Heat Treatment Effects in Multipass Weldments of a High-Strength Steel

Properties of fine low carbon bainitic/martensitic microstructures initially formed are not appreciably altered by the thermal cycling of subsequent passes in multipass welds

BY S. S. STRUNCK A N D R. D. STOUT

ABSTRACT. Weld metal in the 120—150 ksi yield strength range deposited by the GMA process was studied to determine whether control of -the thermal cycles accompanying multipass welding, through selection of the welding process variables, could prove beneficial to weld metals of this strength range, for which the combination of strength and notch toughness is critical and difficult to obtain. The p rog ram inc luded determination of pass-by-pass weld metal cooling rates, measurement of variations of composit ion, observa­t ion of microstructural changes pro­duced in successive passes, and determination of s t r e n g t h and Charpy V-notch toughness; all for a range of welding conditions of 250 to 450 amp and travel speeds of 5 to 25 ipm.

Cooling rates of the individual passes were sensitive to their loca­t ion relative to the bottom and side walls of the groove. In addition, the average weld metal cooling rate of the weld upper-half w a s ap­proximately 25% less than that of the corresponding weld lower-half, irrespective of heat input.

A correlation between the average cooling rate and the notch toughness was found w i th scatter attributed, in part, to the pass-to-pass variation of microstructure resulting f rom pass-

S. S. STRUNCK is affiliated with the Con­struction Engineering Research Labora­tory, Dept. of the Army, Champaign, III., and R. D.. STOUT is Dean of the Graduate School, Lehigh University, Bethlehem, Pa. Paper presented at the AWS 53rd Annual Meeting held in Detroit, Mich., during April 10-14, 1972.

to-pass variation of cooling rate. A weak relation, predominant at med­ium heat input levels (40-60 kJ/ in. ) , was observed between the per­centage of as-cast (columnar) weld metal, un-reheated above the A , , and the notch toughness. For the weld metal studied, it was concluded that the properties of the fine low-carbon baini t ic /martensi t ic micro-structures initially formed were not appreciably altered by the thermal cycling of subsequent passes in multipass welds.

In t roduct ion

For the past several years the Weldabil i ty Committee of the Weld­ing Research Council has sponsored a project at Lehigh University to investigate the influence on weld metal properties exerted by the ther­mal cycles that accompany mult i ­pass welding. As strength level and the accompanying need for notch toughness of constructional steels in­crease, there is an attendant in­crease in the difficulty of duplicating the base metal properties in welded joints. Since the weld metal prop­erties are influenced by the thermal cycles accompanying multipass weld­ing and since those thermal cycles are a function of the process vari­ables the possibility of benefit ing from controlled welding conditions is attractive.

The original as-cast structure is a l ­tered by the reheating cycles result­ing from the deposition of overlying passes. The supposition is that partial grain refinement, alteration of as-cast (columnar) structure, and tempering effects may produce favor­able changes in the ductil ity and notch toughness" of the welded joint.

Welding procedures which maximize these effects could then be used to obtain improved weld metal prop­erties, particularly in high-strength weldments for wh ich the combina­tion of strength and notch toughness becomes crit ical.

In a previous paper1 it was shown that the primary factor controll ing strength and notch toughness of a 100 ksi yield strength weld metal was the heat input level employed

Fig. 1—Welding setup with a weldment ready for an intermediate pass

Fig. 2— Thermocouple sled in position after welding of an intermediate pass

508-s I O C T O B E R 1 9 7 2

for deposition. Except for scatter, there was no change in the notch toughness produced by the range of heat inputs from 30 to 8 0 k J / i n . The notch toughness decreased pre­cipitously at the 80 k J / i n . level and remained at the lower level as the heat input was increased to a max­imum of 155 k J / i n . The microstruc­ture was studied at 60 diameters and the percentage of weld metal wh ich had been "recrystal l ized," i.e., a l ­tered from its original as-deposited columnar structure to a more equiaxed structure, was counted by lineal analysis. Wi th in each of the two groups of data separated by the sudden change at 80 k J / i n . , there were definite improvements of the transit ion temperature w i th increas­ing amounts of recrystallized weld metal. The amount of data did not warrant a quantitative relation be­tween the fraction of recrystalliza­t ion produced in the multipass welds and the resulting notch toughness. However, from a study of the process variables, it was shown that current had little effect upon the percent re­crystallized at a given travel speed, but at a constant current, raising the travel speed increased the amount of recrystallization which in turn was associated wi th improvements in the notch toughness.

It was the purpose of the present investigation to extend this type of. study to a weld metal system wi th a yield strength in excess of 100 ksi. It was desired to determine whether observed mechanical p roper t ies could be optimized by control l ing the welding process var iab les and whether recrystallization, as pro­duced by the heat treatment effects of subsequent passes, could account for those property variations which could not be explained on the basis of composition and cooling rate alone.

Exper imenta l Program

Materials and Welding Procedure

Al l welds were deposited by the gas metal-arc (GMA) process either in a 1 in. deep V-groove mil led in a 2 in. thick A542 plate or, in the case of the bead-on-plate welds, on the top of 2 in. thick surface-ground blocks of A542 steel plate. The variable and fixed welding conditions are listed in Table 1. The welding sequence was standardized to f i l l the groove from the walls into the center. The f inal, or capping passes, were laid adja­cent to each other from one side to the other.

To produce welds w i th no lack of fusion at the sidewalls, it was neces­sary, in the lower half of the V-groove, to ti lt the weldment by 5 deg about its longitudinal axis. At higher heat input levels, the first pass was

Table 1 --Welding Parameters

Weldment Current, code

G 1 * Z5*

zr A 1 * A 2 * B4 Z2 B3 F3* B2 B1 C5 Z3 C3 C2 CI C6 D2 D1 D5

a m p + 1 0 amp

250 300 300 300 300 350 350 350 350 350 350 400 400 400 400 400 400 450 450 450

(a).

Voltage, v + 3 v

25.0 28.5 29.0 26.0 26.5 27.0 28.0 27.0 28.0 26.5 2 6 0 28.0 29.0 26.5 26.0 26.5 25.5 29.0 28.5 28.0

Travel speed. ipm + 0.5 ipm

10.0 5.0 7.5

10.0 15.0

5.0 7.5

10.0 10.0 15.0 20.0

5.0 7.5

10.0 15.0 20.0 25.0 15.0 20.0 25.0

Heat input. k J / i n .

