situation in an
actual welding operation than those
made at lower current levels. As noted
previously, the D / W measurements on
laboratory process material were made
at a fixed strip thickness.
Tests on various thicknesses of strip
made from the same heat of Type 304
stainless steel showed that
D/W
Table 3 - Compositions and Penetration Ratios for Types 316 and 316L Production Heats
Thickness
Percent
D/W
Heat No
Grade
in.
mm
C
Mn
Si
S
P
Cr
Ni
Mo
Cu
Al
Ti
N
70008
70143
70411
90587
91711
91716
91804
91958
92131
70248
90291
90448
90532
90591
90690
91712
91715
91717
91992
92132
SS
316L
316L
316L
316L
316L
316L
316L
316L
316L
316
316
316
316
316
316
316
316
316
316
316
316
0.077
0.075
0.102
0.036
0.062
0.063
0.062
0.060
0.060
0.105
0.036
0.112
0.060
0.103
0.106
0.062
0.062
0.061
0.060
0.060
0.117
1.96
1.91
2.59
0.91
1.57
1.60
1.57
1.52
1.52
2.67
0.91
2.84
1.52
2.62
2.69
1.57
1.57
1.55
1.52
1.52
2.97
0.022
0.015
0.018
0.015
0.020
0.018
0.015
0.017
0.022
0.063
0.048
0.058
0.053
0.037
0.042
0.057
0.049
0.049
0.046
0.059
0.065
1.75
1.78
1.68
1.82
1.72
1.68
1.67
1.70
1.66
1.46
1.54
1.57
1.56
1.60
1.59
1.62
1.68
1.67
1.71
1.72
1.10
0.45
0.62
0.67
0.59
0.50
0.60
0.54
0.46
0.60
0.55
0.61
0.63
0.70
0.62
0.69
0.52
0.69
0.67
0.54
0.68
0.61
0.017
0.016
0.016
0.010
0.005
0.007
0.007
0.006
0.007
0.025
0.023
0.017
0.015
0.018
0.004
0.003
0.004
0.010
0.008
0.001
0.002
0.021
0.027
0.024
0.026
0.027
0.027
0.034
0.031
0.029
0.028
0.025
0.024
0.024
0.026
0.026
0.025
0.028
0.028
0.042
0.031
0.028
17.00
16.74
17.55
16.34
17.33
16.69
16.75
16.67
16.48
16.77
16.99
16.79
16.77
16.28
16.33
16.79
16.98
16.70
16.84
16.78
16.77
10.37
10.24
12.58
11.02
11.24
11.05
11.07
11.05
11.16
10.22
10.36
10.35
10.90
10.92
10.95
10.79
11.11
11.06
10.90
11.07
10.26
2.08
2.14
2.66
2.12
2.32
2.14
2.03
2.06
2.10
2.07
2.10
2.10
2.12
2.05
2.11
2.03
2.09
2.10
2.06
2.17
1.99
0.16
0.37
0.31
0.30
0.30
0.32
0.31
0.27
0.37
0.56
0.24
0.29
0.65
0.33
0.30
0.30
0.31
0.32
0.49
0.24
0.28
0.001
0.005
0.001
0.004
0.001
0.001
0.001
0.001
0.001
0.001
0.001
0.001
0.001
0.002
0.001
0.001
0.001
0.001
0.001
0.001
0.001
ND
0.007
ND
0.016
0.009
0.009
0.010
0.008
0.008
0.005
0.006
0.018
0.012
0.008
0.014
0.010
0.010
0.010
0.008
0.012
0.008
0.084
0.068
0.061
0.017
0.072
0.025
0.040
0.026
0.037
0.045
0.036
0.037
0.014
0.014
0.020
0.030
0.029
0.035
0.043
0.029
0.017
204-s I SEPTEMBER 1988
O, ppm Measured
ND
ND
ND
ND
78
ND
ND
ND
ND
ND
ND
ND
ND
ND
ND
64
ND
ND
ND
ND
ND
0.396
0.376
0.390
0.447
0.346
0.364
0.362
0.360
0.358
0.442
0.455
0.451
0.422
0.413
0.360
0.437
0.377
0.447
0.434
0.391
0.381
Corrected
0.412
0.390
0.429
0.425
0.348
0.367
0.364
0.360
0.358
0.483
0.433
0.499
0.422
0.453
0.400
0.439
0.379
0.448
0.434
0.391
0.433
0.70 r—
Table 4—Welding Conditions
Welding speed
0.60
**
1 1 _
o-
0.50
0.40
decreased with increasing strip thickness—Fig. 1. The slope of the curve for
the welding conditions used was —0.92
i n . - 1 (—2.34 cm - 1 ), and this value was
used to correct the D / W values for
specimens of production material to a
common thickness of 0.060 in. (1.52
mm).
Oxygen Content
Although oxygen content was not
varied systematically, measurements of
oxygen content were required to explain
the difference in penetration characteristics between laboratory and production
heats.
The oxygen contents of the strip were
measured instead of the oxygen contents
of the welds because prior experience
had indicated little difference between
base metal and weld oxygen contents.
Moreover, weld oxygen contents are
more difficult to measure accurately
because of inhomogeneous distribution
of slag particles within the welds.
Since some oxygen present in the steel
as inclusion is lost as slag that floats to the
weld surface, base metal content may, in
fact, be a more appropriate value to use
than weld metal oxygen content. The
latter is characteristic of the weld after it
has solidified, while the penetration characteristics of the weld are determined by
the oxygen content of the molten metal
during weld pool formation.
It seems evident that significant oxidation of the weld pool by the air during the
formation of the weld did not occur,
because the compositions of the slags
were consistent with their formation
from inclusions in the steel. For example,
a slag with a high-calcium content could
not have been formed by air reacting
with the weld pool, because essentially all
of the very low calcium content of the
steel was already present as oxide.
-
Arc length
Current
HEAT 36I96
Gas tungsten arc
0.50 in. (12.7 mm)
25 cfh (11.8 L/min) argon
25 cfh (11.8 L/min) argon
Tungsten —2% thoria,
0.125 in. (3.18 mm)
diameter, 60-deg point
0.10 in. (2.54 mm)
90 to 140 A for full
penetration
20 ipm (0.85 cm/s)
-col
Process
Nozzle diameter
Shielding gas
Backing gas
Electrode
.080
.100
A
SLOPE =-0.92
0.30
0
.020
.040
.060
.120
THICKNESS, INCHES
Fig. 1 —Effect of strip thickness on the penetration ratio, D/W
ganese, 0.5% silicon and 0.014% sulfur
decreased 14% for an increase in aluminum content from 0.001 to 0.015%, but
only an additional 1% for a further
increase to 0.031 % — Fig. 2. It is significant
that most of the change in the penetration ratio occurred at aluminum contents
of less than 0.010%, which had little
effect on the measured oxygen content
(Fig. 3); and conversely, aluminum contents of 0.010 to 0.031%, which had little
effect
on
penetration,
significantly
reduced the measured oxygen content.
