Prevention of Solidification Cracking in Very Low Carbon Steel Welds
Prevention of Solidification Cracking in Very Low Carbon Steel Welds An increase in weld metal C and Ni is effective in preventing cracking caused by S-phase solidification
BY S. OHSHITA, N. YURIOKA, N. MORI AND T. KIMURA
ABSTRACT. Contrary to the common belief that weld solidification cracking occurs in a higher carbon content region, it was found that solidification cracking is more likely to occur as the carbon content in the weld metal decreases. Under the pipe girth-welding conditions with SMAW (cellulosic electrode) and G M A W (100% C0 2 ), the occurrence of this type of cracking is influenced by the following factors: 1. Carbon and nickel content of the weld metal (cracking is prevented with an increase in their amount). 2. Morphology of the solidified dendritic structure of the weld metal. 3. Travel speed (cracking never occurs at the speed less than the critical level). 4. Joint restraint (the higher it is, the more likely cracking occurs). The main cause of this type of solidification cracking is presumably 5-phase solidification. It stems from the experimental facts that: 1. Higher carbon and nickel contents are beneficial to prevent cracking, and 2. Weld shrinkage, which is one of factors of cracking, is raised by b-y transformation of 5-phase solidified weld metal. Based on the experimental results, engineering charts were developed for finding suitable welding conditions and welding materials to avoid low-carbon solidification cracking in weld metal.
programs as part of efforts to develop new steelmaking processes. As verified by a recent report (Ref. 1), the fracture toughness property of crack tip opening displacement (CTOD) at the weld heat-affected zone (HAZ) is improved by reducing the steel's carbon equivalent and especially the carbon content. Low-alloy high strength steel with a lower carbon content is known to furnish higher resistance to cold cracking. Structural steel, which is weldable without preheat and reported to be crack-free, is generally of a low alloy type with a carbon content less than 0.07% or thereabouts (Ref. 2). From a standpoint of improving weldability, reducing the carbon content of steel is an inevitable trend. For steel structures exposed in corrosive environments, the control of weld heat-affected zone hardness to less than a certain level is required. This requirement is imposed to avoid the occurrence of stress-corrosion cracking during service. However, it is not an easy task to reduce HAZ hardness, especially when welding with a very low heat-input as experienced in vertical-down or overhead welding. In such welding, a postweld cooling rate becomes so high that HAZ microstructures are fully martensitic. As the following formula by Beckert, et al. (Ref. 3), indicates, the HAZ hardness of
Introduction
Based on a paper presented at the 63rd A WS Annual Meeting in Kansas City, Missouri, during April 26-30, 1982, under the title, "Solidification Cracking in Welding Very-Low Carbon Steel"
There is an increasing demand for steels having both a higher toughness at low temperatures and excellent weldability—especially steels intended for pipelines and offshore structures in cold climates and for pressure vessels in low temperature service. To meet this demand, steel companies have conducted extensive metallurgical research
S. OHSHITA is Research Engineer, N. YURIOKA is Senior Researcher, and N. MORI is Chief Researcher, Products R&D Laboratories, Nippon Steel Corporation; and 77 KIMURA is with the Kimitsu Works, Nippon Steel Corporation, Japan.
a martensitic structure is determined to be a function of carbon content alone: HV(Vickers hardness) = 939 [%C] + 284
(1)
Therefore, the carbon content in steel should be lowered to meet the HAZ hardness requirement, if postweld heat treatment is not employed. The carbon content in newly developed steel tends to be reduced in order to satisfy the strong demand for high toughness, high resistance to cold cracking, and lower HAZ hardness concurrently. However, the authors recently found that low-carbon steels are somewhat susceptible to solidification cracking when they are welded with filler metals of very-low carbon. The objectives of the study described in this paper were to find the main causes of low-carbon solidification cracking and to develop engineering charts which can be used to determine suitable welding conditions in the welding of low-carbon steel. Experimental Procedure Materials Piping with large diameters as well as flat plates was used in the weld cracking tests. The plate thicknesses ranged between 15.2 and 25.0 mm. Wall thicknesses of the piping ranged between 15.9 and 19.1 mm, and outer diameters were 609 mm (24 in.) and 1219 mm (48 in.). The chemical compositions of the steel plates and pipes are shown in Table 1. These are mainly API-5LX-X60 and X-70 grade of line-pipe steels except for steels FH and FJ which are ordinary structural steels. The carbon contents of these steels ranged between 0.011 and 0.192%. A very low carbon level less than 0.04%
