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Introduction

In the 21st century, the world faces thecritical challenge of providing abundant,cheap electricity to meet the needs of agrowing global population while at thesame time preserving the environment.Worldwide research and development ef-forts have been conducted in the lastdecades to develop and qualify high-temperature materials technology for ul-trasupercritical boiler (USCB) power gen-eration. The goal of the USCB programsis to increase boiler efficiencies andachieve lower emission levels. In theU.S.A., a program funded by the U.S. De-partment of Energy (DOE) and OhioCoal Development Office (OCDO) hasthe goal to develop material technologycapable of operating with steam tempera-tures of up to 732°C (1350°F) and pres-sures as high as 35 MPa (5000 lb/in.2) (Ref.1). These service conditions translate intohigher hot gas path temperatures and the

need for materials possessing the requiredstrength, heat, and corrosion resistancefor long-term, reliable service at highertemperatures. Additionally, the weldabil-ity of these base materials or ability to pro-duce crack-free joints with acceptable me-chanical properties is equally important asit has implications on original fabricationand repair.

A nickel-based alloy, Inconel® Alloy740 (IN740) (Ref. 2), has been identified inthe U.S.A. as a potential candidate for thesevere operating conditions of USC boilersbased on its resistance to fireside corrosionresistance, steamside oxidation, and goodstress-rupture properties (Refs. 3–5). Alloy

IN740 was developed as part of a similarprogram in Europe known as THERMIE.Alloy IN740 is a precipitation-hardenable,nickel-chromium-cobalt alloy with nio-bium, titanium, and aluminum additions. Itis primarily strengthened through precipita-tion of gamma-prime phase (γ’) and sec-ondarily through the precipitation of car-bides. Alloy 740 is a new alloy with nofabrication history that was derived fromNIMONIC® Alloy 263. Therefore, theweldability of this new precipitation-hard-enable alloy needed to be evaluated as partof the material selection criteria for appli-cation in USC boilers for power generation.

In general, nickel-based precipitation-hardenable alloys are susceptible to liqua-tion cracking during welding and postweldheat treatment (PWHT) cracking, alsoknown as reheat cracking or strain-agecracking, in the heat-affected zone (HAZ).Additionally, application of IN740 in theUSC boilers requires wall thicknesses of upto 3 in. The high level of restraint in thickjoints not only increases the susceptibility toliquation and PWHT cracking, but alsomakes the HAZ susceptible to ductility-dipcracking (DDC) during welding. Finally,the mechanical properties of Alloy IN740welded joints are expected to depend on the“as-received” condition of the base metal,welding procedures, and pre- and postweldheat treatments.

The conducted evaluation of the weld-ability of Alloy IN740 included as follows:(1) HAZ liquation cracking susceptibilityand HAZ DDC susceptibility as reportedin this paper; (2) PWHT cracking suscep-tibility of Alloy IN740 and comparisonwith other common nickel-based precipi-tation-hardenable alloys such as IN718and Waspaloy™; and (3) optimization ofweldability or cracking resistance of AlloyIN740 as function of chemical composi-tion along with mechanical properties ofGleeble-simulated HAZs of IN740 andfull-sized IN740 welded plate joints in thepostweld heat treated condition.

Liquation cracking is a high-tempera-ture cracking phenomenon that occurs in

Susceptibility of IN740 to HAZ LiquationCracking and Ductility-Dip Cracking

The susceptibility of Ni-Cr-Co Alloy IN740 to heat-affected zone liquation crackingand ductility-dip cracking was evaluated and correlated to its

solidification behavior

BY J. E. RAMIREZ

KEYWORDS

HAZ Liquation CrackingDuctility-Dip Cracking (DDC)Alloy IN740Ultrasupercritical (USC) BoilersNil-Ductility Range (NDR)

J. E. RAMIREZ (jramirez@ewi.org) is principalengineer, Edison Welding Institute, Columbus,Ohio.

ABSTRACT

Evaluation of susceptibility to liquation cracking and ductility-dip cracking (DDC) ofAlloy IN740 was conducted as part of the material selection criteria for application in ul-trasupercritical (USC) boilers for power generation. The susceptibility to heat-affectedzone (HAZ) liquation cracking and DDC was determined using Gleeble hot-ductilitytesting in the temperature range from 600° to about 1300°C and spot-Varestraint testing.The large on-cooling, nil-ductility range (NDR) of 229°C observed in Alloy IN740 indi-cates that the tested heat of this alloy is highly susceptible to liquation cracking in theHAZ. Welding conditions inducing shallower thermal gradients in the HAZ increase thesize of the HAZ liquation cracking susceptibility region. Phase computational modelingindicates that segregation of B, Nb, and Si during solidification of Alloy IN740 resultsnot only in the increase in the amount of low melting equilibrium phases such as MC andMB2 but in the formation of low-melting, nonequilibrium phases such eta, Laves, and Gphases. Therefore, the solidus temperature of the system is depressed to 1060°C. The re-sulting increase of the solidification range from 169.7° to 293.5°C of the Alloy IN740 heattested explains the large NDR of 229°C observed in the alloy heat tested and its associ-ated high susceptibility to HAZ liquation cracking. A pronounced loss or ductility dipwas not observed in the Gleeble hot-ductility tested alloy heat in the temperature rangefrom 600° to 1050°C. Therefore, Alloy IN740 is not susceptible to DDC cracking.

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the HAZ region adjacent to the weld inter-face. This type of cracking is particularlyprevalent in nickel- and aluminum-based al-loys, plus fully austenitic stainless steels.The metallurgical basis for liquation crack-ing is the presence and persistence of liquidfilms at grain boundaries and the inability ofthose films to accommodate the thermallyand/or mechanically induced strain experi-enced during weld cooling.

