identiﬁcation of process

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P. EdwardsDepartment of Mechanical Engineering,

University of Washington,Seattle, WA 98195;

Boeing Research & Technology,The Boeing Company,

Seattle, WA 98124

M. RamuluDepartment of Mechanical Engineering,

University of Washington,Seattle, WA 98195

Identification of ProcessParameters for Friction StirWelding Ti–6Al–4VAn experimental study was conducted to identify acceptable welding parameters forfriction stir welding Ti-6Al-4V butt joints, ranging from 3 mm to 12 mm in thickness. Theprimary parameters of interest were the spindle speed and feed rate. Welds were pro-duced using spindle speeds of 140–320 rpm and feed rates between 40 mm/min and 125mm/min. Joints were evaluated by macro- and micrometallurgical examination alongwith limited fatigue and tensile testing. The weld parameters were found to influence themicrostructure, penetration, void formation, and tool wear among other things. A processwindow was identified for combinations of the feed rate and spindle speed capable ofachieving defect free joints for a given tooling configuration and thickness. It was foundthat the tensile and fatigue properties of the welds produced in this study were compa-rable to Ti–6Al–4V base material properties. �DOI: 10.1115/1.4001302�

Keywords: friction stir welding, titanium, microstructure, fatigue, tensile behavior

IntroductionFriction stir welding �FSW� has been successfully applied to

elatively soft materials such as aluminum and copper �1�. Onlyecently has this process been implemented on high strength ma-erials such as titanium and steel, primarily due to difficultiesssociated with identifying tooling materials and equipment ca-able of withstanding the high forces and temperatures involvedith welding these alloys. To date, there have only been a limitedumber of reports on friction stir welds in titanium alloys. Here, arief literature review of the work performed in titanium frictiontir welding is presented, followed by the results of a new experi-ental study aimed at identifying the process parameters neces-

ary to produce defect free joints in a variety of material thick-esses in the most widely used titanium alloy, Ti–6Al–4V.

Review of Titanium FSWTrapp et al. �2� were one of the first to report on the FSW of

itanium, specifically the FSW of 12 mm thick Ti–17 and Ti–6Al–V. The investigation included parameter development, metallur-ical examinations, mechanical testing, postweld heat treat evalu-tions, tool material assessments, and thermal management.owever, little detail with respect to the process parameters or

ools used was provided. Leinert �3� recently provided an excel-ent review of titanium FSW. Tool materials, tool designs, weldingarameters, microstructures, microhardness, and mechanical prop-rties were presented for commercially pure titanium �CP Ti�, Ti–Al–4V, Ti–15V–3Cr–3Al–3Sn, and Beta 21-S. With respect toi–6Al–4V, Leinert et al. �4� first referenced his own work ini–6–4, where the tool geometry, tool material, and weld param-ters were provided. A pure tungsten tool was used with a 6.4 mmong cylindrical pin, a 7.9 mm pin diameter, and a 19 mm shoul-er with no other features such as threads. The welding param-ters used were a 275 rpm spindle speed and travel speeds of up to00 mm/min. Tool temperatures, weld microstructures, and me-hanical properties were also evaluated. Microstructural evalua-ions revealed a refined transformed beta/acicular alpha structure,ndicating supertransus welding temperatures. Tool temperature

Contributed by the Materials Division of ASME for publication in the JOURNAL OF

NGINEERING MATERIALS AND TECHNOLOGY. Manuscript received December 8, 2009;nal manuscript received February 15, 2010; published online June 16, 2010. Assoc.
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measurements of 1115°C also indicate supertransus process tem-peratures. The microhardness of the welds was found to be similarto the base metal, but with a hardness increase in the heat affectedzone �HAZ�. Tensile testing showed higher yield and tensilestrengths of the welds and equivalent elongations �%e�, comparedwith the parent material. Leinert also noted the difficulty of iden-tifying acceptable tooling materials for FSW of Ti–6Al–4V be-cause of the temperatures involved. It was stated that pure tung-sten, W–25%Re tungsten-rhenium with HfC, and sintered TiCtools have successfully been used to FSW titanium. Active ther-mal management techniques, such as cooled tool holders, werealso recommended.

Ramirez and Juhas �5� also reported on the FSW of Ti–6Al–4Vmill and beta annealed 6 mm thick plate. The welding parametersused were identical to those reported in Ref. �3�. The only differ-ence was that a W–25%Re tool was used. Microstructures ofwelds in both material conditions were reported to be very similar,consisting of small prior beta grains with thin layers of grain-boundary alpha, also indicating supertransus welding tempera-tures. This implies that the microstructural evolution depends onthe thermomechanical cycles during FSW and is not dependent onthe starting microstructure. This was a key observation.

The FSW of 2 mm thick Ti–15V–3Cr–3Al–3Sn, Ti–6Al–4V,and CP Ti sheets using W–25%Re alloy tool have also been re-ported on �3�. The tool used had a 14 mm shoulder diameter anda 1.9 mm pin length. No threads or other profiles were used on thepin. Tool plunging was performed in position control, and thetraverse was conducted in load control for Ti–15V–3Cr–3Al–3Snand CP Ti, while welds were made in position control for Ti–6Al–4V. A tool tilt of 1 deg, with a spindle speed of 200 rpm and atravel rate of 100 mm/min were used. Successful welds could notbe made in load control for Ti–6Al–4V. Forge loads of either 9.8kN or 10.7 kN were used for Ti–15V–3Cr–3Al–3Sn, and muchhigher loads were needed for CP Ti welds. Travel loads for the CPTi and Ti–6Al–4V welds were higher than those for Ti–15V–3Cr–3Al–3Sn. No significant tool wear was observed. CP Ti was foundto be very difficult to weld. All CP Ti welds have excessive flash,sheet thinning, and poor surface finish. The CP Ti weld nuggetspossessed a refined grain size. Ti–6Al–4V was also found to bedifficult to weld. A better surface finish was achieved in Ti–6Al–4V compared with CP Ti, but a consistent lack of penetration
was evident. Again, a refined grain size in the Ti–6Al–4V weld
JULY 2010, Vol. 132 / 031006-110 by ASME

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ugget was observed. Full penetration, defect free welds wereeported in Ti–15V–3Cr–3Al–3Sn, which possessed a refined betarained microstructure.

