Analytical Bonding Criteria for Joint Integrity Prediction in Friction Stirwelding of Aluminum...

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ARTICLE IN PRESSG ModelROTEC-13902; No. of Pages 10

Journal of Materials Processing Technology xxx (2014) xxx–xxx

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Journal of Materials Processing Technology

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nalytical bonding criteria for joint integrity prediction in friction stirelding of aluminum alloys

ianluca Buffa ∗, Sergio Pellegrino, Livan Fratiniepartment of Chemical, Management, Computer Science and Mechanical Engineering, University of Palermo, Viale delle Scienze, 90128 Palermo, Italy

r t i c l e i n f o

rticle history:eceived 22 July 2013eceived in revised form3 December 2013ccepted 7 February 2014

a b s t r a c t

In this study, two bonding criteria, previously used for porthole die extrusion, are applied to FSW startingfrom the local value of the main field variables calculated through a specifically developed 3D numericalmodel of the process. Their applicability and effectiveness have been assessed through an experimen-tal and numerical campaign carried out with the main process parameters varying in a wide range.The pressure–time–flow criterion was demonstrated to be better suited for FSW processes when large
vailable online xxx
eywords:riction stir weldingluminum alloysEM

welding speed is used.© 2014 Elsevier B.V. All rights reserved.

onding criterion

. Introduction

Friction Stir Welding (FSW) is a solid-state welding process par-icularly suited for joining both light alloys and highly resistant

aterials that are difficult to be welded by conventional fusionelding. In FSW a non-consumable rotating tool with a speciallyesigned pin and shoulder is inserted into the abutting edges ofhe sheets to be joined which are rigidly positioned one against thether. It is possible to incline the tool respect to the normal workingxis by a 2–5◦ angle in the opposite direction of the welding one (tiltngle). The base of the pin must be plunged nearly to the bottomf the joint so that the pin is in contact with all the welding pieceslong the total thickness. Once the plunge phase is finished, theimple translation of the rotating tool along the welding line deter-ines the formation of the joint. The weld seam will present finely

triped bands, and it will be slightly lowered and not so irregularnd protruding as the conventional welding seams.

The relative motion between the tool and the workpiece causes,ue to the friction, the generation of the needed heat: maximumemperature is found between the pin and the workpiece but the

Please cite this article in press as: Buffa, G., et al., Analytical bondingaluminum alloys. J. Mater. Process. Tech. (2014), http://dx.doi.org/10.

elting temperature of the alloy is not reached. The generated heatnd the high temperature cause a decrease of the yield stress of theetal that can easily flow around the tool. Being temperature below

∗ Corresponding author. Tel.: +39 091 23861869; fax: +39 091 7099973.E-mail addresses: [email protected] (G. Buffa), [email protected]

S. Pellegrino), [email protected] (L. Fratini).

ttp://dx.doi.org/10.1016/j.jmatprotec.2014.02.014924-0136/© 2014 Elsevier B.V. All rights reserved.

the melting one, there is no change in volume during the processand this entails low values of residual stresses after cooling. Mishraand Ma (2005) reported the mechanisms responsible for the forma-tion of seam welds in FSW processes and also the microstructuralrefinement and final mechanical properties of the welded joint. Theauthors highlighted also that the maximum temperature reached,in correspondence of the central part of the welding seam, is a func-tion of the ratio between the tool rotation speed – R [rpm] – andtool feed rate – v [mm/min] – along the weld seam: in general, withthe increase of this ratio, maximum temperature increases.

Based on the description of the process, it is immediately notedthat the plastic flow of the material generated by the action ofthe tool gives origin to asymmetric joints. Guerra et al. (2002)focused their attention on a few key aspects of the process as micro-structural issues and material flow analysis. They observed, in across section of the joint, an “advancing side” and a “retreatingside”: in the first case the vectors peripheral speed of rotation andfeed rate of the tool have the same direction; on the contrary, inthe retreating side the two vectors are opposite each other.

