Dissimilar Friction Stir Welds in AA5083-AA6082. Part I: Process Parameter Effects on Thermal History and Weld Properties M.J. PEEL, A. STEUWER, P.J. WITHERS, T. DICKERSON, Q. SHI, and H. SHERCLIFF The aim of this study was to explore the so-called processing window, within which good-quality welds can be produced, for the friction stir welding of AA5083 to AA6082. To that end a systematic set of nine instrumented welds were made using rotation speeds of 280, 560, and 840 rpm and traverse speeds of 100, 200, and 300 mm/min with AA5083 on the advancing side and another nine with the materials reversed. For comparison a smaller series of AA5083–AA5083 and AA6082– AA6082 welds were also made. Thermocouple measurements, tool torque, extent of material mixing, and macrostructural observations all indicate that the temperature under the tool is more strongly dependent on the rotation than the traverse speed. It was found that in the current case, the power (energy/s) and heat input (energy/mm) do not correlate simply with the weld temperature. As a result, such metrics may not be suitable for characterizing the conditions under which welds are produced.
I. INTRODUCTION
FRICTION stir (FS) welding is a relatively new joining technique developed by TWI, Cambridge, in 1991.[1] The technique has been optimized for aluminum alloys, although it has also been applied to the joining of magnesium,[2] titanium,[3] and steel.[4,5] The process takes place in the solid state and appears to offer a number of advantages over conventional fusion welding techniques, such as no need for expensive consumables such as filler wire and gas shields, ease of automation on simple milling machinery, good mechanical properties of the resultant joint, and low distortion.[6,7] In addition, since welding occurs by the deformation of material at temperatures below the melting temperature,[8,9] it is possible to avoid problems commonly associated with the joining of dissimilar aluminum alloys.[10] Successful trials have been performed on the joining of various aluminum alloy series.[11–14] However, few systematic studies have been performed on dissimilar welds, and the relationships between the various weld parameters and the resulting weld properties have not been identified. The aim of this study was to investigate the extent of the so-called processing window—the range of FS welding speeds (rotation and traverse) within which good-quality welds could be produced between the dissimilar alloys AA5083 and AA6082. The approach has been to produce two sets of welds based around identical combinations of rotation and traverse speeds, with the relative positions of the alloys reversed. Furthermore, a limited number of similar material (AA5083–AA5083 and AA6082–AA6082) welds have been produced for comparison. The welds were instrumented to measure the forces and torque acting on the tool as well as the temperature within the backing plate M.J. PEEL, Post-Doctoral Fellow, A. STEUWER, Post-Doctoral Fellow, and P.J. WITHERS, Professor, are with the Materials Science Centre, Manchester University, Manchester, U.K. M.J. PEEL and A. STEUWER, Post-Doctoral Fellow, are with FaME38 at the ESRF-ILL, Grenoble, France. Contact e-mail:
[email protected] T. DICKERSON, Post-Doctoral Fellow, Q. SHI, Post-Doctoral Fellow, and H. SHERCLIFF, Senior Lecturer, are with the Engineering Department, Cambridge University, Cambridge, U.K. Manuscript submitted September 27, 2005. METALLURGICAL AND MATERIALS TRANSACTIONS A
during welding. Several welds were fitted with thermocouples to measure the thermal profile at specific points within the plate. Subsequently, the processing window has been evaluated in terms of (a) the forces acting on the tool, (b) the macroscopic material flow, with emphasis on the absence of defects, (c) the resulting microstructure, and subsequently (d) the mechanical properties of the welds. Simple models have been developed to aid the evaluation of these properties against the rotation and traverse speed variations. The first is a numeric thermal model that has been calibrated against experimental data and provides estimations of thermal data throughout the plate over the weld cycle. In addition, a hardening model has been produced for each of the alloys. The hardening models have been calibrated using material subjected to isothermal holds. In this article, part I, we focus on the results from the instrumented welding trials and the thermal predictions derived from the data. This is because previous work has shown that it is the thermal history rather than the deformation that plays the dominant role in determining subsequent weld properties. This data are evaluated against the macroscopic material flow in the welds. In part II,[15] we will present a detailed analysis of the microstructure and hardness distributions across the welds and compare the results to the predictions of our hardness models. In the future these data will be used to rationalize the accumulation of residual stresses in the welds. II.