WELDING RESEARCH SUPPLEMENT 1129-s
was provided by low-carbon bainitic steels PA, PB, FA and FB in Table 1. The low-carbon bainitic steel possesses excellent weldability even in very low heat input and high hydrogen cellulosic electrode welding; also, it furnishes sufficiently high toughness in its HAZ (Ref. 4). The 3.5% Ni steel, FJ, which is for a low temperature use, was selected to examine an effect of Ni on the weld solidification cracking. Shielded metal arc welding (SMAW) and gas metal arc welding (GMAW) were employed. In SMAW, cellulosic electrodes ranging from AWS E6010 to E9010
grades were used. In GMAW, 100% C 0 2 was used as shielding gas, and electrode diameters were 0.9 mm (0.035 in.). The chemical compositions of the welding materials examined by the all-weld metal tests are shown in Table 2. The cellulosic electrodes and the G M A W electrodes from GA to GD in Table 2 were commercially available. The remaining G M A W electrodes with carbon contents higher than 0.17% were produced from laboratory-melted ingots. Electrode HH was of 2.4% Ni type. Test Piece and Pipe As summarized in Table 3, tests A and B employed SMAW and tests C, D, E and F employed GMAW. A single root weld pass (stringer bead) was made for tests A to E. In tests A and D, a hot pass was also made following root bead deposition, and only a hot pass was deposited in test F. Test A used flat plate specimens with restraint welds as shown in Fig. 1. A 60 deg V-groove was used in test A; the groove had a 1.5 mm (0.06 in.) root face
Restraint weld
Fig. 1 — Test A root-of-weld specimen
and 1.5 mm (0.06 in.) root opening as shown in Fig. 2. This groove shape is normally used in vertical-down welding with cellulosic electrodes. Test B was a girth-weld test on pipes. The groove shape was the same as that used in test A. Two pipes with pipe length of 1.0 m (39.4 in.) were tackwelded, followed by downhill root welding by three welders for pipes with 1.22 m (4 ft) outer diameter and by t w o welders with 0.61 m (2 ft) diameter pipe. The restraint against joint shrinkage in test B was somewhat less than that of test A with restraint welds.
]
1.5—1|—
(mm unit)
Fig. 2 — Groove shape for test A specimen
Table 1 - -Chemical Compositions of Steels Chemical composi ions. Ni Cr Cu
Symbol
Steel
Shape
Thick., mm
C
Si
Mn
P
S
PA PB PC PD PE PF PC FA FB FC FD FE FF FG FH FI FI
BNT (1) BNT (2) Cr-V-Nb (1) Cr-V-Nb (2) V-Nb Ti Mn-V-Nb BNT (1) BNT (2) AF(1) AF(2) V-Nb (1) V-Nb (2) Ti Si-Mn-V Mn-V-Nb 3.5Ni
Pipe do. do. do. do. do. do. Plate do. do. do. do. do. do. do. do. do.
15.9 15.9 15.9 15.9 15.9 15.9 19.1 22.0 15.9 18.3 18.3 15.2 16.5 15.7 20.0 19.1 25.0
.020 .034 .039 .065 .072 .090 .192 .011 .034 .021 .048 .072 .098 .090 .166 .192 .030
.16 .16 .21 .25 .24 .29 .27 .15 .16 .30 .28 .24 .33 .29 .33 .27 .60
1.83 1.61 1.06 1.08 1.57 1.57 1.31 1.87 1.61 1.61 1.57 1.57 1.49 1.57 1.39 1.31 .50
.023 .016 .016 .009 .018 .015 .016 .022 .016 .014 .017 .018 .017 .015 .025 .016 .008
.003 .003 .001 .001 .003 .005 .013 .007 .003 .002 .001 .003 .004 .005 .011 .013 .004
-
—
.29
.17
-
-
—
.48 .42
.22 .17
.29
.17 .20 .18 .22 .27 .17
— — — —
-
-
3.52
% Mo
V
Nb
Ti
B
— — — -
— -
.043 .046 .037 .037 .044
.017 .016 .013 .017 .018 .077
.0013 .0010
.20 .18
.069 .069 .071
—
-
.045
.041 .042 .046 .038 .040 .044 .048
.072 .073 .071 .074
— — -
.028 .045
.041
.10
-
-
—
—
.020 .016 .013 .011 .018
.077
— -
— .0010 .0010
— -
Table 2—All-Weld Metal Chemical Compositions of Welding Materials Chemical compositions, % ^~~
r-
C
Si
Mn
P
S
Ni
Cr
Mo
.12 .18 .24 .11 .14 .14 .09 .12
.14 .24 .20 .21 .12 .11 .14 .20
.35 .98 1.01 .62 .88 .75 .40 .49
.011 .014 .014 .019 .021 .014 .015 .013
.016 .008 .011 .012 .012 .007 .011 .012
.04 .02 .24 .02 .23 .21 .02 2.40
.03 .02 .25 .04 .02 .02 .02
.57
SMAW (Cellulose type — 4.0 mm diameter electrode)
HA HB HC HD HE HF HC HH
.18
GMAW (100% c o 2 0.9 mm diameter electrode)
GA GB GC GD GE GF GG GH
.04 .09 .24 .14 .17 .23 .29 .33
.53 .94 .50 .50 .85 .86 .87 .87
1.01 1.53 1.40 1.40 1.37 1.38 1.38 1.41
.005 .006 .008 .009 .004 .001 .002 .002
.012 .004 .007 .006 .005 .005 .004 .005
— —
— — —
Welding method
Z
LU
:> Q.
