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SA-25619; No. of Pages 7

Materials Science and Engineering A xxx (2009) xxx–xxx

Contents lists available at ScienceDirect

Materials Science and Engineering A

journa l homepage: www.e lsev ier .com/ locate /msea

ffects of thermal boundary conditions in friction stir welded AA7050-T7 sheets

. Upadhyay, A.P. Reynolds ∗

epartment of Mechanical Engineering, University of South Carolina, 300 Main Street, Columbia, SC 29208, USA

r t i c l e i n f o

rticle history:eceived 30 July 2009

a b s t r a c t

A series of friction stir welds was made in laboratory air and with the plates submerged in water toinvestigate how the quenching rate affects properties of the joint and some weld response parameters.

eceived in revised form 12 October 2009ccepted 20 October 2009vailable online xxx

eywords:

Select welds were also made at a sub-ambient temperature of −25 ◦C. Temperature measurements weremade in the probe center and at the minimum hardness location of the weld. Weld response variables,hardness distributions, joint strength and nugget grain size were measured and correlated with boundaryconditions and welding parameters. A consistent decrease in the peak temperature and increase in coolingrate were observed in the submerged welds. Submerged welds show improvement in tensile strength

ut the

XXX series alloysoundary conditionnder-water FSW

and elongation througho

. Background and introduction

Peak temperature and the rate of cooling during the friction stirelding (FSW) process are key parameters that dictate the weldroperties. It has been reported in the literature [1,2] that all otherhings being equal, the peak temperature in the stir zone is pro-ortional to the tool rotational rate, whereas the cooling rate and,ence, length of time of stay above a certain temperature is depen-ent on welding speed. Sato et al. [1] have reported the effect ofelding parameters on the peak temperature in the stirred zone forA6063. The study reported that the peak temperature increasesith the increase in tool rotational rate at constant welding speed,

he rate of increase in temperature being less with high rpm. Tem-erature simulation by Reynolds et al. [2] has shown that transient

ength of thermal history is governed by the welding speed. Higherhe welding speed shorter will be the time for which the stir zonetays above an elevated temperature. In other words, for weldserformed under standard, ambient conditions, the heating andooling rate is dependent on the welding speed. It is intuitive andas been demonstrated [3], that the heat extraction rate can also be

ncreased by employing rapid cooling techniques such as, weldingnder water and in the presence of cold fluids.

Welds have been performed under water in offshore structural

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epair and development using conventional welding methods likehielded metal arc welding for a long time. These methods sufferrom issues like hydrogen embrittlement, oxidation and porosityhich worsen at greater depth [4]. Being a solid state joining pro-

∗ Corresponding author at: Department of Mechanical Engineering, University ofouth Carolina, 300 Main Street, Room A224, Columbia, SC 29208, USA.el.: +1 803 777 9548; fax: +1 803 777 0106.

E-mail address: reynolds@cec.sc.edu (A.P. Reynolds).

921-5093/$ – see front matter © 2009 Elsevier B.V. All rights reserved.oi:10.1016/j.msea.2009.10.039

range of parameters tested.© 2009 Elsevier B.V. All rights reserved.

cess friction stir welding bypasses these problems. Some referencescan be found in the literature concerning friction stir welding per-formed submerged in water and other cold fluids and propertyalterations associated with it [3,5–9].

Early work in submerged friction stir welding performed at lowambient temperature can be attributed to Benavides et al. [5]. Theyreported a significant reduction in stir zone temperature whenAA2024 plates were friction stir welded submerged under liquidnitrogen. Although the joint suffered from worm-hole defect, a sig-nificant reduction in stir zone grain size to 0.8 �m was reported[5]. Sakurada et al. used the inertia friction welding method to joinAA6061 under water [6]. Comparing the weld with regular in-airwelds, they reported an increase in joint strength and decreasein the width of heat affected zone. Nelson et al. [3] studied theeffect of quench rate on 7075 and 2195 aluminum alloys. They per-formed friction stir welding by externally heating and cooling theparent metal plate and anvil. Cold water and mist were used tochill the plates in the wake of the tool. They reported a maximumof 10% increase in tensile strength with respect to conventionalFSW of AA7075 after post-weld natural aging of 1000 h. Staron etal. applied liquid CO2 coolant near the weld seam to investigateresidual stress improvements [7]. They report significant reduc-tion in tensile residual stress for 6.35 mm thick AA2024 plate. Suet al. reported production of nano-scale grain size in 7075 Al sheetby quenching the plate behind the tool with a mixture of water,methanol and dry ice. Similar results were reported by Hoffmanand Vecchio for AA6061 by welding under water [9].