37.5 102.6

69.6 46.8 31.8

113.4 78.4 56.7 58.8 37.1 27.3

134.4 92.8 63.6 41.6 31.8 24.5 52.2 38.5 30.2

Total no. of passes

22 12 13 17 23

9 12 18 15 26 34 10 10 14 21 26 29 20 25 30

(a) Fixed parameters: gas and flow rate — Ar | 28 0 2 at 50 cfh; preheat and interpass temperature — 200 F; polarity and mode: dcrp with spray transfer, electrode diameter — 0.045 in. for weldments designated with [*); 0.0625 in. for all others.

Table 2—Chemical Analyses of Base, Filler, and Weld Metals, (al wt-%

Metal Mn Si Ni Cr Mo .(b)

Base Filler Weld Upper-Half Weld Lower-Half

(a) Average (b) Parts per million (ppm).

0.12 0.09 0.08 0.085

0.60 1.80 1.80 1.68

0.013 0.004 0.008 0.009

0.021 0.007 0.007 0.009

0.25 0.67 0.62 0.53

0.14 2.40 2.40 2.05

2.40 0.99 8.0 0.04 0.53 12.0 0.12 0.56 11.5 0.51 0.59 12.6

placed at the bottom of the V-groove w i th the weldment flat. Figure 1 shows the welding set-up w i th a mounted specimen ready to weld.

Only sound welds, as determined by visual inspection of the saw cut cross sections, were used in this study. Lack of fusion or failure of the arc during welding was cause for scrapping the weldment.

The chemical analyses of the base plate and fi l ler metal, and the aver­age weld metal composition of the weld upper and lower halves, are listed in Table 2. The variation of weld metal composition w i t h loca­t ion in the groove is a consequence of the dissimilar nature of the base and fi l ler metals.

W e l d M e t a l Cool ing Curves

To measure the cooling rates and transformation temperatures of the individual passes of the weldments, the device shown in Fig. 2 was con­structed. Thermocouple wires of 0.015 in. diameter W-5%Re and W-26%Re were led f rom spools at the top, through the body of the holder where the cold junction was located, and out through a cylindrical two-

holed ceramic insulator. Compensat­ing wire was used to transmit the signal from the holder to an X-Y recorder.

When the gun had progressed to the midpoint of the weldment, the thermocouple holder was manually positioned w i th the thermocouple im­mersed in the molten weld puddle along the weld center-l ine, the re­corder started by remote control, and a recording of mill ivoltage vs. t ime was obtained. A typical cooling curve and welding pass sequence is shown in Fig. 3.

Specimen Preparation

The run-on and run-off tabs plus an additional 1VS in. of material from each end of the weldments were saw cut and discarded. The weld reinforcement was ground flush to the level of the base metal. The weld­ment was then saw cut into Vi in. thick slices along the direction of travel. The slice nearest the location of the W-Re thermocouples and the slices immediately adjacent to it on either side were discarded to obtain specimens free of extraneous alloy­ing effects. Each slice was then sur-

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ftSR/tN&B)

- T = a « w - » T

2 3 4 5 i t 8 9 10 U tt

T I M E fegoNsg »

F/<7. 3 — / I typical weld metal center-line cooling curve for one pass and a typical welding pass sequence

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face ground on the 5V2 by 2 in. face "downs t ream" f rom the direction of welding.

To ensure accurate and consistent placement of the Charpy V-notch, the ground faces were etched w i th 20% nital and the notch location and cutting lines were scribed on each slice. A photograph of a typical etched, laid-out, and identified speci­men blank is shown in Fig. 4. As may be noted, two Charpy speci­mens were obtained f rom each slice; one from the weld upper-half and one from the weld lower-half.

Following saw cutting along the appropriate scribed lines, the blanks were ground to final size by remov­ing metal from the side opposite the notch location and equally from the top and bottom and then notched. It should be noted that the Charpy notches were perpendicular to the plate surface, that the metal behind the notch was all weld metal, and that the crack advanced in the direc­t ion of welding or the "downs t ream" direction.

Tensile specimens w i th the tensile axis transverse to the direction of welding were obtained from the weld upper-halves by a similar pro­cedure involving etching to reveal

Fig. 4 — Typical weld cross-section etched, laid out, and identified for Charpy V-notch impact specimens

the location of the weld and machin­ing along scribed lines. There was sufficient weld metal to produce ten­sile specimens wi th a 3A in. long by 0.250 in. diam reduced gage length, terminated in Vs in. radius shoulders, all entirely wi th in weld metal.

Chemical Analyses

The fracture surfaces of Charpy bars broken at low temperatures were end-mil led to a depth of 0.050 in. to obtain weld metal for analysis. Only the bars broken at low tempera­ture were used to avoid specimens wi th large shear lips wh ich would in­troduce metal farther away from the centerl ine. Base metal material was also mil led, and fi l ler metal was-chopped into short pieces for sampling.

Metallography

Standard metallographic prepara­tion techniques, using 6 and 1 micron diamond paste, were used to prepare the metallography specimens for both macro and micro examination. For microscopic examination, the specimens were etched in a 2% solu­tion of nital. Pertinent areas were marked w i th microhardness identa-

tions and photographed at 500 d i ­ameters.

After etching for macrophotogra-phy, fiducial marks were scratched on the weld cross-sections. A center mark, corresponding to the line chosen for the notch of the Charpy specimens, and marks 0.1 in. on either side of the center mark were made. Etching was performed by im­mersion in a 5% nital solution.

The etched cross-sections be­tween the vertical fiducial marks were photographed wi th a Zeiss macroscope at 8 diameters and en­larged to 1 6 diameters. On the photo­graph of each weld half, a grid sys­tem of 19 equally spaced horizontal lines was constructed between the vertical fiducial lines. The as-cast (columnar) regions were outl ined on the photographs w i th the aid of the macroscope. Since the specimens could be rotated and t i l ted under the macroscope to highlight the various areas, the use of the macroscope enabled a more exact discernment ot the borders between various areas. Figure 5-a shows a photograph of a weld cross-section including struc­ture demarcation and scanning lines.

The intercept method was used to ob ta in f r om the pho tog raphs measurements of several microstruc­tural elements including the per­centage of columnar, percentage of equiaxed, and the percentage of as-cast metal not reheated above the A , temperature. The accuracy of the line count method for this applica­tion was checked w i th a polar plan-imeter for several cases and the max­imum discrepancy was found to be less than 2%.

As explained in the discussion, the whi te bands which are only faint ly visible in Fig. 5-a were important to the analysis. To enable them to be more exactly located, the cross-sec­tions were repolished, etched by im­mersion in a special solution (8 ml HN0 3 , 2 gr picric acid, 10 gr am­monium persulfate, 10 gr citric acid, and 10 drops benzalkonium chloride in 1500 ml H 20) and photographed. An example of the etching effects produced by that solution is shown in Fig. 5-b.