o
LU _l LU
> Q X
o oc < LU V) LU
OC
130-S | M A Y 1983
Symbol
— — .22
— — — — —
mm unit)
22.5'
(mm unit)
'Copper backing
Fig. 3 — Test C root-of-weld specimen
Fig. 4 - Groove shape for test C specimen
Fig. 5 — Test D root and hot weld specimen
G M A W with 100% C 0 2 shielding was used to prepare the test C specimen shown in Fig. 3; the groove shape with copper backing is shown in Fig. 4. In order to obtain a good form of root bead in the inner side of a pipe (the back side for a flat plate) in GMAW, a small backside groove with a 45 deg angle and a 2.5 mm (0.1 in.) depth was made as shown in Fig. 4.
F, a flat plate with a 60 deg V-groove was used —Figs. 7 and 8. This test was conducted to examine the occurrence of solidification cracking in hot passes and filler passes in the girth welding. In order to compare the joint restraint in the test methods listed in Table 3, the shrinkage of joints in the direction perpendicular to the welding line was measured by a contact-type strain gauge. The contact balls were inserted in the back side of the plates and the inner side of pipes. The gauge length was 20 mm (0.79 in.), and the amount of shrinkage was obtained by measuring the change in the gauge length after welding.
position in pipe welding. The travel speed in G M A W with 0.9 mm (0.04 in.) electrode ranged from 3.3 to 26.7 mm/s (7.9
Test D also employed a G M A W process; shapes of the test piece and joint geometry are shown in Figs. 5 and 6, respectively. Unlike other tests involving single-groove welding, test D employed double groove welding —that is, a root pass was deposited in a 45 deg V-groove in a flat welding position and then the test specimen was turned over for hot-pass U-groove welding, also in a flat position. Test E was a girth-weld test involving the same groove shape as that of test C and a fully automatic welding process (Ref. 5). This mechanized process used an inner clamper with a 500-ton oil pressure to set two bevelled pipes in a favorite state for root welding. Therefore, groove restraint in test E is comparable with that in test C. Pipes with outer diameter of 610 mm (2 ft) and wall thickness of 15.9 mm (0.63 in.) were used in test E. For test
to 63 ipm).
Welding Conditions The welding conditions for each test method are summarized in Table 4. The SMAW employed cellulosic electrodes with 4 mm (0.16 in.) diameter, and the travel speed ranged from 4 to 9.8 mm/s (9.4 to 23.2 ipm). The higher travel speed was obtained in the completely verticaldown position —that is, the 3 o'clock
Results and Discussion Morphology of Solidification Cracking In pipe-girth welds having very low carbon contents, cracks were detected by radiographic testing after the completion of root welding. Figure 9 shows a macrograph of this crack in a root weld made using a cellulosic electrode (test A). For all practical purposes, root solidification cracks disappeared by subsequent hot pass deposition (Fig. 10) which was taken from a root-and-hot pass bead in test A. Figure 11 shows an example of cracking initiated at a root bead by a GMAW process with copper backing (test C). In this case, heat flowed in t w o dimensions—to the base metal (horizontal) and to the Cu backing (vertical). Thus, dendritic structures of weld metal grew in
Table 3—Solidification Weld Cracking Test Methods SMAW (cellulose electrode)
GMAW (C02) C. Root weld test with Cu backing in plate (Figs. 3 and 4)
A. Root weld test in V-groove in plate (Figs. 1 and 2)
D. Root weld test in pipe girth welding with Cu backing
Root weld test in pipe girth welding 610—1219 mm O.D. X \ 15.9—19.1 mm thick. I Groove shape: Fig. 2 /
(
(
610 mm O.D. X 15.9 mm thick.\ Groove shape: Fig. 4 I
E. Hot pass test in V-groove in plate (Figs. 5 and 6)
Fig. 6 — Groove shape for test D specimen
Table 4—Welding Conditions Welding Method ;MAW ellulose)
3MAW (C0 2 )