Temperature history plays a significant role in determining

ng. A (2009), doi:10.1016/j.msea.2009.10.039

properties within a friction stir weld; therefore, accurate mea-surement of temperature inside the stir zone during the weldingprocess is crucial if the process is to be understood. Unfortunately,temperature measurement inside the FSW process zone is highlyproblematic. Several factors make it nearly impossible to capture

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were obtained from the nugget center. Grain size was measured atthe nugget center using the MLI method. Three views at a magni-fication of 500× were examined. Intersections were counted on atest line of length 100 mm. Five test lines per view were used.

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he actual temperature transient experienced by the weld nuggetaterial. Steep temperature gradients, finite thermocouple size,

eformation, and movement of the material in which the ther-ocouple is embedded combine to make accurate temperature

etermination complicated [10]. Due to these reasons the peakemperature inside the stirred zone has been estimated using indi-ect analytical methods like comparison of micro-hardness, nuggetrain size, and quantification of precipitates in the matrix [11,12].ome investigators have used computational models, normally val-dated with far-field temperature measurements, to deduce theemperature history [11–15].

Numerous studies have been devoted to understand the rela-ionships between properties and welding parameters for 7XXXeries alloys [1,2,12,15–18]. These alloys are precipitation hardenedith �′ (Mg2Zn) as the primary strengthening precipitate phase.hen solution heat treated and subsequently aged using appro-

riate times and temperatures, the alloy microstructure will havedistribution of fine particles in the solid solution matrix with opti-um tensile and other mechanical properties. During the welding

rocess this optimized microstructure is altered at various levelsecause of the temperature cycle. This cycle will always produceome level of dissolution of precipitates, diffusion of solute, andncrease in vacancy concentration. Thus some non-strengtheninghases may form and some strengthening precipitates may beoarsened or dissolved, causing varying degrees of loss of strength.he extent of this modification in microstructure will be primarilyoverned by the temperature history which in turn is dependentn welding parameters used and the thermal boundary conditionsuring the weld.

To obtain the best possible properties in the weld nugget zone its desirable that the nugget reach a peak temperature that is abovehe solution heat treatment temperature for the alloy being welded.he nugget which reaches such a peak temperature would be incondition similar to that obtained after solution heat treatment

nd quenching; therefore, reprecipitation of strengthening phasesay occur during post-weld cooling and subsequent natural and

rtificial aging treatments. Typically, if a weld is made in a pre-ipitation hardening aluminum alloy and the peak temperature isreater than the solution heat treatment temperature for the alloy,characteristic “W” shaped hardness distribution is observed. Thisrises due to solution heat treatment of the nugget and overagingf the heat affected zone (HAZ) [15–17]. If the weld is performed atelatively low power (with a stir zone peak temperature less thanpproximately 350 ◦C) the characteristic W shape in the hardnessrofile will not be observed. With a peak weld temperature near50 ◦C, normally, the hardness in the HAZ and the nugget will beimilar to each other and less than a T6 or T7 base metal. Sincehe kinetics of precipitate coarsening are maximum near 350 ◦Cor 7050 and other 7XXX series alloys, the rate of formation of non-trengthening � phase is at peak near this temperature [19]. Duringts formation, � phase particles take away a significant amount ofolute from the matrix which otherwise would have been availableor reprecipitation of strengthening �′ phase during the post-weldeat treatment. The situation will be aggravated when the weld-

ng speed is lower. Lower welding speed will cause the stir zone toemain in the � phase formation temperature range for a relativelyonger time causing the solute depletion to increase. In general, theepth of the HAZ and/or nugget hardness minimum will dependtrongly on the time spent near 350 ◦C without subsequent solutioneat treatment, hence, the depth of the minimum is highly depen-ent on the welding speed which primarily governs the temporal

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ength of the temperature transient [2,20].In this work we present thermal behavior, torque requirements,

nd resulting mechanical properties of friction stir welds in alloy050 performed in-air, under-water, and under sub-ambient tem-erature conditions for a wide spectrum of welding speed and

PRESSand Engineering A xxx (2009) xxx–xxx

rotation rate. The effects of welding parameters and some thermalboundary conditions on the resulting weld properties, in partic-ular, the grain size, hardness distributions, and transverse tensilestrength are reported and discussed.