The bead-on-plate cross-sections were polished, etched in a 10% solu­t ion of nital, and photographed on the macroscope at 3.5 to 10 diam­eters. Linear dimensions w e r e measured wi th a ruler and areas were measured wi th a polar plan-imeter.

Mechanical Testing

The Charpy specimens were test­ed in accordance w i th ASTM stand­ard E23-66. Values of absorbed en­ergy, lateral expansion, and percent fibrous appearance were deter-

510-s I O C T O B E R 1 9 7 2

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rT/gr. 5 — Cross-section of weld upper half of weldment C5 (400 amp; 25 ipm; 24.5 kJ/in.): (a) — nital etch; (b) — special etch. X8.

mined over the temperature range of 140 to -240 F. The tension testing was performed on a universal ma­chine at a nominal strain rate of 0.05 in . /minute. A Vi in. gage length commercial extensometer was f i t ted to the reduced gage section of the tensile specimens.

R e s u l t s a n d D i s c u s s i o n

Mechanical Tests

From the large number of Charpy specimens, it was possible to deter­mine the upper and lower shelves as wel l as the transit ion region. At ten­t ion was centered on the transit ion region, and a method devised to re­port the data in a convenient form and ensure that the measured transi­tion temperatures were determined in a consistent manner.

In the transit ion region both ab­sorbed energy and lateral expansion varied fairly linearly w i th tempera­ture. Therefore, a linear least squares line was fit to the data in the fol lowing manner. A freehand curve drawn through the plotted data was used to determine the upper and lower shelf values. Al l data 10 ft-lb (or 10 mils) below the upper shelf and 10 ft-lb (or mils) above the lower shelf was utilized for the fit. Ex-cepted were those few points which lay outside those 10 ft- lb l imit lines and, if included, would lend an ab­

normal slope to the line. Such points were excluded only if another point located in the shelf range lay be­tween the transit ion region and the point to be excluded. Once the points to be used for the least squares f i t of the transit ion were determined, none were excluded for intuitive reasons.

Statistical measures of the scatter of the data about the least squares line (1 sigma values) and the degree to which the data were accurately represented by a straight line (correlation coefficients) were ob­tained. The one-sigma values for the individual Charpy tests ranged from 3.8 to 13.0 ft- lb and the correlation coefficients were never less than 82%.

Presented in Table 3 are the values of the upper and lower shelves, the 50% fibrous fracture-ap­pearance transition temperatures, and the values of the 20 and 40 ft- lb (or mils) transit ion temperatures cal­culated from the least squares equa­tions.2 The 40 ft-lb Charpy V-notch impact transit ion temperatures are shown in Fig. 6 as a funct ion of the heat input. The data for each half of the weld are contained wi th in two scatter bands joined by a gradual transit ion in the 60 to 80 k J / i n . range. The data for the weld lower-halves show slightly less scatter than the upper-halves. Because there was a wider variation in the al­

loy content of the weld lower-halves than the weld upper-halves, it might be predicted that a larger amount of scatter would be exhibited by the weld lower-halves than the weld upper-halves. However, such was not the case.

The Charpy data of Fig. 6 exhibit individual trends when plotted as functions of travel speed or current. As shown in Fig. 7, changes in travel speed from 5 to 15 ipm at the 250 and 300 amp levels effected little change in the transit ion tempera­tures. However, at current levels of 350 and 400 to 450 amp, increasing travel speed produced a large im­provement in transit ion temperature wi th the largest change occuring at the 350 amp level.

When the 40 ft-lb transit ion temperatures were plotted as a func­tion of current at constant travel speed, it was noted that at travel speeds of 10, 15, 20, and 25 ipm the transition temperatures initially de­creased w i th increasing current then increased slightly. At travel speeds of 5 and 7.5 ipm there was a slight increase in transit ion tempera­ture w i th increasing current. A com­parison of the effects of current and travel speed on the transit ion temperatures revealed that current affected the transit ion temperatures less at constant travel speed than did travel speed at constant current.

W E L D I N G R E S E A R C H S U P P L E M E N T ! 511 -s

Table 3—Charpy V-Notch Transition Temperatures

Weldment designation

G1 Z5

Z1

A1

A2

B4

Z2

B3

F3

B2

B1

C5

Z3

C3

C2

C1

C6

D2

D1

D5

Section

U U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L

Upper

108 100

86 114 80 98 78 92 78 83 78 90 74

112 86

108 78

100 94 94 86 94 80 92 70

109 74 87 97 97 80 92 72

112 70 92 86 90 78

Shelf f t - lb

Lower

10 14

8 8

10 6 6

10 6

10 8 8 8

10 10 12

6 10 10 10

8 8 8

10 8

10 9

10 12 10 8 8

10 8 8

12 6

10 12

and Shelves

Shelf, mils

Upper

76 80 70 82 66 76 60 72 56 70 62 72 60 84 68 80 62 70 66 70 66 76 66 74 56 80 62 64 72 76 64 68 54 8 0 52 70 60 68 54

Lower

12 12 10 10 12 12 12 2 2 10 18 18 10 10 12 16 12 12 14 12 16 14 12 12 12 1 1 16 16 14 16 16 15 10 10 12 10 14 12 14 12

50% Fibrous FATT,

°F

- 1 2 0 -35 - 2 0 - 9 0 - 6 0 - 6 5 - 8 0

- 1 3 0 - 5 0 - 5 0 - 5 0

- 1 2 0 - 6 0

- 1 1 0 - 1 1 0

- 8 0 - 8 0

- 1 1 0 - 1 0 0 - 1 4 0 - 1 3 0

- 6 0 - 6 0 - 8 0 - 5 0 - 5 0 -35 - 9 0 - 9 0

- 1 1 0 - 1 2 0 -105 - 1 2 0 - 1 1 0

- 8 0 - 1 2 0 - 1 0 0 - 1 1 0

- 8 0

Trai

20

ft- lb

- 1 5 9 - 1 5 5

- 9 9 -143 - 1 0 5 - 1 7 3 - 1 4 0 - 1 7 4 - 1 3 3 - 1 2 2 - 1 2 6 - 1 6 4 - 1 1 3 - 1 9 2 - 1 7 2 - 185 -161 - 1 8 8 - 1 7 3 - 2 2 5 - 1 7 8 - 1 3 8 - 1 0 4 - 1 4 5 -101 - 1 0 6 - 1 3 1 - 1 4 6 - 1 7 9 - 1 7 5 -181 - 1 7 9 -185 - 1 6 0 - 1 1 5 -202 - 1 5 4 - 1 5 3 - 1 6 3