Test Method
Position
Carbon content, % Base Metal Weld Material
Current, A
Travel speed, mm/min
290-560 240-590
A. Root in plate B. Root in pipe
Vertic down All
0.01-0.19 0.02-0.09
0.11-0.24 0.08-0.18
140-190 140-200
C. D. E. F.
Flat Flat All Flat
0.01-0.16 0.01-0.09 0.02-0.09 0.02-0.16
0.04-0.33 0.04 0.04-0.33 0.-0.33
220-300 200 and 250 200-300 250
Root in plate Root and hot in plate Root in pipe Hot in plate
400-1,600 400-1,600 300-1,500 200-1,000
WELDING RESEARCH SUPPLEMENT 1131-s
Q. O
(mm unit)
LU
>
Fig. 7 — Test F hot pass specimen
X
30-
o
X
o oc < LU
tf) LU
o:
i-
z
LU
o _J LU
> X o oc
' / / / 9,//'
" c
/A' V
CD -4-a
0 /
C
^'aA , ^ ' \ x °o'^ (C)o = (C]o
-
O
o • * » ° • * o
GA GB GO GD GE GF GG GH
A/A V/JT'
,s*o
o oc o
/ /
/
/
HA'/ m A"
-O
/•' s
CO
S&' - A/*'*
ro
CD
/ /
i
,
0.10
i
1
"
Carbon Content in Weld Metal Estimated from Dilution R a t e , ( C ] D ( % ) Fig. 21 — Relationship between measured and estimated carbon content in SMAW root bead deposited in plate and pipe using groove shape shown in Fig. 2
Alloying elements such as Al, Cr, Si, Ti, Mo, V, W and Zr are known as 5-phase stabilizing elements in steels. C, Ni, Mn and Cu are 7-phase stabilizing elements. Wada (Ref. 9) proposed the carbon equivalent which expressed the element's contribution to 7-phase solidification of steel welds as:
S
i
0.20
steel welds with higher carbon contents were more susceptible to the solidification cracking. However, the present study showed that the solidification cracking, on the contrary, was more likely to occur as carbon content decreased. The solidification cracking did not occur in weld metals with carbon contents as high as 0.23% which was obtained by using electrode wires with 0.33% carbon. In the equilibrium state, which is never attained in welding with rapid heating and cooling, steels solidify in a 5-phase state when they contain carbon contents less than 0.09%. When the steels transform from b to y (austenitic) phase during cooling, transformation shrinkage occurs. The lateral shrinkage at the transformation from b to y is 0.0011 (Ref. 8). Although it is a very small amount, the transformation raises the shrinkage of solidified metals so that steels with very low carbon content become susceptible to the solidification cracking.
X
83
15
21
76
(2)
Dilution Rate: 5 8 %
[C]D=0.58[C]Base+0.42[C]wire
~~— [C]D=0.89[C]D+0.02
0
:
0
I
I
0.10
:
I
I
[
I
I
1
0.20
Carbon Content in Weld Metal Estimated from Dilution R a t e , ( C ] D ( % ) Fig. 22 — Relationship between measured and estimated carbon content in GMA W root bead in plate and pipe deposited using groove shape shown in Fig. 4
Equation (2) predicts that C and Ni are most beneficial to prevent solidification cracking enhanced by 5-phase solidification. This is verified by the experiments in the present study as shown in Fig. 16. Prevention of Lower Carbon Solidification Cracking Low-alloy high-strength steels of very low carbon content furnish improved fracture toughness at low temperatures, excellent weldability (resistance to cold cracking), and lower HAZ hardness in very low heat-input welding. However, precautions have to be taken to avoid lower carbon solidification cracking. Although this cracking is likely to occur in a root weld pass, it disappears in most cases by subsequent weld pass deposition. The engineering charts in Figs. 16 to 20 are useful for finding crack-free conditions, depending on the welding method to be employed. The relations between carbon contents of weld metal, base metal, and welding material shown in Figs. 21 and 22 are also used to select steel and/or welding material so that welding can be carried out in the crackfree region with respect to the weld metal carbon content. Conclusion
Ni Mn CE7T = C 4- — + + 28 110 Cu Si Mo Cr
/ •
SW»
Dilution Rate: 5 6 %
. a [C]D=0.56[C]Base+0.44[C]a„(1epo
,
^/.s*'
This equation also coincides satisfactorily with the experimental results.
WRC Bulletin 283 February, 1983 A Critical Evaluation of Fatigue Crack Growth Measurement Techniques for Elevated Temperature Applications by A. E. Carden
The report contains a discussion and evaluation of several crack length measurement techniques at elevated t e m p e r a t u r e and presents results f r o m the experimental technique developed at the University of Alabama. Publication of this report was sponsored by the Subcommittee on Cyclic and Creep Behavior of Components of the Pressure Vessel Research C o m m i t t e e of the Welding Research Council. The price of WRC Bulletin 283 is $12.00 per copy, plus $5.00 for postage and handling (foreign 4- $8.00). Orders should be sent with payment to t h e Welding Research Council, 345 East 4 7 t h St., Room 1 3 0 1 , New York, NY 10017.