2. Materials and experimental procedure

All welding was performed on 6.35 mm thick plates of the highstrength aluminum alloy, 7050-T7451 having a nominal compo-sition of 5.6%Zn–2.5%Mg, 1.6%Cu, 0.23%Cr, balance Al. The typicalultimate strength of the alloy in the T7451 temper is 524 MPa. Theincipient melting temperature for homogenized 7050 is 488 ◦C andthe solution treatment temperature is 477 ◦C. The dimensions of thewelded plates were 6.35 mm thick, and 101.6 mm wide (yielding atotal welded width of 203.2 mm).

Welds were produced on an MTS FSW Process Development Sys-tem (PDS) using Z-axis force control. The welding direction wasparallel to the plate rolling direction and the tool rotation axis wasnormal to the plane of the plate. Welds were made with the parentmetal plates in lab air or completely submerged in water: for under-water welds, the depth of water was approximately 25 mm abovethe top surface of the plates (see Fig. 1). For three sets of parameters,welds were also performed with the parent metal plates submergedin a mixture of 50% ethylene glycol, 50% water and 30 lb of dryice resulting in a steady far field plate temperature of approxi-mately −25 ◦C during the weld. . For the sake of convenience, thesethree conditions will henceforward be referred to as IA (in-air),UW (under-water), and SA (sub-ambient). The combinations of toolrotation speed, welding speed, and Z-axis force used are tabulatedin Table 1. The double entries in columns with 300 and 400 rpm forIA and UW welds are because of the use of two welding speeds foreach of those rotational speeds.

The tool used for production of all welds was of a two piecedesign with a 17.8 mm diameter single scroll H13 tool steel shoul-der and a probe fabricated out of MP-159 (a high temperature cobaltbased super alloy) in the shape of a truncated cone (8◦ taper) withthreads and three flats. The probe was 6.1 mm long, with a diam-eter of 7.9 mm at the intersection with the shoulder. Temperatureduring welding was monitored and recorded using a thermocou-ple spot welded into the probe on the axis of rotation at the probemid-plane height. For some select welds, thermocouples were alsoplaced at approximate location of the HAZ hardness minimum onthe advancing side of the weld at the mid-plane height. This corre-sponds to a distance of 5.5–7 mm away from the weld centerline.

Samples for metallographic and hardness measurement wereground, polished and etched using Keller’s reagent. Micrographs

ng. A (2009), doi:10.1016/j.msea.2009.10.039

Fig. 1. Under-water friction stir welding.

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Table 1Welding control variables.

Rotational speed (rpm) 150 200 250 300 400 540 650 800 1,000

33,8039,14–

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Welding speed (mm/s) 1.7 2.54 3.4 3.4, 6.8In-air Z force (N) 25,800 25,800 26,690 24,021,Under-water Z force (N) 35,586 35,586 34,696 31,138,Sub-ambient Z force (N) 35,586,

Hardness and tensile tests were performed after post-weld heatreatment of weld samples at 121 ◦C for 24 h preceded by a weekf natural aging: this approximates an industrial T6 heat treat-ent (peak strength). A Vickers hardness indenter, with a load of

000 g and a load application time of 10 s was used for measure-ent of hardness as a function of distance from the weld centerline

n transverse cross-sections along the plate mid-plane. Full thick-ess (6.35 mm) rectangular transverse tensile samples of 180 mmauge length and 12.5 mm width were machined from completedelds. Tensile tests were performed using an initial strain rate of

.0001 s−1. Three samples were tested for each welding condition.

