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A Study of the Weld Heat-Affected ZoneToughness of 9% Nickel Steel

Although thermal cycling reduces cryogenicimpact toughness and upper shelf energy,

ASTM specifications are still exceeded

BY E. F. NIPPES AND J. P. BALAGUER

ABSTRACT. The weld heat-affected zonetoughness of 9% nickel steel weldmentsat cryogenic temperatures was investi

gated. The material was thermallytreated, using a Gleeble, to produ ce synthet ic weld heat-affected zones. Toughness of this steel, subjected to threedifferent thermal cycles, was measuredusing the Charpy V-notch impact test. Itwas found that the material subjected tohigh peak-temperature thermal cycles stillmet the ASTM toughness requirementsfor this steel at liquid nitrogen temperatures. The upper shelf energy for thissteel was lowest for an intermediatethermal cycle of 1000C (1832F) peaktempera ture .

A fractographic analysis of the materialwas conducted. Charpy V-notch specimens fractured at liquid nitrogen temperatures were examined in the scanningelectron microscope. Material representative of the coarse-grained region of theheat-affected zone failed by a partiallylow-energy mode at cryogenic temperatures.

The retained austenite content of thismaterial, under varying conditions ofthermal t reatments, was determinedusing standard x-ray diffraction techniques. It was found that upper shelftoughness increased as the amount ofretained austenite increased, whiletoughness at liquid nitrogen temperatureswas related to combined effects of prioraustenite grain size and retained austenitecontent. Retained austenite was alsomeasured using the Magne-Gage, modified for the measurement of austenite ina ferritic matrix, and these data werecorrelated with the x-ray diffractionresults.

Introduct ion

Increases in demand for natural gashave made it necessary to construct storage facilities for liquefied natural gas.Because liquid natural gas (LNG) is stored

at , or below, the -162C (-2 60F) boi ling temperature, the vessel must be c o nstructed from a material which possesses

high strength and suitable fracture toughness at cryogenic temperatures.

Nine percent nickel steels were firstproduced in the United States in the early1940's. More recently, Japanese steelcompanies have also begun to producelarge tonnages of 9% nickel steel. Thealloying addition of approximately ninepercent nickel and the subsequent heattreatment are primarily responsible forthe relatively high strength and excellentfracture toughness exhibited by this steel.The superb low-temperature propert iesof 9% nickel steel make it a go od materialcandidate for liquid natural gas storagetanks.

Almos t all the 9% nickel steel pro du cedin the United States is intended for thefabrica tion of we lde d pressure vessels. Itis we ll kno w n that w eldin g can seriouslyalter the metallurgical properties of thematerial immediately adjacent to theweld joint . Therefore, the effect of weldheat-affected zones (HAZ) on the cryogenic fracture toughness of 9% nickelsteel LNG storage tanks must be properlyevaluated.

Previously, the weld heat-affectedzone propert ies of a wide variety ofsteels and nonferrous metals have beenevaluated utilizing weld simulationdevices, Charpy V-notch impact testing,and metallographic investigation. Theobject of the present investigation will beto evaluate the effect of weld thermalcycles on the impact absorption energyof the weld HAZ of 9% nickel steel atcryogenic temperatures.

Metallurgical Characteristics

The fundamental mechanical properties of 9% nickel steel must meet theminimum requirements of ASTM 353 orASTM 553 Type I specifications, as show nin Table 1. ASTM 353 and 533 Type I

specify double-normalize-and-temper(NNT) and quench-and-tem per (QT) heattreatments, respectively. These specifica

tions require measurement of impactabsorption energy at liquid nitrogen temperatures, -196C ( -320F) . However,at the LNC s torage temperature of- 1 6 2 C (- 260F), these steels, in thewrought condit ion, safely exceed theASTM specifications described in Table 1.

The excellent low-temperature toughness of this steel results from the alloyingaddition of approximately nine percentnickel. The nickel acts to suppress theformation of ferr i te/pearl i te high-temperature transformation products; thus, amicrostructure is produced which is higher in strength and notch toughness. It isgenerally recognized that the superiorfracture toughness of 9% nickel steel atcryogenic temperatures is due to thepresence of stable retained austenite(Refs. 1-5). The addition of nickel lowersthe martensite finish temperature suchthat unstable austenite remains aftercooling from the austenit izat ion temperature to room tempera ture .