EXPERIMENTAL PROCEDURE
Two materials form the basis of this study. The first material is the commercial alloy FORMALL* 545 supplied *FORMALL is a trademark of Alusuisse Group Ltd., Chippis, Switzerland.
by EADS. This alloy has a composition equivalent to AA5083 (Table I) and relies primarily on a high solute concentration and associated work hardening for its strength. The plates were originally in a heavily cold-rolled VOLUME 37A, JULY 2006—2183
Table I. Nominal Concentration (Weight Percent) of the Primary Non-impurity Elements and the Mechanical Properties of AA5083 and AA6082 in the Initial Condition
AA5083 AA6082
Mg
Mn
Si
0.2 Pct Yield (MPa)
UTS (MPa)
Reduction in Area (Pct)
Hardness (HV 1)
4.0 to 4.9 0.6 to 1.2
0.40 to 1.0 0.4 to 1.0
0.40 0.7 to 1.3
392 6 4.3 301 6 0.96
457 6 2.3 337 6 1.2
7.3 6 0.6 31 6 2.7
130 105
Table II.
Designations, Materials, and Tool Speeds Used to Produce the Dissimilar Welds
Material Advancing
Retreating
Rotation Speed (rpm)
Traverse Speed (mm/min)
Weld Pitch (mm/rev)
M1 M2 M3 M4 M5 M6 M7 M8 M9 M11 M12 M13 M14 M15 M16 M17
AA5083 AA5083 AA5083 AA5083 AA5083 AA5083 AA5083 AA5083 AA5083 AA6082 AA6082 AA6082 AA6082 AA6082 AA6082 AA6082
AA6082 AA6082 AA6082 AA6082 AA6082 AA6082 AA6082 AA6082 AA6082 AA5083 AA5083 AA5083 AA5083 AA5083 AA5083 AA5083
280 560 840 280 560 840 280 560 840 280 560 840 280 560 840 280
100 100 100 200 200 200 300 300 300 100 100 100 200 200 200 300
0.36 0.18 0.12 0.71 0.36 0.24 1.07 0.54 0.36 0.36 0.18 0.12 0.71 0.36 0.24 1.07
M18 M19
AA6082 AA6082
AA5083 AA5083
560 840
300 300
0.54 0.36
Designation
condition with a high hardness (130 HV). The second material was the age-hardening alloy AA6082. This alloy has the nominal composition outlined in Table I and was provided in the T6 (peak aged) condition. Both alloys were provided in the form of rolled sheets of 3 mm thickness. The welding trials were undertaken at the German Aerospace Center (DLR, Cologne, Germany) on a modified milling machine of the cantilever variety. The welds were produced by butt-welding plates 150 mm long and 60 mm wide. These plates were clamped to a steel backing plate that was itself attached to the traverse table of the FS welding machine. The total weld length was around 105 mm from pin entry to pin exit, with the pin inserted 15 mm from the leading plate edge and extracted at 120 mm. The tool was manufactured from hardened tool steel with a nominal shoulder diameter of 18 mm, although the effective diameter increases slightly as the tool settles further into the weldment due to the chamfered contour of the outer edge. The internal face of the shoulder is not flat but is inclined by 10 deg to create a conical surface. The pin was originally a thread-burnishing tool made of high-speed steel with a titanium nitride coating. The pin diameter was 6 mm in diameter and had a metric thread with a 0.8-mm pitch. The weld direction was parallel to the rolling direction of the plates. In all cases the tool was tilted at 2 deg from the plate normal, such that the rear of the tool was lower than the front and had a shoulder plunge depth of 0.55 mm below the plate surface. During welding the tool was driven 2184—VOLUME 37A, JULY 2006
Comments Thermocoupled Thermocoupled
Thermocoupled Thermocoupled Thermocoupled Thermocoupled, damaged Contains voids Thermocoupled, contains voids Thermocoupled
through a Kistler 9123C rotating dynamometer that allowed the measurement of the reaction torque (M) and the (x,y,z) tool forces. A thermocouple was placed within the backing plate with the junction 1 mm below the anvil surface, on the centerline of the welds and 75 mm from the pin start position. The thermocouple was welded to a metal shim that was cut and shaped to fit tightly into a hole drilled from the backside of the anvil. Aluminum powder-filled adhesive was used to ensure good thermal contact. In certain welds (Table II) K-type thermocouples were used to measure the variation in temperature throughout the weld cycle. They were tapped into small holes (;0.8 mm deep and slightly narrower than the thermocouple width) in the upper surface of the sheets using a hammer and punch and were then peened into place to prevent movement and ensure good thermal contact. Thermocouples were placed on both the advancing and retreating side at a distance of 15 and 30 mm from the original join line. At these distances the thermocouples would not be damaged or moved by the passage of the tool. The data from the thermocouples and dynamometer were fed into a data logger sampling at 10 Hz; sampling began as soon as a 1 kN down force was detected. In this study only the tool rotation and traverse speeds were varied. Welding was performed at three different tool rotation rates (280, 560, or 840 rpm) and three traverse speeds (100, 200, or 300 mm/min), thereby creating a 3 3 3 matrix of welds for each material position. Those welds with AA5083 on the advancing side were labeled M1-9; METALLURGICAL AND MATERIALS TRANSACTIONS A
those with AA6082 on the advancing side were labeled M11-19 (Table II). The advantage of this matrix of weld parameters is that it allows the influence of each speed to be investigated independently and, in addition, permits an analysis of welds produced with the same ratio of traverse and rotation speed but different speeds. This ratio is often referred to as the weld pitch and, in this case, has units of millimeters traveled per revolution of the tool (mm/rev). Operational restraints prevented the completion of a full matrix of trials for the similar material welds. These were restricted to the three traverse rates at a constant rotation speed of 560 rpm. The power (i.e., energy per unit time) generated by the tool is defined as P 5 M.v, where M 5 the torque and v 5 the angular velocity of the tool. The additional energy input produced by the forward motion of the tool (forward force 3 forward velocity) was found to be less than 0.1 pct of the total power and so was neglected in the analysis. The heat input (energy per unit length of weld) is then given as Q 5 P/v, where v 5 the traverse speed of the tool. III.