These results indicate that it is the amount
of oxygen in solution in the molten steel
rather than the total oxygen content that
affects penetration.
The effect of sulfur content on penetration for heats containing 1.5% manga-
nese and 0.5% silicon was found to
depend on the aluminum content —Fig.
4. Sulfur content had no effect on penetration for heats containing no more than
0.004% aluminum. In contrast, penetration ratios for heats containing 0.02 to
0.04% aluminum increased with an
increase in the sulfur content up to
0.028%, beyond which a further increase
to 0.052% had no further effect on penetration. At the 0.005% sulfur level,
0.02% aluminum reduced penetration by
31.5%. Two heats which contained 0.009
and 0.010% aluminum had similar D / W
values as the heats containing 0.02 to
0.04% aluminum, which again emphasizes the fact that to obtain maximum
penetration the aluminum content must
be tightly controlled, preferably to a level
0.60
V
Q
0.50
Results
Laboratory Heats
Penetration results for the laboratory
heats are presented in Table 1. Penetration ratios for heats containing 1.5% man-
0.40
0
.005
.010
.015
.020
.025
.030
.035
WT % ALUMINUM
Fig. 2—Effect of aluminum content on penetration
WELDINC RESEARCH SUPPLEMENT I 205-s
*t(JU
1
1 1
300
Z 200
k ~ -o-
c_
a.
a.
— ' • — « ^ ^
13 — — - c 3 — —*
m
\
5
IOO
«> 80
X
O
v\~
^w c s
Q >o
o \ "^ >
60
Vs
40
30
on
.001
(10)
.004.006
.002
.010
(100)
.020
.040.060
.100
(1000)
for Types 316 and 316L in Table 3. No
difference in penetration was found
between production heats of Types 304
and 304L stainless steels containing less
than 0.004% aluminum, which both had
lower penetration ratios than the best
laboratory heats —Fig. 8. Penetration
increased with increasing sulfur content,
but was not quite equal to the laboratory
heats even at the 0.020% sulfur level.
The Types 316 and 316L heats had
similar penetration ratios to the Types
304 and 304L heats. However, on average the penetration ratio was about 14%
higher for the Type 316 than for the Type
316L heats —Fig. 9. The penetration ratios
for the Types 316 and 316L production
heats also increased with increasing sulfur
content, but, like the 304 and 304L heats,
none had penetration ratios equal to the
best laboratory heats.
W T % ALUMINUM.(PPM)
Fig. 3 - Effect of aluminum content on the oxygen content of the laboratory heats
Discussion
Penetration Mechanism
beyond which further manganese additions reduced penetration slightly —Fig 6.
The increase in penetration associated
with a 0.80% addition to a manganesefree heat was about 20%.
Phosphorus contents in the range
0.005 to 0.019% had no effect on penetration for heats containing 1.5% manganese, 0.5% silicon and 0.014% sulfur - Fig.
7. Likewise, additions of 0.5% molybdenum or 0.5% copper had negligible effect
on penetration —Table 1.
of not more than 0.005%.
The effects of silicon on penetration
for heats containing 1.5% manganese and
0.014% sulfur appear to be more complex than those of aluminum and sulfur,
since penetration first increased with silicon additions up to about 0.50% and
then decreased with further additions up
to 0.92% —Fig. 5. The increase in penetration associated with a 0.50% silicon
addition to a silicon-free heat was about
30%.
The effect of manganese on penetration for heats containing 0.50% silicon
and 0.014% sulfur was similar to that of
silicon. Penetration increased with the
addition of manganese to a maximum
value at about 0.80% manganese,
Production Heats
Penetration test results for production
heats of Types 304 and 304L stainless
steel are presented in Table 2, and those
0.60
0.55
o
o
o
o
D
D
0.50
/ •
1
0.35
0.30
WT % &l
O 0 . 0 0 4 °>'.MAX
• 0 . 0 0 9 - 0.010
/
1
i
.010
.020
1
.030
WT % SULPHUR
Fig. 4 —Effect of sulfur content on penetration
206-s I SEPTEMBER 1988
1
.040
1
.050
.060
Various mechanisms have been proposed to explain the effect of small
variations in composition on the penetration characteristics of stainless steels and
other metals. In general, these mechanisms can be divided into three groups:
those that involve changes in arc characteristics (Refs. 5, 7, 10-12), such as the
size of the anode spot (Ref. 10) or the
anode drop voltage (Ref. 12); those that
involve changes in the liquid-gas or liquidsolid interfacial tensions (Refs. 13-18);
and those that depend on changes in the
pattern of fluid flow in the weld pool
(Refs. 9, 19).
The circulation of molten metal within
the weld pool may be a result of magnetohydrodynamic flow (Refs. 20, 21) or
surface tension driven fluid flow (Ref. 9),
otherwise known as Marangoni convection, or a combination of both processes.
At the current levels typical of most GTA
welding, Marangoni convection appears
to be the dominant driving force, and the
surface tension driven fluid flow model
proposed by Heiple, ef al. (Ref. 9), best
explains the effects of very small amounts
of elements such as sulfur and aluminum
on penetration. The basis of the phenomenon of surface tension driven fluid flow
is that at the surface of the weld pool
molten metal is pulled from a region of
low surface tension to one of high surface tension, and molten metal flows
upwards from the depths of the pool to
maintain continuity —Fig. 10. If the temperature coefficient of surface tension is
positive, then the hot metal immediately
beneath the arc is drawn down into the
depths of the pool, producing deep penetration. Conversely, if the temperature
coefficient of surface tension is negative,
then the hot metal immediately beneath
the arc flows outward to the edges of the
pool, thereby widening the weld pool
and producing shallow penetration. The
temperature coefficients of surface tension for pure metals are negative (Ref. 22)
and the beneficial effects of surface
active elements such as sulfur and oxygen
on weld penetration are believed to be
due to the fact that they produce positive
temperature coefficients of surface tension (Refs. 9, 23). A mathematical model
for convection in weld pools developed
by Oreper, et al. (Ref. 24), is in good
agreement with the Heiple, et al. model,
particularly in regard to the effect of the
temperature coefficient of surface tension on the direction of fluid flow within
the weld pool.
