multi-pass friction stir welding in alloy 7050-t7451: effects on weld response variables and on weld...

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Materials Science and Engineering A 513–514 (2009) 115–121

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Materials Science and Engineering A

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ulti-pass friction stir welding in alloy 7050-T7451: Effects on weld responseariables and on weld properties

ebecca Brown, Wei Tang, 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 2 October 2008eceived in revised form 14 January 2009ccepted 15 January 2009

a b s t r a c t

In some situations it may be necessary or desirable to perform a friction stir weld pass through material,which has already been friction stir welded. Examples include weld repair or crossing weld beads. Ineither case, one would like to know whether the control parameters used for welding require modificationfor passes through existing welded material and what is the effect of the multiple passes on the weld

eywords:riction stir welding050ultiple passesverlapping passes

properties. To examine this issue, multiple weld passes (five) were performed in 6.4 mm thick 7050-T7451plate at a tool rotation speed of 540 rpm and a welding speed of 6.77 mm s−1. Results indicate that forthis material and these conditions weld control parameters require no adjustment. The overall reductionfrom pass 1 to 5 in transverse tensile strength is 7%. In addition, weld metallurgy is only slightly changedin the heat affected zone (HAZ) due to overaging for each weld pass while there is no change in metallurgyin the nugget. Hardness in the HAZ reduces by 14 Vicker’s hardness points from pass 1 to 5. The peak

ss is r

. Introduction

The development of friction stir welding (FSW) [1] has pro-ided a great opportunity for industry to utilize welding of hightrength aluminum alloys, such as 7XXX series alloys, in the produc-ion of high performance aerostructures. When properly specified,SW is capable of producing welds with a very low incidence ofefects; however, it is not possible to completely prevent processpsets. Hence, defects may arise during production of the fric-ion stir welded structure. Rather than scrap an expensive piecef hardware, it would be desirable to be able to repair the defec-ive portion of the weld. One technique that may be used for repairf defects arising from process upsets is simply rewelding usinghe nominal process parameters. In such a case, it is important tonow whether or not the nominal process parameters are capa-le of producing defect free welds in the already welded materialwhich will presumably have somewhat different properties asompared to the base metal) and what additional levels of prop-rty reduction may accompany the additional weld pass(es). Inhis paper the effects of performing multiple friction stir weld

xamined.

∗ Corresponding author. Tel.: +1 803 777 9548; fax: +1 803 777 0106.E-mail address: Reynolds@engr.sc.edu (A.P. Reynolds).

921-5093/$ – see front matter © 2009 Elsevier B.V. All rights reserved.oi:10.1016/j.msea.2009.01.041

educed with increasing number of passes.© 2009 Elsevier B.V. All rights reserved.

2. Background

Aerospace aluminum alloys of the 7XXX series are high strengthmaterials which are generally heat treated in order provide anoptimum balance between strength, toughness, and stress corro-sion cracking resistance. They are precipitation strengthened alloyswith the primary strengthening precipitates being the �′, Mg2Znphase [2]. As for essentially all precipitation hardening aluminumalloys, the heat treatment of the base metal is roughly as follows:a solution heat treatment followed by rapid quenching producesa supersaturated solid solution. The solid solution is decomposedduring an aging treatment which is comprised of one or more stepsof holding the material at a temperature less than the solutiontreatment temperature. The resulting microstructure exhibits anoptimized distribution of fine precipitates. The temperature histo-ries associated with any type of welding operation will modify the“optimized” microstructure to varying extents depending on thetype of welding and the specific parameters (heat input, weldingspeed) applied. It should be noted that many precipitation harden-ing alloys contain a greater level of alloying elements than can bebrought into solid solution: in many 7XXX series alloys, this fact,coupled with the presence of low-melting phases, can result in thephenomenon of local melting at temperatures well below the bulk

A significant amount of work has been published on single passFSW of 7XXX series aluminum alloys [2–9]. The general features ofaluminum alloy friction stir welds have been described in numerouspublications so they will be only briefly reiterated here. Aluminum

1 d Engineering A 513–514 (2009) 115–121

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16 R. Brown et al. / Materials Science an

lloy friction stir welds are normally divided into a central nuggetone, thermo-mechanically affected zones (TMAZ) on either side ofhe nugget, heat affected zones (HAZs) adjacent to the TMAZ’s, andnaffected base metal. The nugget is the region of the weld whichas undergone recrystallization, the TMAZ is deformed but notecrystallized, and the HAZ is not deformed but has a microstruc-ure which has been altered by the temperature history associatedith the welding process.