Ductility-dip cracking is a solid-statecracking that occurs intergranularly in weldmetal and the HAZ. Although its exactmechanism has not been determined, it isassociated with a ductility loss in the tem-perature range from one-half the absolutemelting temperature up to the recrystalliza-tion or solidus temperature. This drop in el-evated temperature ductility may occur ei-ther on heating to an elevated temperatureor cooling from some peak temperature.Most of the materials susceptible to DDCare single phase from room temperature tothe solidus temperature and tend to solidifyas a face-centered cubic (FCC) singlephase. In austenitic stainless steels andnickel-based alloys, there appears to bestrong grain size dependence with coarsergrain sizes promoting a higher susceptibil-ity. Ductility-dip cracking is often associ-ated with welding heavy sections commonlyencountered in the power-generation in-dustry. It has also been identified in otherthermomechanical processes such as hotrolling and forging.

Figure 1 shows a schematic representa-tion of the two temperature ranges inwhich low ductility is conducive to crack-ing during welding. Temperatures near tothe solidus temperature and temperaturesabove half the melting temperature corre-spond to the temperature ranges whereHAZ liquation cracking and HAZ DDC,respectively, can take place.

Materials and ExperimentalProcedures

Base Materials: The Alloy IN740 barstock used for the weldability evaluation

was initially melted using vacuum induc-tion and subsequently electroslagremelted. The cylindrical cast (~21.0 in. indiameter) was then forged to 11.65 in. indiameter at 1180°C. The resulting 11.65-in.-diameter billet stock was then extrudedat 1180°C to 3- × 6-in. bar. Finally, the barwas annealed at 1121°C (2050°F) andwater quenched. A reannealing treatmentof the bar at 1121°C for 3 h prior to waterquenching was necessary due to high hard-ness in the as-produced bar. The providedmaterial was identified as Heat BLT2819.The chemical composition of this heat of

material is listed in Table 1.Evaluation of HAZ Liquation Crack-

ing Susceptibility: The susceptibility toHAZ liquation cracking of Alloy IN740was determined using Gleeble hot-ductil-ity testing and spot-Varestraint testing.

Gleeble Hot-Ductility Testing: TheGleeble hot-ductility test was used to char-acterize the ductility (reduction in area) ofthe material as function of temperatureand to relate ductility recovery to HAZ li-quation cracking susceptibility. Basically,small tensile samples were pulled to frac-ture rapidly in the vicinity of the solidus

Fig. 1 — Schematic illustration of the temperature ranges where liquationcracking and DDC are observed in the HAZ of base metals during welding.

Fig. 2 — MCD as a function of cooling time using saturation strain duringspot-Varestraint testing.

Table 1 — Chemical Composition of Alloy IN740 (Heat BLT2819) Used for Testing

Element C Ni Cr Co Nb Ti AlWt-% 0.029 Balance 24.4 20.0 2.0 1.8 1.0Element Si Mn P S Ag B BiWt-% 0.53 0.26 <0.005 <0.001 <0.1 0.0045 <0.1Element Cu Fe Mo Pb V Zn ZrWt-% 0.02 0.45 0.50 0.6 ppm <0.01 1 ppm 0.225

Table 2 — Conditions of Gleeble Hot-Ductility Testing of IN740 Alloy

On-HeatingParameter Description

Test temperature 950°C and higher (1112°F and higher)Heating rate to test temperature 111°C/s (200°F/s)Holding time at test temperature 0.03 sStroke rate 5 cm/s (2 in./s)Sample-free span (jaw spacing) 19 mm (0.75 in.)Atmosphere Argon

NST Determination

Load Just enough to overcome frictional force [≈20 kg (45 lb)]

Number of samples 3

On-Cooling

Peak temperature NSTHeating rate to peak temperature 111°C/s (200°F/s)Holding time at peak temperature 0.03 sCooling rate to test temperature 40°C/s (72°F/s)Holding time at test temperature 0.03 sStroke rate 5 cm/s (2 in./s)Sample-free span (jaw spacing) 19 mm (0.75 in.)Atmosphere Argon

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temperature during the on-heating andon-cooling portion of a simulated weldthermal cycle in a Gleeble machine. Thetransverse reduction in area of the frac-ture samples was determined, therebyproviding a measure of the ductility of thematerial at high temperature.

All testing was performed under anargon atmosphere using a Gleeble® 1000system. After reaching a vacuum level of40 millitorr, the chamber was backfilledwith argon. The thermal cycles experi-enced by the samples during testing were

measured using fine wire K-type thermo-couples. These thermocouples were per-cussion welded to the samples. Immedi-ately after the thermocouples wereattached, the specimen was clamped be-tween water-cooled jaws, which in addi-tion to serving as grips, provide a meansfor introducing the current to the speci-men for heating and ensuring a rapid cool-ing when the current flow is stopped.

The conditions for the Gleeble hot-duc-tility testing are given in Table 2. The testingtemperatures were 950°C and higher. Stan-

dard specimens 6.35 mm in diameter and100 mm long (0.25 × 4 in.) were used.

Spot-Varestraint Testing: The spot-Varestraint test was used to determine thesize of the crack susceptibility region in theHAZ as function of welding conditions.During spot-Varestraint testing, a gastungsten arc welding (GTAW) spot weldwas produced in the center of a small spec-imen. After a predetermined weld time,the arc was extinguished and the specimenwas forced to conform to the surface of aradiused die block. The strain imposed onthe specimen was controlled through theradius of the die block and the thickness ofthe specimen. In this manner, HAZ liqua-tion cracks were generated on the surfaceof the specimen adjacent to the GTA spotweld. The spot-Varestraint samples were 8× 1.5 × 0.25 in. in size, and three spots weredeposited in each specimen. The parame-ters used during spot-Varestraint testing ofIN740 are listed in Table 3. The number ofcracks and maximum crack distance(MCD) were measured in each sample asa function of augmented applied strain.

