Effect of tool geometry and process condition on static strength of a magnesium friction stir lap linear weld

Materials Science and Engineering A 528 (2011) 2463–2478

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Effect of tool geometry and process condition on static strength of a magnesium friction stir lap linear weld Q. Yang ∗ , X. Li, K. Chen 1 , Y.J. Shi Automotive Products Research Laboratory, Hitachi America Ltd., Farmington Hills, MI 48335, USA

a r t i c l e

i n f o

Article history: Received 13 August 2010 Received in revised form 17 November 2010 Accepted 8 December 2010 Available online 15 December 2010 Keywords: Friction stir lap linear welding Triangular pin tool Horizontal material flow Hook geometry Weld strength Magnesium alloy

a b s t r a c t Friction stir lap linear welding is conducted on overlapped AZ31 magnesium plates with different welding tools. Welds are made mainly with the orientation such that the weld retreating side on the upper plate is to be placed under load. Welding tools consist of a concave shoulder and a pin having a cylindrical, or triangular, or pie shape. This work addresses the effects of tool geometry and process condition on lap shear strength of welds. The shape of the hook formed due to upward bending of the plate interface on the retreating side and the strength of friction stir processed material are quantitatively characterized. Compared to the cylindrical tool, the triangular tool effectively suppresses the hook on the retreating side due to enhanced horizontal material flow. This primarily leads to a 78% increase in optimized weld strength. A ‘pure’ shear surface present on the tool pin significantly reduces weld strength. © 2010 Elsevier B.V. All rights reserved.

1. Introduction Friction stir welding (FSW) is a solid-state process where joining of metal plates is achieved through a thermo-mechanical action exerted by a non-consumable welding tool onto metal plates [1]. During welding, the synergy of tool rotation and tool translation along the butt line of aligned plates imparts severe plastic deformation to the materials (the absolute values of deformation strain and strain rate being up to one or two orders of magnitude [2]) as well as induces flow of the plasticized materials around the welding tool. FSW modifies the original microstructure to various extents in different locations of a weld. A stir zone where a recrystallized grain structure is generated, a thermo-mechanically affected zone, and a heat affected zone are encountered in sequence from the weld center towards the parent material. FSW has been extensively investigated on aluminum, titanium, steel, magnesium, and copper alloys [3–7]. In addition to welding in a butt configuration, friction stir welding is also oftentimes conducted in a lap configuration. The latter process is called friction stir lap linear welding (FSLLW) and shown

∗ Corresponding author at: Automotive Products Research Laboratory, Hitachi America, Ltd., 34500 Grand River Avenue, Farmington Hills, MI 48335, USA. Tel.: +1 248 474 2800x1814; fax: +1 248 473 8420. E-mail address: [email protected] (Q. Yang). 1 Current address: School of Materials Science and Engineering, Shanghai Jiao Tong University, Shanghai 200240, PR China. 0921-5093/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2010.12.030

in Fig. 1. As illustrated, two metal plates are overlapped by a certain width. A rotating tool, plunged into the material to a predetermined depth with the tool shoulder in close contact with the upper plate, is traversed along the centerline of the overlap. A lap linear weld is therefore produced after tool retraction. FSLLW has been widely applied in automotive and aeronautic industries to replace selfpierce rivet joining as well as to manufacture firm panels and beam assemblies where attachment of stringers and frames to the main body of a structure is required [8]. Material selection affects the execution of FSLLW. When FSLLW is conducted on two intrinsically different materials that have significant difference in their melting points, and/or form brittle inter-metallic compounds (e.g., aluminum and magnesium, aluminum/magnesium and steel, aluminum and copper, and so on), the welding tool pin is oftentimes penetrated only into the upper plate rather than through the interface of the overlapped metal plates [9,10]. Avoiding the formation of excessive intermetallic compounds at the plate interface and improving weld integrity by reducing micro-cracks present between the fragments of the higher-melting-point material and the matrix of the lowermelting-point material are two critical concerns for obtaining reasonably high joint strength [9,10]. When FSLLW is conducted on two intrinsically similar materials or the same material, the rotating tool pin is generally penetrated into the lower plate through the plate interface during its translation. Until they are deposited behind the welding tool, the plasticized materials undergo rotational, horizontal and vertical flows which break up thin oxide films at the plate interface around the tool pin and subsequently

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Fig. 1. Schematic illustration of friction stir lap linear welding.

contribute to partial/complete metallurgical bonding between the two overlapped metal plates. Meanwhile, the plate interface close to the pin is bent upwards to accommodate the tool penetration and translation, resulting in the formation of a hook-shaped macrostructure. During linear welding, plastic flows on the two sides of the welding tool are asymmetric: on the advancing side (AS) the tool rotation and translation provide the forces for plastic flow along the same direction (both are the driving forces), while on the retreating side (RS) they provide the forces in opposite directions (the driving force due to the tool rotation and the resistance due to the tool translation). As such, the hook geometries on these two sides also show dissimilarity. The presence of hook reduces the effective thickness of the upper plate, and thereby decreases the load-carrying capacity when a friction stir lap linear weld is loaded. Although the mechanical property of friction stir processed material within the weld region affects strength of a lap joint, increasing the effective thickness of the upper plate (designated as ‘load-carrying thickness’ hereafter) by suppressing the hook could be more effective to enhance joint strength. Hook suppression has been achieved by two approaches in aluminum friction stir lap linear welds. One approach includes lowering welding heat input and employing double weld passes from the perspective of welding operation [8,11]. Lowering welding heat input is achieved through decreasing tool rotation speed and/or increasing travel speed. During double-pass welding, welding orientation is reversed to switch the AS and RS of the two parallel weld passes so that the side (AS or RS) having the higher resistance to failure is placed as a loading side on both the upper and lower plates [11]. The approach of the double-pass welding was proven to effectively improve joint strength of a friction stir lap linear weld when the AS and RS of a friction stir weld have significantly different mechanical property. The other is, from the perspective of tool design, to use specially-profiled tools for welding, such as the Flared-Triflute and Skew-stir pin tools designed by Thomas et al. [12]. Previous investigation by the present authors showed that concave portions present on the tool pin may cause severe material sticking on the pin during friction stir welding of magnesium alloys, resulting in defects within a weld or even un-weldability. Magnesium alloys have high specific strength and good bending stiffness, and thus hold great potential for lightweight structural applications. Joining is one of the key issues that have to be addressed in order to realize this goal. In this paper, different tool pin geometries are designed for FSLLW of a magnesium alloy to investigate how they affect hook formation through varying material flow during welding. The effect of process condition on hook geometry and strength of a lap weld is examined as well. 2. Welding tool A cylindrical threaded pin tool is conventionally used for friction stir welding. Successive rotation of the cylindrical pin primarily causes shearing of the material in its vicinity [13,14]. As illustrated