Most FSW of 7XXX alloys has been performed on base metalshich are in either the T6 or T7XX type tempers, meaning that

he base metal has been aged to some specific strength level corre-ponding to some precipitate distribution. The effects of the FSWrocess on the strength/hardness distribution near the weld haveeen widely studied and it has been shown that this distributionan be manipulated, within limits, by the choice of welding param-ters [6,10]. Typically, the hardness distribution resulting from FSWf precipitation hardened aluminum alloys (including those of theXXX series) has a characteristic “W” shape with a local or absoluteaximum in the nugget and minima in the heat affected zones. In

rder to produce the W shape, the nugget must be heated up to, orearly up to, the solution heat treatment temperature. This enableshe strength of the nugget to approach or exceed that of the base

etal after suitable post-weld aging. Attainment of the solutionreatment temperature in the nugget requires welding with some

inimum power level, which will depend on the welding speednd the tool design. In the HAZ, the hardness is normally reducedelow that in the base metal due to coarsening of the base metalrecipitate distribution. The depth of the hardness minimum in theAZ is closely tied to the welding speed and is not generally related

o the peak temperature experienced in the nugget [11]. Relativelyow welding speed results in relatively low HAZ hardness due toonger time spent in the overaging temperature regime. In certainases, corresponding to low power welding conditions, the char-cteristic W shape is not produced: the hardness is more or lesshe same in the nugget and the HAZ. This results from a peak tem-erature in the nugget, which is in the overaging rather than theolution treatment range; hence, the precipitates in the nugget areot dissolved and re-precipitated, they only coarsen.

Only a minimal amount of research has been performed previ-usly in Portugal involving friction stir welding of multiple passes..M. Leal and A. Loureiro performed four overlapping passes withA5083-O and three overlapping passes with AA6063-T6 [12]. Therocess parameters used on pass 1 were maintained for all passes.A5083 showed an increase in weld hardness and strength due tolastic deformation and dynamic recrystallization in the nugget andMAZ from the base metal to pass 1. The hardness increases from thease metal to pass 1 as a result of plastic deformation. The hardnessurves between passes one through four show little variation whencatter is taken into account. Transverse tensile testing of passeshree and four showed increases in yield strength and ultimate ten-ile strength from base metal to pass 3 to 4. All fracture occurred inhe base metal. AA6063 showed much lower hardness and strengthalues in comparison to the base metal. Hardness is similar betweenass 1 and 2 while tensile properties decrease. Hardness betweenass 2 and 3 increases along with tensile properties. All fracturesccurred in the retreating side of the HAZ. Grain size in the nuggetemains fairly consistent between passes for both alloys indicatingimilar microstructure. Also, a defect found along the entire lengthf pass 1 of AA5083 was completely eliminated by pass 4.

. Materials and experimental procedures

.1. Alloy 7050

All friction stir welding was performed on 6.4 mm thick platesf 7050-T7451. Alloy 7050 is a high strength aluminum alloy with

Fig. 1. Multi-pass FSW.

a nominal composition (weight %) of 6.2Zn–2.3Mg–2.3Cu–0.12Zrand balance aluminum. The typical ultimate strength of the alloyin the T7451 temper is 524 MPa. The incipient melting tempera-ture for homogenized 7050 is 488 ◦C and the solution treatmenttemperature is 477 ◦C [13].

3.2. Welding procedure

Square butt welds were made at the interface between sets oftwo plates. The plates were each 102 mm wide by 914 mm long.Welds were overlapped in order to produce material that had beenwelded 1–5 times. The multi-pass welds are designated by a num-ber and a letter (e.g. 2417A). The weld number is associated witha set of plates and the letter is associated with the pass number.Fig. 1 is a photograph showing a multi-pass weld. The first weld,2417A, extends the length of the plates while 2417B covers two-thirds of 2417A and 2417C covers half of 2417B. The resulting platecontains ≈300 mm each of 1–3 pass welds. A similar techniquewas used to produce another batch of material, which had beenwelded/processed up to 5 times.