Fig. 1—The pseudo-steady state down force during the welding of the similar and dissimilar material welds as a function of weld pitch.
RESULTS
A. Forces, Temperature, and Energy Input as a Function of Weld Pitch The following data are reported as average values measured when the tool was between 50 and 70 mm from the weld start position. This regime can be reasonably approximated to steady-state conditions and corresponds to a region close to the thermocouples. There was very little difference in the data obtained from the welds produced with identical tool rotation and traverse speeds but reversed material locations. Hence, the data presented below are an average of the two welds produced at each combination of speeds. Figure 1 shows the downward force exerted by the tool on the workpiece as a function of weld pitch for each material combination. For the dissimilar welds, it is difficult to distinguish the results by rotation or traverse speed. Increasing the traverse speed or decreasing the rotation speed will increase the down force for a given weld pitch, but the effect is slight, particularly at higher weld pitches. Hence, a general trend has been drawn that clearly indicates a rise in force with increasing weld pitch. This tends toward a limiting value of around 20 kN within the current range of welding parameters. The material combination clearly has a more significant effect, with the down force being around 2 kN greater for the AA5083 welds compared to the dissimilar welds, while the force is lower for the AA6082 welds by a similar amount. The decreased downward force for the AA6082 was observed to correlate with an increased tool depth below the plate surface that led to a wider tool track on the surface, increased flash, and a greater level of thickness undermatching. The torque (M) and power (P) are plotted in Figure 2 as a function of the weld pitch. While the torque acting on the tool is influenced by both tool speeds, the rotation speed has the greater impact. On average, increasing the traverse speed from 100 to 300 mm/min resulted in a ;5 Nm increase in the torque. In comparison, increasing the rotaMETALLURGICAL AND MATERIALS TRANSACTIONS A
Fig. 2—Plots of (a) the reaction torque acting on the tool and (b) the derived power input during the pseudo-steady state period as a function of weld pitch.
tion speed by the same factor (280 to 840 rpm) caused the torque to reduce by ;40 Nm. As a result, the torque varied considerably for the three different parameter combinations corresponding to the same weld pitch. The strong influence of the rotation speed is in agreement with Johnson and Seidel et al. but contradicts the work of Colegrove et al. and Lindner et al.[16–19] The power input appears to increase as both the traverse speed and the rotation speeds are increased. The lowest power input was seen at 280 rpm and 100 mm/min, the highest at 840 rpm and 300 mm/min. The difference between the extremes is around 800 W, or 30 pct of the maximum power input. It is noteworthy that these VOLUME 37A, JULY 2006—2185
welds have the same weld pitch. Figure 2 shows some similarities with the variation in power predicted by Seidel and Reynolds,[17] although there is a greater dependency on the traverse speed in our case, particularly at low rotation speeds. Figure 3(a) shows the variation in the heat input (Q) as a function of weld pitch for the current welds. The heat input is heavily influenced by the traverse speed, but less so by the rotation speed. At 300 mm/min the heat input is 0.55 to 0.6 kJ/mm (increasing with rotation speed), while at 100 mm/min the heat input is 1.3 to 1.55 kJ/mm. A similar dependency can be seen for the temperature measurements in the backing plate (Figure 3(b)), and there is an apparent correlation between the two factors. B. Thermal Model Experimental restrictions on the locations and number of thermocouples that could be placed in the plates meant that to evaluate the effect of parameter changes on the temperature close to the tool, it was necessary to use a thermal model of the welding process. Our model was developed by Shi et al. as part of the JOIN-DMC project.[20] The model has been implemented in ABAQUS/Standard. This approach means that material flow has been neglected, as such extensive material movements would challenge ABAQUS. Besides, it is believed that the thermal excursion is much more important in terms of as-welded residual stress and hardness than the mechanical deformation. There are four main regions that may be included in a thermal model of
FSW: the weldment (which, in dissimilar welds, may be further split into parts for each material), the tool, the backing plate, and the clamps and other supporting structures. In this model only the weldment and the backing plate have been modeled explicitly. The tool is modeled simply as two heat sources representing the shoulder and pin, while the clamps have been excluded entirely. Neglect of heat flow into the clamps was considered acceptable since their location and small area of contact suggests that this is unlikely to be a significant factor. The entire weldment is modeled as a single body that is partitioned down the original joint so that different material properties can be allocated to each side. Heat is free to flow across this interface, even in front of the tool, where in practice the thermal contact may be relatively poor. In any event, the extent of lateral heat flow across the weld line is not expected to be significant, since the temperature gradients in front of the tool are predominantly in the longitudinal direction. The mixing of material around the tool is also neglected. The weldment is represented by a threedimensional mesh composed of brick elements of type DC3D8. In the region