Effects of Sulfur and Aluminum
The results obtained in this investigation for the effects of sulfur and aluminum on weld penetration are in complete
agreement with the Heiple, et al. theory
that surface tension driven fluid flow is
the dominant driving force for metal
circulation within the weld pool. As
explained previously, the positive effect
of sulfur on penetration is due to the fact
that it produces a positive temperature
coefficient of surface tension, while the
negative effect of aluminum is due to it
combining with the oxygen, thereby
reducing the soluble oxygen content of
the molten metal and producing a negative temperature coefficient of surface
tension.
0.60
0.50
U?f--y
0.40
0.2
0.4
0.6
0.8
I.O
2.0
2.5
W T % SILICON
Fig. 5 —Effect of silicon content on penetration
0.60
0.50
Effects of Silicon and Manganese
The results for silicon and manganese
additions to Type 304 stainless steel both
showed penetration increasing up to a
maximum value beyond which further
additions of either element caused a
reduction in penetration. These results
are not inconsistent with the surface
tension driven fluid flow model, but they
do indicate that other factors besides the
temperature coefficient of surface tension must be taken into account. Silicon is
a moderately strong deoxidant, and its
addition to a steel that does not contain
any stronger deoxidants must lower the
soluble oxygen content of the weld pool.
However, the laboratory heats contained
0.014% sulfur, which was apparently sufficient to produce a positive temperature
coefficient of surface tension and a
downward movement of the hot metal
at the center of the weld pool. The
beneficial effect of the initial silicon addition was most likely due to it reducing the
viscosity of the molten steel (Ref. 25). If a
downward flow pattern already existed
because of the 0.014% sulfur content,
then a reduction in the viscosity of the
molten steel would increase the flow
velocity and hence increase penetration.
However, with increasing silicon additions, a point was eventually reached
p
0.40
0.50
1.5
I.O
WT % MANGANESE
Fig. 6 — Effect of manganese content on penetration
where the soluble oxygen content
became the dominant factor, the temperature coefficient of surface tension was
reduced sufficiently to more than offset
any further decrease in viscosity by the
silicon and penetration subsequently
decreased.
The same explanation cannot be used
for the initial increase in penetration
observed with manganese additions to
laboratory heats containing 0.5% silicon
and 0.014% sulfur, since manganese additions up to 1.0% have little effect on the
viscosity of iron (Ref. 26). A more likely
explanation for the beneficial effect of a
0.5% manganese addition on penetration
is that it drastically changes the nature of
the slag formed on the surface of the
weld pool. The patch-type slag formed
on the manganese-free heat produced a
puckered surface on the weld bead,
which was not observed with any other
heat. This puckered surface suggests a
high slag-metal interfacial tension, which,
although not itself detrimental to penetration, may also be associated with a reduction in the temperature coefficient of
surface tension.
Effects of Phosphorus, Molybdenum
and Copper
Additions of 0.005 to 0.019% phosphorus, 0.5% molybdenum or 0.5% copper had no effect on penetration. This is
not surprising since there is no evidence
WELDING RESEARCH SUPPLEMENT 1207-s
to indicate that these elements are surface active, interact in any way with
surface active elements, or have a significant effect on viscosity at the levels
added.
Production Heats
The results of tests on the laboratory
heats suggested that good penetration
characteristics could be obtained with
low-sulfur steel by simply limiting the
aluminum content. However, tests on
production material with aluminum contents in the range of 0.001 to 0.003%
showed penetration ratios that were significantly lower than those obtained with
the laboratory heats, not only at lowsulfur contents, but also to a lesser extent
at high-sulfur levels —Figs. 8 and 9. The
most likely reason for the difference in
penetration characteristics between the
laboratory and production heats was
oxygen content, which was only 40 to 80
ppm for the production heats versus 150
to 200 ppm for the laboratory heats.
Increasing penetration by increasing
the oxygen content of the steel is not a
practical means of compensating for a
low-sulfur content because it would
result in low chromium recoveries, poor
surface quality and increased slagging.
Penetration can be increased by adding
oxygen (Refs. 2, 27) or sulfur dioxide (Ref.
0.6
o
—
V.
Q
cT
<s>
0.5
—
0.4
.004
.008
.012
.016
WT % PHOSPHORUS
Fig. 7—Effect of phosphorus content on penetration
.60
LAB HEATS 150 TO 200 PPM OXYGEN
.50
.40
.020
28) to the shielding gas. However, the
amount required is 500 to 1000 ppm,
which is sufficient to cause accelerated
electrode wear in production welding of
tubing, where uninterrupted runs of one
hour or more are normal practice. Also,
sulfur dioxide is toxic and its use would
create potential health and environmental problems. Therefore, oxygen or sulfur
dioxide additions to the shielding gas are
not likely to alleviate the need for adequate sulfur levels in stainless steel for
welded tubing applications.
The optimum sulfur content is a tradeoff between weld penetration and other
properties such as the hot workability of
the steel and the cold formability of the
finished tubing. For good penetration,
not less than 0.010% and preferably
0.015% sulfur is recommended. A range
of 0.005 to 0.010% sulfur will give
improved hot workability and cold formability at the cost of some decrease in
penetration, while less than 0.005% sulfur
gives a significant improvement in the hot
workability of difficult-to-work grades,
such as Type 316, but penetration is
generally only poor to fair. At the other
end of the sulfur range, sulfur contents of
up to 0.025% have been satisfactory for
thin-walled tubing (e.g., 0.060 in./1.52
mm thick) not subject to severe forming
operations. For heavy-wall tubing, too
high a sulfur content can reduce the
surface tension to the point where it
cannot support the weld pool and excessive melt through occurs. For this type of
application, sulfur content is usually limited to a maximum value of 0.015%.
With a sulfur content of 0.014%, the
effect of aluminum on weld penetration
was found to be greatest at aluminum
contents below 0.010%-Fig. 2. With
appropriate steelmaking practice, most
heats contain less than 0.005% aluminum,
but aluminum contents of 0.005 to
0.010% sometimes occur. A maximum
aluminum content of 0.010% is, therefore, specified. This limit guarantees only
slightly better penetration than would be
obtained at higher aluminum contents,
but prevents the formation of patch-type
slags and uneven penetration, which will
be discussed further in Part II of this
paper.