With respect to microhardness, the CP Ti and Ti–6Al–4V weldshowed an increase in hardness compared with the parent mate-ial, while the microhardness of the Ti–15V–3Cr–3Al–3Sn weldsere more similar to the parent material �3�. The hardness of theeat affected zone of the CP Ti welds was lower than the parentaterial, while the hardness of the Ti–6Al–4V welds were higher

han the base material �BM�. Tensile properties of the CP Ti weldsere much lower than the base material because of weld defects

nd thinning. Failure in the CP Ti tests occurred in the weld nug-et. A 95% joint efficiency was observed for the Ti–6Al–4Velds, but elongations were reduced compared with the baseetal. Failure of the Ti–6Al–4V specimens occurred in the weldhen lack of penetration defects were present. For the Ti–15V–Cr–3Al–3Sn, an increase in the ultimate strength �US� and yieldtrength �YS� was observed in the weld compared with the parentaterial with failure occurring outside the weld zone, but thereas a reduction in the elongation. This elongation reduction was

xpected to be a result of the microstructural gradients in the testpecimen gage length.

FSW of Beta 21-S �6� and CP Ti �7� has also been performed byther researchers. Beta 21-S welds were performed on 1.6 mmheet suing a tungsten alloy tool. A 200 rpm spindle speed wassed at travel speeds of 50–300 mm/min. Defect free welds wereroduced with a refined weld nugget grain size. CP Ti welds in 5.6m plate were welded with a sintered TiC tool in water cooled

older. Defect free welds were made at 1100 rpm spindle speednd feed rate of 500 mm/min. Microhardness and tensile proper-ies of the CP Ti welds were found to be similar to that of the base

etal.Zhang et al. �8� explored the microstructures and mechanical

roperties of 3 mm thick Ti–6Al–4V FSWs made under variedool rotation speeds using a Mo based tool. The plunge depth andravel speed of the tool were held constant at 2 mm and 60 mm/

in, respectively. The spindle speed varied from 300 rpm to 600pm. Defects were observed in the 300 rpm and 600 rpm welds,ut defect free welds were obtained at 400 rpm and 500 rpm.icrostructural investigation of the welds showed a refined grain

ize with evidence of exceeding the beta transus temperature. Fur-hermore, the grain size of the weld nugget increased with increas-ng rotation rate due to the higher processing temperatures asso-iated with higher tool rotation rates. The microstructures of theeat affected zones were all formed below the beta transus tem-erature. In general, it was found that the hardness of the weldsas greater than that of the base material, and softening was ob-

erved in the HAZ. This was the first time a HAZ softening wasbserved. It was also found that the hardness of the welds de-reases with increasing spindle speed. Tensile testing of the entireoint showed that all welds had lower strength and elongation thanhe base material, and all properties decreased with increased toolotation speed. Failures occurred in the HAZ due to the softeningbserved. Tensile testing of the stir zone showed higher strengthsnd elongations than the base material. These properties also de-reased with increasing spindle speed. Zhang et al. �9� also at-empted to weld Ti–6Al–4V with a polycrystalline cubic boronitride �pcBN� tool. These welds were unsuccessful because theeaction between the titanium and pcBN led to extreme tool wear.

Pilchak et al. �10� used friction stir processing �FSP� to modifyhe microstructure of cast Ti–6Al–4V. Welds were made using a

–25%Re tool at 100 rpm and 10 mm/min under a 38.7 kN forgeoad. It was found that FSP can transform the large grained cast

icrostructure to very fine equiaxed alpha grains. Some asymme-ries were also observed in the microstructure, such as from thedvancing side of the weld to the retreating side, due to strain andemperature gradients during the process. Fonda et al. �11� exam-ned the microstructural evolution during the FSW of Ti–5Al–
Sn–1Zr–1V.
31006-2 / Vol. 132, JULY 2010


Little information is available on the durability and damagetolerance properties of friction stir welded titanium. John et al.�12� were the first to report on the crack growth in friction stirwelded titanium. It was found that the crack growth rates weredifferent in compact tension �CT� specimens than in middle ten-sion �MT� specimens, due to the influence of residual stress fields.Pasta and Reynolds �13� reported on the effect of residual stresseson the crack growth rates in 2 mm thick Ti–6Al–4V FSWs. AW–25%Re tool was used with a flat 15 mm diameter shoulder anda cylindrical 5 mm diameter pin. Welds were made under forcecontrol using a load of 36.5 kN. A 150 rpm spindle speed and a100 mm/min fee rate were used. A refined microstructure consist-ing of recrystallized alpha and beta grains, which formed belowthe beta transus, was seen. This is the first report of a weld thatmay have been made below the beta transus temperature. A slighthardness increase was observed in the weld nugget compared withthe HAZ and base material. Residual stresses were found to benearly 200 MPa in the weld nugget and over �100 MPa in thesurrounding base material. The peak tensile residual stresses wereobserved on the weld centerline. The residual stresses were higheron the advancing side of the weld than on the retreating side.These residual stresses led to lower crack growth rates in the weldzone than the base material. Fractography found that the failuresurfaces in the weld were similar to that of the base material, butsmearing of the striations in the weld was seen, because of thepresence of compressive residual stresses in the weld.