Welds were produced on an MTS FSW Process DevelopmentSystem (PDS) using Z-axis force control (28 kN) with a tool rota-tion speed of 540 rpm and a tool traverse speed of 6.8 mm s−1. Thetool rotation axis was normal to the plate surface. The tool wasa two-piece design. The 17.8 mm diameter single scroll shoulderwas machined from H13 tool steel. The MP-159 probe was a trun-cated cone (8◦ taper) with threads and three flats. The probe was6.1 mm long, with a diameter of 7.9 mm at the intersection withthe shoulder. Temperatures during welding were measured usinga thermocouple spot welded into the probe on the axis of rota-tion at the weld midplane height. Temperature data was acquiredat 1 Hz using a Hobo data logger (Onset Computer Corp.), whichwas attached to the spindle of the PDS. All welds were allowed tocool for 1 h prior to performing subsequent passes: clamping wasleft in place during each series of welds. After 2 weeks of natu-ral aging, each completed weldment was heat treated for 24 h at121 ◦C.

3.3. Metallography and hardness testing

Samples for metallography and micro-hardness evaluationswere cut from each weld pass using an abrasive waterjet. Sam-ples were ground on an automatic grinder with silicon carbidepaper of 320, 600 and 800 grit. The samples were then polishedwith 5 and 3 �m aluminum oxide powder and finished with col-loidal silica (<0.05 �m) in an automatic polisher. The samples wereetched for approximately 7 s using Keller’s etchant. Micrographswere obtained from the center of the nugget region at magnifi-cations of 500× to perform grain size measurements. The meanlinear intercept method was used on three images from each

pass.Vicker’s hardness was measured on transverse cross-sections at

the weld midplane, over a distance of 36 mm centered on the weldcenterline. Measurements were made with a spacing of 0.64 mm,using a 1 kg load and a loading time of 10 s.

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.4. Transverse tensile testing

Tensile specimens, 178 mm long and 12.7 mm wide, were cutrom each weld pass using a water jet cutter. The edges were milledat and deburred. Three specimens from each weld pass wereested in a MTS 810 Material Test System, at a displacement ratef 0.0254 mm s−1. Data was acquired at a sampling rate of 2 Hz. Anxtensometer (25.4 mm gage length) was used to measure speci-en strain across the weld region of one specimen from each weld

ass.Two tensile tests from each pass were performed using digital

mage correlation (DIC) to measure full field strain. This methodas been described in detail in previous publications [14]. Images

or correlation were obtained every 2 s during the tensile testingsing an Allied Dolphin F-201B camera with a Nikon Nikkor 28 mm2.8D AF lens. The images were analyzed using VIC2D softwareCorrelated Solutions, Columbia, SC).

.5. Residual stress measurement

Longitudinal residual stress was measured for a one, three andve weld pass using the cut compliance technique [15]. In the cutompliance technique, a slot is extended through a specimen hav-ng the same plan form as a compact tension specimen (see Fig. 2).s the slot is extended, residual stress is relieved and the resultingtrain, ε, is measured with a strain gage on the back face of the spec-men, opposite to the notch. Then the stress intensity factor due toesidual stresses, KIres, can be calculated from the strain distribu-ion as a function of the notch length, a, as shown in Eqs. (1) and2):

Ires = E′

Z(a)dε

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(W − a)3/2(2)

here E′ is the Young modulus, W the specimen width and Z(a) ishe influence function given by Schindler [16] and is only dependentn the sample geometry. The residual stress profile can then beetermined from the stress intensity factor, by inverting Eq. (3) as

Fig. 2. Specimen for cut compliance testing.

ineering A 513–514 (2009) 115–121 117

explained by Prime.

KIres =∫ a

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Where h(x, a) is a weight function depending on the geometry ofthe specimen. This technique provides a through thickness averageof the residual stress in the welding direction.