Maximum penetration for the laboratory heats occurred with a silicon content
of 0.50%, which is the same value that
has been found in practice to provide
good weldability with production materi.30
al. With this silicon content, maximum
4 0 TO 8 0 PPM OXYGEN
O 304L
penetration occurred at a manganese
0 . 0 0 4 % MAX Al
content of 0.80%. However, since man• 304
ganese is an austenite former, it is used as
a low-cost substitute for nickel, and the
.20
.030
.020
.025
manganese content of production heats
.015
.005
.010
of Type 304 stainless steel is usually in the
range 1.50 to 2.00%, the maximum value
allowed by the AISI specification for this
Fig. 8 — Effect of sulfur content on penetration for production heats of Types 304 and 304L stainless
grade. Fortunately, the decrease in the
steels containing 0.004% maximum aluminum
W T % SULPHUR
208-s I SEPTEMBER 1988
penetration ratio, D / W , attributable to
the higher manganese content is only 4 to
9%.
Conclusions
1) The results of this investigation confirm that the effects of small amounts of
sulfur, oxygen and aluminum on weld
penetration in the GTA welding of stainless steels are due to their effects, either
directly or indirectly, on the temperature
coefficient of surface tension and the
resultant pattern of surface tension driven fluid flow within the weld pool.
2) The effects of silicon and manganese on weld penetration are not inconsistent with the surface tension driven
fluid flow model for penetration, but in
the case of silicon, its effect on viscosity
must also be taken into account.
3) Good penetration can be obtained
with sulfur contents of less than 0.005%
provided that sufficient oxygen is
present. However, with oxygen contents
typical of A O D steelmaking, there is a
steady decrease in penetration with
decreasing sulfur content when the latter
is less than about 0.025%.
4) The optimum sulfur content is a
trade-off between hot workability, weldability and formability requirements. For
Type 304 stainless steel tubing, a sulfur
content of 0.010 to 0.015% is a good
compromise for most applications.
5) The effect of aluminum on penetration is greatest in the range 0 to 0.010%,
so the aluminum content should preferably be limited to 0.005%, but steelmaking
limitations dictate a more realistic limit of
0.010% maximum.
6) Phosphorus, molybdenum and copper, in amounts typical of those normally
found in Type 304 stainless steel, have no
effect on penetration.
60
LAB HEATS
I50 TO 200 PPM OXYGEN
.50
.40
.30
03I6L
• 3I6
4 0 TO 8 0 PPM OXYGEN
0 . 0 0 4 % MAX Al
.20
.005
.010
.015
.020
.025
.030
WT % SULPHUR
Fig. 9 — Effect of sulfur content on penetration for production heats of Types 316 and 316L stainless
steels containing 0.004% maximum aluminum
Part II—Slagging Characteristics
Introduction
By comparison with the penetration
problem, slag formation during the GTA
welding of stainless steel has received
much less attention. Nevertheless, excessive slag formation is particularly troublesome in automatic welding operations,
such as making an autogenous weld in
tubing, where it has been reported to
cause arc instability, an irregular weld
bead and reduced penetration (Refs. 6, 7,
29). Also, it can cause weld discontinuities
during postweld rolling and drawing
operations (Ref. 29).
Welding tests in an inert gas chamber
have proved that the source of the slag is
the steel and not oxidation due to incomplete shielding (Ref. 6). Slagging has been
attributed primarily to high melting-point
oxides, particularly the oxides of aluminum, calcium, magnesium, titanium, zirconium and cerium (Refs. 6, 7, 10, 29);
Fig. 10 —Effect of Marangoni convection on penetration
but silicon, chromium, manganese and
iron have also been found in welding
slags (Refs. 10, 29).
Although it is widely accepted that
oxide inclusions in the steel are the
source of the slag, a correlation between
the compositions of the steel, the inclusions and the welding slag has not been
established. The effect of slag characteristics on weldability in tubemaking also
has not been adequately explained.
Therefore, in order to determine how to
minimize the effect of slagging, the
effects of minor elements on the slagging
characteristics of both laboratory and
production heats of 300 Series stainless
steels were investigated.
Experimental Procedure
Specimens for slagging studies were
made using the same specimen size,
cleaning procedure and welding conditions as used in the penetration studies,
except for the current level, which was
constant and about 50 A above the
WELDING RESEARCH SUPPLEMENT 1209-s
Results
Laboratory Heats
Fig. 11 — Typical globular slag particle on a
weld crater. 10X
current level required to just penetrate
the strip. Slag particles formed on the top
surface of the weld crater, sometimes on
the underside of the weld crater, and on
the top and bottom surfaces of the weld
bead. These slag particles were examined
visually and classified according to their
appearance. They were basically t w o
distinct types: more or less hemispherically shaped globules, which were often
glassy in appearance (Fig. 11), and thin
patches-Fig. 12. In Fig. 12A, the patchtype slag is clearly visible as the cometshaped gray area covering the weld
crater and forming a tail downstream.
The weld discontinuity in Fig. 12B is a
hump on the underside of the weld bead
caused by the slag patch. Specimens with
attached slag particles were cut from the
welds. Qualitative energy dispersive
spectroscopic (EDS) analyses were then
performed on the slag particles and on
inclusions in polished longitudinal metallographic sections of the base metal.
The effect of slag on weld penetration
was examined by measuring penetration,
defined in this case as the ratio of the
weld root width to the weld face width,
at intervals along a weld.
As expected, the types of inclusions
present in the laboratory heats and the
resultant weld slag compositions were
found to depend upon the manganese,
silicon and aluminum contents of the
steel-Table 5. With a %Mn/%Si ratio of
0.26 (Heat 1329), the inclusions were
pure silica, but the slag contained primarily chromium oxide and only a small
amount of silica. For heats with % M n /
%Si ratios of 1.0, 2.0 to 2.9 and 4.0 to
5.6, the inclusions identified were 100%
manganese chromium silicates, a mixture
of manganese chromium silicates and
chrome galaxites (Cr203-MnO), and
100% chrome galaxites, respectively, but
all formed manganese chromium silicate
slags. However, Heat 1334, which contained only 0.05% silicon, and hence
chrome galaxite inclusions, formed a slag
of essentially the same composition.