in Fig. 2a, without taking the effect of pin threads into consideration, the material at a location ‘R’ within the plasticized region flows in a direction tangent to the pin surface (Vs ). Beyond the plasticized region, the material undergoes marginal deformation. A new type of tool pin which is in an equilaterally-triangular shape with three flat facets is designed. Successive rotation of this triangular pin enhances horizontal plastic flow of the material in the vicinity of the pin, as compared with the cylindrical pin. As illustrated in Fig. 2b, assuming that there is no slippage between the pin and the material in its immediate vicinity, with the designated pin rotation direction, the motion of the material between location ‘a’ at the vertex and location ‘b’ at the center of the pin surface can be resolved into two components. In one component, the material shears along the pin surface (Vs ). In the other component, the material is pushed outwards by the pin surface (Vp ). From location ‘a’ to location ‘b’, the latter component (Vp ) gradually reduces until it becomes zero at location ‘b’. Immediately in front of the pin surface between location ‘b’ and location ‘c’, the rotation of the pin has created a virtual cavity. Therefore, not only the material that earlier experiences an outward movement (from location ‘a’ to location ‘b’) would be now deflected inwards by the wall of the outer ‘cold’ bulk material (indicated by the dotted circle) towards the virtual cavity, but also the material originally located in front of the surface ‘bc’ flows towards this cavity when the welding tool is translated forwards (i.e., towards the bottom of the page). The instantaneous virtual cavity is quickly filled by the inward-flowing material which further moves to the back of the pin during pin rotation. Hence, a plastic flow forwards and backwards in the horizontal direction in the vicinity of the pin becomes significant when the triangular pin is rotated. Note that the geometry of the plasticized region illustrated in Fig. 2 does not represent the actual geometry developed during friction stir welding. In a real weld, the plasticized region is asymmetric on the two sides of the welding tool. In addition, it should reflect the pin geometry. Only a circular concave profile is selected for tool shoulder (the concavity is 10◦ ). Combining the geometries of tool shoulder and pin, three types of welding tools are designed. They are called cylindrical threaded pin tool (C-type), pie-shaped pin tool (H-type) and triangular pin tool (T-type), as illustrated in Fig. 3. The H-type tool is a transition from the C-type to the T-type, where both the feature of flat facet and the feature of round surface (with grooves) are included. The welding tool has a variable shoulder diameter and pin diameter. For the H-type and T-type tools, the pin diameter is equivalent to the diameter of an inscribed circle of the profiled pin. Here, a parameter called tool geometry ratio (R) is defined as the ratio of shoulder diameter to pin diameter. The geometries and features of the welding tools are listed in Table 1. For the cylindrical tools, Tools C2 and C3 have a larger shoulder size than Tool C1, and the pin diameter of Tool C2 is equal to that of Tool C1 but larger than that of Tool C3. Similarly, for the triangular tools, Tools T2 and T3 have a larger shoulder size than Tool T1, and the pin diameter of Tool T2 is equal to that of Tool T1 but larger than that of Tool T3. All the welding tools are made of the tool steel H13.

3. Experiment 3.1. Material AZ31-H24 magnesium plates 2.0 mm thick were chosen for the present study. The nominal chemical composition of this alloy is Mg–3.0 wt.% Al–1.0 wt.% Zn. Typical microstructure of the asreceived alloy contains recrystallized fine grains and a large number of mechanical twins (Fig. 4a). The as-received alloy has yield strength of 228 MPa and ultimate tensile strength of 295 MPa (Fig. 4b).

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Fig. 2. Schematic illustration of material motion at the periphery of rotating profiled pins: (a) cylindrical pin; (b) triangular pin. Illustrated geometries of the plasticized regions outside the tool pins do not represent the actual geometries developed during friction stir welding.

Fig. 3. Schematic illustration of welding tools: (a) cylindrical threaded tool; (b) pie-shaped tool; (c) triangular tool. Pin threads in (a) and grooves in (b) are not shown.

3.2. Friction stir lap linear welding (FSLLW) Prior to welding, magnesium plates 300 mm long and 100 mm wide were overlapped and tightly clamped on the work table by the welding fixture. The longitudinal direction of the plates was perpendicular to the plate rolling direction (RD). A supporting plate of the same thickness was placed underneath the upper plate to help align and stabilize the plates to be joined. The overlap along the longitudinal direction of the plates was 40 mm wide. One pass of FSLLW was performed along the longitudinal center line of the overlap (i.e., the welding direction, WD was perpendicular to the RD). To enhance weld consolidation, the tool was tilted at 3◦ from the normal direction (ND) of the plate towards the trailing side of the tool during welding. In addition, the tool shoulder was cut into

the upper plate by a depth of approximately 0.25 mm. This depth was defined as the maximum depth from the top surface of the upper plate to the weld surface. FSLLW was conducted at selected rotation speeds of 1500, 2000 and 2500 rpm and selected travel speeds of 100, 200, 350 and 500 mm/min. Strength of welds loaded nominally in lap shear was examined (bending of the upper and lower plates is unavoidably introduced and becomes more severe at the later deformation stage). The details for mechanical testing were described below. Due to the asymmetric characteristic of friction stir welding, two welding orientations are present, thus resulting in two different loading modes. In Mode I, the advancing side of a lap weld on the upper plate is loaded, while in Mode II, the retreating side of a lap joint on the upper plate is loaded, as schematically illustrated in Fig. 5a. Fig. 5b

Table 1 Welding tool geometry and feature. Tool #

C1 C2 C3 H1

T1 T2 T3

Tool shoulder

Tool pin

Profile

Diameter (mm)

Profile

Concave circular shoulder (concavity: 10◦ )

12 14 14 12

Cylindrical shape having right-handed threads on the perimeter Pie shape having right-handed grooves on the perimeter. Two flat facets are inclined at 60◦ . Truncated equilaterally-triangular shape having three flat facets