Sanders et al. �14� recently worked to combine FSW and super-plastic forming �SPF� of Ti–6Al–4V, since the solid state nature ofthe FSW process retains the fine grained, and superplastic, char-acteristics of the wrought parent material. In Ref. �15�, the micro-structures, microhardness, and tensile properties of 2–3 mm thick,standard and fine grained, Ti–6Al–4V FSWs were reported in theas-welded, stress relieved, and superplastically formed conditions.Metallographic examinations noted a refined grain size. It wasalso found that joint hardness and strengths were similar to theparent material, but elongations were degraded. A follow-up onthe study �16� to assess the high cycle fatigue performance of 2.5mm thickness welds showed that if the FSW tool marks were lefton the joint, very low fatigue lives were achieved. By removingthe tool marks, the fatigue life was increased to 80% of the parentmaterial. Burnishing the machined welds increased the life to 90%of the parent material. No significant differences were found be-tween stress relieved and as-welded specimens. Sanders et al. �17�and Edwards and Ramulu �18� also showed that the FSW processparameters can be used to control the microstructures and SPFcharacteristics of the weld in 2.5 mm and 5 mm joint thicknesses,respectively. In general, increasing the spindle speed, and thus, thetemperature of the weld, increased the grain size of the weld nug-get and decreased the superplastic performance �19�.

A preliminary report was provided on the surface and subsur-face characteristics of Ti–6Al–4V friction stir weld in thicknessesranging from 3 mm to 12 mm �20�. It was found that the grain sizeincreased with weld thickness and decreased thru the thickness inwelds of a particular thickness. The 3 mm and 6 mm welds had ahigher microhardness than the base metal, while the 9 mm and 12mm welds had a more uniform hardness. Weld microstructuresindicated supertransus processing temperatures, which was vali-dated in a study, where experimental weld nugget temperatureswere evaluated in 6 mm thickness welds made at spindle speedsbetween 200 rpm and 400 rpm and feed rates of 50–150 mm/min�21�. Peak weld temperatures ranged from 1010°C to 1150°C,and were primarily dependent on the tool rotation speeds. Higherrotation speeds lead to higher temperatures along with larger weldnugget grain sizes, lower torques, and lower forge loads. Travelspeed had a minimal effect on the peak temperatures, microstruc-tures, and loads, but did influence the time the welds were ex-posed to the high temperatures. A summary of the work performedon FSW of titanium alloys to date is given in Table 1.

The purpose of the present study is to evaluate the effect of

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eld parameters on producing defect free Ti–6Al–4V friction stireld butt joints in thicknesses ranging from 3 mm to 12 mm. Therimary processing conditions of interest will be the spindle speednd feed rate, while the tool designs and materials will be heldonstant and were chosen based on previous research �21�. Weldsill be produced under a variety of these process conditions and

valuated via metallographic analysis in order to determine ancceptable processing window for each thickness and to accesshe effect of processing conditions on the resulting weld nugget

icrostructure. Furthermore, there is limited literature availablen the mechanical properties of titanium FSWs, and it is almosttrictly limited to tensile property evaluations. Thus, once the pro-essing parameters that are capable of producing high qualityelds in the thicknesses of interest, tensile, and fatigue propertyave been identified, evaluations will be conducted to quantify theechanical performance of titanium friction stir welds in material

hicknesses used commonly in the aerospace industry.

Experimental Procedure

3.1 Material. The material used in this investigation was 3m thick Ti–6Al–4V titanium alloy sheet, and 6 mm, 9 mm, and

2 mm thick Ti–6Al–4V alloy plate. The chemical compositionor this material is 6.28% Al, 0.1% C, 0.21% Fe, 0.010% N,.13% O, 4.05% V, with Ti as the balance. Representative micro-raphs of the base material for the 3 mm, 6 mm, 9 mm, and 12m materials are given in Fig. 1.

3.2 Friction Stir Welding. All materials were machinedlong the abutting edges and chemically etched prior to welding.elds were produced at the Edison Welding Institute with guid-

nce from Boeing Research & Technology. FSWs were made withvariety of process parameters and then evaluated to determine

he optimal conditions for processing. The parameters used areiven in Table 2. For all test welds, two 125 mm�300 mmieces were joined together to make a 250 mm�300 mm part.he FSW welding tool used was a tungsten-lanthanum �W–La�lloy with a small shoulder and a large tapered pin. The exact toolimensions are proprietary, but approximate values are given inable 2. All welds were also made on a W–La backing anvil.

3.3 Metallurgical Characterization. Macro- and microstruc-ural evaluations were performed on each weld nugget �WN�.

elds were sheared perpendicular to the welding direction and

Table 1 Summary

Ti alloyThickness

�mm�Spindlespeed

Feedrate

6Al–4V, Ti–17 12 - -

6Al–4V 6 275 1006Al–4V 6 275 100

15V–3Cr–3Al–3Sn,CP Ti, 6Al–4V

2 200 100

Beta 21S 1.6 200 50–300CP Ti 5.6 1100 500

6Al–4V 3 300–600 60Cast 6Al–4V - 100 10

5Al–1Sn–1Zr–1V 12 140 51

6Al–4V 6.35 - 100

6Al–4V 2 150 100

Std. and FG 6Al–4V 2–3 150–400 50–3006Al–4V 5 200–750 50–1506Al–4V 3–12 140–320 40–1306Al–4V 6 200–400 50–150

ounted for metallographic examination. All samples were pol-

ournal of Engineering Materials and Technology


ished on 5 �m, 1 �m, and finally 0.3 �m alumina-oxide polish-ing wheels. The samples were then etched using Kroll’s to exposethe microstructure and examined under an optical microscope.Micrographs were taken at multiple locations in the weld nuggetto observe the grain structure variation across the face of the weldand through the thickness. Micrographs were taken at variousdepths in the weld nugget along the center line of the joint. Thepositions are identified as the top of the weld nugget �WNT�,center of the weld nugget �WNC�, and bottom of the weld nugget�WNB�. Micrographs were also taken across the midthickness of

Ti FSW literature

Toomaterial Evaluations Ref.