around the weld line it is necessary to use a fine element size to capture the large thermal gradients around the tool. The computational requirements are minimized by increasing the element size further from the weld line, where the thermal gradients are lower. The backing plate is modeled by four-noded rectangular shell elements of type DS4 with five integration layers through the 25-mm thickness. The use of shell elements places a lower demand on computer resources than brick-type elements. The distribution of heat flux during welding is determined by an implementation of the ABAQUS usersubroutine, DFLUX. This represents the heating action of the tool in two sections: a circular surface interaction representing the shoulder and a volume interaction equal in size to the swept area of the pin. The cylindrical volume does not include any threads or other features. Since the pin is modeled simply as a heat source rather than as a physical entity, the material within this volume is allocated the thermal properties of aluminum rather than steel. During each time increment, the subroutine determines whether any given node is in contact with the tool and applies an appropriate heat flux. The magnitude of this flux depends on the location of the node with respect to the tool, primarily whether the node lies in the pin region or under the shoulder. Under the shoulder the heat flux at a given radial distance from the center of the tool (r) is described by: qs ðrÞ 5
ðm 1 1Þ Ps :r m :f 2p ðRs2 1 m " Rp2 1 m Þ
for Rp # r # Rs and z 5 0
Fig. 3—Plots of (a) the heat input per mm of travel (power/welding speed) and (b) the temperature at the location of the backing plate thermocouple as a function of weld pitch. 2186—VOLUME 37A, JULY 2006
[1]
where Ps 5 the fraction of the total power contributed by the shoulder, z 5 the depth below the surface of the weld, Rs and Rp 5 the shoulder and pin radius, f 5 the efficiency factor (see below), and m is a constant. By changing the value of this constant, the heat flux can be uniform under the shoulder (m 5 0), increase linearly with radius (m 5 1), or show a stronger dependence (m . 1). When m 5 1, the equation reduces to the heat flux equation described by METALLURGICAL AND MATERIALS TRANSACTIONS A
Chao and Qi.[21] Increasing the value of m will result in a greater portion of the heat flux being allocated to the outer edge of the shoulder. Thus, the exponent can be used as a fitting parameter to reduce the temperature on the weld line and even out the temperature under the shoulder. The physical basis for this parameter is difficult to quantify but essentially allows for a nonuniform distribution of pressure and friction coefficient (for sliding friction) or flow stress (for sticking friction) around the tool. The heat flux at the surface of the pin is given by: Pp :f for r , Rp and 0 # z # t [2] qp ðrÞ 5 pR2p :t where t 5 the thickness of the weld plate (and the length of the pin), Pp 5 the fraction of total power provided by the pin, and f 5 the efficiency factor. In this case the heat flux is uniformly distributed. The small radius of the pin, compared to the shoulder, and material mixing around the pin, which should act to even out temperature variations in this area,[18] suggest that this assumption is unlikely to lead to a significant error. There are three major sources of heat loss from the weldment: conduction into the tool, convection from the regions of the welded plate exposed to the surrounding atmosphere, and conduction into the backing plate. To account for convection, all of the surfaces exposed to the atmosphere were allocated a uniform convection coefficient of 15 W/m2K.[20] The conduction of heat into the tool has been accounted for by reducing the total energy input into the weld by some efficiency factor f. Previous work suggests that 5 to 20 pct of the total heat may conduct into the tool during welding, depending on factors such as the weld length and tool shape/material.[9,22,23,24] However, the modification of the heat input also allows us to account for other sources of energy loss such as vibration, excess flash production, or the storage of energy in the deformed material. Hence the fraction of heat input into the model can be used as a fitting parameter. The level of conduction into the backing plate is uncertain, as the large downward pressure below the tool will increase the actual area of contact at the interface and so increase the local rate of heat transfer. The high temper-
ature under the tool will also increase the rate of heat transfer, as the corresponding decrease in the yield stress will allow greater conformity between the weldment and the backing plate. Since the region of high temperature coincides with the region of high pressure, the variation in conductivity has been simulated in an approximate manner by using a temperature-dependent conductivity (Figure 4). The values used by Shi et al. were initially based on published data for heat transfer between aluminum and steel[25] and were modified on an iterative basis to achieve a good fit with experimental thermocouple data.[20] Further modification of the conductivity was required to achieve a good fit with the measured thermal data from this project. Temperature-dependent thermal properties for the two alloys have been taken from the literature (Figure 5). The thermal conductivity and specific heat capacity were taken from ‘‘Recommended values of thermo-physical properties for selected commercial alloys.’’[26] The data for AA6082
Fig. 4—The thermal conductance between the weldment and the backing plate used in the current study compared to those used by Shi et al.[20] These values were determined iteratively to obtain best fits of the model to the thermocouple data.