The mechanism of formation of the weld through FSW can beassimilated to an extrusion and forging of the plates to be welded.Buffa et al. (2006a,c) investigated the operating parameters of theFSW process, as optimal tool geometry and welding speed, forimproving nugget integrity of aluminum alloys as well as the bond-

criteria for joint integrity prediction in friction stir welding of1016/j.jmatprotec.2014.02.014

ing mechanics. Due to the rotation, the pin moves the material incontact with it towards the area by creeping, namely the metal isextruded around the pin and immediately forged by the shoulder.The longitudinal axis of the tool will then be the forging axis, while

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ARTICLEROTEC-13902; No. of Pages 10

G. Buffa et al. / Journal of Materials P

he welding direction will be the extrusion axis. The mixed mate-ial is usually dragged many times around the tool before beingeposited. Reynolds et al. (2000) and Heinz et al. (2000) investi-ated the mechanical properties and the microstructure of a FSWedoint showing the characteristics of the welded joints. They demon-trated that the material flows also vertically, in an almost circularattern as seen in a longitudinal section.

The most common welding defects that occur in a FS weldedoint are interruptions in the physical structure of the junction dueo incorrect material flow (tunnel defects). Additionally, at the bot-om of the joint swirl phenomena may be observed, thus resultingn an ineffective material flow and the possible insurgence of inter-al folding defects due to the geometrical discontinuity inducedy the tool pin shape. Chen et al. (2006) pointed out the condi-ions that cause the typical defects of the FSW and the influencen the mechanical properties of the joint. They found that the loweat input leads to the generation of tunnel and/or “kissing bond”.he latter defect is particularly difficult to be spotted with non-isruptive tests. Finally, besides tool rotation and welding speed,he selection of the correct tilt angle influences the heat input.

The material flow induced by the tool, which determines theffectiveness of the weld, is dramatically influenced by materialroperties, such as yield strength, tool design, and FSW processarameters. Once the material to be welded is fixed and the tool

s designed, process parameters as tool feed rate and rotation areesponsible for the distribution of the main field variables that gov-rn the solid bonding phenomenon. Proper values of temperature,train and strain rate are needed in order to get an effective solidonding. Balasubramanian (2008) studied how the FSW processarameters influence the weld quality. It is worth noticing that, athe moment, only empirical relationships can be used to predicthe FSW process parameters to produce defect free joints. A weld-ng criterion, depending on the local values of the above cited fieldariables and embedded in a proper process model, represents anmportant tool for sound joints design. As far as the authors know,o solid bonding criterion has been used for FSW.

A few researches can be found in literature characterizing theolid bonding phenomena for processes as Roll Bonding (RB) ororthole Die Extrusion (PDE), also including the use of a properonding criterion. Donati and Tomesani (2005) studied the rela-ion between product quality and die design in extruded aluminumrofiles with seam welds. The effectiveness of the obtained jointtrongly depends on several operating parameters, both geomet-ical and technological. Using different combinations of processarameters different weld quality is obtained for the joints (defec-ive or sound). Donati et al. (2007) also analyzed the processarameters and the different die geometries for the extrusion of

hollow profile with a seam weld. Ceretti et al. (2009a,b) showedow it is possible to correctly simulate the bonding phenomenaccurring in extrusion porthole dies. They investigated the solidonding phenomenon through modular matrices in which it wasossible to change the geometric parameters of the referenceatrix, in order to vary the welding conditions. In this way, the vari-

bles having a major influence on the process could be identified.he analysis was conducted from a purely phenomenological pointf view, identifying one by one of the critical parameters for thebtainment of a sound bond, in order to derive general design rulesor the process. Donati and Tomesani (2004) and Donati (2004) haveeveloped numerical campaigns of extrusion processes in ordero obtain the local conditions of the welds experimentally pro-uced deriving, as a result, a bonding criterion valid under differentelding conditions. In fact, once the welding surface is identified,


t is possible to calculate, through a numerical analysis, the fieldariables involved in the bonding process, as contact pressure, tem-erature, local flow stress, strain, nodal speed. In the following thehree most utilized bonding criteria are briefly described.