In order to evaluate the susceptibilityto liquation cracking of IN740 during theon-heating and on-cooling parts of thethermal cycle, the cracking susceptibilityenvelope ABC as shown schematically in

Fig. 3 — Gleeble hot-ductility of Alloy IN740 in the liquation cracking temperature range. A — Fracture strength; B — reduction of area as a function of test-ing temperature.

A B

Fig. 4 — Spot-Varestraint testing results of IN740 during the on-heating partof the cycle.

Fig. 5 — Comparison of the behavior of IN740 as a function of arc-on timeof the spot-Varestraint testing.

Table 3 — Parameters Used during the Spot-Varestraint Testing

Parameter Description

Welding process GTAW, DCENElectrode 3⁄32-in. diameter, 2%CeElectrode preparation 30-deg included angle, 0.02-in. bluntElectrode extension (beyond cup) 0.25 in. (6.35 mm)Arc gap 0.080 in. (2.0 mm)Voltage 11 ± 0.5 VCurrent 150 ± 5 AArc-on time 10 and 20 sShielding gas Argon at 30 ft3/hShielding gas cup size 7Bending rate 8 in./sHold time after bending 10 sSpot diameter 11 mm

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Fig. 2 was determined. This required thetesting of specimens at different levels ofstrain and at zero cooling time (the timebetween arc extinction and specimenbending) until the saturation strain isreached (section AB of the envelope). Thesaturation strain is the strain above whichthe MCD remains constant. After that,specimens were tested using the satura-tion strain and different cooling times todetermine the remaining part of the enve-lope (section BC). Section AB character-izes the susceptibility of the material dur-ing the on-heating part of the cycle.Section BC characterizes the susceptibil-ity of the HAZ to cracking during the on-cooling part of the welding cycle.

Additionally, the effect of differentwelding conditions on the dimensionalcracking susceptibility envelope was char-acterized by changing the arc-on timefrom 10 to 20 s during the on-cooling spot-Varestraint testing.

Evaluation of HAZ DDC Susceptibil-ity: The susceptibility to HAZ DDC ofIN740 was evaluated using Gleeble hot-ductility testing. The same procedure,specimen size, and geometry along withmeasurement of ductility as described forthe Gleeble hot-ductility testing in theHAZ liquation cracking evaluation wereused for the evaluation of DDC cracking.For DDC evaluation, the small tensilesamples were fractured rapidly during theon-heating and on-cooling portions of asimulated weld thermal cycle in a Gleeblemachine at specific temperatures in the600° to 1000°C temperature range.

Experimental Results andDiscussions

HAZ Liquation CrackingSusceptibility

Gleeble Hot-Ductility Testing: The re-sults of the Gleeble hot-ductility testing ofAlloy IN740 are listed in Table 4 andshown in Fig. 3. As part of the evaluationof the susceptibility to liquation crackingof Alloy IN740, the nil-ductility tempera-ture (NDT), nil-strength temperature(NST), and ductility recovery temperature(DRT) were determined. As shown in Fig.3B, the IN740 material loses all ductilitywhen the testing temperature reaches1205°C (NDT) on heating. The ductilityremains low at higher temperatures andthe fracture strength decreases to almostzero as the testing temperature is in-creased to 1280°C (NST), as shown in Fig.3A. On cooling, after being exposed to apeak temperature, the alloy starts recov-ering ductility at 1050°C (DRT). There-fore, there are two temperature rangeswithin which the material exhibits negligi-ble ductility (ND), the on-heating NDtemperature range, and the on-cooling

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Table 4 — Results of Hot-Ductility Testing during On-Heating and On-Cooling of IN740 Alloyin the Liquation Cracking Temperature Range

Specimen Temperature (°C) RA (%) FractureNumber ID Peak Test Strength (ksi)

HT-17(a) HT950 – 952 82.2 60.3HT-8 HT1000 – 1000 84.9 55.7HT-1 HT1100 – 1099 77.7 40.0HT-2 HT1150 – 1149 68.3 33.0HT-3 HT1200 – 1200 13 27.7HT-5 HT1200 – 1200 2 22.6HT-6 HT1205 – 1205 3.6 24.8HT-7 HT1210 – 1210 1.1 2.1HT-4 HT1210 – 1211 0 19.1

HT-13 HT1220 – 1222 0 19.1HT-14 HT1240 – 1243 0 15.4NST-1 NST-1 – 1281 0 0.9NST-2 NST-2 – 1282 0 0.9NST-3 NST-3 – 1271 0 0.9NST-4 NST-4 – 1284 0 0.9

CT-26(b) CT1240 1272 1240 0 3.4CT-22 CT1200 1271 1200 0 2.7CT-23 CT1175 1271 1174 0 2.4CT-17 CT1150 1271 1150 0 1.3CT-18 CT1120 1268 1120 0 2.4CT-25 CT1110 1272 1111 4.8 30.8CT-16 CT1100 1271 1100 6.2 19.9CT-19 CT1100 1272 1099 1.4 2.1CT-28 CT1075 1273 1077 3.2 2.8CT-29 CT1075 1273 1077 3.0 3.1CT-15 CT1050 1271 1049 5.9 38.9CT-20 CT1050 1268 1050 0 20.0CT-24 CT1025 1272 1023 33.0 50.1CT-27 CT1025 1272 1024 44.3 53.4CT-1 CT1000 1267 999 11.5 2.1CT-14 CT1000 1271 1001 46.7 56.8CT-21 CT1000 1271 1000 46.4 55.7

(a) H: On-heating test specimens; T: Test temperature.(b) C: On-cooling test temperature; T: Test temperature.