Previous investigations (Refs. 3, 5, 6)have shown that if tempering is performed in the two-phase a + i regionabove the low er crit ical temp erature, i.e.,above approximately 580C (1076F),the high nickel content stabilizes the austenite, and the final product contains 5 to10 volume percent (v-%) ret ain ed austenite. This second phase increases low-temperature toughness primarily by scavenging interstitials, and this scavenging

Based on a paper presented at the 66th A WSAnnual Meeting, held April 28 to M ay 3, 1985,in Las Vegas, Nev.

E F. NIPPES is Professor of Metallurgical Engi-

neering, and J P. BALAGUER is Research Fel-low, Department of Materials Engineering,Rensselaer Polytechnic Institute, Troy, N. Y.

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Table 1ASTM 553-

Composi t ion(v-%)

TensileRequirements(20C)

Impact AbsorptionEnergy (CharpyV-Notch) at- 1 9 6 C

Heat Treatment

Specifications for 9% Nickel Steels

C M n P S Si Ni

0.13 0.90 0.035 0.040 0.15-0.4 0 8.5-9.5max. max. max. max.

Tensile Strength Yield Stren gth, min. Elongation inMPa (ksi) MPa (ksi) 50 m m , min.%

690 -825 (100-12 0) 585 (85) 20.0

Longitudinal, min. Transverse, min.J (ft-lb) | (ft-lb)34 (25) 27 (20)

Austenitizing Tem pering Tem peratureTempera ture C (F)

C(F) 5 6 5 - 6 3 5 ( 1 0 5 0 - 11 7 5 )8 0 0 -925

(1475-1700)

Table 2Chemical Analysis of Base Plate Used in This Investigation (Ref. 19)

Element C

v-% 0.05

Table 3Mechanical

TensileStrengthMPa (ksi)

742 (107.5)1005 (145.7)

1118.5(162.1)

M n P S Si Ni Fe

0.71 0.018 0.019 0.41 8.90 bal.

Properties of Base Plate Used in This Investigation (Ref. 19)

Yield TestStrength Elongation Tem peratureMPa (ksi) % C (F)

695 (100.7) 26.4 20 (68)905 .5 (131 .2 ) 2 7 .1 -162 (-260)

984 (14 2 .6 ) 29 .3 -196 (-321)

Tensi le properties longi tudinal plate or ie ntat ion .Values reported represent average ot two tes ts .

action prevents the formation of embrit tling carbides and nitrides. The resultingdecrease in yield strength and increase inwork-hardening ability increases fracturetoughness. During tempering, the austenite also gathers uncombined carbon,which is thought to initiate cleavagecracks in martensite (Ref. 2). Kim (Ref. 6)and Schwa rtz (Ref. 7) have p ropo sed thatadditional mechanisms, including the

blunting of cleavage crack tips uponentering regions of retained austenite,may operate to a lesser degree than thescavenging effect. However, recentwork by Kim and Morris (Ref. 8) on a5.5% nickel steel suggests that the crackblunting hypothesis is unlikely. They propose that the reduc ed effective grain sizeof a composite microstructure is responsible for the high observed toughness atcryogenic temperatures.

Simulation and Analysis of WeldHeat-Affected Zones

Welding can seriously alter the metallurgical properties and the stress state of

the material immediately adjacent to theweld joint. The intense heat of a weldingarc can subject the surrounding materialto severe thermal cycles. The thermalcycles always alter, to some extent, themetallurgical structure of the base material near the w e l d . This region, referred toas the we ld hea t-affected z one (HAZ),has been studied extensively by Nippes(Ref. 9) and others. The methodology of

simulating the weld HAZ used in thisstudy, i.e., reproduction of measuredweld thermal cycles in a small specimenusing the Gleeble machine, is an intrinsicpart of the welding literature. A detaileddiscussion of these techniques can befound elsewhere (Refs. 9-11).

The effect of weld heat-affected zoneson the cryogenic fracture toughness of9% nickel steel LNG storage tanks hasbeen the subject of recent study. Thefracture toughness of this material hasbeen evaluated using both traditionaltesting methods, such as Charpy V-notch(CVN) and crack-opening displacement(COD) tests , and more advanced methods, such as J-integral and R-curve tech

niques. The variety of testing methodsand the inconsistencies of weld HAZstructures have resulted in contradictingdata on the fracture toughness of theHAZ in 9% nickel steel welded joints(Refs. 12-18).