PRESSsing Technology xxx (2014) xxx–xxx

1.1. Maximum pressure criterion

This criterion considers as a discriminating parameter only themaximum pressure inside the welding chamber. Ackeret (1972)stated that when this value exceeds a critical threshold, whichdepends only on the local and instantaneous conditions of thematerial, the weld can be considered sound. This criterion is by farthe most applied in practice because of its simplicity. The effective-ness of this criterion for RB processes has been demonstrated byAzushima et al. (2008) which produced lightweight parts by usinghigh strength metal for the safety and reliability of micro-parts. Itis worth noticing that, due to its extreme simplicity, this criterion isnot particularly suited for manufacturing processes characterizedby more complex material flow, as PDE and FSW. Consequently,this criterion has not been considered in the present study.

1.2. Pressure–time criterion

This criterion, proposed by Plata and Piwnik (2000), is based onthe integral in time of the ratio between the contact pressure andthe flow stress of the material. The value obtained must exceed acritical threshold. Jo et al. (2003), in their paper, studied the influ-ence of bearing length, product thickness and billet temperature inporthole die extrusion of hollow section tubes on the pressure at theinterface of the welding plane. The best mechanical properties areobtained when the pressure at the welding plane is approximately3.5–5.8 times the average flow stress. Donati et al. (2007) investi-gated the effectiveness of the criterion in the PDE process throughboth mechanical and metallurgical surveys in order to validatethe numerical approach. The obtained results demonstrated thatit is possible to correctly simulate the solid bonding phenomenaoccurring during the porthole die extrusion. Ceretti et al. (2009a)determined the critical value of the criterion through roll bonding(RB) experiments on AA6061. The threshold curve obtained wasused to model the extrusion of complex hollow profiles. Finally,D’Urso et al. (2012), through a coupled experimental-simulativestrategy on aluminum alloy AA6082, implemented a new proce-dure for the identification of the pressure–time bonding criterionas a function of the temperature. Flat rolling experimental tests andFEM simulations of the rolling process for the same conditions werecarried out and an exponential limit curve was identified.

Although the above cited authors found satisfying results,Donati and Tomesani (2004) pointed that this approach is some-times unable to give correct predictions when large gradients ofthe node velocity are observed.

1.3. Pressure–time–flow criterion

As previously observed, Donati and Tomesani (2004) showedthat the pressure–time criterion over emphasizes the role deadzones of the material in a PDE matrix, in which the residence timestend to infinity. Consequently, they introduced the speed correc-tion factor, pointing out that the material flow passing througha generic point should be considered. The criterion validity wasdemonstrated starting from experimental PDE tests. In particular,special H profiles were produced by changing the geometry of theleg and the width of the central section, so as to create a variety ofoperating conditions in order to obtain joints with different weldingquality. The obtained results showed that it is possible to pre-dict the solid bonding phenomena occurring during PDE using thepressure–time–flow approach. The effectiveness of this criterionhas not been tested on industrial case studies yet.


In this paper, two solid bonding criteria are selected, prop-erly applied to FSW and finally compared. Experimental weldingtests have been performed with varying tool feed rate and rota-tion in order to obtain both defective and sound welds. Then, the


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Table 1Experimental case studies.

Test ID R [rpm] V [mm/min]

A 500 400B 500 200C 500 100

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D 1000 400E 1000 200F 1000 100

xperimental tests have been simulated through a pre-existing FEModel for FSW. The objective of this work is to assess the reli-

bility and the effectiveness of bonding criteria, already presentn literature and developed for other processes, for FSW throughxperimental test, microstructural investigation and FEM analysis.dditionally, the two criteria are compared and the most suited forSW process is identified.

. Methods

.1. Experiments

Butt joints were obtained out of aluminum alloy AA6061-T600 mm × 200 mm sheets, 2.4 mm in thickness. The base materialas characterized by a yield stress of 260 MPa and an ultimate ten-

ile stress (UTS) of 305 MPa. A Mazak Nexus 410A milling machineroperly equipped with clamping fixture for FSW was used. Thetilized tool was made in H13 steel quenched at 1020 ◦C, charac-erized by a 52 HRc hardness. A conical pin was designed, withonical angle of 30◦, major diameters equal to 3.5 mm and heightqual to 2 mm. The shoulder diameter was equal to 10 mm.