The majority of the inclusions in heats
containing 0.010% or more aluminum
were alumina, but manganese chromium
aluminum oxides and manganese chromium silicates were also present. With
these heats, alumina patch-type slags
were formed on the top and bottom
faces of the weld crater and the weld
bead, but several small manganese chromium silicate globules containing varying
amounts of
aluminum
were
also
observed on the rear half of the weld
crater top surf ace — Table 5. The patchtype slag present on the top surface of
the weld bead was extremely variable in
composition. It contained chromium, aluminum and silicon, varying amounts of
manganese and, with some heats, a small
amount of calcium. The variable composition of this slag was due to the fact that
it consisted of a mixture of the alumina
and manganese chromium aluminum silicate slags in various proportions. In addition, the manganese content of the slag
decreased with an increase in the aluminum content of the steel. The small
amount of calcium present, which was
not added intentionally and not identified
in the inclusions, was probably from the
furnace lining.
Globular slags were formed when the
%Si/%AI ratio was greater than about 50
and patch-type slags when it was less
than 50. The only exceptions were heats
that were very low in manganese (Heat
1329) or silicon (Heat 1334), which
formed chromium-rich, patch-type slags
even though they had %Si/%Al ratios of
94 and 50, respectively.
Production Heats
The compositions of the weld slags
formed with production heats are in
Table 6. They varied more widely than
those formed with laboratory heats, even
though their range of manganese, silicon
and aluminum contents was more
restricted. These slags contained as few
as t w o or as many as six different oxides.
Their wide range of compositions
undoubtedly reflects differences in steelmaking practices since the samples were
from many different sources. Another
notable difference in regard to slag compositions between laboratory and production heats was the presence of significant amounts of calcium, from the lime
used in steelmaking, in the latter.
The slags formed with production
heats were classified on the basis of their
appearance in the same manner as the
slags formed with the laboratory heats.
The high-silica slags generally formed distinct globules, whereas the high-alumina
slags formed thin patches. However,
there were a few exceptions, most
noticeably the manganese chromium
oxide and calcium titanium oxide slags,
which were the patch type, and the
calcium titanium aluminum oxide and calcium chromium aluminum oxide slags,
which were the globular type —Table 6.
As with the laboratory heats, globular
slags were formed when the silicon-toaluminum ratio was greater than 50 and
patch-type slags when less than 50.
Effect of Slag on Weld Penetration
Globular slags had no significant effect
on weld penetration. Yet, when a patchtype slag was present, although width of
the weld face, Wp, was fairly constant,
local increases in root width, W R , were
observed — Fig. 13. This variation in penetration was more clearly shown by a plot
of W R / W F versus distance along the
weld. Penetration peaks were found to
coincide with the locations of slag
patches, which also produced humps on
the surface of the weld bead.
Discussion
Slagging Mechanism
The results of this investigation confirm
that the source of the slag formed in the
Fig. 12 — Typical patch-type slags on: A — the weld crater; B — the underside of the weld bead.GTA welding of stainless steels is primarily
the oxide inclusions in the steel. A slag is
2X
210-s I SEPTEMBER 1988
generated by the motion of molten metal
in the weld pool transferring inclusions
from the interior to the surface of the
pool. Not all slag particles, however,
remain on the surface of the weld pool.
Some are subsequently swept back into
the depths of the pool by the same
forces that brought them to the surface.
Nevertheless, in the process of making a
long continuous weld, as in tubemaking,
there is a net transfer of inclusions to the
surface of the weld pool, and the slag
particle formed increases in size until it is
caught on the surface of the advancing
weld bead. A new slag particle then
begins to form, and the whole process is
repeated.
The motion of inclusions within the
weld pool is a combination of an upward
motion resulting from buoyancy forces
and the pattern of fluid flow established
by Marangoni convection. Alumina slags
were formed on both the top and bottom surfaces of the weld pool, which
indicates that the effect of buoyancy
forces on the alumina inclusions was
small. This result is a consequence of the
high melting point of the alumina inclusions (2020°C/3668°F) (Ref. 30), which
were solid in all regions of the weld pool
except the hot zone immediately
beneath the arc and, therefore, did not
coalesce and grow to a size where the
buoyancy force was significant. In contrast, silicate inclusions were found primarily on the top surface of the weld
pool. Since most silicates have melting
points in the range 1291° to 1723°C
(2350° to 3133°F) (Ref. 30), they were
molten in the weld pool and were able to
coalesce and grow by collision with other
inclusions until they attained a size such
that the buoyancy force in combination
with Marangoni convection propelled
them to the surface of the weld pool.
Slag Composition
A comparison of inclusion and slag
compositions for the laboratory heats
(Table 5) showed that not all the heats
had slag compositions similar enough to
the inclusion compositions for the slags to
be considered simply as mixtures of all
the inclusions in the steel. There are a
number of possible explanations for differences between the slag and inclusion
compositions. In most cases, the smallest
inclusions analyzed were those visible at
500X magnification in an optical microscope, which was used to locate areas
for subsequent EDS analysis in a scanning
electron microscope. It is quite possible
that smaller inclusions of different compositions could have been present in
some specimens. However, although
these inclusions might modify the slag
composition to some extent, it is unlikely
that they would be the major constituents of the slag when many larger inclusions were present.
A possible source of higher chromium
contents in some slags, as compared to
the inclusions from which they were
formed, is the passive film present on the
surface of stainless steel. However, this
film is usually very thin and would only be
significant in cases where the volume
Table 5—Inclusio n and Slag Analyses for Laboratory Heats
Percent
Elements in Inclusions
Heat
No.