Tsition temperatures,

40

ft- lb

- 121 - 1 0 9

- 5 9 - 1 1 0

- 5 2 - 1 2 3

- 8 5 - 1 1 6

-71 - 7 7 - 6 6

- 1 0 0 - 5 0

- 1 5 3 - 1 1 6 - 1 3 9

- 7 7 - 1 4 9 - 1 1 9 - 1 6 4 -111

- 8 9 -41

- 1 0 3 - 3 8 - 7 6 - 5 7

- 1 0 2 - 1 2 3 - 1 3 5 - 1 0 9 - 1 1 8 - 1 0 3 - 1 2 4

- 5 8 - 1 3 4 - 1 0 6 -121

- 8 7

20

mils

- 1 7 8 - 1 5 8 - 1 0 6 - 1 5 0 - 1 0 5 - 1 7 6 - 1 4 8 - 1 8 4 -131 - 1 6 6 -131 - 163 - 1 1 4 - 2 0 5 - 2 1 2 - 1 7 5 - 1 6 2 - 1 7 0 -191 - 2 1 8 - 1 8 9 - 1 6 3 -111 - 1 4 3 - 1 0 2 -131 - 1 3 9 - 1 6 1 - 1 9 0 - 1 9 3 - 1 9 3 -161 - 1 6 9 -171 - 1 2 3 - 2 0 6 -151 - 1 5 8 - 1 6 2

°F 40

mils

-105 - 8 7 - 3 9 - 9 3 - 2 7

- 1 0 4 - 5 6 - 9 4 - 2 9 - 7 4 - 5 4 - 7 9 - 2 3

- 1 4 0 - 9 7

- 113 - 4 2

- 1 1 7 - 79

- 1 2 8 - 6 7 - 8 0 -21 - 8 2 -01 - 7 5 - 3 3 - 8 7 - 93

- 1 1 5 - 7 9 - 75 - 35

-107 - 2 6

-106 - 8 4

- 105 - 1 8

Table 4—All-Weld Metal Tensile Properties (a)

Weld desig­nation

G1 Z5 Z1 A1 A 2 B4 Z2 B3 F3 B2 B1 C5 Z3 C3 C2 C1 C6 D2 D1 D5

Current, amp

250 300 300 300 300 350 350 350 350 350 350 400 400 400 400 400 400 450 450 450

Travel speed,

ipm

10 5 7.5

10 15

5 7.5

10 10 15 20

5 7.5

10 15 20 25 15 20 25

Heat input

kJ/ in.

37.5 102.6

69.6 46.8 31.8

113.4 78.4 56.7 58.8 37.1 27.3

134.4 92.8 63.6 41.6 31.8 24.5 52.2 38.5 30.2

0.2% yield strength,

ksi

140.3 124.8 140.4 138.6 143.5 113.6* 129.0 141.8* 134.8 152.3* 162.2* 121.7 122.2 134.2* 145.0* 153.6* 159.0 146.0 156.6 157.7

Ult imate tensile

strength, ksi

151.6 144.7 161.6 155.6 159.5 155.6* 156.0 158.9* 153.2 168.8* 171.5* 148.6 153.4 157.0* 165.0* 168.3* 170.2 162.5 174.9 168.9

Reduction in area,

% 66.3 70.6 71.2 70.5* 68.3 67.7* 67.2 69.7* 68.9 68.9* 66.8* 69.8 68.5 70.5* 67.0* 65.5* 65.9 69.4 67.4 66.6

(a) Average of 3 unless average of 2. denoted by (*).

512-s I O C T O B E R 1 9 7 2

These findings parallel those of the previous study1 over a wider range of the process variables.

The data from the tension tests of the weld upper-halves are presented in Table 4. The variation of the yield and tensile strengths w i th respect to the cooling rates are discussed later.

Weld Metal Chemical Compositions

The chemical analyses of the in­dividual weldments were statistically analyzed to determine whether there were correlations between the ob­served distributions of the elements and the welding process variables. The distributions of carbon, man­ganese, nickel, chromium, and molyb­denum were found to be nearly ran­dom wi th respect to current, travel speed, and heat input.

Cooling Rates and Transformation Temperatures

To process the cooling curve data, 20—30 data points per curve in the temperature range from approximate-

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- 6 0

- 1 0 0

- 1 4 0

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KEY o 300A Q 4 0 0 A A G1

ft 350A • 450A O F3 ~\—i—i—i i — i — i — i — i — i — r

H—I—I—h H—I—h

J I I L

WELD LOWER HALF

J L

2 0 4 0 6 0 8 0 100 120 140

HEAT INPUT (KJI)

Fig. 6 — Variation of the 40 ft-lb transition temperature with heat input

- 6 0

100

140

- 4 0

- 8 0

-120

KEY

I

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- D \

O

" -

1

- D

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1 1

D400A

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1 1 WELD

•

WELD

• o

4 0 0 A

1 1

1 UPPER

& G1

O F 3

1 1 HALF

4 0 0 / 4 5 0 _

y •

a

350 A

i i i

LOWER HALF

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L I

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e 5 10 15 2 0 25

TRAVEL SPEED (IN/MIN)

Fig. 7 — Variation of the 40 ft-lb transition temperature with travel speed

ly 1700 to 100 F above the transforma­tion temperature were recorded and punched on computer data cards. As suggested by Dorschu,3 the mathe­matical expression:

t=C [exp (DT)] (D

was first chosen to represent the data. In that exponential equation, t is arbitrary t ime in seconds starting wi th 1.0 at the highest temperature, T is temperature in °F and C and D are constants to be determined.

Excellent values of the correlation coefficient and one-sigma statistics were obtained for the data by the method of linear least squares. How­ever, w h e n d i sc repanc ies of 50—150 F were found between the actual and the computed tempera­tures wi th a systematic curvature to the supposedly straight l ine, a poly­nomial of the fol lowing form was tried

T=E+Ft+Gt2+Ht3 (2)

The constants, E, F, G, and H were determined by computer for each weld metal cooling curve. The dif­ference over the entire range be­tween values of temperature pre­dicted by eq (2) and the measured values were always less than or equal to 10 F. Those differences.

commonly referred to as "residuals," were randomly distributed.

To determine the cooling rate at a specified temperature, eq (2) was dif­ferentiated w i th respect to t ime. It was then necessary to determine the t ime at wh ich that temperature oc­curred. That value of t ime was found by the Newton-Raphson iterative process, entered into eq (2), and the resultant cooling rate determined and reported by another computer sub-routine.