4. Results and discussion

4.1. FSW response variables

Table 1 shows the response variables for weld passes 1–5. Theprobe temperature is very nearly the same for all passes: the low-est measured temperature is 466 ◦C and the highest is 471 ◦C. Thetemperature measured inside the probe is some average of the tem-perature of all of the material in contact with the probe and cannotbe directly related to the temperature at any specific position; how-ever, it should provide a reasonable gage of the relative temperaturefrom weld to weld. In this case, it is apparent that the weld tem-perature does not vary greatly from one pass to the next. It is alsonoteworthy that the temperature measured inside the probe oneach pass is close to the solution treatment temperature for alloy7050 so, presumably, after welding, the material in the weld nuggetshould be in a solution treated or nearly solution treated condition.

The average torque and power increase very slightly withincreasing pass number (6.5% from pass 1 to 5). It is not clear thatthis small effect is significant.

The most salient effect of weld pass number is on the x-axis force.The x-axis force is significantly higher for pass 1 than for subsequentpasses. This observation may be rationalized by the differences inthe temper, hence hardness, of the material being welded. On thefirst pass, the material in front of the tool is in the T7451 conditionwith a hardness of HV ≈ 170. While the hardness of the material inthe weld path was not measured in between passes, it is likely thatthe solution treated, but not aged, material which is welded on eachpass subsequent to the first is substantially softer than the T7451material. It is important to keep in mind that the welded materialis not artificially aged between passes and will be in essentially ashort time (1 h) naturally aged, W condition. Hence, the “freshly”welded material may provide less resistance to the passage of thewelding tool than does the original base metal.

Interestingly, the y-axis force changes relatively little comparedto the x-axis force. The maximum percentage variation of the x-axisforce relative to pass 1 is −34% while the y-force varies from −19%to +15% of the pass 1 value indicating that the y-force variation maybe a variation around an average which is independent of the passnumber, while the x-force variation is due to changes in the mate-rial being welded after the initial pass. Fig. 3 is a polar plot of thein-plane force vector acting on the welding tool for each pass. In thefigure, a positive x-force is acting opposite the welding direction anda positive y-force is pushing the tool toward the advancing side. Thelength of the lines represents the total magnitude of the in-planeforce and the resultant angle is measured relative to the weldingdirection with the positive direction indicating deviation towardthe advancing side. The relatively lower and non-monotonic devia-tion of the y-force from the pass 1 value may, perhaps, be explainedbased on the kinematics of the FSW process. In several previouspublications, both experimental and finite element based simula-tion studies have indicated that the volume of material displaced

118 R. Brown et al. / Materials Science and Engineering A 513–514 (2009) 115–121

Table 1Process response variables.

FSW no. Pass # Probe temp. (◦C) Torque (N m) Power (kW) Ave. X force (kN) Ave. Y force (kN)

2418 A 1 470 54.0 3.06 7.38 4.442417 B 2 471 55.4 3.12 5.26 4.592417 C 3 466 55.7 3.14 5.29 5.092418 D 4 471 56.2 3.16 5.09 4.462418 E 5 470 57.7 3.25 4.85 3.59

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Fig. 5. Typical nugget grain appearance.

Fig. 6. Nugget grain size by MLI method at nugget center for each weld pass. Errorbars indicate 95% confidence intervals.

ig. 3. Polar plot of the in-plane forces acting on the tool. Welding direction is fromight to left, and the advancing side is up.

onsidered nearly incompressible, so the passage of this materialround the retreating side of the probe will provide a pressure loadn the tool which is (1) causing a force on the tool in the directionf the advancing side and (2) dependent primarily on the volumeer revolution which is displaced by the probe. As stated above, thisolume is the same for each weld pass; hence, the y-axis force mayot be expected to vary a great deal.

.2. Metallography and microstructure

Fig. 4 shows a typical weld cross-section from pass 1. All otherasses were quite similar in appearance. No volumetric defectsere found in any of the passes when observing their cross-sectionsith the light optical microscope. A representative micrograph of

he nugget center grain structure from pass 1 is shown in Fig. 5.ll other passes appeared similar. The grain size measured at theugget center for each pass is plotted in Fig. 6 (error bars are 95%onfidence limits). Only slight and non-monotonic variation wasbserved from pass to pass, consistent with the similar measuredrobe temperatures. It has been observed in several previous stud-

es that the welding temperature is closely tied to the final grainize in the weld nugget [21–23].