Mn
Si
Al
1329 0.12 0.47 0.005
94
1330 0.50 0.49 0.002 245
1331 0.96 0.49 0.001 490
1332 1.38 0.47 0.001 470
1333
1334
1335
1336
1337
1346
2.02
1.46
1.45
1.48
1.42
1.53
0.50
0.05
0.26
0.73
0.92
0.50
Major
Si/AI Mn/Si
0.001 500
0.001 50
0.005 52
0.001 730
0.001 920
0.010 50
0.26
Minor
Si
Mn,
1.02
1.96 1 ) M n ,
2)Mn,
2.94 1 ) M n ,
2) Mn,
Mn,
4.04
Mn,
29.20
Mn,
5.58
Mn,
2.03
1.54
Mn,
3.06 1) Al
Cr,
Cr
Cr,
Cr
Cr,
Cr
Cr
Cr
Cr,
Cr,
Si
32.7
Si
Si
3.00 1) Al
3 ) M n , Cr, Si
17.1
Major
Slag Type
Weld bead
top surface
Ti
Crater top surface
Ti, Al, Si Crater top surface
Patch
Cr
Globular
Globular
Mn Cr, Si
Mn Cr, Si
Ti, Al, Si Crater top surface
Globular
Mn Cr, Si
Ti,
Ti,
Ti,
Ti,
Ti,
Globular
Patch
Globular
Globular
Globular
Patch
Mn
Mn
Mn
Mn
Si
Al
Globular
M n Cr, Al, Si
Minor
Si
Trace
Ca, Al
Ca
Si
2) Al
1348 1.55 0.53 0.031
Elements in Slag
Slag Location
Si
2) M a Cr, Al
1347 1.47 0.49 0.015
Trace
2.92 1) Al
2 ) M n , Cr, Si
Si
Al, Si Crater top surface
Al, Si Crater top surface
Al. Si Crater top surface
W
Crater top surface
Al
Crater top surface
Crater top
surface, front
Ti
Crater top
surface, rear
Weld bead top
surface
Crater bottom
surface
Weld bead
bottom surface
Cr
Crater top
surface, front
Cr
Crater top
surface, rear
Weld bead
top surface
Crater bottom
surface
Weld bead
bottom surface
Crater top
surface, front
Crater top
surface, rear
Weld bead
top surface
Crater bottom
surface
Weld bead
bottom surface
Patch
Patch
Cr, Si
Cr
Cr, Si
Cr, Si
1)Cr, Al
2 ) M n Cr, Al, Si
Al
Patch
Al
Patch
Al
Globular
Mn Cr, Al, Si
Patch
Patch
2) Al,
2)Cr, Al, Si
Al
Ca, Al
Ca, Al, Ti, Si
Mn Cr
Si, Cr
Al
Ti
Si
Ca
Cr
Cr
Cr
Ca
Mn Cr, Ca
Cr
Patch
Al
Cr
Patch
Al
Ca
Globular
Mn Cr, Al, Si
Patch
Cr, Al, Si
Patch
Al
Patch
Al
Cr
WELDING RESEARCH SUPPLEMENT!211-s
0.8
HEAf 1348 ( 0 . 0 3 1 % A I )
W E L D I N G SPEED 3 0 " / M I N
^
o
SLAG PATCHES
\
er, silicon, r e d u c e d part o f t h e m a n g a nese a n d c h r o m i u m oxides in the m a n g a nese c h r o m i u m a l u m i n u m o x i d e inclusions t o f o r m a manganese c h r o m i u m
a l u m i n u m silicate slag.
0.6
Slag Physical Characteristics
CO
UJ
X
0.4
£
0.2
u
ft
tx
4
6
8
DISTANCE, INCHES
Fig. 13 — Effect of slag patches on weld
penetration
fraction of o x i d e inclusions w a s v e r y l o w
or w h e r e v e r y f e w o f t h e m b e c o m e a
part of t h e slag. Heat 1329 is a g o o d
e x a m p l e of the latter situation.
M o r e f r e q u e n t l y , the d i f f e r e n c e in
c o m p o s i t i o n b e t w e e n t h e slag and inclusions consists o f the slag being richer in
t h e m o r e stable oxides, such as t h o s e of
calcium, a l u m i n u m , titanium a n d silicon,
t h a n the inclusions. T h e m o s t likely reason for this d i f f e r e n c e is the m u c h higher
t e m p e r a t u r e s that exist near t h e surface
o f a w e l d p o o l , as c o m p a r e d to those
e n c o u n t e r e d in steelmaking. H o w d e n
a n d Milner estimated t h e e f f e c t i v e react i o n t e m p e r a t u r e of the hot z o n e at the
surface o f an i r o n w e l d p o o l u n d e r an
a r g o n arc t o b e 2 1 0 0 ° C (3812°F) (Ref.
31). A t this t e m p e r a t u r e , a d i f f e r e n t e q u i librium exists b e t w e e n metal and slag
t h a n that w h i c h existed at steelmaking
t e m p e r a t u r e s o f a b o u t 1 6 0 0 ° C (2912°F).
This h i g h - t e m p e r a t u r e equilibrium f a v o r s
t h e f o r m a t i o n o f the m o r e stable oxides.
T h e r e f o r e , it is p o s t u l a t e d that, p r o v i d e d
t h e y are present in sufficient quantities,
elements such as calcium, a l u m i n u m , titan i u m a n d silicon will t e n d t o r e d u c e the
less stable oxides, such as t h o s e of m a n ganese a n d c h r o m i u m , d u r i n g slag f o r m a t i o n . T h e simplest examples o f w e l d slags
f o r m e d in this m a n n e r w e r e t h e l a b o r a t o ry Heats 1333 and 1335, w h i c h b o t h
c o n t a i n e d a p p r o x i m a t e l y 0.5% silicon a n d
f o r m e d manganese c h r o m i u m silicate
slags a l t h o u g h t h e y c o n t a i n e d
only
c h r o m e galaxite inclusions — Table 5. In
contrast, l a b o r a t o r y Heat 1334, w h i c h
also c o n t a i n e d c h r o m e galaxite inclusions,
f o r m e d a manganese c h r o m i u m o x i d e
slag of similar c o m p o s i t i o n because it
c o n t a i n e d o n l y 0.05% silicon, w h i c h w a s
clearly insufficient t o r e d u c e the c h r o m e
galaxites. It is t o be e x p e c t e d that m o r e
2 1 2 - s | SEPTEMBER
1988
c o m p l e x slags f o r m e d in t h e same m a n ner will b e richer in calcium, a l u m i n u m
a n d t i t a n i u m , a n d p o o r e r in manganese
a n d c h r o m i u m than the inclusions in the
steel. T h e results f o r the a l u m i n u m - c o n taining l a b o r a t o r y heats s h o w this t r e n d .
Heats 1347 a n d 1348, w h i c h c o n t a i n e d
alumina a n d manganese c h r o m i u m silicate inclusions, f o r m e d a mixture o f alumina and manganese c h r o m i u m alumin u m silicate slags, w i t h a l u m i n u m replacing part of t h e manganese a n d c h r o m i u m . Heat 1346, w h i c h c o n t a i n e d only
0.010% a l u m i n u m , c o n t a i n e d alumina and
manganese c h r o m i u m a l u m i n u m o x i d e
inclusions, b u t f o r m e d a mixture of alumina a n d manganese c h r o m i u m a l u m i n u m
silicate slags. In this case, essentially all the
a l u m i n u m w a s p r o b a b l y in t h e f o r m o f
oxide so that the next strongest deoxidiz-
The m e l t i n g p o i n t a n d surface tension
o f t h e m o l t e n slag at t e m p e r a t u r e s just
a b o v e the melting p o i n t of steel are
clearly the t w o p r o p e r t i e s w h i c h m o s t
strongly influence the b e h a v i o r of w e l d ing slags. High-silica slags typically h a v e
l o w melting points a n d high surface t e n sions at t e m p e r a t u r e s close t o the melting
p o i n t of steel, so that t h e y solidify as
hemispherically shaped globules o n t h e
w e l d surface. In contrast, p u r e alumina
slags have high melting points ( 2 0 2 0 ° C /
3668 C F), a n d o n the b o t t o m of t h e w e l d
a n d the o u t e r regions o f t h e w e l d p o o l
surface, t h e y f o r m p a t c h - t y p e slags,
w h i c h are actually rafts of solid alumina
inclusions. Nevertheless, w h e n t h e slag
patch o n the surface of the w e l d p o o l
g r o w s t o a size such that it c o v e r s m o s t
o f the w e l d p o o l surface, t h e p o r t i o n
closest t o the center of t h e w e l d p o o l will
be m o l t e n . H o w e v e r , since alumina slags
a p p a r e n t l y have l o w e r surface tensions
than high-silica slags, t h e solid a n d liquid
p o r t i o n s o f the slag f o r m a c o n t i n u o u s
thin layer o v e r the w e l d p o o l surface.