A wide range of variation was selected for both the rotationalpeed, varying form 500 rpm to 1000 rpm, and the feed rate, varyingorm 100 mm/min to 400 mm/min. A constant tilt angle of 2◦ wasdopted for all the tests. In Table 1 the analyzed case studies areeported:

Each test was repeated three times and from each jointpecimens were cross-sectioned perpendicularly to the weldingirection. Macrographs were used to analyze the material area

nvolved in the process mechanics; furthermore, the presence ofow defects was investigated through micrographs. In order tobtain such results the specimens were hot mounted, polished andnally etched with Keller reagent and observed by a LM.

.2. Numerical simulations

The commercial FEA software DEFORM-3DTM, Lagrangianmplicit code designed for metal forming processes, was used. Buffat al. (2006b) proposed a continuum based FEM model for frictiontir welding process. This model was first calibrated by comparingalculated force and temperature distribution with experimentalesults. Then it was used to investigate the distribution of the maineld variables in the heat affected zone and the weld nugget. Aoupled thermo-mechanical analysis with rigid-viscoplastic strain,train rate and temperature dependent material behavior was per-ormed. The material data were taken from the ASM Handbook1990) and in house experiments. The tool was modeled as a rigidbject. The FSW modeling was divided into two stages: the plungetage and the welding stage. In the first phase the tool, which has

tilt angle of 2◦, moves down vertically with the assigned rotatingpeed. In this way the temperature field needed for the beginningf the actual welding is reached. In the next step the rotating tool


oves along the welding line with proper feed rate.The tool was meshed, for the thermal analysis, with about 3000

etrahedral elements. The two sheets to be welded were modeleds a single-block. This is because in a double block model instability

Fig. 1. The figure shows the utilized mesh for tool and sheet.

problems may arise due to the sheet-sheet contact. However, Buffaet al. (2006b,c) have demonstrated that this approximation doesnot significantly affect the distribution of the main field variables.In other words, Friction Stir Processing (FSP) is simulated insteadof FSW.

For the plate material, i.e. the aluminum alloy AA6061, the fol-lowing values of thermal conductivity k = 180 [N/(s ◦C)] and thermalcapacity c = 2.4 [N/(mm2 ◦C)], taken from literature, were used. Novariation of k and c with temperature was taken into account.This assumption makes the thermal problem linear, speeding upthe numerical solution at each time increment. A constant inter-face heat exchange coefficient of 11 [N/(mm s ◦C)] was utilized forthe tool sheet contact surface. The sheet blank, 2.4 mm thickness,was meshed with about 12,000 tetrahedral elements. Because ofthe large gradients of the calculated variables, it was necessary tointroduce a finer discretization along the welding line. Addition-ally, a re-meshing referring volume was identified all along the toolfeed movement. In this area, each tetrahedral element had mini-mum single edge of about 0.5 mm; in this way, about five elementswere placed along the sheet thickness. An interface penalty con-stant equal to 1E9 was used to penalize the penetration velocityof the nodes of the sheets through the tool master surface. In thisway, elements folding, especially in the contact area between thetool pin and the workpiece, where the material “closes” behind thetool, was avoided reducing the risk of unexpected simulation stopdue to failure in convergence. A constant shear friction factor of0.46 was used for the tool-sheet interface on the basis of a previousexperimental thermal characterization and of a numerical sensi-tivity analysis for the shear friction factor m conducted by Fratiniet al. (2005). In particular, the shear factor was optimized basedon experimental temperature measurements performed with boththermo-camera and thermocouples. Fig. 1 shows the meshed singleblock and tool.

At the end of the simulation, the material flow was investigatedthrough the analysis of the nodes movement and the main fieldvariables history that they experience. The “node tracking” optionof the software DEFORM-3DTM was utilized, highlighting, for a setof nodes initially placed along the sheets separation line in a trans-verse section, their final position after deformation. The identifiedpoints were monitored throughout the process: six points equallyspaced along the joint thickness were identified as shown in Fig. 2.

The reference transverse section was taken after 40 mm of weldlength, when the process has already entered a steady state and theobtained data are free from transient effects.