12 14 14

Tool geometry ratio, R Equivalent diameter (mm) 5 5 4 6.5

2.4 2.8 3.5 1.85

6.5 6.5 5.8

1.85 2.15 2.41

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Fig. 4. (a) Microstructure of as-received AZ31 Magnesium plate; (b) Engineering strain-stress relationship of as-received AZ31 magnesium plate (tensile direction along the RD).

exemplifies the effect of loading mode on lap shear strength of welds. The welds were made at a process condition of 2000 rpm and 350 mm/min with the tools C1 and T1, respectively. Higher weld strength is attained at loading Mode II, in particular for the welds made by the T-type tools. Hereafter, FSLLW was conducted with the retreating side of a weld on the upper plate placed on the loading side. During welding, the tool pin was penetrated through the interface into the lower plate. Fig. 6 shows the variation of lap shear strength with pin length which is in the range of 2.2–3.7 mm. Friction stir lap linear welds were produced with Tool T1 at a rotation

speed of 2000 rpm and variable travel speeds. At travel speeds higher than 200 mm/min when reasonably high joint strengths are obtained, lap shear strength first increases with the pin length. As the pin is longer than 2.6 mm, further increasing the pin length does not cause a significant strength variation. Therefore, a welding tool that has a pin length more than 1.3 times the thickness of the upper plate (but less than the entire thickness of the overlapped plates) is suggested for FSLLW. A shorter pin was previously reported to help suppress hook and increase weld strength in Al alloy [11]. For the present study, the pin length of 3.2 mm was chosen. 3.3. Characterization of welds Test specimens approximately 12 mm wide were cut from friction stir lap linear welds along the transverse direction (TD) of welds. Spacers were attached onto both ends of a specimen prior to testing to ensure that an initial pure shear load was applied to the interfacial plane (Fig. 5a). Lap shear testing was conducted at a rate of 2.0 mm/min. Weld strength (also called ‘lap shear strength’) was computed as the ratio of load-at-failure to initial specimen width. No fewer than three specimens were tested to obtain average weld strength

500

Lap Shear Strength, N/mm

AZ31; Tool T1

2000RPM

400

300

200

100

0

2

100mm/min

350mm/min

200mm/min

500mm/min

2.5

3

3.5

Pin Length, mm Fig. 5. (a) Two different loading modes and (b) Effect of loading mode on lap shear strength of welds.

Fig. 6. Effect of pin length on lap shear strength of welds.

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for a process condition. Standard deviation of strength was given. Macrostructural and microstructural examinations were performed on as-received alloy, and as-welded and mechanicallytested specimens. As-received alloy was sectioned in a plane containing the ND and RD. As-welded specimens were sectioned in a plane containing the ND and TD. Mechanically-tested specimens were sectioned in a plane containing the plate ND and the tensile direction (i.e., weld TD). All specimens were subjected to mechanical grinding and polishing with 0.05 ␮m silica suspension. The specimens were then etched with an acetic-picral solution (4.2 g picric acid, 70 ml ethanol, 10 ml acetic acid and 10 ml distilled water). Macrostructure and microstructure were examined using optical microscope (OM). Hook geometry and load-carrying thickness were measured on planar macrostructures. 4. Results 4.1. Weld strength Fig. 7 shows the effect of pin shape on lap shear strength of welds. Take the welding tools C1, H1 and T1 for example. The tool shoulders have the same size, and the effect of the equivalent pin diameter is neglected tentatively. The cylindrical threaded pin leads to weld strength less than 225 N/mm (Fig. 7a). The pieshaped pin having two flat facts leads to a slight increase in weld strength at proper process conditions, but the magnitude of weld strength is less than 290 N/mm (Fig. 7b). The triangular pin leads to a remarkable increase in weld strength when the tool travel speed is higher than 200 mm/min (Fig. 7c). In addition, high weld strengths around 400 N/mm can be maintained within a wide range of rotation speeds from 1500 rpm to 2500 rpm. Compared to the peak weld strength given by the cylindrical tool, the peak weld strength given by the triangular tool is increased by 78%. The presence of ‘pure’ shear surface on the tool pin significantly reduces lap shear strength of a friction stir lap linear weld. Fig. 8 shows the effect of tool dimensions on lap shear strength of welds. Consider the cylindrical tool and triangular tool only. Comparing plots (a–c) with plots (d–f), pin shape dominates weld strength. The welding tools having a triangular pin constantly lead to higher weld strengths than the welding tools having a cylindrical threaded pin, regardless of the tool dimensions. For the cylindrical tool, increasing the tool shoulder size (and therefore the R value) generally reduces weld strength at lower tool travel speeds and particularly higher rotation speeds (Fig. 8a–c). For the triangular tool, increasing the tool shoulder size increases weld strength at the travel speed of 100 mm/min but does not affect weld strength at the travel speeds higher than 200 mm/min, thus making weld strength less sensitive to process condition (Fig. 8d–f). The limited variation of pin diameter does not result in an appreciable weld strength change. 4.2. Weld macrostructure Fig. 9 shows the effect of pin shape on macrostructure of welds. Take the welding tools C1, H1 and T1 for example. A pore (often in an elongated shape) is present close to the weld bottom at lower rotation speeds and/or higher travel speeds. The process window for producing defect-free welds by Tools C1, H1 and T1 becomes narrower in sequence. This indicates that rotation of pin threads/grooves in a beneficial direction contributes to driving the plasticized material downwards until the material is deposited behind the welding tool. The hook on the advancing side (AS) of welds is directed upwards at an inclination angle towards the weld surface and then arrested at the stir zone extremity. For the cylindrical pin (Tool C1),