. Results and discussion

.1. Torque and probe temperature

Fig. 2 shows the representative thermal history from the ther-ocouple welded into the probe mid-plane during a two parameter

et weld: 400 rpm, 5.1 mm/s and a 540 rpm, 6.8 mm/s. The probeemperature is observed to reach a reasonably steady state condi-ion and to react quickly to changes in the welding parameters (theeld parameters were changed at approximately 140 s of welding

ime). The temperature measured inside the probe is some aver-ge of the temperatures of the material in contact with the probend does not represent the maximum temperature; however, thiseasured temperature is quite repeatable [21] and provides an

xcellent representation of the trends in nugget temperature withhanging weld parameters.

The maximum measured probe temperature and average mea-ured torque have been graphed against the rpm in Fig. 3 foroth IA and UW welds. Note that welding speed is not a con-

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tant for this graph (see Table 1). With increasing rotational speed,robe temperature increases and the torque decreases. Two gen-ral observations can be made for both the thermal conditions. (1)eak probe temperature is inversely correlated with the measured

ig. 2. Representative temperature transients from the thermocouple welded intohe probe mid-plane during a two parameter set weld: 400 rpm, 5.1 mm/s and40 rpm, 6.8 mm/s. Parameter switch takes place approximately at 140 s.

3.4, 5.1 6.8 6.8 6.8 10.27 22,241, 25,800 28,024 30,693 33,362 37,8105 29,359, 34,252 39,145 39,145 40,479 43,593

–, 40,479 41,369

torque. (2) Both the quantities approach a limiting value as rpm isincreased. This may be rationalized in the way proposed originallyby Tang et al. [22]: the higher the rotational speed the higher willbe the temperature causing the material flow stress to decrease.This decrease in flow stress in turn will limit the power generationby plastic dissipation and, hence, temperature increases. A similartrend in peak temperature was reported by Sato et al. for alloy 6063in the rpm range of 800–3600 [16]. Yan et al. [20] and Long et al.[23] have also reported similar observations based on experimentsand CFD simulations.

Considering Fig. 3 again, all other things being equal, the peakprobe temperature is consistently lower for UW welds. This is dueto enhanced heat transfer from the tool and plate into the surround-ing water. The torque and hence power requirement for otherwiseequivalent UW welds is higher than that for IA welds. That thepower is higher while the peak temperature is lower is a reflectionof the relatively higher heat transfer to the environment for the UWwelds as compared to the IA welds. The higher torque observed forunder-water welding is no doubt related to lower temperature ofthe material in contact with the tool. The lower temperature willcorrespond to higher flow stress, thus the tool would require moretorque (and, for a given rpm, more power) to “stir” the material.

3.2. Thermal cycle

Representative weld thermal cycles for IA, UW, and SA weld con-ditions for two different welding speeds are shown in Figs. 4 and 5.The data in these figures are obtained from thermocouples placedon the advancing side of the weld at the respective locations ofminimum hardness which were determined from hardness testingof identical welds made previously. There are several features ofnote: firstly, the peak temperature measured in the HAZ minimumhardness location is close to 350 ◦C for all six conditions. This is

ng. A (2009), doi:10.1016/j.msea.2009.10.039

consistent with the temperature associated with maximum over-aging kinetics for 7XXX type alloys [19,24]. In Fig. 4 (weld made at3.4 mm/s), it can be observed that the heating rate for the IA andUW welds are similar, but the UW weld experiences a faster coolingrate than the IA weld. The temperature transient for the SA weld

Fig. 3. Measured torque and peak probe temperature vs. rpm for under-water andin-air welds for several welding sets.

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ig. 4. Thermal history of 3.4 mm/s welds for in-air, under-water and sub-ambientonditions.

s different from both the IA and UW welds. Firstly, the initial andnal temperatures are significantly lower for the SA weld and sec-ndly, the temporal length of the temperature transient is shorter.he same types of data are shown in Fig. 5 except that the weldingpeed is 6.8 mm/s. Here, the UW and SA welds exhibit very similaremperature transients except at the low temperatures which havessentially no effect on properties: the initial temperature for theA weld transient is approximately −25 ◦C. The temporal lengthsf the UW and SA weld temperature transients are much shorterhan that of the IA weld. The heating portions of the transients forll three welds are quite similar and the cooling rates for the UWnd SA welds are much higher.