Fig. 6 — Schematic representation of the hot-ductility response of the HAZ.

Table 5 — NST of the IN740 Alloy

Specimen NST (°C) Average NST (°C)

1 12812 12823 1271 1279.54 1284

Table 6 — NDT, NST, and DRT Temperaturesof IN740 Alloy

NDT (°C) NST (°C) DRT (°C) (NDR) (°C)

1205 1279 1050 229

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ND temperature range.The susceptibility to HAZ liquation

cracking of Alloy IN740 was evaluated byusing the on-cooling nil-ductility tempera-ture range criteria (NDR). This is definedas the temperature difference between the

peak temperature of the thermal cycle,which corresponds to the NST found duringthe on-heating portion of the simulatedweld thermal cycle, and the on-coolingDRT, the temperature at which the ductilitystarts to recover after exposure to the NST.

The NDR indicates how fast the materialrecovers its ductility during cooling. A fasterrecovery of the ductility or smaller NDR in-dicates lower susceptibility to HAZ liqua-tion cracking under similar welding condi-tions and, therefore, better weldability. TheNST temperature obtained from differentsamples of Alloy IN740 are listed in Table 5.The NDT, NST, and DRT temperatures andthe NDR determined for Alloy IN740 arelisted in Table 6.

Spot-Varestraint Testing: The resultsof the spot-Varestraint testing during theon-heating part of the cycle are listed inTable 7 and shown in Fig. 4. The saturationstrain for Alloy IN740 is about 4% asshown in Fig. 4. The maximum crack dis-tance (MCD) was used as an index to char-acterize the size of the crack susceptibilityregion and, therefore, the susceptibility toliquation cracking. The MCD at satura-tion strain was about 0.66 mm. The resultsof the spot-Varestraint test during the on-cooling part of the cycle are listed in Table8 and shown in Fig. 5.

The size of the crack susceptibility re-gion determined by the spot-Varestrainttesting is related to the NDR determinedwith the hot-ductility Gleeble testing, and

Fig. 7 — A — Example of microfissures observed in cross sections of the IN740 full-size welded joint; B — general fractographic characteristics of Gleeble on-cooling hot-ductility sample showing the presence of liquated phase remaining in the fracture surface.

A B

Fig. 8 — Predicted segregation of elements in the liquid as a function of thefraction solid.

Fig. 9 — Predicted partitioning of Nb into different phases as function ofthe fraction solid.

Table 7 — Results of Spot-Varestraint Tests during the On-Heating Part of the Welding Cycle(0-s Cooling Time)

Sample ID Augmented Strain (%) Number of Cracks MCD (mm) Average MCD (mm)

0-0.25-1(a) 0.25 13 0.38 0.420-0.25-2 0.25 12 0.460-0.5-1 0.5 19 0.48 0.470-0.5-2 0.5 20 0.460-1-1 1 33 0.51 0.510-1-1 1 35 0.510-2-1 2 39 0.56 0.560-2-2 2 38 0.560-3-1 3 41 0.64 0.630-3-2 3 43 0.610-4-1 4 45 0.66 0.660-4-2 4 44 0.660-5-1 5 48 0.66 0.660-5-2 5 49 0.660-7-1 7 54 0.66 0.66

(a) Cooling time – percent strain – specimen number.

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their relationship depends on the temper-ature profile induced in the HAZ by thewelding conditions. Figure 5 also showsthe change in the size of the liquationcracking susceptibility envelope of IN740as a function of the arc-on time. Increas-ing the arc-on time from 10 to 20 s in-creased the size of the susceptibility enve-lope by inducing a shallower thermalgradient in the HAZ, a wider HAZ, ascompared to a shorter arc-on time condi-tion. The same effect on the thermal gra-dient in the HAZ would be induced by in-creasing the heat input or by increasingpreheat and/or interpass temperature dur-ing welding.

Discussion of Observed HAZLiquation Behavior

Macroscopic Behavior: A schematicthree-dimensional representation of thehot-ductility response of the HAZ of analloy near the solidus temperature isshown in Fig. 6. Regions in the HAZ areinitially heated and then cooled as theweld pool passes. The peak temperaturewill depend on the distance from the weldinterface, decreasing as distance from theweld interface increases. For the condi-tions shown in Fig. 6, a point at a distanceD4 from the weld interface will be heatedto a peak temperature of T4 and shows noloss in ductility on heating or cooling. Apoint at a distance D3 from the weld in-terface will reach a peak temperature ofT3. At this point in temperature, ductilitydrops to zero but recovers immediatelyon cooling. At a distance D2 from theweld interface, peak temperature is T2.However, ductility drops to zero on heat-ing at T3 and does not return until a sig-nificant amount of cooling below T2 oc-curs. Closer to the weld interface, andconsequently at higher peak tempera-ture, the zero-ductility region (ZDR) be-comes larger.

Additionally, during welding, materialtoward the front side of the weld pool isheated and expands, creating compressive

stresses. As the weld poolmoves in the welding direc-tion, the material moving to-ward the back of the weldpool cools down and con-tracts, compressive stressesare relaxed, and tensionstresses are developed. As il-lustrated in Fig. 6, adjacentto each moving weld poolthere is a zero-ductilityplateau. Therefore, thestrain required to accommo-date the generated tensilestresses may eventually reach

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Fig. 10 — Predicted fraction of phases present in the solid as a function oftemperature during solidification of the tested IN740 alloy heat.