Fig. 5—(a) The conductivity and (b) the heat capacity of the two aluminum alloys and the tool steel as a function of temperature.[26,27,28] METALLURGICAL AND MATERIALS TRANSACTIONS A
VOLUME 37A, JULY 2006—2187
actually refer to AA6061, a similar alloy with a slightly lower silicon and manganese content. The AA5083 data refer to AA5182-O (fully annealed), which has a slightly lower manganese content. The density and latent heat for both alloys were taken from the literature, as were the property data for the steel tool.[27,28,29] Three welds were chosen for data fitting and a fourth weld was used as a test to assess the predictive capability using the heat input and loss parameters determined from the other three welds. These welds represent the greatest spread of tool speeds used in this project. In each case the weld was produced with AA5083 on the advancing side. The fractional heat input and conductivity were altered to achieve the best fit across all three thermocoupled welds; they were not varied individually for each weld. Figure 6 compares the predicted and measured thermal profiles for welds M3 and M9, which were produced with the same rotation speed (840 rpm) and different traverse
Fig. 6—The correlation between the predicted (dotted) and measured (solid) thermal histories for the thermocouple 15 mm from the weld line on the advancing side of the weld for dissimilar welds M3 (100 mm/min) and M9 (300 mm/min), both of which were produced using a rotation speed of 840 rpm.
speeds (100 and 300 mm/min, respectively). These profiles correspond to the thermocouple placed 15 mm from the weld line on the advancing side of the weld (AA5083). In the model, the predicted temperature was recorded by nodes on the upper surface of the plate and at the midthickness of the plate. A comparison of data from the nodes at the upper surface and at the midthickness of the welds, at a distance of 15 mm from the weld line, indicated that the predicted temperature was around 40 °C lower at the surface due to the convection of heat. Since the thermocouples are embedded in the plate to a depth of ;0.8 mm (the center of the thermocouple will be ;0.55 mm from the upper surface), the recorded temperature will be intermediate between the values reported for the surface and midthickness. The predicted temperature of the thermocouple was therefore estimated by linear interpolation between the surface and midthickness values. This resulted in a comparatively minor decrease of ;20 °C in the thermocouple temperature compared to the midthickness temperature. The interpolation was applied to all of the welds. The match is generally good, which is not surprising since the model was fitted to the experimental data, although the predicted peak temperature is slightly too low and the real welds cool slightly more slowly than the modeled ones. A thermal history according to the position of the backing plate thermocouple was estimated by linear interpolation between FE data points, as shown in Figure 7(a). Since the thermal field is not steady state in the backing plate, this method will probably overestimate the temperature at the thermocouple position. An example of the predicted and measured temperatures in the backing plate is shown in Figure 7(b). The predicted peak temperature is noticeably overestimated and the rate of cooling is higher in the model. This could indicate that the estimated thermal conductivity into the backing plate (Figure 4) is too high at high weld temperatures and too low at lower temperatures. Figure 8 shows predictions of the peak temperatures reached during welding as a function of distance from the weld line for each of the welds. These profiles correspond to the temperature at the midthickness and the midline of the plate (i.e., after the tool has traveled 50 pct of the total distance) and should represent approximately steady-state
Fig. 7—(a) The predicted anvil temperature as a function of distance below the upper surface, illustrating the linear interpolation of the data to estimate the temperature 1 mm below the surface, and (b) the predicted and measured temperatures in the backing plate 1 mm below the surface. In both graphs the data correspond to weld M9 (840 rpm and 300 mm/min, AA5083 advancing side). 2188—VOLUME 37A, JULY 2006
METALLURGICAL AND MATERIALS TRANSACTIONS A
Fig. 8—The predicted and measured peak temperature at the midlength of the weld as a function of distance from the weld line. For clarity, only the advancing side of the weld is shown.