In all cases, the pass-to-pass cool­ing rates measured at 1000 F were smaller in magnitude than those measured at 1300 F. However, on a percentage basis, the scatter in the individual measurements made at 1300 and 1000 F was about the same regardless of heat input. This is illustrated in Fig. 8.

It should also be noted from Fig. 8 that the cooling rates measured w i th the weldment in the ti l ted position tend to be higher than those meas­ured wi th the weldment in the flat position. Since the weldment was ti lted to permit the arc to penetrate into re-entrant angles formed be­tween previous passes and the side-walls, the increased cooling rates were attributed to an increase in the weld nugget area in contact w i th the metal through which heat was con­

ducted. That situation approached the thermal coupling obtained for welds deposited in a tee joint. The heat f low condition observed in bead-on-plate welds is similar to that of multipass welds, made wi th the weldment flat, and are those associ­ated wi th less effective thermal coupling. The observations of Naka-bayashi,4 who showed that the fused area was smaller for bead-in-groove than bead-on-plate welds made wi th the same GMA welding conditions, are in concert w i th the observations of cooling rate of this study since more rapid cooling rates would tend to decrease the fused area.

The cooling rates of those passes closest to the notch of the Charpy specimens were averaged to obtain a center-l ine cooling rate for the up­per and lower halves of the mult i ­pass welds. Those average weld metal cooling rates are listed in Ta­ble 5, plotted as a function of the in­verse of the heat input in Fig. 9 and fitted to a linear least squares line. Despite large pass-to-pass variations in the individual values used to compute the average, excel­lent correlation was obtained be­tween the average cooling rate at 1300 F and the inverse of the heat input.

It is important to note that there

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were pass-to-pass variations in the cooling rate as a function of the pass location in the groove. There was ap­proximately a 25% difference in the average cooling rates between the upper and the lower halves, irrespec­tive of heat input. Passes in the bot­tom of the groove that were de­posited w i th the weldment in the ti lted position showed the highest cooling rates of all.

As shown in a later section of this discussion, the mechanical prop­erties were primarily a function of the average cooling rate, but there is considerable scatter in the correla­tion w i th the transit ion tempera­tures. One source of the scatter in the correlation w i t h t r a n s i t i o n temperatures may wel l be attributed, in part, to the pass-to-pass varia­tions in cooling rate wh ich are re­flected in the metallurgical struc­tures formed. These pass-to-pass variations in cooling rates in mult i -

Table 5—Average Weld Metal Cooling Rates

Weld desig­nation

G1

Z5

Z1

A1

A2

B4

Z2

B3

F3

B2

B1

C5

Z3

C3

C2

C1

C6

D2

D1

D5

Sect­ion

U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L U L

Current, amp

250

300

300

300

300

350

350

350

350

350

350

400

400

400

400

400

400

450

450

450

Travel speed, ipm

10

5

7.5

10

15

5

7.5

10

10

15

20

5

7.5

10

15

20

25

15

20

25

Heat input,

kJ/in.

37.5

102.6

69.6

46.8

31.8

113.4

78.4

56.7

58.8

37.1

27.3

134.4

92.8

63.6

41.6

31.8

24.5

52.2

38.5

30.2

Average weld metal cooling rate (°F/sec) and range of values

used in average

At 1300 F

125 154 48 56 60 68 100 121 147 169 49 61 61 70 97 109 78 98 143 167 173 208 41 50 51 61 73 99 106 133 149 180 165 199 101 107 141 161 141 178

144/108 173/144 52/44 66/49 66/54 85/66 114/91 137/114 166/130 197/157 51/47 70/56 63/59 83/59 98/95

117/100 85/68 105/81 160/131 195/151 204/157 254/183 42/40 58/45 57/44 70/54 80/64 124/90 113/102 148/113 178/127 225/148 182/158 235/170 106/97 115/92 146/133 187/139 156/129 193/161

At 1000 F

67 89 24 29 32 38 52 67 79 93 26 30 32 38 55 62 51 54 78 94 96 115 22 26 29 32 41 58 63 76 84 104 92

114 56 62 77 89 77 102

81/57 98/81 27/22 35/24 38/28 44/35 60/46 76/60 88/71 105/84 27/24 35/28 35/31 42/32 58/53 64/58 45/37 61/45 86/76 107/85 109/89 136/99 23/21 31/24 31/27 36/30 46/36 72/52 66/59 81/67 96/75 131/86 100/86 135/99 53/60 67/55 80/72 101/76 82/71 110/88

Average temperat

range used ir

856 835 912 898 914 894 888 852 834 813 902 880 896 889 856 845 902 872 816 781 806 768 927 907 914 896 864 814 814 786 791 749 819 794 842 819 823 793 837 784

transition jre (°F) and of values average

875/842 847/818 922/903 913/883 917/913 913/880 903/865 865/847 795/847 828/790 908/894 903/847 913/865 913/865 870/847 865/828 922/884 898/851 828/804 813/752 828/752 809/696 936/917 931/884 927/903 903/884 884/847 837/752 828/809 809/762 813/752 813/686 832/790 847/762 847/828 847/790 842/795 828/757 851/809 809/757

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Fig. 10 — Variation of the average weld metal transformation temperature with the average weld metal center-line cooling rate measured at 1300 F

pass welds deposited in heavy sec­tions have not been adequately rec­ognized in the literature.

The weld metal cooling curves re­vealed that a transformation oc­curred over a temperature range of 930 to 700 F. For the purposes of this investigation, the transforma­tion temperatures were determined visually by selecting the point of min­imum slope caused by the heat of transformation. The weld metal cool­ing curve of Fig. 3 indicates such a transformation temperature.

The measured values of the trans­formation temperatures were aver­aged in the same manner as the cooling rates. Those averages are also listed in Table 5 and are plotted as a function of the average weld metal cooling rate at 1300 F in Fig. 10. The linear variation in the aver­age transformation temperature w i th the average cooling rate was deter­mined by multiple linear regression analysis to be primari ly a function of the average cooling rate, but it was also a statistically signif icant ( 1 % level) secondary function of a com­position factor. For the weld upper-half that factor was the product

of Mn and Mo and for the weld low­er-half that factor was the product of C and Mn. Only the linear least squares line and two-sigma limits for the variation wi th average cooling rate are drawn in Fig. 10. It is reasonable that of the cooling rate and the transformation temperature, only the latter should be a function of composit ion, since the metal­lurgical reactions are strongly inf lu­enced by composition whereas the thermal properties are little inf lu­enced.