Fig. 11 — Gleeble hot-ductility tests of Alloy IN740. A — Fracturestrength; B — reduction of area as a function of testing temperature.

Table 8 — Results of Spot-Varestraint Tests during the On-Cooling Part of the Welding Cycle atSaturation Augmented Strain (4%)

Sample ID Cooling Number of MCD Average MCD Time (s) Cracks (mm) (mm)

10-s Arc-on time

0-4-1(a) 0 45 0.660-4-2 0 44 0.66 0.660.05-4-1 0.05 42 0.380.05-4-2 0.05 43 0.38 0.380.1-4-1 0.1 31 0.30.1-4-2 0.1 38 0.33 0.3150.2-4-1 0.2 26 0.250.2-4-2 0.2 25 0.28 0.2650.25-4-1 0.25 25 0.250.25-4-2 0.25 22 0.25 0.250.3-4-1 0.3 21 0.130.3-4-2 0.3 14 0.25 0.190.4-4-1 0.4 5 0.150.4-4-2 0.4 0 0 0.070.5-4-1 0.5 0 00.5-4-2 0.5 0 0 0.01-4-1 1.0 0 0 0

20-s Arc-on time

0-4-1 0 49 0.89 0.890.3-4-1 0.3 23 0.38 0.380.5-4-1 0.5 18 0.28 0.280.8-4-1 0.8 7 0.2 0.21.0-4-1 1.0 1 0.13 0.131.2-4-1 1.2 0 0 0

(a) Cooling time – percent strain – specimen number.

A

B

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the strain capability of the HAZ. Hence, thesusceptibility to liquation cracking of analloy will increase with the size of the nil-ductility plateau.

The maximum size of the nil-ductilityplateau of heat BLT2819 of Alloy IN740 is

characterized by the on-cooling NDR(229°C) determined through the Gleeblehot-ductility testing and the MCD at satu-ration strain (0.66 mm) determinedthrough the spot-Varestraint testing. TheNDR and MCD represent the tempera-

ture difference and distance, respectively,between the isotherms and location of theweld interface and the tip of the liquationcracks in the HAZ. The large NDR ofabout 230°C observed in Alloy IN740 in-dicates that this alloy is highly susceptibleto liquation cracking in the HAZ. Table 9shows a comparison of the susceptibility toliquation cracking of different alloys de-termined by Gleeble hot-ductility testingbased on the NDR criteria as reported inthe literature (Refs. 6–10).

The actual size of the dimensional en-velope of the nil-ductility plateau ob-served in the HAZ of a particular materialdepends on the thermal profile induced bywelding conditions (heat input, preheat,etc.), as shown in Fig. 5 for Alloy IN740.Welding conditions creating a shallowerthermal profile will increase the size of thecracks (nil-ductility enveloped) as the dis-tance between the isotherms of the weldinterface and the tip of the liquationcracks in the HAZ will increase.

Reported welding field testing ofIN740 (Ref. 11) indicates the use of dif-ferent combinations of welding processesincluding pulsed gas metal arc welding (P-GMAW) and cold- and hot-wire GTAW,shielding gases like argon and argon-helium mixtures, welding wires, heat in-puts, single-V and U-narrow groove jointbevel designs, number of weld passes tocomplete the joint by changing the beadsize, and plate thicknesses up to 1.0 in. todevelop and optimize the welding proce-dures. Even though some of those weldingprocedures showed promising results inthin-walled applications, most of the weld-ing techniques that were tested resulted insome degree of liquation cracking in theHAZ of Alloy IN740, as shown in Fig. 7.Additionally, variation of the degree ofcracking or number of microfissures pres-ent in the welded joints from heat to heatof Alloy IN740 was also reported.

The reported welding results confirmsthe high susceptibility to HAZ liquationcracking of heat BLT2819 as determinedby Gleeble hot-ductility and spot-Vare-straint testing. Additionally, it also indi-cates that if an alloy such as IN740 is verysusceptible to HAZ liquation cracking,the window of opportunity to develop a ro-bust welding procedure that allows a nor-mal welding parameter variation found ina production environment without induc-ing HAZ liquation cracking is very lim-ited. Therefore, a better understanding ofthe fundamental causes of the observed li-quation cracking susceptibility of AlloyIN740 is of outmost importance and fur-ther analysis of the solidification behaviorof Alloy IN740 was conducted.

Microscopic Behavior: Liquid metal isinvariably associated with the HAZ be-cause the HAZ stretches from the basemetal to the edge of the fusion zone and

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Table 9 — Comparison of the Susceptibility to Liquation Cracking of Different Alloys as Determined by the Gleeble Hot-Ductility Testing using the NDR Criteria

Alloy NDT (°C) NST (°C) DRT (°C) (NDR) (°C)

IN740 (Ref. 6) 1205 1279 1050 229800H (Ref. 6) – 1365 – 115AC66 (Ref. 6) – 1331 – 94926 (Ref. 6) – 1315 – 40825 (Ref. 6) – 1343 – 71A-286 (Ref. 7) 1200 1350 1050 300Type 310 (Ref. 7) 1325 1350 1325 25Alloy 242 (annealed) (Ref. 8) – – – About 120Alloy 242 (aged) (Ref. 8) – – – About 120Waspaloy (different 1232–1253 1302–1330 1038–1204 52–232conditions) (Ref. 9)IN718 (different conditions)(Ref. 9) 1190–1199 1272–1276 1149–1171 103–127IN718 (different conditions)(Ref. 10) 1085–150 1283–1301 1010–1090 193–280