The temporal length of the temperature transients are com-actly represented in Fig. 6 which is a plot of the length of timepent above 200 ◦C for each of nine welding conditions: UW, IA,nd SA at three welding speeds (3.4 mm/s, 5.1 mm/s, and 6.8 mm/s).he choice of 200 ◦C is arbitrary. From the graph it is seen that, forhe three welding conditions shown, the time above 200 ◦C at theAZ minimum hardness location is much greater for the IA weldshile the differences between the UW and SA welds are small. Also

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here is hardly any decrease in the time of stay beyond 5.1 mm/sf welding speed. This indicates that there exists a finite limit tohe highest cooling rate achievable for any given welding condi-ion. This idea can be illustrated by considering a highly simplified

ig. 5. Thermal history of 6.8 mm/s welds for in-air, under-water and sub-ambientonditions.

Fig. 6. Time spent above 200 ◦C at HAZ vs. welding speed for in-air, under-waterand sub-ambient conditions.

model of welding in which a point heat source moves across the sur-face of a thick plate at a given speed. The temperature field solutionfor this model was first presented by Rosenthal [25] according towhich the temperature at any given point on the plate is given by

T = T0 + Q

2�k· 1

r· exp

(−V(r − x)2˛

)

where T0 is initial plate temperature, Q is total heat produced bythe source, r is the radial distance of the point from the source, V isthe welding velocity, ˛ is the thermal diffusivity of the plate and xis the coordinate in the direction of welding.

Using an appropriate power value, Q, from a typical FSW and thethermal diffusivity of AA7050 at room temperature, temperaturefields were obtained for a series of welding speeds. The times above200 ◦C are plotted with respect to welding speeds in Fig. 7. The timespent above 200 ◦C is observed to follow a power law relationshipwith the welding speed thus illustrating the diminishing returns inthe increase in cooling rate with the increase in the welding speed.

ng. A (2009), doi:10.1016/j.msea.2009.10.039

3.3. Hardness

Representative hardness distributions after post-weld aging forthe three thermal boundary conditions are shown in Fig. 8a. The

Fig. 7. Time spent above 200 ◦C vs. welding speed obtained using the Rosenthal’ssolution for a moving heat source.

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further increased. For the UW welds, almost no increase in mini-mum hardness is observed above a welding speed of 5.1 mm/s. Thisobservation is consistent with the data in Fig. 6 which indicate thatthe time spent above 200 ◦C in the minimum hardness location isnot reduced substantially by welding at ever increasing rates and

ig. 8. (a) PWHT micro-hardness trends for in-air, under-water, sub-ambient wel.54 mm/s weld for in-air and under-water.

ardness distributions shown are for welds made at 6.8 mm/s and40 rpm. All of the distributions exhibit the “W” shape typical ofXXX type alloys welded under conditions for which the nuggetemperature is close to the solution heat treatment temperature.he nugget regions exhibit hardness which is equivalent to a T6emper and the base metal regions exhibit T7 (slightly overaged)ardness. In the HAZ, substantial overaging leads to a hardness min-

mum. The shape indicates that the nugget region is in a nearlyolution heat treated condition prior to post-weld aging for thiset of welding parameters. The hardness minima in the SA andW welds are not as deep as that observed in the IA weld, consis-

ent with the differences in the temporal length of the temperatureransient as shown in Fig. 6. Also, the near equivalence betweenhe minimum hardness levels in the UW and SA welds is consis-ent with the similar lengths of time spent above 200 ◦C at the HAZ

inimum locations. Fig. 8b illustrates an interesting effect of theubmerged welding for relatively lower power welds. Here, it cane seen that, for IA and UW welds made at 200 rpm and 2.54 mm/sery different hardness distributions are observed. The UW weldhown in Fig. 8b exhibits a flat hardness profile through the nuggethile the IA weld exhibits a W shaped distribution. Compared to

he profiles shown in Fig. 8a, produced at higher rpm and weldower, the peak hardness in the IA weld made at 200 rpm is less.eferring to Fig. 3, it is seen that the probe temperature for the00 rpm IA weld is near 400 ◦C while the probe temperature forhe 200 rpm UW weld is near 350 ◦C. Hence, in the UW weld, over-ging takes place throughout the nugget and in the IA weld, someissolution followed by reprecipitation occurs in the nugget. Theolution treatment is not complete in the IA nugget; therefore theardness is less than in the welds shown in Fig. 8a. On the otherand, the time for overaging of the UW nugget (Fig. 8b) is less thanhat for the IA HAZ leading to a higher minimum hardness in theW weld as compared to the IA weld.