Effect of Cooling Rate and Composition on Mechanical Properties

The Charpy data are plotted as a function of the average cooling rate at 1300 F in Fig. 1 1 . The linear least squares lines fitted to the data and the two-sigma lines are also plotted in Fig. 11 . For the weld upper-halves, a linear multiple regression analysis including composit ion varia­bles indicated a slight additional ef­fect of carbon on the behavior of the notch toughness. However, since that effect was statistically marginal and since weld composit ion was not shown to correlate w i th any of the

process variables, no metallurgical significance was attached to the weak multiple linear correlation which was found.

The 0.2% offset yield strengths and the tensile strengths listed in Ta­ble 4 were also plotted as a function of the average cooling rates at 1300 F. Linear least squares lines were fit-j ted to both sets of data. The yield strength was a signif icant function ( 1 % level) only of the average cooling rate. In the case of the tensile strengths, the multiple linear regres­sion analyses indicated that a com­position factor, the product of the Cr and Ni contents, was a signif icant ( 1 % level) factor along w i th the aver­age cooling rate. However, no metal­lurgical interpretation was drawn from that observation.

Bead-On-Plate Welds

The bead-on-plate welds were made to obtain reference data on cooling rates and stringer bead dimensions una f fec ted by the presence of the groove. The cooling rates measured from the bead-on-plate welds corresponded well w i th those measured for the capping

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AVERAGE WELD METAL COOLING RATE AT 1300 °F (°F/SEC)

11 — Variation of the 40 ft-lb transition temperature the average weld metal center-line cooling rate meas-at 1300F

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TRAVEL SPEED (IN/MIN)

Fig. 12 — Effect of increasing travel speed on the penetration area and the ratio of the heat-affected zone area to the penetration area

passes of the multipass welds. Of the parameters measured from

the cross-sections of the bead-on-plate welds, only the "penetrat ion area," the area of fused metal below the plate surface, and the ratio of the penetration area to the area of the heat-affected-zone was of any direct importance to this investigation. Fig­ure 12 shows that the penetration area decreases linearly w i th increas­ing travel speed wi th lower currents naturally tending to produce less penetration area than higher cur­rents. When the heat-affected-zone area is divided by the penetration area, also shown in Fig. 12, the ratio is non-linear w i th travel speed.

Figure 12 suggests that weld beads deposited at lower levels of current tend to produce smaller proportions of columnar weld metal than do beads deposited at higher current levels; Fig. 12 also suggests that the welds deposited at lower current levels should tend to heat treat more of the previously de­posited structure in relation to the amount of columnar structure they create. As shown in a later part of this discussion, multipass welds made w i th lower levels of current tend to have less as-cast columnar metal, unreheated above the A , ,

than do those made at higher levels of current.

As shown in Fig. 13, there was a good correlation between the in­verse of the total nugget area, meas­ured from the bead-on-plate welds, and the average cooling rates of the lower and upper halves of the mult i ­pass welds.

Weld Macrostructure and Motch Toughness

A principal objective of this in­vest igat ion was to d e t e r m i n e whether improvements in impact properties could be attained by the heat treatment effects of the in­dividual passes. Despite the gen­erally recognized fact that increased heat input is detrimental to impact properties, there were several weld­ments welded at heat inputs in the 40—60 k J / i n . range which were as tough or tougher than welds made wi th heat inputs in the 25—40 kJ / i n . range. Also, w i th in each heat input range there were large tough­ness variations. An example of such toughness variation was the -118 F transition temperature of weld upper-half C6 (24.5 kJ / in . ; 400 amp; 25 ipm) compared to the -153 F transi­tion temperature of weld upper-half B3 (56.7 kJ / i n . ; 350 amp; 10 ipm).

Attempts to correlate toughness variations w i th the percentage of as-cast structure showed some useful results. The 40 ft- lb transit ion tem­peratures of the welds listed in Table 6 were plotted in Fig. 14 as a function of the percent as-cast struc­ture. For both the upper and lower halves, most of the data were in­cluded wi th in bands which exhibited decreasing toughness wi th increas­ing as-cast structure. Whi le the scat­ter is appreciable, it is less when this factor is considered than that ob­served when transit ion temperature is plotted as a function of heat input alone as in Fig. 6. The difference be­tween the lowest and highest values in the scatter bands of Fig. 14 is about 50 F. In Fig. 6 the same data points fall w i th in limits of about 75 F, or 25 F larger than the scatter bands of Fig. 14.

The transit ion temperatures, cool­ing rates, compositions, and per­centage of as-cast structure of the data points wi th in the limits of the two scatter bands of Figs. 6 and 14 were analyzed separately by linear multiple regression in an attempt quantitatively to determine the inf lu­ence of the various factors on transi­tion temperature. The results of those analyses were statistically in-

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AS-CAST STRUCTURE UNHEATED ABOVE THE A1

Fig. 14 — Effect of the percentage of as-cast structure not reheated above the A, on the 40 ft-lb transition temperature

conclusive. They did not quantify the 25 F reduction in scatter observed between Figs. 6 and 14 in terms of the programmed variables, in gen­eral, or the percentage of as-cast structure, in particular.

These results suggest that the detrimental effects of increased heat input on transit ion temperature might be partially relieved by choos­ing combinations of process varia­bles which result in a lower ratio of as-cast structure to metal reheated above the A, by subsequent passes. The bead-on-plate weld data indicate that over a wide range of travel speeds the 350 amp current level had less penetration area, a shal low­er depth of penetration, and a higher heat-affected-zone area to penetra­t ion area ratio than the 400 and 450 amp current levels. These factors combined to give the generally low values of percent as-cast structure, unreheated above the A , , present in welds made at 350 amp or less shown in Fig. 14.

If the data of Fig. 14 are grouped according to medium (40—60 kJ/ in. )

and low-to-medium (25—40 k j / i n . ) heat input, they suggest that the influence of the percentage of as-cast structure on the impact strength is stronger in the medium heat input range. An important deduction is that the heat treatment effects of subsequent passes benefit notch toughness more as heat input is in­creased. A similar conclusion can be deduced from the data of Horvath5

on plain carbon weld metals. He ran tests on half-size Charpy specimens from multipass and single pass welds. As shown in Table 7, in the range of heat inputs from 20 to 80 kJ / i n . there was no benefit from the heat treatment effects at the low heat input levels; however, as the heat input increased the benefits derived became appreciable.