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the inclination angle is reduced as the rotation speed decreases and/or the travel speed increases (Fig. 9a). For the pie-shaped pin (Tool H1) and the triangular pin (Tool T1), the inclination angle does not show a noticeable change with process condition (Fig. 9b and c). The hook geometry on the retreating side (RS) is more complex. For the cylindrical pin, the hook first runs gradually upwards. It is then deflected downwards into the stir zone and even goes further towards the advancing side (Fig. 9a). Within the stir zone, the hook becomes discontinuous and indistinct. For the pie-shaped pin and the triangular pin, the hook enters the stir zone in a manner similar to that for the cylindrical pin. However, after passing the weld center, the hook moves again upwards towards the end of the hook which extends from the advancing side. This phenomenon becomes pronounced at lower travel speeds and higher rotation speeds (Fig. 9b and c). No intersection exists between the hook extending from the advancing side and the hook from the retreating side. Hook morphology and weld consolidation affect failure mode of a lap weld, which is described below. In addition, the hook from the advancing side is generally pointed upwards at a sharper angle and without deflection compared to the hook from the retreating side. This could be the main reason for the higher weld strength attained under the loading Mode II instead of the loading Mode I (Fig. 5b). Look into the hook geometry on the retreating side first in a qualitative way, since it affects the magnitude of loading-carrying thickness. The hook height (hh ) is defined as the depth from the initial plate interface to the apex of the hook, as illustrated in Figs. 9c and 13b. As shown in Fig. 9, for all the pin shapes, the hook height decreases as the tool travel speed increases. Beyond 200 mm/min, the hook height does not show a prominent decrease. The hook height does not display a significant variation with tool rotation speed at medium to high travel speeds for Tools C1 and T1. Very critically, the hook heights of the welds produced by the triangular pin are dramatically reduced, as compared to those by the pie-shaped pin and the cylindrical pin. This finding indicates that enhanced horizontal plastic flow due to flat facets suppresses hook formation during FSLLW, and that flat facets, if combined with a ‘pure’ shear surface, significantly reduce the effectiveness of hook suppression. Fig. 10 shows the effect of tool dimensions on the macrostructure of welds made by the cylindrical pin tool and the triangular pin tool, respectively. Exemplary welds were made at 2000 rpm and different travel speeds. For the cylindrical tool, increasing the tool shoulder size generally increases the hook height (Fig. 10a), and this increase is apparent for the welds made at medium travel speeds. (For brevity, weld macrostructures at medium travel speeds of 200 and 350 mm/min are not shown here. The magnitude of hook height is shown in Fig. 15a–c). For the triangular tool, increasing the shoulder size remarkably reduces the hook height at a low travel speed of 100 mm/min, and this reduction becomes less prominent when the travel speed is higher than 200 mm/min (Fig. 10b). The limited variation of pin size does not cause an appreciable change in hook height with process condition. Furthermore, the tool rotation speed has a much smaller effect on the variation of hook height with tool dimensions for the triangular pin tool than for the cylindrical pin tool. 4.3. Weld microstructure As shown in Fig. 11, a basin-shaped stir zone is developed in a friction stir lap linear weld. Microstructure analysis shows that (1) friction stir welding changes the original twinned microstructure into recrystallized microstructure; (2) no obvious elongated and aligned grain structure is formed in the transition zone between the base material and the weld stir zone. This indicates that the deformed microstructure is readily subject to dynamic recovery

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Fig. 7. Effect of pin shape on lap shear strength of welds: (a) cylindrical pin from Tool C1; (b) pie-shaped pin from Tool H1; (c) triangular pin from Tool T1. Ds = diameter of tool shoulder, Dp = diameter of tool pin, R = Ds /Dp .

and recrystallization during welding as well as subsequent static recrystallization. Microstructure evolution is found to be primarily associated with process condition. Fig. 11b and c show the grain structures for Region I (above the hook apex) and Region II (at a central location) in the weld made with the cylindrical tool C1 at 2000 rpm and 350 mm/min, respectively. Regions I and II are marked on the cross-sectional macrostructure in Fig. 11a. Grain size was measured using the linear intercept method. The average grain size is 10 ␮m for Region I and 10.1 ␮m for Region II. Fig. 11e and f show the grain structures for similar locations, Regions III and IV in the weld made with the triangular tool T2 at 2000 rpm and 100 mm/min, respectively. Regions III and IV are marked on

the cross-sectional macrostructure in Fig. 11d. The average grain size is 16.1 ␮m for Region III and 14 ␮m for Region IV. A higher travel speed (and also a lower rotation speed) along with a smaller shoulder size leads to a finer grain structure. Furthermore, grain structure becomes more uniform from the transition zone towards the center of the stir zone in such a process condition. 4.4. Failure modes of welds As illustrated in Fig. 12, five failure modes are observed when friction stir lap linear welds are lap shear tested. They are: failure initially propagating along the hook and then linking with a large

Fig. 8. Effect of tool dimensions on lap shear strength of welds for: (a–c) cylindrical pin tool; (d–f) triangular-shaped pin tool.

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Fig. 9. Effect of pin shape on macrostructure of welds: (a) cylindrical pin from Tool C1; (b) pie-shaped pin from Tool H1; (c) triangular pin from Tool T1. Metallographic specimens were not etched.

Table 2 Failure modes of lap welds. Process condition Rotation speed (RPM)

Failure mode Travel speed (mm/min)

Tool C1

Tool C2

Tool C3

Tool T1

Tool T2

Tool T3

1500

100 200 350 500

FM5 FM5 FM2 FM1

FM3, FM5 FM3 FM3 FM3

FM3 FM3 FM3 FM3

FM3 FM1, FM3 FM1, FM3 FM1

FM3 FM3 FM3, FM5 FM3, FM5

FM4 FM3 FM3 FM3

2000

100 200 350 500

FM3 FM3 FM3 FM2

FM3 FM3 FM3 FM3

FM3 FM3 FM3 FM3

FM3 FM3 FM3 FM1, FM3

FM3 FM3 FM3 FM4, FM5

FM3 FM3 FM4, FM5 FM3

2500

100 200 350 500

n/a n/a FM3 n/a

FM3 FM3 FM3 FM3

FM3 FM3 FM3 FM3

FM3 FM3 FM3 FM4

FM3 FM3 FM3 FM3

FM3 FM3 FM3 FM3

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Fig. 10. Effect of tool dimensions on weld macrostructure for: (a) cylindrical pin tool; (b) triangular-shaped pin tool. Exemplary welds were made at 2000 rpm and different travel speeds. Metallographic specimens were not etched.

pore at the weld bottom (FM1); fracture within the stir zone and on the advancing side of the lower plate (FM2); fracture within the stir zone and on the retreating side of the upper plate (FM3); failure propagation into the stir zone along the hook and a final fracture within the stir zone on the advancing side of the upper plate (FM4); plate separation along the hook throughout the stir zone (FM5). The failure propagation along the hook is evidently seen in Fig. 12f. Table 2 lists all the failure modes of friction stir lap welds under external lap shear loading. The welds are made with the cylindrical and triangular tools at different process conditions. As shown, FM3 is the most dominating failure mode. FM1 occurs when a weld is produced at a lower rotation speed and/or a higher travel speed (i.e., a process condition of low heat input), and contains a sufficiently large pore. A small-sized pore present in the stir zone (for example, a weld made with Tool T1 at 2000 rpm and 350 mm/min) may not influence the weld strength since the failure mode FM3 takes effect. The failure mode FM2 is seldom observed only in the welds made by the cylindrical tool at process conditions of low heat input. Microscopic examination of fracture surface indicates that large oxide particles most probably contribute to such a weld failure. Insufficient heat and deformation due to low heat input are expected not to fully break up the original interfacial oxide film, leaving large-sized particles in the stir zone. The failure mode FM4 is similar to the failure mode FM5, and their occurrence is attributed to the hook extension into the advancing side of the stir zone. The slight difference in the hook morphology and the upward direction of the extended hook towards the advancing side makes one of them operative. The welds made by the triangular tools and with the failure modes FM4 and FM5 have high lap shear strengths.