Fig. 9 shows the trends of average nugget hardness with mea-ured probe temperature for IA, UW, and SA welds. The averageugget hardness is taken at the plate mid-plane over the distance

rom edge to edge of the recrystallized region in the weld. Nuggetardness generally increases with the increase in peak probe tem-erature; however, the increase is not linear. At the lowest probeemperatures the nuggets are apparently in an overaged condition

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nd the hardness is similar to that observed in the HAZ min-ma (near HVN = 120). In the intermediate temperature range, theugget hardness rises rapidly with increasing temperature, pre-umably because post-weld, there is a greater amount of solutevailable for reprecipitation. In the highest temperature range, the

de at 800 rpm, 6.8 mm/s and (b) PWHT micro-hardness distribution for 200 rpm,

effect of increasing temperature on hardness diminishes becauseabove some temperature, the additional solute obtained by fur-ther temperature increase becomes insignificant. For a given probetemperature, the UW and SA welds generally exhibit higher nuggethardness after PWHT. This can be rationalized on the basis of ahigher quench rate in the UW and SA welds which will serve to trapmore solute and more vacancies in the as welded nugget, therebyenhancing the response to the post-weld aging treatment.

Fig. 10 shows the trends in HAZ minimum hardness for theSA, UW and IA welds. The minimum hardness values are plottedvs. the welding speed. It has previously been shown that the HAZminimum hardness in precipitation hardening aluminum alloys isalmost solely dependent on the welding speed [2,20]. The UW andSA minimum hardness is consistently higher than that of IA weldsmade with the same welding speed: as for the nugget hardnessvalues, this is most likely due to increased quench rates in the UWand SA welds (see Fig. 6). The increase in minimum hardness withincreasing welding speed is not linear. The increase is rapid at lowwelding speeds and then slows substantially as welding speed is

ng. A (2009), doi:10.1016/j.msea.2009.10.039

Fig. 9. Average nugget hardness vs. peak probe temperature.

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wile

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Fig. 14 shows % elongation in transverse tension (solid symbols)

Fig. 10. HAZ minimum hardness vs. welding speed.

ith the results of the Rosenthal analysis presented in Fig. 7. Thiss a useful observation from a practical standpoint, indicating thatittle improvement in weld properties is likely to arise from greatffort to increase the welding speed above some level.

.4. Nugget grain size

The nugget grain size is plotted against peak probe tempera-ure with corresponding welding speed indicated by symbols inig. 11. Notice that at the welding speed of 6.8 mm/s welds wereerformed at three rotational speeds, hence three data points. Therrows indicate corresponding grain sizes for in-air and under-ater welds made with the same parameters. The grain sizes ofelds performed below 400 rpm, with probe temperatures below

00 ◦C, were not resolvable by optical means and are not reportedere. For the rest of the weld parameters the UW grain size isonsistently smaller than the IA. This can be attributed to theower peak temperature in the UW welds. This should be true

hether one takes the stance that grain size is due to post-weld

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ecrystallization and growth [1] or due to geometric dynamicecrystallization [26–28] or continuous dynamic recrystallization18,29].

ig. 11. Average nugget grain size vs. peak probe temperature for different weldingpeeds. The arrows show equivalent welds done in air and under water.

Fig. 12. Ultimate tensile strength against tool rpm for in-air and under-water cases.