Table 10 — Phase Computational Modeling Predictions of Phase Transformation Characteristicsof IN740 Alloy Heat BLT2819 in the Solidification Range (Based on Nominal Chemical Composition)

Temperature (°C) Description1353.5 Liquidus temperature (L)/start of solidification of gamma phase1283.0 Start formation of MC carbides1202.8 Start formation of MB2 phase1183.8 Solidus temperature (S)169.7 Solidification range (L-S)

Temperature (°C) Phase fraction (wt-%)Gamma MC MB2 Liquid

1183.8 (solidus) 99.78 0.22 0.01 –1120.0 (annealing) 99.76 0.24 0.01 –1205.0 (NDT) 98.76 0.2 – 1.041280.0 (NST) 86.8 0.03 – 13.17

Table 11 — Phase Computational Modeling Prediction of Segregation Behavior of Different Elements during Solidification of Alloy Heat BLT2819 (Based on Scheil-Gullier SolidificationModel and 0.02 wt-% Liquid Fraction Solidification Cut-Off)

Element Concentration in Liquid (wt-%) Segregation Ratio (Cf/Cs)Start Solidification Finish Solidification

(Cs) (Cf)

Ni 48.8 59.0 1.2Cr 24.4 7.44 0.3Co 20.0 3.8 0.19Zr 0.225 9.82 43.6Nb 2.0 10.71 5.36 (*)Ti 1.8 4.94 2.7Si 0.53 1.56 2.94 (*)Mo 0.50 1.07 2.14Mn 0.26 1.06 4.1Al 1.0 0.76 0.76Fe 0.45 0.18 0.4B 0.0045 0.053 11.78 (*)C 0.029 0.004 0.13

(*) Elements associated with liquation cracking behavior.

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will include all or part of the mushy zone.However, the mushy zone liquid generallydoes not contribute to poor weldability,because during the normal course of so-lidification, it is always open to the fusionzone. The fusion zone acts as a source ofliquid to backfill or heal the shrinkage orcracks that might form.

However, phases that form during cast-ing or processing of the alloy such as MCcarbides and Laves phases have the po-tential to initiate melting by constitutionalliquation or equilibrium melting in theHAZ on-heating during welding andspread along the grain boundaries (Refs.12–14). Constitutional liquation can occurbelow the system solidus and be locatedaway from the HAZ mushy zone. Addi-tionally, the liquation of MC carbides mayproduce such great local solute enrich-ment at the grain boundary that the reso-lidification of these boundaries can begreatly delayed, increasing the nil-ductilityplateau size, which increases the chancesof HAZ liquation cracking.

Anything that increases the life of theintergranular liquid relative to the onset ofresidual tensile stresses accentuates liqua-tion cracking. Additionally, chemical seg-regation at grain boundaries can also pro-mote grain boundary liquation wheremelting or resolidification of equilibriumphases is controlled by their equilibriumsolidus temperature (Refs. 12, 15, 16).

Therefore, in order to gain under-standing of the solidification behavior ofAlloy IN740 and identify the factors thatcontrol the high susceptibility to liquationcracking observed in heat BLT2819, phasecomputational modeling was used to pre-dict the formation of thermodynamic sta-ble phases as well as segregation of ele-ments, partition of elements, and phasetransformations during solidification.

Table 10 lists the thermodynamic stablephase transformation characteristics ofheat BLT2819 predicted by computationalmodeling based on nominal chemicalcomposition. The model predictions indi-cate a solidus and liquidus temperatureequal to 1183.8° and 1353.5°C, respec-tively, and the formation of MC and MB2as stable phases. Taking into account thatthe tested alloy heat was processed undera controlled condition and it was annealedtwice before testing, it can be assumed thatthe degree of segregation present in theas-received material was low and the be-havior during the on-heating part of thewelding cycle is controlled by thermody-namically stable phases present in the ma-terial. The solution-annealed heat treat-ment cleans the grain boundary of residualelements and dissolves small particlesavailable for constitutional liquation, re-ducing the potential for liquation cracking(Ref. 15). The results of computationalmodeling pointed out that at 1205°C

(NDT of the tested alloy heat), the alloyconsists of about 98.67 wt-% gammaphase, 0.2 % MC, and 1.04% liquid phase.This suggests that the NDT of the alloycorresponds to the temperature where liq-uid starts to form at the grain boundariesduring heating. At 1280°C (NST of thealloy heat), 86.8 wt-% gamma phase,0.03% MC, and 13.17% liquid are presentin the material. Therefore, these resultsindicate that the NST of the material dur-ing heating corresponds to the tempera-ture where most of the stable secondphases are dissolved and the amount ofliquid phase is enough to envelop most ofthe grain, which results in almost zerostrength of the material.

On the other hand, the DRT (1050°C)determined during cooling of the testedheat is much lower than the solidus temper-ature of 1183.8°C predicted by the compu-tational model. Therefore, a different con-dition needed to be considered to explainthe behavior observed during the on-cool-ing part of the testing. The additional mod-eling was conducted to evaluate segregationof elements, partitioning of some elements,and phase transformations during solidifi-cation on cooling. Solidification was mod-eled based on the Scheil-Gulliver approach,and a 0.02 wt-% liquid fraction solidifica-tion cutoff was assumed.