WRC Bulletin 282 November, 1982 Elastic-Plastic Buckling of Axially Compressed Ring Stiffened Cylinders—Test vs. Theory by D. Bushnell
Concern for the safety of nuclear plants and offshore structures has stimulated efforts to determine buckling characteristics of stiffened cylindrical steel shells. In this paper, BOSOR 5 c o m p u t e r programs were used to predict buckling loads of f o r t y axially compressed mild steel cylindrical shells previously tested at Chicago Bridge & Iron Co. Publication of this report was sponsored by the Subcommittee on Shells of the Pressure Vessel Research C o m m i t t e e of the Welding Research Council. The price of WRC Bulletin 2 8 2 is $10.75 per copy plus $3.00 for postage and handling (foreign + $5.00). Orders should be sent with payment to the Welding Research Council, 345 E. 4 7 t h St., Room 1301, New York, NY 10017.
136-s | M A Y 1983
(A-1)
BY S. OHSHITA, N. YURIOKA, N. MORI AND T. KIMURA
ABSTRACT. Contrary to the common belief that weld solidification cracking occurs in a higher carbon content region, it was found that solidification cracking is more likely to occur as the carbon content in the weld metal decreases. Under the pipe girth-welding conditions with SMAW (cellulosic electrode) and G M A W (100% C0 2 ), the occurrence of this type of cracking is influenced by the following factors: 1. Carbon and nickel content of the weld metal (cracking is prevented with an increase in their amount). 2. Morphology of the solidified dendritic structure of the weld metal. 3. Travel speed (cracking never occurs at the speed less than the critical level). 4. Joint restraint (the higher it is, the more likely cracking occurs). The main cause of this type of solidification cracking is presumably 5-phase solidification. It stems from the experimental facts that: 1. Higher carbon and nickel contents are beneficial to prevent cracking, and 2. Weld shrinkage, which is one of factors of cracking, is raised by b-y transformation of 5-phase solidified weld metal. Based on the experimental results, engineering charts were developed for finding suitable welding conditions and welding materials to avoid low-carbon solidification cracking in weld metal.
programs as part of efforts to develop new steelmaking processes. As verified by a recent report (Ref. 1), the fracture toughness property of crack tip opening displacement (CTOD) at the weld heat-affected zone (HAZ) is improved by reducing the steel's carbon equivalent and especially the carbon content. Low-alloy high strength steel with a lower carbon content is known to furnish higher resistance to cold cracking. Structural steel, which is weldable without preheat and reported to be crack-free, is generally of a low alloy type with a carbon content less than 0.07% or thereabouts (Ref. 2). From a standpoint of improving weldability, reducing the carbon content of steel is an inevitable trend. For steel structures exposed in corrosive environments, the control of weld heat-affected zone hardness to less than a certain level is required. This requirement is imposed to avoid the occurrence of stress-corrosion cracking during service. However, it is not an easy task to reduce HAZ hardness, especially when welding with a very low heat-input as experienced in vertical-down or overhead welding. In such welding, a postweld cooling rate becomes so high that HAZ microstructures are fully martensitic. As the following formula by Beckert, et al. (Ref. 3), indicates, the HAZ hardness of
Introduction
Based on a paper presented at the 63rd A WS Annual Meeting in Kansas City, Missouri, during April 26-30, 1982, under the title, "Solidification Cracking in Welding Very-Low Carbon Steel"
There is an increasing demand for steels having both a higher toughness at low temperatures and excellent weldability—especially steels intended for pipelines and offshore structures in cold climates and for pressure vessels in low temperature service. To meet this demand, steel companies have conducted extensive metallurgical research
S. OHSHITA is Research Engineer, N. YURIOKA is Senior Researcher, and N. MORI is Chief Researcher, Products R&D Laboratories, Nippon Steel Corporation; and 77 KIMURA is with the Kimitsu Works, Nippon Steel Corporation, Japan.