3.5. Tensile strength and correlation with weld hardnessdistributions

Fig. 12 is a plot of the ultimate tensile strength (UTS) in trans-verse tension vs. tool rotation rate for the IA and UW cases. Errorbars indicate the total range of test results for three specimensper condition. Generally, the UW welds are substantially strongerthan the corresponding IA welds. At 200 rpm (welding speed of2.54 mm/s), the UW weld is weaker than the corresponding IA weld.This is the only situation in which any of the welds without a defectbroke in the nugget: the hardness distribution for the 200 rpm UWweld (shown in Fig. 8b) is essentially flat in the HAZ and nuggetareas leading to failure through the nugget. It is evident from exam-ination of all the transverse tensile data that the rpm cannot fullycorrelate the transverse tensile strength. Fig. 13 shows average ten-sile strength plotted against the measured HAZ minimum hardnesson the retreating side of the weld for those welds which exhibited a“W” shaped hardness distribution. Obvious and well known directcorrelation between tensile strength and the minimum hardness isexhibited by all IA, UW and SA welds.

ng. A (2009), doi:10.1016/j.msea.2009.10.039

and the difference between the average nugget hardness and theHAZ minimum values (�VHN, open symbols) both plotted againstthe probe temperature. The solid and dotted lines are fit by eye

Fig. 13. Correlation between ultimate tensile strength and HAZ minimum hardnessin the retreating side of the weld.

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ig. 14. Percentage elongation and difference of hardness values between nuggetnd HAZ minimum plotted against the peak probe temperature. The dotted and solidines are fit by eye.

o the elongation and �VHN data respectively. Firstly, it can beoted that the elongation values are similar for the IA and UWonditions. Elongation is relatively high at low values of probe Tnd declines to a plateau value near 6% at probe T > 400 ◦C. TheVHN increases with increasing probe T, from a value near zero

or a probe T of 320 ◦C to a high of near 60. For probe T substantiallyreater than 350 ◦C, precipitate dissolution in the nugget is occur-ing simultaneously with coarsening in the HAZ leading to higherugget hardness and lower HAZ hardness after PWHT. One sideffect of high nugget hardness is that deformation during trans-erse tensile testing tends to be concentrated in the HAZ. As theVHN becomes larger and larger, the strain concentration in theAZ becomes more prominent and the % elongation declines. Once

he �VHN reaches a high enough value, essentially no deformationccurs in the nugget during transverse loading: here the criticalalue of �VHN is between 35 and 40. The % elongation is notctually a reflection of the ductility of the various weld regions:t is related to variations in strength which cause strain localiza-ion.

.6. Summary and conclusions

In this study the effects of changing the thermal boundaryondition and the FSW control variables on weld properties andome FSW response variables have been examined. Several generalrends have been observed and some useful correlations provided.pecifically:

1) All other things being equal the increased surface convectionresulting from welding under water compared to in-air welding

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results in:(a) Reduced probe temperature.(b) Increased torque and power consumption.(c) Decreased nugget grain size.(d) Increased cooling rates in the HAZ.

[[[[[

PRESSand Engineering A xxx (2009) xxx–xxx 7

(2) While not measured, increased cooling rates in the weld nuggetare implied by the higher hardness observed for UW vs. IA weldswith equivalent probe temperature (Fig. 9).

(3) Increased cooling rates in the HAZ due to enhanced convectionresult in higher HAZ minimum hardness values for UW and SAwelds vs. otherwise equivalent IA welds.

(4) Transverse tensile strength is enhanced by UW welding exceptfor welding conditions for which the UW weld temperature istoo low to enable substantial reprecipitation in the weld nuggetduring PWHT. The transverse tensile strength exhibits a strong,direct, correlation with the HAZ minimum hardness.

(5) Temperature measurement in the HAZ minimum hardnessregion indicates that, for the experimental setups under consid-eration, little reduction in time at temperature can be achievedby welding at speeds greater than 5.1 mm/s.

(6) Under the conditions studied, pre-cooling to −25 ◦C does notprovide a significant benefit compared to welding under waterat ambient temperature.

Acknowledgements

The authors acknowledge the financial support of the Centerfor Friction Stir Processing which is a National Science FoundationI/UCRC supported by Grant No. EEC-0437341. The authors thankDr. Wei Tang and Daniel Wilhelm, Department of Mechanical Engi-neering, University of South Carolina, Columbia, SC, USA for theirhelp in preparing the weld joints.

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