3. Results


3.1. Welding criteria implementation

In this study two welding criteria were considered, namelythe pressure–time (W) and the pressure–time–flow (W′). Both the


ARTICLE ING ModelPROTEC-13902; No. of Pages 10

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cttt

W

W

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Fig. 2. The identified points for the “node tracking” option.

riteria are based on integrals in time, that, in order to performhe calculation, have been approximated with sums over smallime intervals. Eqs. (1) and (2) represent the pressure–time andhe pressure–time–flow criteria, respectively.

=∫ t

0

P

�dt ∼=

∑j

pj

�j�tj (1)

′ =∫ t

0

P

�vdt ∼=

∑j

pj

�jvj �tj (2)

here p is the contact pressure at the interface, � is the flow stressf the material in the given temperature, strain and strain rate con-itions and v is the velocity of the considered node. As it can beoticed, W is measured in s, while W′ in mm. Hence, W can be con-idered as the “equivalent time”, calculated with proper weights ofressure and temperature (i.e. the flow stress) for the occurrencef solid bonding. Similarly, W′ can be considered as the “equiva-ent length” of material flow. In other words, a particle of material

ust possess a certain minimum velocity for a proper time onceppropriate temperature and pressure are reached.

In order to calculate, for each of the observation points high-ighted in Fig. 2, the value of the welding criterion, the materialow occurring during the FSW process must be properly predicted.uffa et al. (2013) proved the effectiveness of the developed model

or the prediction of the material flow by comparing the calculatedesults with experimental measurements of the zig-zag line dueo the oxides particles dispersed in the transverse section. Follow-ng the approach proposed in the above cited paper, the referenceoints were tracked during the process. It is worth noticing that thebservation time interval begins as the points experience non-zeroressure values, i.e. before the tool reaches the reference transverseection. By the same token, the observation period ends after theool leaves the reference section and pressure drops to zero. Thisehavior is due to the peculiar FSW process mechanics: the mate-ial close to the sheets separation line rotates together with the toolor a certain number of rounds, depending on the selected combi-ation of process parameters; the rotation begins in the retreatingide. Then the material is left at the advancing side of the joint. Inact, the latter is the area at which fracture occurs during tensileests of sound joints. Fig. 3 shows the evolution of the tracking line,nitially laying on the ideal separation surface between the sheets,


ill the final position reached by the six reference points (test C).rocess time and the x coordinate (welding direction) are providedor each figure. The origin x = 0 is set when the tool axis correspondo the considered transverse section. It is necessary to observe that


at the beginning the reference line is in the leading edge of the jointwhile at the end it lays on the trailing edge.

The local value of the flow stress for a given reference point andtime was calculated starting from the local values of temperature,strain and strain rate. Figs. 4a, 5a and 6a show, for case study C,the distribution of the above cited field variables in the referencetransverse section, right after the material closed the crack behindthe tool; Figs. 4b, 5b and 6b show the field variables histories ofpoint P5, as indicated in Fig. 3.

The field variables distribution highlights some of the charac-teristics of the FSW process: temperature is symmetric respect tothe welding line; in turn, strain and strain rate peaks are shiftedtowards the advancing side. Looking at the field variables variationwith time, it arises that the strain rate assumes non-zero valuesfrom about t = 18 to t = 26. When the material is rotating aroundthe tool, the accumulated strain increases quickly till a maximumvalue of about 26. Temperature start increasing as the welding toolapproaches the considered transverse section and starts decreasingduring the last part of the deformation, i.e. when the tool pin left thereference section and only the effect of the tool shoulder is present.It should be observed that, as the analyzed case study is character-ized by a welding speed of 100 mm/min, during this time intervalthe tool covers a distance of about 12 mm, corresponding to a fewmm more than the shoulder diameter. Then, for the P5 point, pos-itioned in the upper half of the joint thickness, it can be stated thatthe tool shoulder (and not just the pin) has a significant influenceon the evolution in time of the main field variables. Smaller timeintervals are observed for points located closer to the bottom of thejoints indicating that, in those areas, the main mechanical effect isdue to the pin. Besides, also the final values of the accumulatedstrain is lower for points closer to the bottom of the joints, startingfrom about 3.5 (see Fig. 6a).