a martensitic structure is determined to be a function of carbon content alone: HV(Vickers hardness) = 939 [%C] + 284
(1)
Therefore, the carbon content in steel should be lowered to meet the HAZ hardness requirement, if postweld heat treatment is not employed. The carbon content in newly developed steel tends to be reduced in order to satisfy the strong demand for high toughness, high resistance to cold cracking, and lower HAZ hardness concurrently. However, the authors recently found that low-carbon steels are somewhat susceptible to solidification cracking when they are welded with filler metals of very-low carbon. The objectives of the study described in this paper were to find the main causes of low-carbon solidification cracking and to develop engineering charts which can be used to determine suitable welding conditions in the welding of low-carbon steel. Experimental Procedure Materials Piping with large diameters as well as flat plates was used in the weld cracking tests. The plate thicknesses ranged between 15.2 and 25.0 mm. Wall thicknesses of the piping ranged between 15.9 and 19.1 mm, and outer diameters were 609 mm (24 in.) and 1219 mm (48 in.). The chemical compositions of the steel plates and pipes are shown in Table 1. These are mainly API-5LX-X60 and X-70 grade of line-pipe steels except for steels FH and FJ which are ordinary structural steels. The carbon contents of these steels ranged between 0.011 and 0.192%. A very low carbon level less than 0.04%
WELDING RESEARCH SUPPLEMENT 1129-s
was provided by low-carbon bainitic steels PA, PB, FA and FB in Table 1. The low-carbon bainitic steel possesses excellent weldability even in very low heat input and high hydrogen cellulosic electrode welding; also, it furnishes sufficiently high toughness in its HAZ (Ref. 4). The 3.5% Ni steel, FJ, which is for a low temperature use, was selected to examine an effect of Ni on the weld solidification cracking. Shielded metal arc welding (SMAW) and gas metal arc welding (GMAW) were employed. In SMAW, cellulosic electrodes ranging from AWS E6010 to E9010
grades were used. In GMAW, 100% C 0 2 was used as shielding gas, and electrode diameters were 0.9 mm (0.035 in.). The chemical compositions of the welding materials examined by the all-weld metal tests are shown in Table 2. The cellulosic electrodes and the G M A W electrodes from GA to GD in Table 2 were commercially available. The remaining G M A W electrodes with carbon contents higher than 0.17% were produced from laboratory-melted ingots. Electrode HH was of 2.4% Ni type. Test Piece and Pipe As summarized in Table 3, tests A and B employed SMAW and tests C, D, E and F employed GMAW. A single root weld pass (stringer bead) was made for tests A to E. In tests A and D, a hot pass was also made following root bead deposition, and only a hot pass was deposited in test F. Test A used flat plate specimens with restraint welds as shown in Fig. 1. A 60 deg V-groove was used in test A; the groove had a 1.5 mm (0.06 in.) root face
Restraint weld
Fig. 1 — Test A root-of-weld specimen
and 1.5 mm (0.06 in.) root opening as shown in Fig. 2. This groove shape is normally used in vertical-down welding with cellulosic electrodes. Test B was a girth-weld test on pipes. The groove shape was the same as that used in test A. Two pipes with pipe length of 1.0 m (39.4 in.) were tackwelded, followed by downhill root welding by three welders for pipes with 1.22 m (4 ft) outer diameter and by t w o welders with 0.61 m (2 ft) diameter pipe. The restraint against joint shrinkage in test B was somewhat less than that of test A with restraint welds.
]
1.5—1|—
(mm unit)
Fig. 2 — Groove shape for test A specimen
Table 1 - -Chemical Compositions of Steels Chemical composi ions. Ni Cr Cu
Symbol
Steel
Shape
Thick., mm
C
Si
Mn
P
S
PA PB PC PD PE PF PC FA FB FC FD FE FF FG FH FI FI
BNT (1) BNT (2) Cr-V-Nb (1) Cr-V-Nb (2) V-Nb Ti Mn-V-Nb BNT (1) BNT (2) AF(1) AF(2) V-Nb (1) V-Nb (2) Ti Si-Mn-V Mn-V-Nb 3.5Ni
Pipe do. do. do. do. do. do. Plate do. do. do. do. do. do. do. do. do.
15.9 15.9 15.9 15.9 15.9 15.9 19.1 22.0 15.9 18.3 18.3 15.2 16.5 15.7 20.0 19.1 25.0
.020 .034 .039 .065 .072 .090 .192 .011 .034 .021 .048 .072 .098 .090 .166 .192 .030
.16 .16 .21 .25 .24 .29 .27 .15 .16 .30 .28 .24 .33 .29 .33 .27 .60
1.83 1.61 1.06 1.08 1.57 1.57 1.31 1.87 1.61 1.61 1.57 1.57 1.49 1.57 1.39 1.31 .50
.023 .016 .016 .009 .018 .015 .016 .022 .016 .014 .017 .018 .017 .015 .025 .016 .008
.003 .003 .001 .001 .003 .005 .013 .007 .003 .002 .001 .003 .004 .005 .011 .013 .004
-
—
.29
.17
-
-
—
.48 .42
.22 .17
.29
.17 .20 .18 .22 .27 .17
— — — —
-
-
3.52
% Mo
V
Nb
Ti
B
— — — -
— -
.043 .046 .037 .037 .044
.017 .016 .013 .017 .018 .077
.0013 .0010
.20 .18
.069 .069 .071
—
-
.045
.041 .042 .046 .038 .040 .044 .048
.072 .073 .071 .074
— — -
.028 .045
.041
.10
-
-
—
—
.020 .016 .013 .011 .018
.077
— -
— .0010 .0010
— -
Table 2—All-Weld Metal Chemical Compositions of Welding Materials Chemical compositions, % ^~~
r-
C
Si
Mn
P
S
Ni
Cr
Mo
.12 .18 .24 .11 .14 .14 .09 .12
.14 .24 .20 .21 .12 .11 .14 .20
.35 .98 1.01 .62 .88 .75 .40 .49
.011 .014 .014 .019 .021 .014 .015 .013
.016 .008 .011 .012 .012 .007 .011 .012
.04 .02 .24 .02 .23 .21 .02 2.40
.03 .02 .25 .04 .02 .02 .02
.57
SMAW (Cellulose type — 4.0 mm diameter electrode)
HA HB HC HD HE HF HC HH
.18
GMAW (100% c o 2 0.9 mm diameter electrode)
GA GB GC GD GE GF GG GH
.04 .09 .24 .14 .17 .23 .29 .33
.53 .94 .50 .50 .85 .86 .87 .87
1.01 1.53 1.40 1.40 1.37 1.38 1.38 1.41
.005 .006 .008 .009 .004 .001 .002 .002
.012 .004 .007 .006 .005 .005 .004 .005
— —
— — —
Welding method
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130-S | M A Y 1983
Symbol
— — .22
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mm unit)
22.5'
(mm unit)
'Copper backing
Fig. 3 — Test C root-of-weld specimen
Fig. 4 - Groove shape for test C specimen
Fig. 5 — Test D root and hot weld specimen
G M A W with 100% C 0 2 shielding was used to prepare the test C specimen shown in Fig. 3; the groove shape with copper backing is shown in Fig. 4. In order to obtain a good form of root bead in the inner side of a pipe (the back side for a flat plate) in GMAW, a small backside groove with a 45 deg angle and a 2.5 mm (0.1 in.) depth was made as shown in Fig. 4.