High-alumina
slags c o n t a i n i n g
other
oxides h a v e significantly l o w e r melting
points than p u r e alumina, b u t f o r m p a t c h t y p e slags, p r e s u m a b l y because t h e y also
h a v e l o w surface tension.
Effect of the Slag Type on the Weld
Bead Geometry
O t h e r investigators h a v e stated that
w e l d slagging results in p o o r p e n e t r a t i o n
(Refs. 7 , 1 2 , 29). In this study, it w a s f o u n d
that p a t c h - t y p e slags p r o d u c e d irregular
Table 6—Summary of Major Elements Found in Weld Slags Formed on Production Heats of
Type 304 Stainless Steel
Patch Type
Globular Type
Mn Cr
Ca, Al
Ca, Ti
Ti, M
Mn Cr, Al
M n , Si
Ca, Si
Al, Si
Ca, Ti, Al
M n , Cr, Si
Ca, M n , Si
Ca, Al, Si
M n , Cr, Al, Si
Ca, M n , Cr, Si Ca, M n , Al, Si
Ca, Cr, Al
M
Mn Cr, Al Si
Ca. Cr, Al, Si
Si
Al
•— Ca, Cr, Al, Si
Ca, M n , Cr, Al, Si
Si
Al
-<— Ca, M n , Cr, AI, Si
Ca, M n , Cr, Ti, Al, Si
Ca, Mn, Cr, Mg, Al, Si
Ca, Cr, Ti, Si
penetration with localized regions of high
penetration occurring beneath slag
patches. The apparent disagreement with
other investigators is due to the fact that
elements that cause poor penetration,
e.g., aluminum and magnesium, are also
major causes of slagging. A possible
explanation for the effect of a patch-type
slag on weld penetration is that it constricts the arc root so that magnetohydrodynamic flow within the weld pool
increases. When a slag patch is caught on
the surface of the advancing weld bead,
weld penetration abruptly decreases.
However, a new slag patch soon starts to
form, penetration increases and the
whole cycle is repeated.
Patch-type slags also cause the formation of humps on the weld bead surface.
These humps are due to the fact that, at
temperatures just above the freezing
point of the steel, the metal-slag interfacial tension is higher than the metal-gas
interfacial tension.
making. Provided that they are present in
sufficient quantities, elements such as calcium, aluminum, titanium and silicon tend
to reduce the less stable oxides, such as
those of manganese and chromium, during slag formation.
3) Slags formed with production material contain from two to six oxides of
calcium, aluminum, titanium, magnesium,
silicon, manganese and chromium. However, in general, the slagging characteristics of stainless steel were found to be
determined primarily by the ratio of silicon to aluminum in the steel. Patch-type
slags were formed when % silicon/%
aluminum was less than 50 and globular
slags when it was greater than 50.
4) To prevent the formation of patchtype slags, which cause an uneven weld
bead and irregular penetration, the aluminum content must not exceed 0.010% for
a silicon content of 0.50%.
Practical Significance of the Slag Type
in Tubemaking
The author wishes to thank D. ).
Modrak for performing the welding and
P. I. Henderson for the EDS analysis.
The results of this investigation have
confirmed tubemill experience that
patch-type slags are to be avoided
because they produce an uneven weld
bead and irregular penetration. It was
also shown that in most cases globular
instead of patch-type slags will be formed
if the %Si/%AI ratio is greater than 50,
i.e., if for a silicon content of 0.50% the
aluminum content is less than 0.010%.
Globular slags are generally less harmful
than patch-type slags, except for critical
applications such as dairy tubing. Some
welding lines utilize rollers downstream
from the arc to contour the weld while it
is still hot. Slag globules are rolled into the
weld surface and then etched out during
pickling forming pits, which are unacceptable on dairy tubing since they are potential sites for the growth of bacteria.
Although weld slagging cannot be completely eliminated, for this critical application it must be minimized by using clean
steel and welding under conditions that
produce a narrow weld bead, so that the
volume of metal melted and hence the
slag volume are held to a minimum.
Conclusions
1) Slag is formed in the CTA welding
of stainless steel when oxide inclusions
are transferred from the interior to the
surface of the weld pool by a combination of buoyancy forces and Marangoni
convection.
2) The composition of the slag is not
always equivalent to a mixture of all the
inclusions present in the steel because it
tends to approach an equilibrium corresponding to temperatures near the surface of the weld pool, which are much
higher than those encountered in steel-
A ckno wledgments
References
1. Oyler, G. E., Matuszesk, R. A., and Carr,
C. R. 1967. W h y some heats of stainless steel
may not weld. Welding journal 46(12)40061011.
2. Willgoss, R. H., and Yeo, R. B. G. 1975.
Union Carbide Corp., Linde Division, Research
report.
3. Majetic, |. C , and Yeo, R. B. G. lune 8,
1971. Method of Welding Stainless Steel. U.S.
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4. Bukarov, V. A., and Ishchenko, Yu. S.
1974. Melting of metal and seam formation in
the welding of a steel of Type 18-8 with an
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22-25.
5. Bennett, W. S., and Mills, G. S. 1974. GTA
weldability studies on high manganese stainless
steel. Welding lournal 53(12):548-s to 553-s.
6. Franklin, I. E. June 25, 1975. Effects of
Addition Alloys on the Weldability of 304
Stainless Steel. INCO Research report 600.5.
7. Metcalf, ]. C , and Quigley, M. B. C.
1977. Arc and pool instability in GTA welding.
Welding lournal 56(5): 133-s to 139-s.