Fig. 7 shows the hardness distributions on the midplanes of the

ent). Each pass exhibits the typical W shaped hardness profile ofrecipitation hardened alloys welded with sufficient power to puthe nugget region into solid solution. The base metal hardness is

Fig. 4. Scanned transverse cross-section of pass 1.Fig. 7. Transverse hardness profiles at weld pass midplanes. The horizontal, doubleended arrow is discussed in reference to Fig. 9.

R. Brown et al. / Materials Science and Eng

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ig. 8. HAZ minimum hardness on advancing and retreating sides as a function ofass number.

ear 170 HV while the hardness of the nugget of each pass is near90 HV. This difference is due to the fact that, after post-weld aging,he nugget is essentially in a T6 temper while the base metal is inn overaged T7451 condition. The hardness minima are observedn the HAZs approximately 7–8 mm from the weld centerline.

The nugget grain size and nugget hardness of each of the passeso not vary a great deal. This indicates that each time the FSW pro-ess is performed on the material, the microstructure in the weldugget is “reset”. This resetting shows that the deformation andemperature history associated with the FSW process are sufficiento override any temper or grain structure effects related to the con-ition of the material as it enters the processing zone. However,his is not true of the material in the HAZ. In the HAZ, the effects of

ultiple passes accumulate. This effect is illustrated in Figs. 8 and 9.Fig. 8 shows that there is a general trend toward lower HAZ min-

mum hardness with increasing pass number. It is to be expectedhat there will be significant scatter in these results (e.g. pass 4) ashe spatial resolution of the measurement is limited and, as can beeen in Fig. 7, the gradient in hardness near the HAZ minimum isuite steep. Nevertheless, it is apparent that the overaging in the

ig. 9. Hardness “well” width on advancing and retreating sides as a function ofumber of weld passes.

ineering A 513–514 (2009) 115–121 119

sides of each weld pass. This width is represented approximately bythe length of the double ended, horizontal arrow shown in Fig. 7.The “well width” is determined arbitrarily at a hardness value ofHV = 150: that is, the measurement represents the width of the HAZfor which the hardness level is below HV = 150. As for the mini-mum hardness, some scatter should be expected and the reductionin well width observed for pass 4 on the advancing side can onlybe explained by scatter (or by a variation in the height within theweld at which the traverse was performed); nevertheless, in gen-eral, a cumulative effect of increasing pass number on HAZ width isobserved. As the pass number increases, the width of the hardnesswell increases. Also notable in this data is a definite trend for greaterwell width on the retreating side than on the advancing side.

4.3. Tensile test results

Fig. 10 shows the average tensile properties (tensile strength,global yield strength, and elongation in 25.4 mm) from the trans-verse tensile tests for each weld pass. Also shown is the local 0.2%offset yield strength in the retreating side HAZ at the midplane (HAZ(R) yield strength). The local yield stress is obtained by mappingthe local strain to the global stress as described by Lockwood et al.[14,24]. The global yield strengths were determined by a 0.2% strainoffset method using elongation data from a 25.4 mm gage length.The transverse tensile samples are not homogeneous as the variousweld zones are arranged in series for this type of loading; hence,the reported global yield strength is higher than the actual yieldstrength of the lower yielding material in the HAZ by an averageof 45 MPa (for the local measurement location used here). This isbecause the actual plastic strain in the first material to yield will begreater than 0.2% when a 0.2% plastic offset is observed on the globalstress–strain curve. As a general rule, the transverse tensile proper-ties, yield strength, tensile strength, and elongation to fracture, alldeclined with increasing number of passes.

Without exception, all 12 transverse tensile specimens frompasses 1–4 failed in the retreating side HAZ. Fig. 11 shows imagecorrelation strain maps at load levels close to final fracture and thecorresponding broken tensile specimens for passes 1, 4 and 5. Ascan be seen from the strain maps, the strain localizes in both HAZ’s(advancing and retreating sides) but is more strongly localized inthe retreating side HAZ. The softened HAZ regions in which thestrain is localized are inclined to the plate normal because the HAZ

Fig. 10. Transverse tensile properties as a function of pass number.

120 R. Brown et al. / Materials Science and Engineering A 513–514 (2009) 115–121

re load and corresponding, fractured tensile specimens.

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he tool shoulder and narrower at the root) one expects that higheremperatures will be experienced farther from the centerline at therown than at the root.