Figure 8 shows the predicted evolutionof segregation of different elements on theliquid phase as solidification progresses.Table 11 lists the predicted degree of segre-gation of different elements in the liquidphase from the start of solidification (Cs) tothe completion of solidification (Cf), ascharacterized by a segregation ratio (Cf/Cs).Alloying elements that experience a largedegree of segregation during solidificationof Alloy IN740 in decreasing order of thecorresponding segregation ratio include zir-conium, boron, niobium, manganese, andsilicon. It is important to indicate that sul-fur, which may depress the melting point ofmicrostructural constituents like Lavesphases (Ref. 15), is not included in the data-bases that support the computational modelused. Of these elements, boron, niobium,and silicon are normally associated with thesusceptibility to liquation cracking by in-creasing the solidification range or modify-ing the surface tension of the liquid phaseand its ability to wet the grain boundaries(Refs. 12, 13, 17–22).

The partitioning of these three elementswas evaluated as a function of the advance-ment of the solidification front. Table 12 de-scribes the partitioning of B, Nb, and Si be-tween the different phases, in addition tothe gamma phase, as the IN740 solidifies.Boron partitions mainly between MB2

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Table 12 — Phase Computational Modeling Prediction of Partitioning Behavior of Different Elements during Solidification of Alloy Heat BLT2819 (Based on Scheil-Gullier SolidificationModel and 0.02 wt-% Liquid Fraction Solidification Cut-Off)

Element Phase Concentration (wt-%) Fraction Solid Range (wt-%)

B MB2 31.0 0.95–0.98Liquid 0.053 0.98

Nb MC 58.0–70.0 0.75–0.98Laves 38.0 0.93–0.98G-Phase 24.0 0.96–0.98Liquid 11.0 0.98

Si G-phase 12.0 0.96–0.98Laves 4.5 0.93–0.98Liquid 1.6 0.98

Table 13 — Phase Computational Modeling Predictions of Phase Transformation Characteristicsof IN740 Alloy Heat BLT2819 in the Solidification Range (Based on Scheil-Gullier SolidificationModel and 0.02 wt-% Liquid Fraction Solidification Cut-Off)

Temperature (°C) Description

1353.5 Liquidus temperature (L) and start of solidification of gamma phase1276.0 Start formation of MC carbides1140.0 Start formation of Eta phase1130.0 Start formation of Laves phase1120.0 Start formation of MB2 phase1110.0 Start formation of G-phase1060.0 Solidus temperature (S)293.5 Solidification range (L-S)

Temperature (°C) Phase fraction (Vol-%)Gamma MC Eta/Laves MB2/G

1060.0 (S) 97.3 0.3 1.2/1.0 0.1/0.1

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phase and the remaining liquid phase dur-ing the late stages of solidification (0.95 to0.98 fraction solid range). Niobium parti-tions to MC carbides starting at about 0.75fraction solid range while Nb and Si parti-tion into Laves, G phases, and remainingliquid toward the end of the solidificationprocess (0.93 to 0.98 fraction solid range).Figure 9 shows the predicted partitioningbehavior of Nb during solidification as func-tion of fraction solid.

Segregation during solidification ofAlloy IN740 causes not only the amountof low melting equilibrium phases such asMC and MB2 to increase but the forma-tion of low melting nonequilibriumphases such as eta, Laves, and G phases,as listed in Table 13. Figure 10 shows thepredicted fraction of phases present inthe solid as a function of temperatureduring solidification. In this system, thesequence of solidification during coolingseems to be as follows:

L → L + γ → L + γ + MC → L + γ +MC + Eta phase → L + γ + MC + Etaphase + Laves phase → L + γ + MC +Eta phase + Laves phase + MB2 → L+ γ+ MC + Eta phase + Laves phase + MB2+ G phase.

As a result, the solidus temperature ofthe system is depressed to 1060°C, whichis close to the DRT (1050°C) observed forthe tested alloy heat. Therefore, the seg-regation of boron, niobium, and siliconduring solidification results in the increaseof the solidification range from 169.7° to293.5°C of the IN740 alloy heat tested.This result explains the large nil-ductilityrange (NDR) of 229°C observed in thealloy heat tested and its associated highsusceptibility to HAZ liquation cracking.

Experimental Results of HAZDDC Susceptibility of Alloy IN740

The results of the Gleeble hot-ductilitytesting of Alloy IN740 in the DDC temper-ature range are listed in Table 14 and shownin Fig. 11. Figure 11 shows the general hot-ductility behavior of Alloy IN740 in the tem-perature range between 600° and 1300°C.The actual ductility of the samples testedon-heating at 600°C is higher than indicatedin Fig. 11 because the samples did not frac-ture at the maximum displacement pro-grammed for the testing. The general trendof the data shown in Fig. 11 indicates thatduring the on-cooling part of the weldingthermal cycle, the strength and ductility ofthe alloy start to recover and continually in-

crease as it cools down from 1050° to 600°C.A pronounced loss or ductility dip was notobserved in the tested alloy heat. Therefore,the results of the Gleeble hot-ductility test-ing of Alloy IN740 indicate that this alloyheat is not susceptible to DDC cracking.

Conclusions andRecommendations

• On-heating, Alloy IN740 (HeatBLT2819) showed a nil-ductility tests tem-perature (NDT) and nil-strength temper-ature (NST) equal to 1205° and 1280°C,respectively. On cooling, after experienc-ing a peak temperature equal to the NST,the alloy starts recovering ductility at1050°C (DRT).

• Alloy IN740 (Heat BLT2819) showedan on-cooling nil-ductility temperaturerange (NDR) of about 230°C. This largeNDR indicates that Alloy IN740 is highlysusceptible to HAZ liquation cracking.

• Under spot-Varestraint testing, the sat-uration strain for Alloy IN740 is about 4%and the maximum crack distance (MCD) atsaturation strain was about 0.66 mm. As ex-pected, increasing the arc-on time from 10to 20 s increased the size of the HAZ liqua-tion cracking susceptibility envelope.