As far as the contact pressure is considered, it is worth notic-ing that the separation line between the sheets (see again the redline in Fig. 3) changes both in shape and direction during the weld-ing process. Consequently, the contact pressure must be calculatedconsidering the change in the direction normal to this line foreach of the considered points. It is worth noticing that the utilizedsoftware does not allow the extrapolation of the contact pressurebetween two “moving flows” of the same deformable object. Forthis reason, the values of stress in the x and y directions were usedto obtain the contact pressure value. The final results of this proce-dure are shown in Fig. 7a in which, consistently with the previousFigs. 4–6, the evolution of the normal pressure of P5 is consideredfor test C. In Fig. 7b the evolution of the nodal velocity, needed forthe calculation of the W′ parameter, is illustrated. In Fig. 7c the flowstress characterizing P5 during the bonding process and derivingfrom the simultaneous effect of temperature, strain and strain rate,is shown.

Large values of pressure are observed during the deformation.At the same time, the flow stress is below 100 MPa. Starting fromt = 18 s, the flow stress decreases with increasing temperature; thena maximum, corresponding to the strain rate peak, and a minimum,when temperature reaches the highest peak, are observed. Finally,it starts increasing again after strain rate drops to zero and tem-perature starts decreasing. In this way, the welding parameters Wand W′ can assume values larger than 1. Finally, as expected, nodalvelocity shows a trend consistent with the strain rate one.

3.2. Welding criteria comparison


In order to assess the applicability of the considered bondingcriteria to FSW welding experiments were carried out as indicatedin Table 1. Fig. 8 shows the macrographs of the etched cross sectionsfor all the analyzed case studies.


Please cite this article in press as: Buffa, G., et al., Analytical bonding criteria for joint integrity prediction in friction stir welding ofaluminum alloys. J. Mater. Process. Tech. (2014), http://dx.doi.org/10.1016/j.jmatprotec.2014.02.014



Fig. 3. The identified points for the “node tracking option” – test C.

Fig. 4. (a) Temperature distribution in the transverse section and (b) evolution in time for P5 – test C.

Fig. 5. (a) Strain distribution in the transverse section and (b) evolution in time for P5 – test C.



6 G. Buffa et al. / Journal of Materials Processing Technology xxx (2014) xxx–xxx

erse se

tasLwtdr

Fig. 6. (a) Strain rate distribution in the transv

As it can be seen, the three welds characterized by rotation equalo 500 rpm are characterized by flow defects. In particular, test A has

very large void area at the bottom of the pin and a crack is visibletarting from the bottom of the joint and reaching the top surface.ooking at test B a tunnel defect is visible at the bottom of the pin,


hile a very small void can be observed in the same area for test C. Inurn, the three welds produced with rotation equal to 1000 rpm areefect free. In Fig. 9 the micrographs corresponding to the a, b, and ced dotted squares highlighted in Fig. 8 are reported. The small void

Fig. 7. Evolution in time of (a) pressure, (b) noda

ction and (b) evolution in time for P5 – test C.

observed at the bottom of the pin for test C is due to an insufficientheat conferred to the weld. As a consequence the material belowthe pin was too “cold” and was not properly mixed by the tool. Asthe weld becomes colder, i.e. with increasing welding speed, thevoid becomes larger assuming the name of tunnel or “wormhole”.


Finally a sound weld, corresponding to test E, is shown in Fig. 9c.The produced weds underwent tensile tests in order to assess

the effectiveness of the bonding obtained during the process. InFig. 10 the obtained results are reported as percentage of the ratio

l velocity and (c) flow stress for P5 – test C.





Fig. 8. Etched cross section of t

Fig. 9. Micrographs of the red dotted squares highlighted in Fig. 8.

he analyzed case studies.

between the UTS of the tested joint and the one of the base material.The three defect free welds, i.e. the ones obtained with rotationequal to 1000 rpm, have a similar resistance, ranging between 72%and 74%. On the other hand, a decreasing trend is found for thecase studies characterized by rotation of 500 rpm. In particular, theresistance of test C is only slightly lower than the one of the soundjoints. As the welding speed increases, the defect becomes largerand a dramatic drop of the UTS is observed till an almost zero valueis observed for test A.