-Microstructure, tensile

properties �2�

CP WMicro, tensile, hardness,

tool temperatures �3,4�–25%Re Micro, hardness �5�–25%Re Micro, tensile, hardness �3�

–25%Re Micro �6�TiC Micro, hardness, tensile �7�

o, pcBN Micro, hardness, tensile �8,9�–25%Re Micro �10�

W based Micro �11�

-Residual stress,crack growth �12�

–25%ReMicro, residual stress,

crack growth �13�

W–LaMicro, SPF, hardness, tensile,

modeling, fatigue �14–16�W–La Micro, SPF, modeling �17–19�

- Micro, hardness �20�W–La Micro, temperatures, modeling �21�

Fig. 1 Representative Ti–6Al–4V base material microstruc-tures for each thickness

Table 2 Range of the FSW process parameters tested

Thickness�mm�

Spindle speed�rpm�

Feed rate�mm/min�

Shoulder diameter�mm�

Pin length�mm�

3 300 50–130 19 2.86 250–320 45–100 24 5.99 250–285 65–100 24 9.0

12 140–190 40–75 32 13.3

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he weld. In addition to the WNC location, micrographs wereaken on the advancing side of the weld nugget at midthicknessWNA� and on the retreating side �WNR�.

3.4 Tensile Testing. A limited number of tensile tests wereonducted on the 3 mm welds in order to assess the quality of theoint under static loading. This was the only thickness tested inhis manner at this time. Specimens were machined from weldedheets. The dimensions of the specimens were in accordance withSTM E8. Each tensile specimen was fully machined on all sur-

aces to remove any surface features, such as the tool marks, thatould adversely affect the test. All tests were conducted underisplacement control at a constant cross head extension speed of.27 mm/min. Every sample was tested to failure. The load dataere monitored by a load cell built into the load frame. The axial

train data were monitored by a clip on an extensometer, whichas removed at the onset of yielding to prevent damage. A total of5 specimens were tested, and in each case, the yield and ultimatetrengths were recorded in addition to the elongation to failure andailure location. All welds were stress relieved at 730 C for 30 minrior to machining specimens and testing.

3.5 Fatigue. Several welds were also made for high cycleatigue life testing. Testing was performed on the 3 mm sheet and

mm plate thicknesses. For each thickness, four 600 mm longoints were made. Specimens were extracted from the plates viabrasive water jet. The fatigue dog-bone specimens were 200 mmn total length. The width of the gauge section was 25 mm, and theidth of the specimen in the grip area was 45 mm. The length of

he gauge section was 50 mm, and a 100 mm radius transitionedrom the gauge section to the grip area. All edges had a radius of.75 mm. In all cases, the specimen was oriented such that theeld was transverse to the loading direction through the center of

he gauge length.Each specimen was fully machined and polished to a 2.5 �m

a finish, or better, along the edges and on all surfaces to removeny surfaces features, such as the tool marks. The weld tool marksere removed because they would cause a large degradation inroperties if left in tact because of the notch sensitivity of titanium16�. It is expected that most fatigue critical industrial applicationsould need to be fully machined to remove the tool marks prior to

ervice even though they may be acceptable for static design ap-lications. Machining was done to ensure parallelism between theaces of the specimen of 0.1 mm or better. The final thickness ofhe 3 mm and 6 mm welds after machining was approximately 2.2

m and 5 mm, respectively. All welds were stress relieved at 730for 30 min prior to machining specimens and testing.

Fig. 2 Macrographs for 3 mm Ti–6Al–4V FSand „c… 125 mm/min; „d… micrograph from th

Testing of the samples was carried out on a 90 kN load frame at

31006-4 / Vol. 132, JULY 2010


a frequency of 20 Hz and a load ratio of R=0.1. A few additionalsamples were also tested at a higher load ratio R=0.6. The maxi-mum stress applied was 830 MPa, and the specimens were cycledto failure. Three samples were tested at each load level. After eachsample set, the maximum stress was reduced by 35 MPa incre-ments, and the next batch of samples was tested. This was re-peated until a stress level that resulted in a run out fatigue life of1.5�106 cycles was achieved. Results could then be reported asstress �S� versus cycles to failure �N� in order to develop S-Ncurves for the welds that could be compared with known hand-book values.

4 ResultsAll butt welds in the 3 mm material thickness were processed at

300 rpm, but the travel speed was varied from 50 mm/min to 130mm/min. Typical weld cross sections are shown in Figs. 2�a�–2�c�for the low, mid, and high feed rate cases, respectively. At 50mm/min, the joint was fully penetrated, but in every trial, thereappeared to be excessive undercut in the joint. This is expected tobe a result of the relatively low travel speed, which allowed thetool to overprocess the material and cast it out of the weld zone inthe form of excessive flashing. At the 125 mm/min condition,there was incomplete penetration and deformation of the tool tip,because the travel speed was too fast and the workpiece was notallowed sufficient time to heat, soften, and flow under the pin. The75 mm/min feed rate produced the most consistently full penetra-tion and defect free welds.

Micrographs from the weld nugget of the 300 rpm at 75 mm/min weld are given in Fig. 2�d�. The grain size in these 3 mmwelds is quite small compared with the base material �Fig. 1�. Ingeneral, the grain size is relatively homogeneous through thethickness and across the width of the weld. Thin sheet thicknessallows uniform heat input and cooling, which relates to the result-ing uniform grain size distribution. All of the microstructures ap-pear to consist of small primary alpha grains with some trans-formed beta. This would imply processing temperatures near thebeta transus temperature. Remnant primary alpha grains in theweld nugget microstructure indicate that the weld was not signifi-cantly over the beta transus temperature, or the transus tempera-ture did not exceed for a long enough time to allow the entiremicrostructure to be changed to transformed beta.