conditions. It has been found that the thermal field is essentially symmetrical around the weld line, indicating that the small difference in thermal properties between the two alloys does not have a significant effect of the thermal profile. As a result, only the advancing side of the weld has been shown to enhance the clarity of the plot. The temperature under the tool varies between 425 °C for the coolest weld (M7, highest weld pitch) and 540 °C for the hottest weld (M3, lowest weld pitch). Although the other two welds have the same weld pitch, the peak temperature under the tool is significantly different. The weld produced with the higher rotation speed (M9) has a significantly higher temperature than the low-rpm weld (M1) and is only slightly cooler than the hottest weld (which was produced using the same rotation speed). The same situation is observed for the two welds produced at 280 rpm, where the change in traverse rate has a comparatively small effect on the peak temperature. This result is consistent with the relative insensitivity of the steady-state torque to the traverse speed. Although the temperature under the tool is dominated by the rotation speed, the temperature beyond the shoulder radius is more strongly affected by the rate of traverse. At a higher traverse speed, the temperature decreases more rapidly with distance from the tool. For this reason, the two welds produced with the same weld pitch have the same peak temperature at around 30 mm from the weld line. At greater distances from the weld line, the high rotation/ traverse speed weld is actually slightly cooler than the low rotation/traverse weld, despite the higher predicted temperature at the weld line. Hence, thermocouple measurements in the mid- to far field may not accurately reflect the actual peak temperature around the tool during the weld cycle. C. Estimation of the Interfacial Condition Schmidt et al. have proposed that the two basic models for the interfacial contact shear stress between the tool and the work material—sliding friction and sticking friction— represent two extremes and that the real situation can be METALLURGICAL AND MATERIALS TRANSACTIONS A
Fig. 9—Plots of (a) the interfacial shear stress and (b) the friction coefficient as a function of the weld pitch during steady-state conditions.
described by a mixture of the two models.[24] Sliding friction uses a friction coefficient and predicts that the interfacial stress will be proportional to the down force. In sticking friction, there is no relative movement between the tool and weld material and heat generation is through the mechanical working of the material; hence, the contact stress is determined by the flow stress of the material. Schmidt et al. have presented an analytic estimate of the interfacial shear stress t contact or the friction coefficient m at the surface of an idealized tool:[24] 3M [3] t contact ¼ 3 3 2pððRs " Rp Þð1 1 tan aÞ 1 R3p 1 R2p H m¼
2FððR3s
"
3MR2s tan aÞ 1 R3p 1 R2p H
R3p Þð1 1
[4]
where M 5 the total torque, F 5 the down force, Rs 5 the shoulder radius, Rp 5 the pin radius, a 5 the shoulder cone angle (the tool shoulder is slightly concave), and H 5 the probe height. The calculated interfacial shear stress and friction coefficient are shown in Figure 9. The estimated shear stress is in the range 15 MPa (840 rpm and 100 mm/min) to 38.5 MPa (280 rpm and 300 mm/min). The flow stress of the two alloys at a single strain rate of 0.1 s"1 has been summarized in Figure 10. This strain rate was chosen because it had the greatest overlap between the sources.[30–33] Examination of Figure 9 indicates that between 400 °C and 500 °C, AA5083 has a flow stress of 30 to 80 MPa. VOLUME 37A, JULY 2006—2189
Fig. 10—A comparison of the flow stress of AA5083 and AA6082 at constant strain rate (data from[30–33]).
pffiffiffi Using a von Mises’ yield criterion (t y 5 sy / 3 ) results in a shear yield stress in the range of 17 to 46 MPa. At the same temperatures and strain rate, A6082 has a shear yield stress of around 10 to 32 MPa. Although the accuracy of these estimates is debatable, the estimated interfacial shear stress is consistent with the flow stress of the materials within the temperature range predicted by the thermal model. The estimated friction coefficient lies in the range 0.25 to 0.57, with the highest values occurring for the lowest rotation speed. These values lie within the range commonly expected for hot metal working processes.[34] At low rpm the friction coefficient appears to decrease with increasing weld pitch, and hence increasing downward force (Figure 1). This could occur if the friction was predominantly sticky in character, since the contact shear stress is then independent of pressure and the apparent friction coefficient will reduce if sliding friction is inappropriately assumed.