For the failure mode FM3, load-carrying thickness (tl ), which is defined as the vertical distance from the failure starting point to the weld surface, is measured on as-welded macrostructure. It is shown in Fig. 13b that the apex of the hook ‘a’ on the retreating side of a weld is not exactly but close to the weld failure starting point ‘f’ (in other words, the failure of a weld is not initiated from the hook apex). For the welds made by the cylindrical tool in which the hook height hh is significant, the failure starting point ‘f’ is readily identified from the cross-sectional macrostructure of the failed weld (Fig. 13a). By locating the position on the as-welded macrostructure which corresponds to the failure starting point ‘f’, the load-carrying thickness tl can thus be measured (Fig. 13b). For the welds made by the triangular tool in which the hook is significantly suppressed, resulting in a small hook height, apparent localization of deformation (i.e., necking) takes place near the shear fracture surface when they are loaded to failure (Fig. 13c and d). Given that the detached portion on the retreating side shows less necking and macroscopic shape change than the other side of the fracture surface during tension, this portion is selected for locating the failure starting point. As shown in Fig. 13d, the position ‘f ’ which is the intersection of the shear fracture surface and the top surface of the upper plate and the intersection angle ˇ are determined on the detached portion. With the intersection ‘f ’ and the intersection angle ˇ, the segment f f which corresponds to the shear fracture surface is drawn on the as-welded macrostructure to obtain the failure starting point ‘f’ (Fig. 13e). The load-carrying thickness tl is therefore measured. As shown, the failure starting point ‘f’ is located close to the apex of the hook as well.

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Fig. 11. Microstructures of welds: (a–c) weld made with Tool C1 at 2000 rpm and 350 mm/min; (d–f) weld made with Tool T2 at 2000 and 100 mm/min.

Fig. 12. Failure modes of welds after lap shear testing. See the text for the details.

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Fig. 13. Illustration of measuring hook height and load-carrying thickness for welds: having an obvious failure starting point (a, b); showing apparent necking near the fracture surface (c, d, e) (a, c, d) macrostructures of failed welds; (b, e) as-welded macrostructures. The welds display the failure mode FM3.

Fig. 14 shows microstructures on the two sides of the fracture surface (i.e., Regions V and VI) in a failed weld. Regions V and VI are marked in Fig. 13c. Intensive twinning is observed immediately near the fracture surface (Locations 3 and 6), and twinning is progressively reduced as the distance from the fracture surface increases on both sides. For the currently examined weld which was made with Tool T2 at 2000 rpm and 100 mm/min, twinning becomes almost invisible 1.7 mm from the fracture surface towards the weld center (Fig. 14a), and 2.6 mm from the fracture surface towards the base material (Fig. 14b). This indicates that plastic deformation is localized when a friction stir lap linear weld is loaded to failure. 4.5. Quantitative assessment of hook geometry and strength of friction stir processed material The effects of process condition and tool dimensions on the hook height (hh ) of the weld made by the cylindrical tool are summarized in Fig. 15a–c. Generally, the hook height increases as the rotation speed increases and/or the travel speed decreases. Increasing the shoulder size apparently increases the hook height at medium travel speeds, and this increase becomes more significant at higher rotation speeds. Fig. 15d–f shows the effects of process condition and tool dimensions on the load-carrying thickness (tl ) of the weld having the failure mode FM3. A decrease of hook height leads to an increase in load-carrying thickness. Re-plot weld strength as a function of load-carrying thickness in Fig. 15g–i. The slopes of the fitting lines which pass through the origin are 200 MPa for all the dimen-

sions of the cylindrical tool. The data points for Tool C2 appear to be quite scattering (Fig. 15h). The slope is the strength of friction stir processes material, which determines the weld strength along with the load-carrying thickness. As indicated, the influence of process condition and tool dimensions within the present investigation ranges on the strength of friction stir processed AZ31 magnesium can be ignored. The effects of process condition and tool dimensions on the hook height of the weld made by the triangular tool are summarized in Fig. 16a–c. Generally, the hook height decreases as the travel speed increases, and this decrease becomes less prominent when the travel speed is higher than 200 mm/min. Beyond 350 mm/min, the hook height is close to zero in the welds made by Tools T2 and T3, indicating that the hook on the retreating side is nearly completely suppressed. The hook height does not exhibit a significant variation with the tool rotation speed, in particular when the travel speed is higher than 200 mm/min. Increasing the shoulder size decreases the hook height at a lower travel speed of 100 mm/min, making the hook formation less sensitive to the process condition. Fig. 16d–f shows the effects of process condition and tool dimensions on the load-carrying thickness of the weld having the failure mode FM3. The load-carrying thickness has larger magnitudes at higher travel speeds. Re-plot weld strength as a function of load-carrying thickness in Fig. 16g–i. The slopes of the fitting lines are about 225 MPa for all the dimensions of the triangular tool. Again, the strength of friction stir processed material indicated by the slope is not evidently affected by process condition and tool dimensions within the present investigation ranges.

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Fig. 14. Microstructures of a failed weld (shown in Fig. 13c) near the fracture surface. The weld was made with Tool T2 at 2000 rpm and 100 mm/min.