The bonding parameters W and W′ were calculated, for the sixobservation points highlighted in Figs. 2 and 3, using the numericalresults shown in the previous paragraph. As far as the W parameteris regarded, i.e. the one based on the Pivnik and Plata criterion, athreshold value was taken from literature. In particular, as brieflydiscussed in the introduction paragraph, Ceretti et al. (2009a) usedroll bonding tests on AA6061 to determine the critical value as afunction of temperature. A regression was carried out obtainingthe following analytical expression:

Wlim = 4.9063e−0.0017T T > 320 ◦C (3)

Fig. 11 shows the calculated W parameter and the thresholdcurve for the 36 points investigated. The temperature correspond-


ing to each point is the average value of the temperatures calculatedduring the time interval corresponding to non-zero pressure val-ues. The points for which the criterion gives a wrong prediction arehighlighted in red. In particular, all the points of test A are below the

Fig. 10. Tensile test results for the considered case studies.



8 G. Buffa et al. / Journal of Materials Proces

lAdbt

TE

Fig. 11. Calculated W parameter and threshold value limit curve.

imit curve, correctly indicating that solid bonding was not reached.


satisfying prediction is obtained also for test B and C. The tunnelefect of test B corresponds to the points P1 and P2, which areelow the curve, as well as P1 for test A. The remaining observa-ion points are above the curve indicating a sound weld in those

able 2xperimental observation and calculated and threshold values for the W and W′ paramet

Test Point W [s]

Calculated value Threshold value

A P1 1.556 5.372

P2 0.697 5.297

P3 1.040 5.334

P4 0.694 5.269

P5 0.715 5.304

P6 1.107 5.341

B P1 2.552 4.810

P2 3.586 4.740

P3 5.565 4.673

P4 6.095 4.702

P5 4.849 4.583

P6 1.280 4.635

C P1 3.917 4.095

P2 7.018 4.106

P3 8.679 4.079

P4 10.898 4.079

P5 10.915 4.057

P6 7.835 4.128

D P1 1.797 3.928

P2 1.756 3.869

P3 1.760 3.847

P4 2.038 3.771

P5 1.069 3.752

P6 1.161 3.757

E P1 4.726 3.492

P2 7.523 3.443

P3 4.379 3.473

P4 3.590 3.496

P5 12.092 3.344

P6 5.558 3.550

F P1 10.059 3.073

P2 12.034 3.045

P3 16.326 3.038

P4 6.419 3.026

P5 5.279 2.989

P6 14.774 2.977


areas. The only exception is P6 for test B, which is well below thecurve. For all the performed tests, maximum pressure in the direc-tion orthogonal to the sheets separation line is observed at aboutmid thickness. Being P6 close to the top surface of the joint, smallvalues of pressure are calculated inhibiting reaching the expectedbonding parameter value. The network correctly predict test E andF, being all the observation points above the limiting curve. Finally,a completely wrong prediction is obtained for test D. Although theexperimental evidence clearly shown that a sound weld is obtained,all the six points are well below the threshold values. Again, the rea-son is the insufficient pressure and time due to the large weldingspeed used for the test. It should be observed that the points notshown in Fig. 11, namely C-P4, C-P5, E-P5, F-P1, F-P1, F-P2, F-P2and F-P6, exceeds the maximum value of the scale selected for thegraphic in order to focus on the wrong predictions. Finally, it can benoted that, even using a specifically defined limiting curve insteadof the one taken from literature, no advantage in the predictionaccuracy could be obtained. In fact, the six points of test D and B-P6have a too low W value. Using a lower limiting curve would resultin wrong prediction for B-P1, B-P2, C-P1 and many points of testA. In this way, the curve taken from literature and defined for adifferent process appears as the best possible thus confirming therobustness of proposed approach.