[35] The similar material welds have rather different friction coefficients to the dissimilar welds, with the AA5083 and AA6082 welds having lower and higher friction coefficients, respectively. The difference in the calculated friction coefficient is a direct result of the fact that the down force, and hence pressure, changes between the two materials while the torque is essentially constant for a given set of welding speeds. D. Macrostructure and Cross-Weld Tensile Properties Although the details of the variation in the microstructure and mechanical properties of the welds will be discussed in detail in part II of this paper,[15] it is useful to examine the macrostructures here for comparison with the dynamometer/thermocouple data and the thermal model predictions. Figures 11 and 12 contain a matrix of macrographs of the stir zone in the dissimilar welds produced under each of the different conditions, with AA5083 on the advancing and retreating side, respectively. Changing the traverse and rotation speeds can have a significant effect on the flow of material within the stir zone. Generally speaking, the 2190—VOLUME 37A, JULY 2006
extent of mixing and interface disruption increases as the rotation speed is increased or the traverse rate is decreased. In this regard the rotation speed appears to have a significantly greater impact than the traverse speed. As a result, the welds produced with an identical weld pitch (M1/11, M5/15, and M9/19) do not exhibit the same level of material mixing. For a given combination of weld parameters, the welds produced with AA6082 on the advancing side exhibit a significantly lower level of mixing in the stir zone than those with the materials reversed. This is in contrast to the results of Larsson et al.[12] but corroborates those obtained by Tanaka,[13] both of whom joined AA5083 to AA6082 using FS welding. The reduction in mixing results in the formation of defects (voids/tunnels) in welds M14 and M17, which do not appear in welds M4 and M7, despite those welds being produced at the same tool speeds. The uniaxial tensile properties of the FS welds have been determined by cross-weld tensile testing. The main aim of these tests was to determine whether the relatively low level of mixing in some of the dissimilar welds has any effect on the failure stress or failure location in the welds. Such variations have been previously observed for AA5083 similar welds.[36] The ultimate tensile strength (UTS) and elongation to failure of the cross-weld tensile test specimens is shown in Table III. During tensile testing the deformation was concentrated in the heat-affected zone in the AA6082, and in most cases failure was confined to this region. The exceptions were welds M14 and M17, which both failed at the weld line. As noted above, these welds contained defects at the weld line, and these appear to have determined the failure location. However, this did not reduce the UTS or elongation to failure, at least for M14, in comparison to the other welds. Unfortunately, tensile test data are not available for welds M11, M14, and M17 as a result of a software failure after testing. The UTS varies between 200 to 240 MPa, which is rather less than that in unwelded AA6082-T6 (;330 MPa) and in good agreement with previous data for the tensile testing of AA6082-T6 similar material FSr welds.[37] The UTS is also in good agreement with values for the gas tungsten arc welding of AA6061-T4 to AA5083-O, where failure was also predominantly in the 6xxx series HAZ.[10] An examination of the data indicates that there is a slight but consistent increase in the UTS (200 to 240 MPa) and elongation to failure (2.3 to 3 mm) when both the rotation and traverse speeds are increased. Clearly the lack of a statistically relevant number of tensile tests limits the interpretation of the data, but the increase is consistent and there is a good agreement between the welds produced with the same tool speeds but with the material positions reversed. This differs from the results by Reynolds et al., who found that the mechanical properties of AA6082 FS welds were somewhat insensitive to changes in the welding speed, but since the tool speeds were not reported it is not possible to make a meaningful comparison with the current study.[8] IV. DISCUSSION AND SUMMARY The aim of this paper has been to investigate the effect of changing the rotational and traverse welding speeds on the tool forces, power input, and thermal history throughout the welding cycle. This has been achieved by investigating a METALLURGICAL AND MATERIALS TRANSACTIONS A
Fig. 11—Macrographs showing the stir zone/thermo-mechanically affected zone (TMAZ) in the dissimilar welds M1–M9. AA5083 is on the advancing side (right-hand side) and appears darker. The dotted line on weld M5 shows the approximate size and position of the pin (6-mm diameter).
Fig. 12—Macrographs showing the stir zone/TMAZ in the dissimilar welds M11–M19. AA5083 is on the retreating side (left-hand side) and appears darker. The dotted line on weld M15 shows the approximate position of the pin (6-mm diameter). Weld M13 adhered to the backing plate, and the act of removal resulted in significant damage.