F, a flat plate with a 60 deg V-groove was used —Figs. 7 and 8. This test was conducted to examine the occurrence of solidification cracking in hot passes and filler passes in the girth welding. In order to compare the joint restraint in the test methods listed in Table 3, the shrinkage of joints in the direction perpendicular to the welding line was measured by a contact-type strain gauge. The contact balls were inserted in the back side of the plates and the inner side of pipes. The gauge length was 20 mm (0.79 in.), and the amount of shrinkage was obtained by measuring the change in the gauge length after welding.
position in pipe welding. The travel speed in G M A W with 0.9 mm (0.04 in.) electrode ranged from 3.3 to 26.7 mm/s (7.9
Test D also employed a G M A W process; shapes of the test piece and joint geometry are shown in Figs. 5 and 6, respectively. Unlike other tests involving single-groove welding, test D employed double groove welding —that is, a root pass was deposited in a 45 deg V-groove in a flat welding position and then the test specimen was turned over for hot-pass U-groove welding, also in a flat position. Test E was a girth-weld test involving the same groove shape as that of test C and a fully automatic welding process (Ref. 5). This mechanized process used an inner clamper with a 500-ton oil pressure to set two bevelled pipes in a favorite state for root welding. Therefore, groove restraint in test E is comparable with that in test C. Pipes with outer diameter of 610 mm (2 ft) and wall thickness of 15.9 mm (0.63 in.) were used in test E. For test
to 63 ipm).
Welding Conditions The welding conditions for each test method are summarized in Table 4. The SMAW employed cellulosic electrodes with 4 mm (0.16 in.) diameter, and the travel speed ranged from 4 to 9.8 mm/s (9.4 to 23.2 ipm). The higher travel speed was obtained in the completely verticaldown position —that is, the 3 o'clock
Results and Discussion Morphology of Solidification Cracking In pipe-girth welds having very low carbon contents, cracks were detected by radiographic testing after the completion of root welding. Figure 9 shows a macrograph of this crack in a root weld made using a cellulosic electrode (test A). For all practical purposes, root solidification cracks disappeared by subsequent hot pass deposition (Fig. 10) which was taken from a root-and-hot pass bead in test A. Figure 11 shows an example of cracking initiated at a root bead by a GMAW process with copper backing (test C). In this case, heat flowed in t w o dimensions—to the base metal (horizontal) and to the Cu backing (vertical). Thus, dendritic structures of weld metal grew in
Table 3—Solidification Weld Cracking Test Methods SMAW (cellulose electrode)
GMAW (C02) C. Root weld test with Cu backing in plate (Figs. 3 and 4)
A. Root weld test in V-groove in plate (Figs. 1 and 2)
D. Root weld test in pipe girth welding with Cu backing
Root weld test in pipe girth welding 610—1219 mm O.D. X \ 15.9—19.1 mm thick. I Groove shape: Fig. 2 /
(
(
610 mm O.D. X 15.9 mm thick.\ Groove shape: Fig. 4 I
E. Hot pass test in V-groove in plate (Figs. 5 and 6)
Fig. 6 — Groove shape for test D specimen
Table 4—Welding Conditions Welding Method ;MAW ellulose)
3MAW (C0 2 )