The parameters used for the 6 mm thickness joints ranged from250 rpm to 320 rpm and 45 mm/min to 125 mm/min. This repre-sents a relatively small “process window” evaluation comparedwith the 5 mm thickness case in �19�, but some significant obser-

: 300 rpm at „a… 50 mm/min, „b… 75 mm/min,5 mm/min weld

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pm, regardless of the travel speed used, i.e., 65 mm/min, 100m/min, or 125 mm/min, there was lack of penetration. This im-

lies that the spindle speed was not high enough to sufficientlyeat and stir the material, regardless of the feed rate used. Whenhe spindle speed was increased to 280 rpm and the travel speedas reduced to 45 mm/min, the joint became full penetration, but

he root was overstirred, and an extremely fine grained micro-tructure and defects were developed at the root �Fig. 3�a��. Thisverstirred root issue was consistent for the 280 rpm condition at65 mm/min feed rate as well. It was not until the travel speedas increased to 100 mm/min and 125 mm/min that the root de-

ects were eliminated and full penetration was maintained �Fig.�b��. The 300 rpm welds possessed root defects similar to thosen Fig. 3�a� at the 65 mm/min condition, but these were eliminatedy going to a higher travel speed of 125 mm/min. The 320 rpmondition possessed the root defects for all feed rates attempted,etween 65 mm/min and 125 mm/min.

Micrographs from the weld nugget of the 280 rpm at 100 mm/in weld are given in Fig. 3�c�. The grain sizes in this 6 mm weld

re significantly larger than those in the 3 mm case and even thease material. The grains appear to be fairly similar in size andtructure cross the width of the weld, from the advancing side,hrough the center to the retreating side. However, there appears to

Fig. 3 Macrographs for 6 mm Ti–6Al–4V FSat 100 mm/min; „c… micrographs from 280 rp

Fig. 4 Macrographs for 9 mm Ti–6Al–4V FS
at 65 mm/min; „c… micrographs from the 270 rp


be a grain size gradient through the thickness of the weld. Thegrains are slightly larger at the top of the weld compared withthose in the center, and there is a sharp grain size drop at the rootof the weld because of the thermal gradients during the process. Itis be expected that the heat input is greater at the top of the weldwith the large tool shoulder, and the heat input decreases throughthe thickness as the pin tapers down in size. The low thermalconductivity of the titanium could also contribute to this thermalgradient. With the exception of the very bottom of the weld, allmicrographs show a transformed beta microstructure implying su-pertransus processing temperatures. It is possible that the root ofthe weld was held to subtransus temperatures.

The 9 mm weld thickness spindle speeds used included 250rpm, 270 rpm, and 285 rpm, and the travel speeds used were 65mm/min and 100 mm/min. The 250 rpm spindle speed conditionwas not sufficient to produce a full penetration joint �Fig. 4�a��, sothe speed was increased to 280 rpm. A spindle speed of 280 rpmat 100 mm/min was also not sufficient to produce full penetration,but lowering the feed rate to 65 mm/min allowed the root of thejoint to be stirred, enough to achieve full penetration �Fig. 4�b��.The spindle speed-feed rate combination of 285 rpm and 100 mm/min also produced a full penetration weld with no defects.

Micrographs from the weld nugget of the 270 rpm at 65 mm/

: „a… 280 rpm at 45 mm/min and „b… 280 rpmat 100 mm/min weld

„a… 250 rpm at 100 mm/min and „b… 270 rpm

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in weld are given in Fig. 4�c�. Again, there appears to be a grainize increase with respect to the thinner welds and the parentaterial, and as with the 6 mm case, there is a grain size gradient

hrough the thickness of the weld but not across the width. Also,imilar to the 6 mm weld case, in all positions within the weldugget, there is an acicular alpha, or transformed beta, microstruc-ure indicating processing temperatures above the beta transusemperature. The grains are larger than the 6 mm case, indicatinghat the time at a high temperature in the 9 mm case is longer,llowing for further grain growth after being stirred. This obser-ation can be attributed to the thicker material, which causeslower cool down rates, in addition to the lower feed rate, whichxposes the weld material to longer times at the elevated tempera-ures during welding. Both of these characteristics would promotehe growth of larger grains.

The spindle speeds evaluated for the 12 mm thick case wereetween 140 rpm and 190 rpm. These low spindle speeds weresed because the larger tool size used produces more heat at aiven rotation speed than the smaller tools used for the thinnerauges. These thicker sections also retain heat more than the thin-er gages, thus requiring less heat input via lower spindle speeds.his was the most challenging thickness to weld. Breaking the

ool due to high torques was common. It was also difficult tochieve a weld with a defect free root. The 140 rpm at 40 mm/minondition contained a significant amount of irregular flow patternefects �Fig. 5�a��. Increasing the spindle speed to 170 rpm at 40m/min improved the overall joint quality, but root defects were

bserved �Fig. 5�b��. Increasing the feed rate to 65 mm/min pro-ided a full penetration and defect free weld �Fig. 5�c��. A 75m/min feed rate at 170 rpm led to the lack of penetration, and

ncreasing the spindle speed to 190 rpm at 65 mm/min also led tooot defects, similar to those in Fig. 5�b�.

Micrographs from the weld nugget of the 170 rpm at 40 mm/in weld are given in Fig. 5�c�. The trends observed in the 3 mm,mm, and 9 mm cases continue in the 12 mm welds. The grain

ize is relatively constant across the width of the weld nugget butaries through the thickness. There are large grains in the top ofhe weld compared with the other weld thicknesses, but the sizeecreases through the thickness to the point where relatively finerains are observed in the root. Again, these observations are re-ated to the slower feed rates, causing longer temperature expo-ure, and the thicker material, leading to slower cool down rates,

Fig. 5 Macrographs for 12 mm Ti–6Al–4V F40 mm/min, and „c… 170 rpm at 65 mm/min; „weld

oth of which enable larger grain growth. For every micrograph

31006-6 / Vol. 132, JULY 2010


location in the 12 mm weld, the structure is made up of acicularalpha, indicating supertransus processing temperatures.