By comparison, the strength of friction stir processed material due to the triangular pin is slightly higher than that due to the cylindrical pin. Both values are 68–76% of the ultimate tensile strength of the base material. 5. Discussion 5.1. Effect of process condition and tool geometry on hook formation The plate interface decorated by oxide film can be regarded as a tracer. An appropriate interpretation of the geometrical change of the plate interface helps understand material flow during welding. As shown in Figs. 9 and 10, the initial plate interface undergoes upward bending to form a hook-shaped macrostructure near the weld stir zone. The hook on the retreating side extends into the stir zone towards the advancing side, and progressively becomes discontinuous within the stir zone. Furthermore, decreasing tool rotation speed (for the cylindrical tool) and/or increasing travel speed, as well as employing a triangular pin are beneficial to hook suppression on the retreating side

(Figs. 15 and 16). A simple but rational analysis is given as follows. Friction stir welding is a synergy of tool translation and tool rotation. Tool translation primarily drives upward motion of the material in front of the tool as well as on the sides [15,16]. The upward flowing material or its majority has to be moved to the back of the tool via any method to ensure the continuity of the welding process. Tool rotation is effective in transporting the plasticized material through the retreating side to the back of the welding tool. Meanwhile tool rotation alters the flow caused by tool translation on the sides of the tool. As believed, on the advancing side, the upward flowing material enters a rotational shear zone and undergoes severe shear deformation before it is mainly deposited on the same side [17,18]. The interfacial oxide film is therefore expected to be fully broken up inside the shear zone, and its connection to the surrounding interfacial oxide is interrupted. As such, the hook on the advancing side is ended at a sharp angle at the extremity of the rotational shear zone which is also the stir zone of a weld, and the initial plate interface becomes discontinuous or even invisible within the stir zone on this side (Figs. 9 and 10). On the retreating side, the upward flowing material is rather extruded between

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Fig. 15. Effect of tool geometry on: (a–c) hook height; (d–f) load-carrying thickness. (g–i) plot the relationship between weld strength and load-carrying thickness. All the welds were made with cylindrical pin tools (Tools C1, C2, and C3). Load-carrying thicknesses were measured on welds having the failure mode FM3.

the base material and the shear zone towards the back of the tool [18]. Less deformation exerted onto such material causes the interfacial oxide film not to be significantly disintegrated, so that the hook appears to extend into the stir zone from the interface of the base plates on the retreating side. During deposition of the plasticized material behind the welding tool, the rear of the tool shoulder (or shoulder-driven material) forges the pin-driven material [19]. Therefore, a downward deflection of material flow is expected on the retreating side to form the hook-shaped macrostructure. It is well known that rotation of a cylindrical pin produces shear deformation onto the surrounding material and thus drives the material backwards. In comparison, material transport by the triangular pin is distinct. Two factors could result in the hook suppression on the retreating side. During tool translation, the

plasticized material is expected to move upwards as well as horizontally into the area between the pin surface and the inscribed circle in front of the tool and on the sides (Fig. 2b). Furthermore, in addition to the ‘extrusion’ process, tool rotation provides a free path for material flow towards the back of the tool on the retreating front side. As described in Fig. 2b and Section 2 where a most severe upward material flow is illustrated, the plasticized material in front of the pin surface ‘bc’ partially moves into a virtual cavity which exists instantaneously but is continually re-created during the tool rotation. This part of material is then directly transported to the back of the tool. Overall, the material deposition on the retreating side is achieved by the ‘extrusion’ process and virtual cavity-assisted material flow. The above two factors enhance the horizontal material flow and arrest the

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Fig. 16. Effect of tool geometry on: (a–c) hook height; (d–f) load-carrying thickness. (g–i) plot the relationship between weld strength and load-carrying thickness. All the welds were made with triangular-shaped pin tools (Tools T1, T2, and T3). Load-carrying thicknesses were measured on welds having the failure mode FM3.

upward material flow on the retreating side, thereby decreasing the hook height (Figs. 9,10 and 16). The quantity of the upward-flowing material stacked on the front advancing side is expectably larger than that on the front retreating side during friction stir welding. This may cause the characteristic geometry of the hook segment within the stir zone which goes downwards from the retreating side and then upwards towards the advancing side (Figs. 9 and 10). Process condition (i.e., a combination of rotation speed and travel speed) could also affect the hook height on the retreating side (Figs. 15 and 16). This is believed to be associated with welding heat input (or welding temperature). The effect of process condition on welding heat input can be characterized by heat index, HI,

as follows [20]: HI =

ω2

v

(1)

where ω is the rotation speed, rpm; v is the travel speed, mm/min. The higher the HI-value, the higher the welding heat input. Note that the effect of tool geometry is not integrated into this expression. From Eq. (1), a higher rotation speed and/or a lower travel speed lead to a higher heat index. Fig. 17a shows hook height as a function of heat index for the cylindrical tools. In general, hook height increases first significantly and then less prominently with heat index. A higher welding heat input reduces the work hardening and viscosity of the plasticized material, which would predomi-

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Fig. 17. Hook height as a function of heat index for: (a) cylindrical pin tool; (b) triangular pin tool.

nantly enhance the upward flow of the material in front of the tool and on the sides rather than the transport of the material towards the back of the tool. As a result, the hook height of the weld is increased. The friction of the tool shoulder against the material is a major heat source. Increasing the shoulder size increases the welding temperature and thereby the hook height. Furthermore, with increasing the shoulder size, the increase in the hook height is more outstanding at medium HI-values which are obtained at medium travel speeds and high rotation speeds (Fig. 15). Fig. 17b shows hook height as a function of heat index for the triangular tools. Hook height does not exhibit an apparent increase with heat index until heat index reaches high values (≥40,000 rpm2 min/mm). These high HI-values are obtained at the travel speed of 100 mm/min. On the other hand, hook height is reduced to some extent by increasing the shoulder size. Therefore, a higher heat input due to increased heat index and/or shoulder size may not increase the hook height. Considering the distinct features of material transport by the triangular tool, it can be imaged that increasing the shoulder size or the tool rotation speed facilitates not only the upward flow but also the horizontal flow of the plasticized material. The overall result of these two flows may cause insignificant variation of the hook height at the tool travel speeds higher than 200 mm/min (Fig. 16). From the effects of tool geometry and process condition on hook formation, it can be concluded that the hook height is more likely related to the volume of material transported towards and discharged at the back of the tool in a specific time period instead of the welding temperature. In summary, the triangular pin enhances horizontal material flow as it is translated and rotated. The horizontal flow dramatically reduces or even arrests the upward flow of the material on the retreating side, suppressing the hook on this side. The hook height on the retreating side is not significantly dependent upon the process condition unless extremely high welding heat is input. In contrast, rotation of the cylindrical pin produces ‘pure’ shear deformation, and the upward material motion due to tool translation cannot be effectively reduced or arrested through tool rotation. The hook height becomes strongly related to the process condition and the tool dimension.