Dh oo ge , et al. (Ref. 16), have suggested that the thermal cycles to which theHAZ is subjected can reduce the impactenergy of this region by almost fiftypercent. Similarly, Syn, et al. (Ref. 5), ha vefound that the retained austenite contentof 9% nickel steel is reduced to immeasurable levels by elevated temperature treatments. Dhooge, et al. (Ref. 16),have also found that a postweld heattreatmen t at 60 0C (1112F), foll ow edby rapid air cooling or water quenching,would increase the toughness of theweld HAZ.

Materials and Procedure

Characterization of Base Material

The base material used in this investigation is a quench-and-tempered 9% nickelsteel. The material was obtained in theform of 12-mm (0.47-in.) thick plate. Thechemical analysis of the plate, shown inTable 2, meets the compositional requirements of ASTM 553-I. The fundamentalmechanical properties of this alloy areshown in Table 3.

The impact absorption energy of thebase material as a function of temperatu re , measured using the Charpy V-notchimpa ct tes t, is sho wn in Fig. 1. It is difficultto determine the ductile-to-brittle transi

tion temperature of this material fromthese data because the lower energyshelf is not clea rly de fine d. If it is assumedthat the lower energy shelf begins at 196 C, the transition tem pera ture,defined as one-half the sum of the upperand lower shelf energies, appears to beapproximately -120C ( -184F). Theaverage impact absorption energy of thismaterial at -196C was about 110 J (86ft-lb), conside rably higher than the 34 J

- 1 5 0 . - 1 0 0 . , - S O .

TEMPERATURE (C)Fig. 1 Impact energy vs. temperature formaterial in the as-received condition

238-s I SEPTEMBER 1986

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(25 ft-lb) minimum specified by ASTM553-1. The average impact absorptionenergy at the LNG service temperature(-1 62C) was about 135 J (101 ft-lb).

The macrohardness and microhardness of the as-received material wereHRC 20 and HV 259, respectively. Grainsize estimates of the as-received andheat- treated material were prepared inaccordance with ASTM El 12, using theintercept method.

Simulation of Weld Heat-Affected Zones

Oversized Charpy V-notch specimenblanks were machined from the basematerial plate and thermally cycled in aGleeble to produce synthetic weld heat-affected zones. The oversized dimensions allowed removal of surface oxideresulting from thermal cycling by finalmachining prior to notching of the specimens to specified dimensions. A 0.254-mm (0.010-in.) d iameter, two-wire ,chromel-alumel thermocouple, percussion welded to the side of the specimenblanks, was used for temperature contro lin the Gleeble. The thermocouple-basemetal joint was later removed during themachining of the notch in the Charpyspecimen.

The specific welding thermal cyclesused in this investigation were calculatedutilizing the data of Nippes, Merrill, andSavage (Ref. 20) for arc welds in 12.5-mm(0.5-in.) plate. Therm al cycles for weld s o f15.75 kj/cm (40 kj/in.) energy input andan initial plate temperature of 22 CC(72F) wer e calculated. Distances of 10.4,7.9, and 7.4 mm (0.41, 0 .31 , and 0.29 in.)from the weld centerl ine were chosen toproduce peak temperatures of 500,1000, and 1300C (932, 1882, and2372F), respectively.

Impact Absorption Energy Testing

Following the thermal cycling treatment, specimen blanks we re machined tofinal dimensions of 10 X 10 X 55 mm(0.394 X 0.394 X 2.165 in.) and notched.Charpy V-notch specimen preparat ionprocedures and impact testing were per

formed in accordance with ASTM E-23.Impact absorption energy testing wasperfo rme d using a Riehle Charpy V-notchtesting machine. Prior to testing, theRiehle machine was calibrated with specimens provided by the United StatesArmy Materials and Mechanics ResearchCenter, Watertown, Massachusetts .Test ing temperatures of 0C to -1 6 2 Cwe re ob tained using a bath of 2-methyl-butane cooled with l iquid ni trogen. Testing temperatures of 196C we reobtained with liquid nitrogen. Specimenswere maintained in a liquid bath at thetest temperature for 5 min prior toimpact testing. Three specimens w eretested at each temperature for the