4.1 Tensile Testing. Tensile tests were performed on 3 mmthickness welds made under the weld parameters, which wereidentified as optimal for the thickness in Sec. 3, which was aspindle speed of 300 rpm at a feed rate of 75 mm/min. YS, UT,%e to failure, and failure locations for the 15 tensile specimensextracted from a 3 mm butt weld are given in Table 3. It wasfound that the average yield and ultimate strengths of the weldswas 974 MPa and 1025 MPa, respectively. The average elongationto failure was found to be nearly 10%. A majority of the failuresoccurred in the BM, with a few occurring in the WN or near theHAZs �Fig. 6�. Since the failure location was relatively random, itappears that there is no one week region in the weld. Since veryfew failures occurred in the weld nugget, the strength of the welditself is probably higher than the base material. This is related tothe refined weld nugget grain size.

Scanning electron microscopy �SEM� images of fracture sur-faces where failure occurred in the weld zone and in the base

s: „a… 140 rpm at 40 mm/min, „b… 170 rpm aticrographs from the 170 rpm at 65 mm/min

Table 3 Tensile test results of 3 mm Ti FSW produced at 300rpm and 75 mm/min

Sample No.UTS

�MPa�YS

�MPa� %eFailurelocation

1 1028.0 983.9 10.0 BM2 1035.6 978.4 9.9 HAZ3 1009.4 960.4 10.0 HAZ4 1013.5 961.8 11.1 BM5 1026.6 976.3 10.7 WN6 1033.5 984.6 8.4 HAZ7 1034.2 980.4 10.7 BM8 1015.6 950.8 10.7 HAZ9 1037.0 986.6 10.8 WN10 1028.0 970.1 9.8 BM11 1025.9 988.0 9.8 BM12 1039.7 993.5 8.4 WN13 1028.0 966.6 9.3 BM14 1008.7 966.0 8.1 BM15 1010.8 956.3 8.1 BMAverage 1025.0 973.6 9.7STDEV 10.7 12.9 1.0

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etal are shown in Fig. 7. Analysis of the failure surfaces showedo significant differences in behavior between the welds thatailed in the nugget and base metal. Both samples failed by mi-rovoid coalescence, which is typical of ductile materials.

Compared with the tensile properties of standard, mill annealedi–6Al–4V �22�, the weld properties are quite comparable. Theeld ultimate strength of 1025 MPa is nearly 8% better than thatf unwelded material at 880 MPa. The yield strength of 974 MPas approximately 11% better than that of the unwelded materialhat possesses a yield strength of 950 MPa. The elongation of thepecimens at failure was 9%. This is 30% lower than the parentaterial, which exhibits a 14% elongation to failure. However, the

ig. 6 Typical failed FSW tensile specimens noting failureocation

ig. 7 SEM images of tensile specimen failure surfaces: „a…eld nugget and „b… Base metal

Fig. 8 Macrographs of „a… 3mm and „b… 6 m„280 RPM at 100 mm/min… and 3 mm „300 R
two load ratios and compared with the handbo


elongation of these specimens is significantly higher than previ-ously reported values, where elongations were on the order of50% compared with the parent material �15�. The standard devia-tion in the data of these test results is also quite low comparedwith the previous studies. This is attributed to fully machining thespecimens prior to testing in this study where welds were left inthe as-welded condition in the previous works. The relativelyrough surface finish of the welds could impact the tensile proper-ties and would absolutely affect fatigue performance. The lowerstandard deviations seen in this work could also be attributed tothe more stable and robust welding process that has been devel-oped since preliminary welds were produced in Ref. �15�.

4.2 Fatigue Performance. Fatigue testing was also per-formed on 3 mm and 6 mm thickness welds made with parametersthat were identified as optimal in Sec. 3. The 3 mm welds wereproduced at a spindle speed of 300 rpm and a feed rate of 75mm/min, while the 6 mm welds were made at a spindle speed of280 rpm and a feed rate of 100 mm/min. The results from thestress life test are given in Fig. 8. The data from the weld fatiguetests are plotted against the S-N curves derived from the MILHandbook �23� data for comparison. The handbook data used arefor fatigue performance of standard Ti–6Al–4V base metal at thesame stress ratio R=0.1, with a stress concentration factor Kt of1.0 and 3.0. The results of the weld tests are reasonably close tothe stress concentration free values given by the handbook. Thehigher stress ratio data are lower, as expected, because the meanapplied stress during testing is higher. These data show a signifi-cant improvement compared with fusion welded joints �24�. Mac-rographs of typical weld specimens that had been fully machinedand tested for both thicknesses are also given in Fig. 8.

A photograph of a typical fatigue specimen with a crack devel-oping in the gauge area during testing is given in Fig. 9�a�, alongwith typical failure surface photographs for both 3 mm and 6 mmspecimens in Figs. 9�b� and 9�c�, respectively. For the 3 mmspecimen, the crack initiation point is at the edge of the specimen.It does not appear to originate from any type of weld defect ormicrostructural flaw. In the 6 mm specimen, the crack initiationpoint is at the surface of the specimen.

fatigue specimens; „c… S-N curve for 6 mmat 75 mm/min… welded specimens tested at

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SEM images of representative fracture surfaces are given inig. 10. As noted in Fig. 9, the SEM evaluations of the 3 mmpecimen showed the crack initiated from the corner of the speci-en. Higher magnification views of the crack propagation region

hows small striations, which is a characteristic of stable fatiguerack growth. Evaluation of the 6 mm specimen shows that therack did initiate form the surface of the weld, but no defects ornclusions were identified. Higher magnification examinations ofhe crack propagation region also showed striations indicative oftable crack propagation prior to overload failure. The overloadailure fracture surfaces were similar to those observed in theensile specimen failure examinations �Fig. 7�, showing microvoidoalescence, which is typical of ductile overload failure.