facilitate the plate separation throughout the stir zone due to the shearing effect. The integrity of metallurgical bonding between the overlapped plates (i.e., the degree of fragmentation of the interfacial oxide), and the length of metallurgical bonding affect the shear load-carrying capability. The tensile load on the retreating side of the upper plate is carried by the remaining thickness of the upper plate (i.e., load-carrying thickness), and the tensile load-carrying capability is determined by multiplication of the load-carrying thickness (tl ) and the strength of friction stir processed material. Tension and shearing compete against each other during loading to failure, as during loading a resistance spot weld to failure [21]. When the shear load-carrying capability is larger than the tensile load-carrying capability, the weld fails on the retreating side of the upper plate (i.e., the failure mode FM3). When the shear load-carrying capability is lower than the tensile load-carrying capability, the plate separation throughout the stir zone (i.e., the failure mode FM4 or FM5) takes place. Both failure modes are attainable when the shear load-carrying capability is comparable to the tensile load-carrying capability. The present study shows that FM3 is the dominant failure mode (Table 2), indicating that friction stir welding creates sound metallurgical bonding between the overlapped plates. Although the triangular pin increases the strength of friction stir processed material to some extent, the weld strength is improved primarily through increasing the load-carrying thickness (tl ) due to hook suppression on the retreating side (Figs. 15 and 16). It is worth noting that the joint strength of a friction stir lap weld cannot be further improved when the length of metallurgical bonding between the overlapped plates exceeds a critical length, since the tensile load-carrying capability is lower than the shear loadcarrying capability when the critical length is exceeded. The pore existing in the weld stir zone may not affect joint strength of a lap weld. When the path that failure propagation follows is connected to the existing pore (usually in a large size), the weld fails in the mode FM1 (Fig. 12a), and the weld strength is quite low. Otherwise, an existing pore does not influence the joint strength of a magnesium lap weld since the failure mode FM3 takes effect. 5.3. Strength of friction stir processed material

5.2. Failure mode and weld strength When a friction stir lap linear weld is lap shear loaded in the loading mode II, the retreating side on the upper plate (and the advancing side on the lower plate) primarily undergoes tension, and the weld region primarily undergoes shearing. The hook morphology on the retreating side and its extension into the stir zone

The strength of friction stir processed material is only 68–76% of the tensile strength of the base material (Figs. 15 and 16). Mechanical property of wrought magnesium alloys displays high crystallographic texture dependence [22,23]. During friction stir welding, the action of the welding tool on the material develops a strong shear texture in the stir zone [13,14]. Fig. 18 shows the varia-

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Fig. 18. Variation of grain orientation in characterized regions from weld center towards the base material on the RS. The {0 0 0 2} pole figure with basal texture intensity for each region is shown below. (a) represents the weld center; (c) represents the stir zone extremity.

tion of grain orientation with respect to the welding direction (WD) and the corresponding {0 0 0 2} pole figure in an AZ31 magnesium friction stir weld from the weld center towards the base material on the retreating side [data from a previous unpublished research]. Welding was performed with a cylindrical tool on a 2.0 mm thick plate. The material in the characterized area is deformed primarily by the rotating pin. As shown in Fig. 18a, the weld center has a single basal fiber texture in which the (0 0 0 2) basal poles are roughly aligned with the WD (i.e., the basal planes are parallel to the pin surface). The region depicted in Fig. 18c is the stir zone extremity where the basal texture splits into two components of similar intensities. In one component, the basal planes are approximately parallel to the pin surface but have a rotation around the plate normal direction (ND). In the other component, the basal planes show a further rotation around the WD from the orientation of the former component. The region in Fig. 18d is immediately close to the base material, and has a basal texture in which the basal planes are parallel to the plate surface. Therefore, from the weld center to the base material, the basal planes aligned with the pin surface, experiences two rotations around the ND and WD in sequence, to reach the orientation aligned with the plate surface. When a tensile load is applied in the transverse direction (TD), the average Schmid factors for basal slip are 0.13, 0.31, 0.41 and 0.23 in the four characterized regions in sequence from the weld center towards the base material on the retreating side. According to the Schmid’s law, the yield stress  y at which dislocation slip is initially activated in a crystalline material is: CRSS (2) y = m where  CRSS is the critical resolved shear stress for a slip system, and m is an average Schmid factor. Therefore, plastic deformation through basal slip in a magnesium weld is initiated in the location which has a large average m value, that is, within the stir zone but close to its extremity. In addition, the texture with the basal pole tilted in the tensile direction (i.e., the TD in the present  study) enhances the occurrence of mechanical twinning  (mainly 1 0 1¯ 2 tensile twinning) and disfavors the activation of non-basal slips [24,25]. The above analysis on the effect of texture

can be used to explain the response of a friction stir lap linear weld subject to the loading mode II which has the failure mode FM3. Figs. 13 and 14 show the failure of a lap weld on the retreating side. Fracture occurs within the stir zone but close to its extremity as analyzed above. The distribution of mechanical twins on the sides of the fracture surface indicates that plastic deformation starts from a location of a large average Schmid factor and then concentrates at this location. In magnesium, mechanical twins are barriers to dislocation motion, and the induced stress concentration is generally unlikely to be relaxed by other slip, kinking or twinning at an ambient temperature. Therefore, the localized deformation accelerates the failure of the friction stir processed material with a relatively low strength. Different pin geometries are supposed to result in somewhat different texture distributions in the weld zone, and therefore different fracture locations (Fig. 13). The friction stir processed material in a weld produced with the triangular tool has a higher strength but could have a larger grain size than the processed material in a weld produced with the cylindrical tool (Figs. 11, 15 and 16). This indicates that the grain size effect is not as important as the texture effect in magnesium when different grain sizes are in a comparable range. 6. Summary Friction stir lap linear welding was conducted on overlapped AZ31 magnesium plates mainly with the orientation such that the weld retreating side on the upper plate was placed under load. The effects of tool geometry and process condition on lap shear strength of magnesium friction stir lap linear welds were addressed. Weld macrostructure and microstructure, and strength of friction stir processed material were quantitatively characterized. The following important results were obtained. (1) Higher lap shear weld strengths are obtained when the retreating side of the upper plate is loaded (i.e., loading mode II) than when the advancing side of the upper plate is loaded (i.e., loading mode I). (2) The hook on the advancing side of a weld is directed upwards towards the weld surface and finally arrested at the stir zone