It is possible that there is, in general, a lower hardness in theetreating side HAZ (this is weakly supported by the data in Fig. 8),ut this may not be a requirement for the more severe strain local-

zation and final fracture which occurs in the retreating side HAZ.ased on the HAZ width measurements (Fig. 9) the advancing sideAZ is significantly narrower than the retreating side HAZ, hence,

he constraint of the soft material in the HAZ by the adjacent, harderaterial (nugget and base metal) will be greater in the advancing

ide than in the retreating side. This constraint difference couldotentially lead to an inhibition of yielding and strain concentra-ion in the advancing side HAZ, leading to higher strain and fracturen the retreating side HAZ. The potential effect of the narrowerdvancing side HAZ is analogous to the notch strengthening effect: aarrower notch will produce a greater level of triaxial tensile stresshence elevation of load required for yielding) than will a widerotch [25]. Also of note are the actual strain levels in the HAZs justrior to fracture: in every case, the local strain in the retreating sideAZ is greater than 0.1 (10%).

The fracture behavior of all three of the pass 5 specimens was dif-erent from that observed in passes 1–4. The strain localization prioro fracture in pass 5 was similar to that observed in the prior passes;owever, the fracture initiated on the weld root in the retreatingide HAZ and then deviated from the “normal” mode and fractureccurred through the weld nugget. The strain map and fracture pathor pass 5 are also shown in Fig. 11. At the present time it is notertain why this deviation occurred but it may be linked to thencreasing depth of the hardness well. This subject will be exploredurther in research to follow.

.4. Residual stress

Fig. 12 shows the through-thickness, average, longitudinal,esidual stress for passes 1, 3 and 5. As has been observed previ-usly [26–29], the center of the weld is surrounded by a region of

inimum. In these results, the maximum stress on the retreatingide is greater than that on the advancing side. As required for equi-ibrium, there are compressive stresses outside of the weld regionn the base metal region.

Fig. 12. Through thickness longitudinal residual stress.

The measured probe temperatures are similar for each pass, andeach pass is performed at the same welding speed; hence, it islikely that the thermal cycle responsible for the residual stress isthe same on each pass. However, the peak values of residual stress,both tensile and compressive, decline with increasing number ofpasses. The stress reduction is greater between passes 1 and 3 thanbetween 3 and 5. This trend is mimicked by the observed reduc-tions in HAZ hardness and transverse tensile properties: that is,the observed reductions in HAZ hardness and transverse tensilestrength are greater between passes 1 and 3 than between passes3 and 5. It seems likely that the minimum hardness and the peakvalues of residual stress are linked as the maximum value of theresidual stress will be effectively limited by the yield strength ofthe material [30]. Hence, as the HAZ strength is reduced, the maxi-mum attainable residual stress is also lower. In addition to the dropin minimum hardness, there may also be an effect of the wideningof the HAZ.

5. Summary and conclusions

1. For the welding conditions in this study, multiple friction stirweld passes may be performed through the same material with-out the need to adjust the control parameters.

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R. Brown et al. / Materials Science an

. The x-force is significantly reduced when welding through pre-viously welded material as compared to that in the base metalT7451 condition.

. The grain size and hardness of the nugget are essentiallyunchanged by application of multiple weld passes: this indicatesan insensitivity of this alloy to initial temper for the weldingparameters used. Additionally, the temperature during weldingis not affected by number of passes.

. Metallurgical changes in the HAZ are cumulative with weldingpass number.a. HAZ hardness and transverse tensile strength are reduced

with increasing number of passes.b. The softened region associated with the HAZ becomes wider

with increasing pass number.. Transverse tensile failure always initiated in the retreating side

HAZ. For passes 1–4 the actual fracture occurred through theHAZ. This may be an effect of the wider softened region in theretreating side HAZ in comparison to the advancing side HAZ.

. Residual stress is reduced by multiple welding passes for theconditions used in this study.

Some issues related to multi-pass welding of 7050 remainutstanding. The anomalous fracture path observed in transverseension after pass 5 is not understood. Nor is it clear why the retreat-ng side HAZ is substantially wider than that on the advancingide, although the fact that it is may help to explain some otherbservations. These issues will be examined in subsequent studies.

cknowledgement

This research was supported by the NSF/IUCRC Center for Fric-

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