The reasons explaining the wrong predictions of the W param-eter leaded to include the effect of nodal velocity in the bondingcriterion, as shown in Eq. (2). In Fig. 12, the values calculated forthe six observation points of each case study are reported. A few

ers.

W′ [mm] Experiments

Calculated value Threshold value

1.039 3.644 not welded1.058 3.474 not welded2.337 3.558 not welded1.936 3.411 not welded1.743 3.489 not welded3.452 3.574 not welded

1.548 2.501 not welded2.158 2.380 not welded4.112 2.268 welded5.178 2.316 welded9.505 2.122 welded3.391 2.205 welded

0.749 1.448 not welded1.943 1.460 welded3.020 1.428 welded4.223 1.428 welded1.965 1.402 welded0.644 1.487 welded

1.412 1.256 welded1.960 1.193 welded2.954 1.170 welded4.619 1.093 welded5.832 1.075 welded5.411 1.080 welded





G. Buffa et al. / Journal of Materials Process

otctiflcWFWmnopitPt

epmseevc

W

3ai

4

oomdrA

Fig. 12. Calculated W′ parameter and threshold value limit curve.

bservations on the obtained values can be made. First, for all theests, P1 is the point characterized by the lowest W′ value. This isonsistent with what experimentally observed: at the bottom ofhe joints the least favorable conditions for solid bonding occurs,n terms of temperature, pressure and velocity. That is why if aow defect is observed, it is located in that area. Additionally, moreonsistent results are obtained for test D, showing the six points a

′ value larger than the corresponding values of test A, B and C.inally, focusing on test C, it is noted that P6 has an extremely low′ value, due to too low velocity values calculated by the numericalodel. This observation could lead to deduce that a sound weld is

ot obtained in that area, which is not confirmed by experimentalbservations. Numerical instabilities may have led to an incorrectrediction of the field variables by the FEM model. Hence, the bond-

ng criterion cannot give a correct prediction and underestimateshe W′ parameter. As for the previous Fig. 11, points E-P5, E-P6, F-5 and E-P5 are not shown in the graph as their W′ value exceedshe scale selected for better readability.

The better consistency of the results with respect both to thexperimental evidences and to the expected effect of the processarameters, in terms of heat input conferred to the joints, per-its to state that the W′ parameter is better suited to predict the

olid bonding phenomena occurring in FSW processes. Based on thexperimental evidences and the calculated values of the W′ param-ter, a regression curve can be obtained representing the thresholdalues, as a function of temperature, of the pressure–time–flowriterion for the considered AA6061 aluminum alloy:

′ = 1327661974400T−4.797 (4)

Finally, Table 2 summarizes the obtained results: for each of the6 points considered, the calculated and threshold W and W′ valuesre reported together with the experimental observation indicatingf, at a given area, material continuity can be observed.

. Conclusions

An experimental and numerical campaign was performedn FSW of AA6061 aluminum alloy. The applicability to FSWf analytical binding criteria, originally developed for different


anufacturing processes, was assessed. FSW experiments wereeveloped with tool rotation and welding speed varying in a wideange in order to obtain different conditions of joint integrity.n already developed and verified numerical model for FSW was

PRESSing Technology xxx (2014) xxx–xxx 9

used to calculate the main field variables values and the occur-ring material flow needed to implement the considered criteria.The following main conclusions can be drawn:

• In FSW the material close to the welding line flows around thetool in the retreating side and is eventually left at the advancingside after a few turns depending on the ratio between the toolrotation and welding speed;

• The Pivnik and Plata criterion can be used for FSW but problemsarise when large welding speed is selected; in those cases lowpressure is obtained and values of the W parameter lower thatthe threshold are calculated, indicating that no bonding occurredeven when sound joints are experimentally obtained;

• The pressure–time–flow criterion is better suited for FSW as theeffect of velocity is considered. In this way results more consistentwith what expected based on the process input parameters arecalculated;

• For the same criterion it was possible to identify a threshold curveand a power function of temperature was calculated through aregression. Using this limit curve, just one of the 36 analyzedpoints, i.e. P6 of test C, gives back a wrong prediction.

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Analytical Bonding Criteria for Joint Integrity Prediction in Friction Stirwelding of Aluminum...

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