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(3)

(4)

(5)

(6)

(7)

(8) (9)

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extremity. The hook on the retreating side first runs gradually upwards. It is then deflected downwards into the stir zone, and even goes further towards the advancing side. The hook within the stir zone becomes indistinct. The triangular pin tool effectively suppresses the hook and decreases the hook height on the retreating side, as compared to the cylindrical pin tool. Decreasing the hook height leads to an increase in the load-carrying thickness in a weld. The hook suppression by the triangular pin is attributed to enhanced horizontal material flow on the retreating side, which reduces or even arrests the upward flow of the plasticized material caused by tool translation. For the cylindrical tool, a higher rotation speed and/or a lower travel speed increase the hook height on the retreating side of a weld. Increasing the tool shoulder size apparently increases the hook height at medium travel speeds, and this increase becomes more significant at higher rotation speeds. For the triangular tool, the hook height decreases as the travel speed increases, and the decrease becomes less prominent when the travel speed is higher than 200 mm/min. The hook height does not show a significant variation with the tool rotation speed. Increasing the shoulder size decreases the hook height at a lower travel speed, making the hook formation less sensitive to the process condition. Five failure modes are observed when a friction stir lap linear weld is loaded in lap shear. Among them, the failure mode FM3, by means of which the weld fractures within the stir zone but close to the stir zone extremity on the retreating side of the upper plate, is dominant. The presence of a pore in the stir zone does not influence the weld strength when the failure mode FM3 is operative. For the failure mode FM3, weld strength is determined by multiplication of the load-carrying thickness and the strength of friction stir processed material on the retreating side. The triangular pin tool increases the optimized lap shear weld strength by 78%, compared to the cylindrical pin tool. Although the triangular pin increases the strength of friction stir processed material, the weld strength is improved primarily through increasing the load-carrying thickness. The presence of a ‘pure’ shear surface on the tool pin significantly reduces weld strength. Deformation localization, indicated by the distribution of mechanical twins near the fracture surface, causes lap welds to fail in the failure mode FM3. Texture analysis presented justifies this observation.

Acknowledgements This material is based upon work supported by the Department of Energy National Energy Technology Laboratory under Award Number No. DE-FC26-02OR22910.

This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof. Such support does not constitute an endorsement by the Department of Energy of the work or the views expressed herein. The authors would like to thank Alan A. Luo and James F. Quinn (from General Motors Corporation) for their comments. The authors would also like to express gratitude to Dr. Sergey Mironov and Dr. Yutaka S. Sato (from Tohoku University, Japan) for their technical assistance. References [1] W.M. Thomas, E.D. Nicholas, J.C. Needham, M.G. Murch, P. Templesmith, C.J. Dawes, G.B. Patent Application No. 9125978.8, Dec. 1991. [2] R. Nandan, G.G. Roy, T. Debroy, Metall. Mater. Trans. A 37A (2006) 1247–1259. [3] J.Q. Su, T.W. Nelson, R. Mishra, M. Mahoney, Acta Mater. 51 (2003) 713–729. [4] P. Wanjara, M. Jahazi, Metall. Mater. Trans. A 36A (2005) 2149–2164. [5] S.H.C. Park, Y.S. Sato, H. Kokawa, K. Okamoto, S. Hirano, M. Inagaki, Scripta Mater. 49 (2003) 1175–1180. [6] W. Woo, H. Choo, M.B. Prime, Z. Feng, B. Clausen, Acta Mater. 56 (2008) 1701–1711. [7] G.M. Xie, Z.Y. Ma, L. Geng, Scripta Mater. 57 (2007) 73–76. [8] L. Dubourg, A. Merati, M. Jahazi, Mater. Des. 31 (2010) 3324–3330. [9] A. Abdollah-Zadeh, T. Saeid, B. Sazgari, J. Alloy Compd. 460 (2008) 535–538. [10] Y.C Chen, K. Nakata, Scripta Mater. 58 (2008) 433–436. [11] L. Cederqvist, A.P. Reynolds, Weld. J. 80 (2001) 281–287. [12] W.M. Thomas, K.I. Johnson, C.S. Wiesner, Adv. Eng. Mater. 5 (2003) 485–490. [13] Y.S. Sato, H. Kokawa, K. Ikeda, M. Enomoto, S. Jogan, T. Hashimoto, Metall. Mater. Trans. A 32A (2001) 941–948. [14] S.H.C. Park, Y.S. Sato, H. Kokawa, Metall. Mater. Trans. A 34A (2003) 987–994. [15] K. Colligan, Weld. J. 78 (1999) 229s–237s. [16] B. London, M.W. Mahoney, W. Bingel, M. Calabrese, R.H. Bossi, D. Waldron, in: K.V. Jata, M.W. Mahoney, R.S. Mishra, S.L. Semiatin, T. Lienert (Eds.), Friction Stir Welding and Processing II, TMS, 2003, p. 3. [17] H.N.B. Schmidt, T.L. Dickerson, J.H. Hattel, Acta Mater. 54 (2006) 1199–1209. [18] M. Guerra, C. Schmidt, J.C. McClure, L.E. Murr, A.C. Nunes, Mater. Charact. 49 (2002) 95–101. [19] K. Kumar, S.V. Kailas, Mater. Sci. Eng. A 485 (2008) 367–374. [20] W.J. Arbegast, in: Z. Jin (Ed.), Hot Deformation of Aluminum Alloys III, TMS, 2003. [21] Y.J. Chao, J. Eng. Mater. Technol. 125 (2003) 125–132. [22] Q. Yang, A.K. Ghosh, Acta Mater. 54 (2006) 5170–5179. [23] S.R. Agnew, J.A. Horton, T.M. Lillo, D.W. Brown, Scripta Mater. 50 (2004) 377–381. [24] J. Bohlen, M.R. Nurnberg, J.W. Senn, D. Letzig, S.R. Agnew, Acta Mater. 55 (2007) 2101–2112. [25] X. Huang, K. Suzuki, A. Watazu, I. Shigematsu, N. Saito, Mater. Sci. Eng. A 488 (2008) 214–220.