DiscussionIn this study a variety of Ti–6Al–4V thicknesses were friction

tir welded using a range of process parameters. In general, forigh spindle speeds and low feed rate combinations, void typeefects can form in the cross section of the weld. At low spindlepeeds and relatively high feed rates, the heat generated by theool during the process is insufficient to produce a joint. However,hen the correct combination of spindle speed, feed rate, and toolesign are selected, a defect free joint can be produced in virtuallyny thickness of Ti–6Al–4V.

In general, the range of parameters, spindle speed and feed ratepecifically, which can be used to produce a defect free joint for aiven thickness and tooling configuration, is relatively small, onhe order of 25 rpm and 25 mm/min. One of the hardest issues tovercome is the lack of penetration defect as the interaction be-ween the pin tool tip and the backing substrate when attempting

ig. 9 „a… Friction stir welded titanium fatigue specimen with aeveloping crack and typical failure surfaces for „b… 3 mm andc… 6 mm specimens

ig. 10 „a… Crack initiation point and „b… crack propagationurface for the 3 mm specimen; „c… crack initiation point and
d… crack propagation surface for the 6 mm specimen
31006-8 / Vol. 132, JULY 2010


to achieve full penetration weld becomes complex. The differencebetween a weld with full penetration and no tool wear, and a weldwith a defect, is very small with respect to the parameters beingused. A slightly higher spindle speed can achieve full penetrationby increased heating and material flow, but too slow a travel speedcan lead to overstirring of the root, cavitations, and voidlike de-fects. Conversely, decreasing the spindle speed or increasing thetravel speed only slightly can cause insufficient heating of theweld root, resulting in inadequate stirring and ultimately, a lack ofpenetration defect.

Out of all the thicknesses welded, most of the optimal condi-tions were in the 250–280 rpm at 65–100 mm/min parameterrange, with the exception of the 12 mm thick welds. The optimalprocessing conditions for each joint thickness are given in Table 4.All of the conditions tested for each thickness are plotted in Fig.11, and the optimal conditions found for each material thicknessare circled. The process window for Ti–6Al–4V is highlighted inthe box on this plot to illustrate the rough combinations of spindlespeed and feed rate, which will lead to the best chances of pro-ducing an acceptable joint for a given tooling design.

It is important to note that the process window identified forthese thicknesses are dependent on the tooling material, thermalmanagement, tool designs, weld machine, fixturing equipment,and so on. In these experiments, a tungsten based backing anvilwas used for all trials. It was found that achieving full penetrationwelds with this anvil material was difficult. It is thought that thiscould be due to the relatively high thermal conductivity of tung-sten, which rapidly extracts heat, keeps the root of the joint cold,prevents material flow, and limits penetration. By using otherlower conductivity materials, such as ceramics and steels, moreheat is retained in the root of the weld, greatly enhancing penetra-tion �Fig. 12� and widening the process window. However, thereare limitations to using these alternate materials, such as alloyingbetween the hot titanium and steel, or cracking of ceramics. Thisis avoided with tungsten based anvils because they do not reactwith titanium at elevated temperature, but the high thermal con-ductivity is a draw back for penetration. These relationshipsshould be considered when choosing the tooling materials andsubsequent parameter identification.

Table 4 Optimal FSW processing parameters for a given tooldesign and thickness

Joint thickness�mm�

Spindle speed�rpm�

Feed rate�mm/min�

3 300 756 280 1009 270 65

12 170 65

Fig. 11 Schematic for Ti FSW process window



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In a previous study �18�, an analytical model was presented forredicting energy inputs during the Ti FSW process as a functionf the tool geometry, weld parameters, and workpiece materialroperties �Eq. �1��. For 5 mm thick Ti–6Al–4V FSW butt joints,he energy inputs were evaluated for a variety of welding condi-ions. It was found that the optimal welding condition was 300pm at 100 mm/min, which corresponded to a 2.1 kW energynput. Using this model to predict the energy inputs for the weldsroduced in this study can help validate the utility of such model.he optimal welding conditions as a function of the material

hickness are given in Table 4. The energy input values �Q� cor-esponding to these welding conditions are given in Table 5 andre plotted in Fig. 13. The contact condition and workpiece mate-ial properties were assumed to be the same as those given in Ref.18�. Since the energy input is a function of the tool geometry, thenergy input of the thicker welds will be significantly higher be-ause the length of the weld tool is longer. To normalize this, thenergy values �Q� can be divided by the thickness of the materialt� to obtain an energy input per unit thickness value �Q / t�. Thesere also given in Table 5

Q = ��C1 + �C2� �1�

here � is the shear yield strength as a function of the temperaturend strain rate, � is the contact condition coefficient between theool and workpiece, � is the rotation speed of the tool, � is theool feed rate, and C1,2 are the tool geometry constants.

The energy input equations predict that the energy input willncrease with thickness, as expected. Thicker materials requirearger tools and will, thus, have greater energy inputs for relatively

ig. 12 6 mm weld produ ced on a steel based anvil showingmproved penetration profile

able 5 Energy input values as a function of the materialhickness

Thickness�mm�

C1�mm3�

C2�mm2�

Q�kW�

Q / t�kW/mm�

5a 1283 426 2.1 0.43 483 226 0.8 0.36 1817 542 2.8 0.59 3435 770 5.1 0.612 10309 1599 9.6 0.8

Data from Ref. �18�.

ig. 13 Predicted energy input as a function of the material
hickness


similar welding parameters. However, the energy per unit thick-ness predictions also shows an increase with thickness. This isbecause the thicker joint weld tools have a larger diameter overthe length of the pin, and these larger tool diameters cause thegreater energy inputs from the tool rotation per unit thickness. Thepredicted higher energy inputs per unit thickness for the thickerjoints could explain the increased grain sizes observed metallur-gically between the thin and thick gauge welds.