8. Heiple, C. R„ Cluley, R. ]., and Dixon,
R. D. 1980. Effect of aluminum on CTA geometries in stainless steel. Physical Metallurgy of
Metal joining. Eds. R. Kossowsky and M. E.
Glicksman. The Metallurgical Society —AIME,
Warrendale, Pa., pp. 160-165.
9. Heiple, C. R., and Roper, ]. R. 1982.
Mechanism for minor element effect on GTA
fusion zone geometry. Welding Journal 61(4):
97-s to 102-s.
10. Ludwig, H. C. 1968. Current density and
anode spot size in the gas tungsten arc.
Welding journal 47{5):234-s to 240-s.
11. Glickstein, S. S., and Yeniscavich, W .
May 1977. A Review of Minor Element Effects
on the Welding Arc and Weld Penetration.
Welding Research Council Bulletin 226.
12. Savage, W . F., Nippes, E. F., and Goodwin, G. M. 1977. Effect of minor elements on
fusion zone dimensions of Inconel® 600.
Welding lournal 56(4): 126-s to 132-s.
13. Hazlett, T. H., and Parker, E. R. 1956.
Effect of individual coating ingredients on surface tension of iron electrode. Welding journal
35(3):113-S to 114-s.
14. Bradstreet, B. |. 1968. Effect of surface
tension and metal flow on weld bead formation. Welding journal 47(7):314-s to 322-s.
15. Ishizaki, K. 1966. Interfacial tension theory of arc welding phenomena — formation of
welding bead. Welding Research
Abroad
12(3):68-86.
16. Ishizaki, K. 1962. Interfacial tension theory of the phenomena of arc welding —
mechanism of penetration. Physics of Welding
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17. Roper, j . R., and Olson, D. L. 1978.
Capillarity effects in the GTA weld penetration
of 21-6-9 stainless steel. Welding
journal
57(47): 104-s to 107-s.
18. Friedman, E. 1978. Analysis of weld
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19. Mills, G. S. 1977. Analysis of high manganese steel weldability problem. Welding
lournal 56(6): 186-s to 188-s.
20. Woods, R. A., and Milner, D. R. 1971.
Motion in the weld pool in arc welding.
Welding lournal 50(4): 163-s to 173-s.
21. Willgoss, R. A. 1980. Electromagnetic
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Physics and Weld Pool Behavior, The Welding
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22. Allen, B. C. 1972. The surface tension of
liquid metals. Liquid Metals— Chemistry and
Physics. Ed. S. Z. Beer, M. Dekker, Inc., New
York, N.Y., pp. 161-212.
23. Keene, B. )., Mills, K. C , Bryant, I. W.,
and Hondros, E. D. 1982. Effects of interaction
between surface active elements on the surface tension of iron. Canadian Metallurgical
Quarterly 21(4):393-403.
24. Oreper, G. M., Eagar, T. W., and Szekely, ]. 1983. Convection in arc weld pools.
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25. Romanov, A. A., and Kochegarov, V. G.
1964. Viscosity of binary iron-silicon melts in
the low concentration range of the second
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26. Romanov, A. A., and Kochegarov, V. G.
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27. Heiple, C. R„ Burgardt, P., and Roper,
|. R. 1984. The effect of trace elements on
GTA weld penetration. Modelling of Casting
and Welding Processes II. Eds. |. A. Danzig and
|. T. Berry. T M S - A I M E , Warrendale, Pa., pp.
193-205.
28. Heiple, C. R., and Burgardt, P. 1985.
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162-s.
29. Linnert, G. F. 1967. Weldability of austenitic stainless steels as affected by residual
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WELDING RESEARCH SUPPLEMENT I 213-s
WRC Bulletin 328
November 1987
This bulletin contains two reports covering related studies conducted at The University of Kansas
Center for Research, Inc., on the CTOD testing of A36 steel.
Specimen Thickness Effects for Elastic-Plastic CTOD Toughness of an A36 Steel
By G. W. Wellman, W. A. Sorem, R. H. Dodds, Jr., and S. T. Rolfe
This paper describes the results of an experimental and analytical study of the effect of specimen size
on the fracture-toughness behavior of A36 steel.
An Analytical and Experimental Comparison of Rectangular and Square CTOD Fracture Specimens of an
A36 Steel
By W. A. Sorem, R. H. Dodds, Jr., and S. T. Rolfe
The objective of this study was to compare the CTOD fracture toughness results of square specimens
with those of rectangular specimens, using equivalent crack depth ratios.
Publication of these reports was sponsored by the Subcommittee on Failure Modes in Pressure Vessel
Materials of the Pressure Vessel Research Committee of the Welding Research Council. The price of
WRC Bulletin 328 is $20.00 per copy, plus $5.00 for postage and handling. Orders should be sent with
payment to the Welding Research Council, Suite 1301, 345 E. 47th St., New York, NY 10017.
WRC Bulletin 330
January 1988
This Bulletin contains two reports covering the properties of several constructional-steel weldments
prepared with different welding procedures.
The Fracture Behavior of A588 Grade A and A572 Grade 50 Weldments
By C. V. Robino, R. Varughese, A. W. Pense and R. C. Dias
An experimental study was conducted on ASTM A588 Grade A and ASTM A572 Grade 50 microalloyed
steels submerged arc welded with Linde 40B weld metal to determine the fracture properties of base
plates, weld metal and heat-affected zones. The effects of plate orientation, heat treatment, heat input,
and postweld heat treatments on heat-affected zone toughness were included in the investigation.
Effects of Long-Time Postweld Heat Treatment on the Properties of Constructional-Steel Weldments
By P. J. Konkol
To aid steel users in the selection of steel grades and fabrication procedures for structures subject to
PWHT, seven representative carbon and high-strength low-alloy plate steels were welded by shielded
metal arc welding and by submerged arc welding. The weldments were PWHT for various times up to 100
h at 1100°F (593°C) and 1200°F (649°C). The mechanical properties of the weldments were
determined by means of base-metal tension tests, transverse-weld tension tests, HAZ hardness tests,
and Charpy V-notch (CVN) impact tests of the base metal, HAZ and weld metal.
Publication of these reports was sponsored by the Subcommittee on Thermal and Mechanical Effects
on Materials of the Welding Research Council. The price of WRC Bulletin 330 is $20.00 per copy, plus
$5.00 for postage and handling. Orders should be sent with payment to the Welding Research Council,
345 E. 47th St., Suite 1301, New York, NY 10017.
214-s I SEPTEMBER 1988