1300 C and 10 00C therm al cycles; tw ospecimens were tested at each temperature for the 500 C thermal cycle.

Metallography

All samples in this investigation wereprepared for metal lographic examinationusing standard techniques. Samples wereembedded in 31.75-mm (1.25-in.) diame

ter bakeli te mounts and prepared bysuccessive stages of lapping, grinding,and mechanical polishing. The followingchemical etchants were used to preparethe sample for viewing under the opticalmicroscope: Etchant A 2% Nital; EtchantB - 2 g p ic r ic acid, 100 cc H 2 0 .

Measurement of Retained Austenite

X-ray diffractometry was used todetermine the amount of retained austenite present in thermally cycled specimens of this steel. The method of Miller(Ref. 21), wh ich com pares the integratedpeak intensities of ferritic and austeniticphases, was used.

A n Aminco Model 5-660 Magne-Gagewas also used to estimate the retainedaustenite content of specimens subjectedto the various heat treatm ents used in thisstudy. The Magne-Gage is normally usedto measure a small amount ( ~ 0 - 1 5 v-%)of ferrite in largely austenitic materials.However, Kotecki (Ref. 22) has shownthat the Magne-Gage can be modified tomeasure volume percent ferrite in partially and fully ferritic steels. Ferrite andmartensi te , both ferromagnetic phases,and retained austenite, a non-magneticphase, are normally present in the micro-structure of 9% nickel steel. Therefore,the attractive magnetic force of the sample, measured by the Magne-Gage,should vary inversely with the volumepercent of retained austenite.

In the present s tudy, the Magne-Gagewas not physically modified. A 0.25-mm(0.010-in.) thick plastic shim was placedbetween the specimen and the magnetduring measurement. The purpose of theplastic shim was t o redu ce the strength ofthe Magne-Gage magnet. This alterationresulted in the increased sensitivity necessary to measure large amounts of ferrite.The Magne-Gage measurements weretaken on three separate areas of eachspecimen, and the measurements wererepeated ten times per area. Data wererecorded from the silver dial of theMagne-Gage.

Results and Discussion

Microstructural Analysis

The microstructure of the 9% nickelsteel base plate, in the as-received co ndit ion , is shown in Fig. 2. The microstructure of this material consists largely oftem pered martensite (TM). The TM is a

3>''MMm

mmmV- i ^ i - , ,

Fig. 2 Photomicrograph o f the material in theas-received condition, E tch A (500X). A Parallel to the plate surface; B Perpendicularto the plate surface

mixture of ferrite (a) and carbide whichhas precipi tated during the temperingheat t reatment of as-quenched martensite. Nine percent nickel steel, in thequench-and-tempered condit ion, alsocontains approximately 5-10 volume percent of retained austenite (7). Figure 2Bshows the extensive banding present inthe as-received material, the lingeringresult of segregation during the originalsolidification.

In orde r to verify the heat treatm ent ofthe as-received material, a small piece ofthe plate was austenitized, quenched,and tempered. The steel was austenitizedat 816C (1500F) for 2 hr, waterquenched and tempered at 593C(1100F) for 30 min, and air coole d. Theas-quenched and quench-and-temperedhardness of this material are listed inTable 4. The microstructure of the mate-

Table 4Results of Macrohardness andMicrohardness Measurements

Heat Treatment

As-receivedAs-quenchedQuench & tempered500C peak

tempera turethermal cycle

1000C peaktempera turethermal cycle

1300C peaktempera ture

thermal cycle

Mac rohardness

(HRC)

20 - 21362023 0

37-38

36

M i c r ohardness

(HV, 200-gload)

256357246255

367

353

(s> Ca.culated values.

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My*

Fig. 3 Photom icrograph of the material in theas-quench ed condition, E tch A (1000X)

Fig. 4 Photom icrograph of the material in thequench-and -tempered condition, Etch A(1000X)

A.:?>.

7 , 7

Fig. 5 Microstructure of materialcycled to a peak tem perature of 500(500X)

thermally'C Etch A

i X? > '^;yiyiX * . '