With respect to microstructural examinations, several interest-ing observations can be made. In comparing the microstructuresfrom each weld thickness, it is clear that the grain size increaseswith weld thickness �Fig. 14�. As the thickness of the weld in-creases, the heat input also increases, and the cool down of theweld nugget after stirring becomes slower, leading to largergrains. Also, for thicker welds, slower feed rates are used. Slowerfeed rates will increase the exposure time of the weld to the pro-cess temperatures and also allow for grain growth.

There is also a grain size gradient through the thickness of thewelds in each case. The grains at the top of the weld are consis-tently larger than those in the bottom because of the tapered toolgeometry and the low thermal conductivity of the workpiece. Atthe top of the weld, the heat input is large because of the shoulderand relatively large pin diameter. In the weld, as the pin tapersdown, the heat input is lower. The root is also exposed to less ofthe heat generated by the shoulder of the tool due to the lowthermal conductivity of titanium. The grain size gradient is moreprominent in the thicker welds. In the thin section weld, such asthe 3 mm case, the pin length to shoulder diameter ratio is smaller,and the thermal gradient thru the thickness of the joint is smaller.There appeared to be no significant grain size gradient across thewidth of the welds, from the advancing to the retreating side. Thisimplies a fairly uniform temperature gradient across the face ofthe welds.

For the 3 mm case, the grain structure in the weld showedevidence of primary alpha and transformed beta. This implies thatthe processing temperature was near the beta transus, but not com-pletely over the point where the microstructure transforms to afully beta structure. However, in all other cases, the structure con-sisted of acicular alpha, which indicates supertransus tempera-tures. As the thickness increased, the size of the acicular alphagrains increased, but the morphology was similar.

The tensile properties of the welds produced in this study weresuperior to the parent material with respect to the strength. Elon-gation properties were consistent with previous evaluations given

Fig. 14 Weld nugget grain size versus thickness comparison

in Refs. �2,3,15�. The results of the high cycle fatigue and tensile

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valuations were quite exceptional. The fatigue tests of FSWs in 3m and 6 mm thick Ti–6Al–4V are virtually equivalent to the

ublished data for the parent material. This is a significant im-rovement over the fatigue performance of fusion welding pro-esses, such as laser beam welding �24�.

ConclusionsFriction stir welding of titanium has been demonstrated on

–12 mm thickness welds. The combination of the process param-ters, mainly the spindle speed and feed rate, must be correct inrder to produce a full penetration, defect free joint. Evaluation ofll the welding trials conducted has lead to a rough process win-ow for friction stir welding Ti–6Al–4V for a given tool material.

The microstructure of the welds varies with joint thickness.hicker welds possess larger grains from increased heat input,

ower tool feed rates, and slower cool down times. The micro-tructure also varies thru the depth of a particular weld nugget.rains are larger near the top of the weld, and the size decreases

hru the thickness. This thru thickness variation is related to thehermal gradient caused by the heat input from a tapered pin toolnd the low thermal conductivity of titanium.

For thin welds less than 3 mm, it is likely that the process isubtransus since primary alpha and only a small amount of trans-ormed beta were observed in the weld nugget microstructure. Forhicker welds, the microstructure consists of coarser acicular alpharains, implying supertransus process temperatures.

The high cycle fatigue performance of titanium friction stirelds is comparable to the parent material properties. Tensileroperties of titanium FSWs possess higher ultimate and yieldtrengths compared with the parent material, but elongations areeduced. However, the elongations measured in the welds pro-uced during this study are equivalent to those reported in therevious researches.

Analytical energy input models can be used to predict the en-rgy inputs required to produce Ti–6-4 FSW butt joints in variedhickness of material. Energy inputs were predicted to be withinpproximately 1–10 kW. Energy inputs per unit joint thicknessere found to be between 0.4 kW/mm and 0.8 kW/mm. Thicker

oints require higher energy inputs and higher energy inputs pernit thickness.

cknowledgmentThe authors of this paper would like to thank The Boeing Co.

or the support and Dr. D. Sanders for his encouragement through-ut this research project. Our sincere thanks are also extended to. Bernath of the Edison Welding Institute for conducting theelding of the specimens and assistance with the process devel-pment.

eferences�1� Mishra, R. S. and Mahoney, M. W., eds., 2007, Friction Stir Welding and

Processing, ASM International, Materials Park, OH.

31006-10 / Vol. 132, JULY 2010


�2� Trapp, T., Helder, E., and Subramanian, P. R., 2003, “FSW of Titanium Alloysfor Aircraft Engine Components,” Friction Stir Welding and Processing II:Proceedings of the TMS Annual Meeting, pp. 173–178.

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�11� Fonda, R. W., Knipling, K. E., Feng, C. R., and Moon, D. W., 2007, FrictionStir Welding and Processing IV, TMS, Warrendale, PA, pp. 295–301.

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�13� Pasta, S., and Reyonds, A. P., 2008, “Residual Stress Effects on Fatigue CrackGrowth in Ti-6Al-4V Friction Stir Weld,” Fatigue Fract. Eng. Mater. Struct.,31, pp. 569–580.

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�15� Sanders, D., Ramulu, M., Klock-McCook, E., Edwards, P., Reynolds, A. P.,and Trapp, T., 2008, “Characterization of Superplastically Formed Friction StirWeld in Titanium 6Al-4V: Preliminary Results,” J. Mater. Eng. Perform.,17�2�, pp. 187–192.

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�17� Sanders, D., Edwards, P., Ramulu, M., and Grant, G., 2010, “Optimization ofthe Friction Stir Welding Process for Superplastic Forming and Improved Sur-face Texture for Titanium Aerospace Structures,” Key Eng. Mater., 433, pp.153–167.

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identiﬁcation of process

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