• Phase computational modeling,based on nominal chemical compositionof heat BLT2819, predicts a solidus andliquidus temperature equal to 1183.8° and1353.5°C, respectively, and the formationof MC and MB2 as stable phases. At1205°C (NDT) and 1280°C (NST) about1.0 and 13.2 wt-% liquid, respectively, ispredicted to be present. Therefore, NDTand NST may correspond to the initiationof liquation and to the presence of enoughliquid to surround the grains, respectively,on heating. The predicted liquidus tem-perature (1183.8°C) is much higher thanthe DRT (1050°C) observed on-cooling.

• Phase computational modeling of so-lidification of Alloy IN740, based on theScheil-Gulliver approach and a 0.02 wt-%liquid fraction solidification cut-off, pre-dicts segregation ratios for zirconium,boron, niobium, manganese, and siliconranging from 44 to 2. Boron, niobium, andsilicon tend to partition to MB2, MC,Laves, G phase, and remaining liquid dur-ing the last stages of solidification.

• Segregation during solidification ofAlloy IN740 results not only in the in-crease in the amount of low melting equi-librium phases such as MC and MB2 but inthe formation of low melting nonequilib-rium phases such as eta, Laves, and Gphases. Therefore, the solidus tempera-ture of the system is depressed to 1060°C,which is close to the DRT (1050°C) ob-served for the tested alloy heat.

• The segregation of boron, niobium,and silicon during solidification results inthe increase of the solidification range from

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Table 14 — Results of Hot-Ductility Testing during On-Heating and On-Cooling of IN740 Alloy,in the DDC Temperature Range

Specimen Temperature °CNumber ID Peak Test RA (%) Fracture Strength (ksi)

HT-15(a) HT600 – 602 21.2(c) 92.0HT-12 HT600 – 602 28.5(c) 91.7HT-11 HT700 – 700 53.9 89.1HT-10 HT800 – 801 61.2 70.4HT-16 HT850 – 850 61.8 83.9HT-9 HT900 – 901 70.4 71.9HT-17 HT950 – 952 82.2 60.3HT-8 HT1000 – 1000 84.9 55.7CT-1 CT1000 1267 999 11.5 2.1CT-14 CT1000 1271 1001 46.7 56.8CT-21 CT1000 1271 1000 46.4 55.7CT-30 CT975 1268 977 52.7 59.9CT-31 CT975 1271 974 48.1 60.3CT-2 CT950 1275 953 3.8 10.8CT-3 CT950 1273 952 2.9 10.9CT12(b) CT950 1271 950 38.1 59.3CT-32 CT925 1272 925 52.9 59.6CT-33 CT925 1270 925 45.4 62.1CT-4 CT900 1270 900 0 3.2CT-9 CT900 1272 901 38.7 59.5CT-10 CT900 1270 901 4 3.1CT-11(b) CT900 1271 900 50 63.7CT-34 CT875 1271 874 55.4 65.4CT-35 CT875 1272 875 48.4 64.3CT-5 CT850 1270 847 51.2 64.0CT-13 CT850 1270 848 50.0 68.1CT-6 CT800 1271 802 53.5 68.3CT-7 CT700 1270 699 56.4 76.7CT-8 CT600 1272 608 52.6(c) 78.8

(a) H: On-heating test specimens; T: Test temperature.(b) C: On-cooling test temperature; T: Test temperature.(c) These samples did not break during testing. Therefore, the actual reduction of area at fracture is higher than indicated.

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169.7° to 293.5°C of the Alloy IN740 heattested, which explains the large nil-ductilityrange (NDR) of 229°C observed in the alloyheat tested and its associated high suscepti-bility to HAZ liquation cracking.

• A pronounced loss or ductility dip wasnot observed in the Gleeble hot-ductility tested alloy heat in the tempera-ture range from 600° to 1050°C. Therefore,Alloy IN740 is not susceptible to DDCcracking.

Acknowledgments

This publication was prepared withsupport of the U.S. Department of En-ergy, under Award No. DE-FG26-01NT41175 and the Ohio Coal Develop-ment Office of the Ohio Department ofDevelopment under Grant AgreementNumber CDO/D-0020 (now D-05-02A).However, any opinions, findings, conclu-sions, or recommendations expressedherein are those of the author(s) and donot necessarily reflect the views of theDOE and/or the Ohio Coal DevelopmentOffice of the Ohio Department of Devel-opment.

Legal Notice/Disclaimer

This report was prepared by Jose E.Ramirez (EWI) pursuant to a Grant par-tially funded by the U.S. Department ofEnergy (DOE) under Instrument NumberDE-FG26-0 1 NT 41175 and the Ohio CoalDevelopment Office/Ohio Department ofDevelopment (OCDO) under GrantAgreement Number CDO/D-OO-20 (nowD-05-02A). No warranty or representation,express or implied, is made with respect tothe accuracy, completeness, and/or useful-ness of information contained in this re-port. Further, no warranty or representa-tion, express or implied, is made that theuse of any information, apparatus, method,or process disclosed in this report will notinfringe upon privately owned rights. Fi-nally, no liability is assumed with respect tothe use of, or for damages resulting fromthe use of, any information, apparatus,method or process disclosed in this report.

Reference herein to any specific com-mercial product, process, or service bytrade name, trademark, manufacturer, orotherwise, does not necessarily constituteor imply its endorsement, recommenda-tion, or favoring by the Department of En-ergy and/or the State of Ohio; nor do theviews and opinions of authors expressedherein necessarily state or reflect those ofsaid governmental entities.

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