It should also be noted that for al­most all of the weldments listed in Table 6, the weld lower-halves had a larger percentage of as-cast metal than their upper-half counterparts. Wi th the exception of weldment C2 (41.6 kJ/ in. ) the weld lower-halves were less tough than the correspond­

ing upper-halves. The upper-half of weldment C2 contained 14% more as-cast metal than the lower half and it was less tough than the lower half.

As a further step in the analysis of the impact properties the transition temperature, composition, cooling rate, and percentage as-cast struc­ture data from the upper and lower halves of Table 6 were grouped to­gether and analyzed by linear mult i­ple regression. The result of the anal­ysis was significant only w i th respect to cooling rate and the statistics of the regression analysis were less satisfactory than those obtained when the halves were analyzed separately. The statistical interpreta­tion of that result is that the upper and lower halves represented two separate populations.

Weld Microstructures Lyttle, et al,6 reported that the

cumulative effect of several thermal cycles, w i th peak temperatures in the range 2000 to 400 F, on as-de­posited (upper bainitic) microstruc-

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tures was to produce a tougher but weaker structure. The as-deposited structure was characterized by fer­rite laths w i th continuous carbide stringers at the boundaries and a heavily tangled high dislocation den­sity. After the cumulative thermal cy­cling the dislocation structure had changed to fixed arrays at sub-bound­aries and the carbide network was more randomly oriented and more agglomerated. Decreases in yield strength were attributed primarily to a change in the dislocation arrays and the increases in impact strength were attributed to changes in the car­bide dispersion. The authors did not, however, fol low the changes step-by-step wi th the thermal cycles, so it is unknown what peak temperature (s) caused the majority of the changes.

A more recent study7 indicates that the largest changes in the ten­sile properties of a weld metal s im­ilar to the weld metal of this study oc­curred when the as-deposited weld metal was cycled to a peak tempera­ture of 1900 F fol lowed by cycling to 1350 F. The as-deposited fully martensitic weld metal was changed to an acicular tempered martensit ic structure w i th agglomerated and spheroidized carbides in a ferrite matrix. Both studies indicated that changes not characterized by re­crystallization on a macro level were important to the overall mechanical behavior of multipass weldments.

Representative weld microstruc­tures taken from the 400 ampere ser­ies are shown in Fig. 15. The areas chosen were the as-cast weld metal in a penetration area in the weld up­per-half and the recrystallized weld metal directly below the location of the first area. Because of the com­plexity and non-uniformity of the microstructures, no attempt was made to quantitize the structures in a manner which could be used to re­late microstructure to the observed mechanical properties. However, several qualitative observations can be made.

As the travel speed was increased and the heat input decreased, the as-cast structure changed from the coarse structure shown in Fig. 15A to the finer structure shown in Fig. 15E. Also, the amount of pro-eutectoid ferri te decreased as the heat input decreased. In Fig. 15A, the proeutectoid ferrite appears as the white elongated clearly defined regions interspersed throughout the microstructure. In the weld of Fig. 15D, the average cooling rate was 149 F/sec, rapid enough to force fer­rite to precipitate in the widmanstat­ten form. In the weld of Fig. 15C, where the cooling rate is 106 F/sec, the proeutectoid ferrite still tends to precipitate in a more blocky form.

The remainder of the structure changes from ferr i t ic /upper bainitic in nature, as in Fig. 1 5A, to one con­sisting primari ly of lower bainite, in Fig. 15E. The transformation tem­perature (as defined in Procedure) decreased from 927 F for the weld metal shown in Fig. 1 5A, to approx­imately 800 F for those shown in Figs. 15C-15E.

Wi th the possible exception of an increased tendency for accentuation of prior austenite grain boundaries and a general coarsening of the structures, the structures in the re­crystallized areas are similar to their counterparts in the as-cast regions. The structures of the weld lower-halves, wh ich have not been in­cluded, appeared f iner than their counterparts in Fig. 15. This observa­t ion may be due to the more rapid cooling rates in the weld lower-halves, or to the difference in com­position, or a combination of the two.

No photomicrographs of the re­gions wi th in the A3 to A , tempera­ture range have been included since, in general, they are similar in ap­pearance to either the recrystallized or as-cast material. Without prior knowledge of location, the structures at X500 throughout the weld deposit are so similar that it is virtually im­possible to identify the origin of any given structure. It is this aspect of these structures which would make quantitative analysis of the structural constituents tedious and of doubtful value.

Interrelation of Variables and Properties

Wi th respect to the impact and ten­sile properties, the average cooling rate was shown to exert a f irst-order influence. For a constant composi­t ion, it is the cooling rate which determines the nature of the trans­formation product, its f ineness or ef­fective grain size, and the amount of proeutectoid ferrite precipitated, if any. In general, the variations en­countered in composition produced only a second-order effect on the continuous cooling transformation behavior and the metallurgical char­acteristics of the transformation products. As the cooling rate increased, the amount of pro­eutectoid ferrite precipitated de­creased and became less blocky in appearance. These changes were ac­companied by an almost continuous improvement in notch toughness. The matrix coarseness and columnar appearance also decreased.

It is known that metallurgical changes such as alterations of dislocation structure, changes in car­bide morphology, and diffusion of ele­ments segregated in cell boundaries into the matrix affect the properties. The extent of these alterations is a

function of the thermal cycles pro­duced by the subsequent passes. Changes such as grain size and morphology resulting from the ther­mal cycles can be observed in these weld metals wi th the naked eye. As the magnification required to detect the structural details increases, there is a concomitant decrease in the area of observation, wh ich limits the investigation and characteriza­t ion of the properties of multipass weldments. This study has shown, however, that certain readily ob­servable structural features of mult i ­pass weldments may be of use in interpreting their mechanical be­havior over a wide range of currents and travel speeds at commercial levels of heat input.

Variations in notch toughness associated w i th the percentage of as-cast structure observed at low magnification (X15) were small in magnitude. It was concluded, there­fore, that the character of the fine l ow-ca rbon b a i n i t i c / m a r t e n s i t i c microstructures initially formed were not appreciably altered by the ther­mal cycling of the subsequent passes, despite the fact that from 56 to 80% of the total structure was reheated above the A temperature and that from 28 to 64% of the total structure was heated to a high enough temperature long enough to form equiaxed grains from the as-cast columnar grains.

An implication of this investigation is that improvement of strength in combination wi th notch toughness of the present generation of com­mercially available h igh-st rength steel weld metals wi l l have to come from the development of new weld­ing processes (or a refinement of existing processes) as wel l as control of composition. The importance of that implication increases wi th re­spect to the development of future generations of base metal-weld metal systems, wi th acceptable levels of notch toughness, at yield strength levels greater than 140 ksi.