Test Method
Position
Carbon content, % Base Metal Weld Material
Current, A
Travel speed, mm/min
290-560 240-590
A. Root in plate B. Root in pipe
Vertic down All
0.01-0.19 0.02-0.09
0.11-0.24 0.08-0.18
140-190 140-200
C. D. E. F.
Flat Flat All Flat
0.01-0.16 0.01-0.09 0.02-0.09 0.02-0.16
0.04-0.33 0.04 0.04-0.33 0.-0.33
220-300 200 and 250 200-300 250
Root in plate Root and hot in plate Root in pipe Hot in plate
400-1,600 400-1,600 300-1,500 200-1,000
WELDING RESEARCH SUPPLEMENT 1131-s
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Carbon Content in Weld Metal Estimated from Dilution R a t e , ( C ] D ( % ) Fig. 21 — Relationship between measured and estimated carbon content in SMAW root bead deposited in plate and pipe using groove shape shown in Fig. 2
Alloying elements such as Al, Cr, Si, Ti, Mo, V, W and Zr are known as 5-phase stabilizing elements in steels. C, Ni, Mn and Cu are 7-phase stabilizing elements. Wada (Ref. 9) proposed the carbon equivalent which expressed the element's contribution to 7-phase solidification of steel welds as:
S
i
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steel welds with higher carbon contents were more susceptible to the solidification cracking. However, the present study showed that the solidification cracking, on the contrary, was more likely to occur as carbon content decreased. The solidification cracking did not occur in weld metals with carbon contents as high as 0.23% which was obtained by using electrode wires with 0.33% carbon. In the equilibrium state, which is never attained in welding with rapid heating and cooling, steels solidify in a 5-phase state when they contain carbon contents less than 0.09%. When the steels transform from b to y (austenitic) phase during cooling, transformation shrinkage occurs. The lateral shrinkage at the transformation from b to y is 0.0011 (Ref. 8). Although it is a very small amount, the transformation raises the shrinkage of solidified metals so that steels with very low carbon content become susceptible to the solidification cracking.
X
83
15
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(2)
Dilution Rate: 5 8 %
[C]D=0.58[C]Base+0.42[C]wire
~~— [C]D=0.89[C]D+0.02
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Carbon Content in Weld Metal Estimated from Dilution R a t e , ( C ] D ( % ) Fig. 22 — Relationship between measured and estimated carbon content in GMA W root bead in plate and pipe deposited using groove shape shown in Fig. 4
Equation (2) predicts that C and Ni are most beneficial to prevent solidification cracking enhanced by 5-phase solidification. This is verified by the experiments in the present study as shown in Fig. 16. Prevention of Lower Carbon Solidification Cracking Low-alloy high-strength steels of very low carbon content furnish improved fracture toughness at low temperatures, excellent weldability (resistance to cold cracking), and lower HAZ hardness in very low heat-input welding. However, precautions have to be taken to avoid lower carbon solidification cracking. Although this cracking is likely to occur in a root weld pass, it disappears in most cases by subsequent weld pass deposition. The engineering charts in Figs. 16 to 20 are useful for finding crack-free conditions, depending on the welding method to be employed. The relations between carbon contents of weld metal, base metal, and welding material shown in Figs. 21 and 22 are also used to select steel and/or welding material so that welding can be carried out in the crackfree region with respect to the weld metal carbon content. Conclusion
Ni Mn CE7T = C 4- — + + 28 110 Cu Si Mo Cr
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Dilution Rate: 5 6 %
. a [C]D=0.56[C]Base+0.44[C]a„(1epo
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This equation also coincides satisfactorily with the experimental results.
WRC Bulletin 283 February, 1983 A Critical Evaluation of Fatigue Crack Growth Measurement Techniques for Elevated Temperature Applications by A. E. Carden
The report contains a discussion and evaluation of several crack length measurement techniques at elevated t e m p e r a t u r e and presents results f r o m the experimental technique developed at the University of Alabama. Publication of this report was sponsored by the Subcommittee on Cyclic and Creep Behavior of Components of the Pressure Vessel Research C o m m i t t e e of the Welding Research Council. The price of WRC Bulletin 283 is $12.00 per copy, plus $5.00 for postage and handling (foreign 4- $8.00). Orders should be sent with payment to t h e Welding Research Council, 345 East 4 7 t h St., Room 1 3 0 1 , New York, NY 10017.
WRC Bulletin 282 November, 1982 Elastic-Plastic Buckling of Axially Compressed Ring Stiffened Cylinders—Test vs. Theory by D. Bushnell
Concern for the safety of nuclear plants and offshore structures has stimulated efforts to determine buckling characteristics of stiffened cylindrical steel shells. In this paper, BOSOR 5 c o m p u t e r programs were used to predict buckling loads of f o r t y axially compressed mild steel cylindrical shells previously tested at Chicago Bridge & Iron Co. Publication of this report was sponsored by the Subcommittee on Shells of the Pressure Vessel Research C o m m i t t e e of the Welding Research Council. The price of WRC Bulletin 2 8 2 is $10.75 per copy plus $3.00 for postage and handling (foreign + $5.00). Orders should be sent with payment to the Welding Research Council, 345 E. 4 7 t h St., Room 1301, New York, NY 10017.
136-s | M A Y 1983
(A-1)