S u m m a r y

The results of the investigations are summarized as fol lows:

1. The cooling rates of t he individual passes of the multipass welds were sensitive to their loca­tion relative to the bottom of the groove and the side-walls. Higher cooling rates were observed for passes made at the bottom or near the sidewalls compared to those made nearer the center and closer to the top. For compositions particularly sensitive to cooling rate, these varia­tions can increase the nonuniformity of structures in a multipass weld­ment.

W E L D I N G R E S E A R C H S U P P L E M E N T ! 519 -s

Table 6—Weld Metal Structural Feat from Photomacrographs

Weld designation

G1 A2

B3

F3

B2

B1

C3

C2

C1

C6

D2

D1

D5

Section location in

weld (upper or lower)

U U L U L U L U L U L U L U L U L U L U L U L U L

ures Measured

Totally re­crystall ized.

% 40.9 43.3 39.0 46.6 38.2 32.6 40.0 44.2 37.9 29.3 31.1 39.7 63.9 32.8 39.4 37.4 40.1 42.2 37.3 34.1 32.5 29.4 33.6 31.1 27.9

As-Cast (columnar

and not reheated

above A , )

% 17.1 20.5 34.9 20.0 28.7 29.4 34.4 22.5 27.9 27.7 36.3 38.0 16.8 44.7 30.6 27.1 22.1 19.8 32.1 23.6 35.0 32.7 28.9 39.4 34.9

40 ft-lb transit ion

temperature °F

-121 - 1 1 6 - 71 - 1 5 3 - 1 1 6 - 1 3 4 - 77 - 1 4 9 - 1 1 9 - 1 6 4 -111 - 76 - 57 - 102 - 1 2 3 - 1 3 5 - 1 0 9 - 1 1 8 - 1 0 3 - 1 2 4 - 58 - 1 3 4 - 1 0 6 -121 - 87

Table 7—Half-Size Charpy 15-Mi l Transition Temperature — from Horvath5

V2-size Charpy 15 mil t ransit ion temperature, °F

Heat input, k j / i n .

20 40 60 80

Current, amp

260 260 350 350

Travel speed,

ipm

20 10 10

7

Mu l t i ­pass

- 7 0 - 4 8 - 4 4 - 5 2

Single pass

- 8 0 - 6 0 - 3 2 - 1 5

AT by multipass

+10 + 12 - 1 2 - 3 7

structure and the notch toughness more than did changes in welding current.

6. The weld lower-halves were generally less tough than their upper-half counterparts despite their faster average cooling rates. The lower notch toughness of the lower-halves is attributed to the generally larger percentage of as-cast structure and to increased percentages of C, Cr, and Mo. The compositional dif­ferences observed between the up­per and lower halves were a conse­quence of the compositional dif­ferences between the electrode and base plate.

7. The variat ion in toughness associated wi th the percentage of as-cast structure was small in magni­tude. It was concluded that the prop­erties of the f ine low carbon ba in i t ic / martensitic microstructures initially formed were not appreciably altered by the thermal cycling of the subse­quent passes despite a large fraction of structure reheated above the A, and an appreciable refinement of the original columnar grains.

Acknowledgements

The chemical analyses were performed at the Paul D. Merica Research Lab­oratory of the International Nickel Co. under the supervision of Dr. J . H. Weber. This program was supported by grants from the National Science Foundation and the Welding Research Council. The research work was completed prior to the affi l iation by one of the authors w i th the Department of the Army, thus the opin­ions expressed herein are solely those of the authors and are not to be construed as official or reflecting the views of the Army.

References

2. The average cooling rates cal­culated for the weld upper and lower halves and corresponding to the Charpy specimen locations varied wi th travel speed, current, and the in­verse of the heat input, in a manner similar to that predicted by heat f low theory and determined experimental­ly for the cooling rates of single-pass welds on semi-inf inite plates. Ir­respective of heat input, the average cooling rate of the weld lower-half was approximately 25% faster than the average cooling rate of the weld upper-half.

3. The transit ion temperatures showed a consistent linear correla­tion wi th the average weld-metal center-l ine cooling rates measured at 1300 F. There was about a 55 F improvement in t he t r a n s i t i o n temperatures of both the lower and upper weld halves as the average cooling rate increased from 40 F/sec (high heat input) to 200 F/sec (low heat input). The scatter observed in the plots of transit ion temperature

versus average cooling rate is attrib­uted, in part, to the variations of microstructure which resulted from the pass-to-pass variation of cooling rate.

4. From metallographic analysis at low magnification (X15), it was found that as the percentage of as-cast columnar metal not reheated above the A , temperature by the subse­quent passes increased, the tough­ness decreased. The relative effect on toughness of the as-cast structure was greater at medium heat input levels (40—60 kJ/ in. ) than at low levels (25—40 kJ/ in.) .

5. At medium heat input levels (40—60 kJ/ in.) , combinations of welding process variables which re­sulted in small penetration areas, shallower depths of penetration, and a high ratio of heat-affected-zone area to penetration area, generally produced low values of percent as-cast structure in the multipass weld­ments. Changes in travel speed af­fected the percentage of as-cast

1. Stout, R. D., McLaughl in, P. F., and Strunck, S. S., "Heat Treatment Effects in Multiple-Pass Welds-Part I," We/ding Journal, 48 (4), Research Suppl., 155-s to 160-s, 1969.

2. Strunck, S. S., "Heat Treatment Ef­fects In Mult i-Pass Weldments Of A High-Strength Stee l , " Ph.D. Dissertation, Le­high University, 1971 .

3. Dorschu, K. E., "Control of Cooling Rates in Steel Weld Meta l , " Welding Journal, AT (2), Research Suppl., 49-s to 62-s, 1968

4. Nakabayashi, M., "Effects of Weld­ing Parameters on Geometry and Re­crystallization of Weld Zone," M. S. The­sis, Lehigh University, 1970.

5. Horvath, P. J. , Jr., "The Influence of Weld Metal Impact Dissertation, Lehigh

Microstructure on Propert ies," Ph.D. University, 1964.

6. Lyttle, J . E., Fragetta, W, A.,

Dorschu, K. E., and "Some Metal lurgical

Characteristics of Tough, High Strength Welds," Welding Journal. 48 (11), Re­search Suppl., 493-s to 498-s, 1969.

7. Krantz, B. M., "Factors Affect ing the Strength Of Multi-Pass Low-Al loy Steel Weld Meta l , " Ibid., 50 (6), Research Suppl., 2 3 5 - s t o 2 4 1 - s , 1971 .

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