Table III. Ultimate Tensile Strengths and Elongations to Failure for the Dissimilar Material Welds AA6082 Advancing Side
AA5083 Advancing Side Rotation Speed (rpm) 280 560 840 280 560 840 280 560 840
Traverse Speed (mm/min)
UTS (MPa)
Elongation to Failure (mm, Over 50 mm Gage Length)
UTS (MPa)
Elongation to Failure (mm, Over 50 mm Gage Length)
100 100 100 200 200 200 300 300 300
204 209 199 216 223 222 228 233 241
2.3 2.44 2.2 2.7 2.68 2.89 2.6 3.17 3.05
— 206 — 212 222 227 — 224 —
— 2.45 — 2.7 2.79 2.67 — 2.9 —
systematic set of similar and dissimilar material welds produced under well-characterized and repeatable conditions. The force, torque, and temperature at various distances from the weld line have been measured throughout the weld cycle. In addition, a simple thermal FE model has been used to estimate the variation in the temperature during the steady-state region of the welding cycle. The results show that generally speaking, good welds can be obtained across a very wide range of conditions, with only those welds undertaken at the lowest rotation speeds and highest METALLURGICAL AND MATERIALS TRANSACTIONS A
traverse speeds with AA6082 on the advancing side producing poor weld properties and macrostructures. It is notable that both the torque and the extent of material mixing in the stir zone display a much stronger dependence on the rotation speed than the traverse speed. In each case the variation is indicative of an increase in the temperature of the material near the tool, as predicted by the current model. However, it is difficult to see any simple correlation between the temperature and the power or heat input. Although the heat input is commonly considered in VOLUME 37A, JULY 2006—2191
fusion welding, the current results suggest that it is misleading to use such a value as an indicator of the temperature of the material surrounding an FS welding tool, at least for the joining of thin plates of aluminum. It is likely that when the traverse speed is reduced, much of the additional heat is conducted into the backing plate, as evidenced by a correlation between the heat input and the backing plate temperature. The calculated interfacial shear stress around the tool is broadly consistent with known flow stress values for the two alloys being joined. While the calculated friction coefficients are also within reasonable bounds, the absence of any obvious relationship between the down force and the torque does not seem to support such a model, which predicts a proportional relationship between the friction force and pressure. Intuitively, a relationship between the torque, down force, and flow stress may be expected, since a low torque implies a low flow stress that would allow the tool to settle more easily into the workpiece. However, the torque will be dependent only on the flow stress of a comparatively small region of material in contact with the tool’s surface, while the down force will depend on the flow stress of the material underneath the entire width of the tool. Further, the down force will also depend on the capacity for the colder material surrounding the hot zone to constrain the hot material and maintain the hydrostatic stress state (which will not alter the torque). AA6082 relies on age hardening to provide its strength, and since the precipitates are unstable above ;200 °C, the flow stress decreases rapidly above this temperature (Figure 10). In contrast, AA5083 is predominantly strengthened by a solid solution of Mg and displays a more gradual decrease in flow stress. This means that even if the thermal contours are the same, a weld in AA6082 will have a wider soft zone than one in AA5083. As a result, material is more easily displaced, allowing the tool to settle into the workpiece with less force. This was evidenced in the current case by the deeper weld track and more copious flash in the AA6082 welds compared to the AA5083 welds (with the dissimilar welds being intermediate). This also explains why AA6082 can be welded faster than AA5083, since the thermal contours can be closer to the tool before a critical flow stress is exceeded and flaws are produced (or the tool breaks). The energy input, as determined from the measured torque, can be described in terms of the power (P) and the heat input (Q 5 P/v). The increase in power with rotation speed can be attributed to the increase in the relative interfacial velocity between the workpiece and the tool, which will increase the dissipation of energy at the interface. However, the increase in power is not proportional to the rotation speed, since an increase in temperature will reduce the capacity to generate heat by lowering the interfacial contact stress (due to a lower flow stress or friction coefficient, depending on the model used). Colegrove et al. have postulated that there are essentially two regimes in FS welding: a cold regime, where interfacial melting does not occur, and a hot regime, characterized by a molten film at the interface.[18] In the former there is likely to be a relationship between the welding speeds and the energy input, since the interfacial flow stress or friction coefficient will be able to change substantially as the rate of heat input alters. If interfacial melting occurs, then the friction coefficient will become very low, 2192—VOLUME 37A, JULY 2006
and changes in rotation or traverse speed will tend to have little effect on the heat input. If this hypothesis is correct, then the current results appear to lie within the cold regime, in contrast to the results by Colegrove et al., who found that the power was effectively independent of the rotation speed in the aluminum alloy AA7075. In the current set of experiments it has been found that the best fit occurred when 55 pct of the total power generated by the tool entered the workpiece as heat (i.e., the efficiency factor f 5 0.55). This value is clearly very much lower than would be expected if the predominant loss of energy were heat flow into the tool (5 to 20 pct[9,22–24]). Chao et al. estimated that up to 20 pct of the energy could be retained in the weld as stored energy due to the deformation of material during the FS welding of AA2195.[29] However, the main basis for this hypothesis appeared to be that their model provided a power estimate around 80 pct that of the experimentally determined power, and the reliability of this estimate has not been determined. In the most extreme case, where 20 pct of the heat is lost into the tool and 20 pct is retained as defects in the material, then the value would begin to approach the fraction of power input as heat used in the current model. This is, however, tentative at best, and it remains possible that the error is due to the assumptions made in the model or a systematic error in the recorded torque data. Clearly there is potential to develop the model and to perform further experimental studies to validate the output, but our primary reason for developing it is to explain the hardness and, eventually, residual stress distributions across and between the different welds. In part II of this paper,[15] the analysis of the influence of the welding parameters is extended to the development of the microstructure and mechanical properties of the welds. It is shown that the strong influence of the rotation speed also applies to factors such as the grain size and the extent of natural aging in the AA6082 alloy. The thermal model will be used as the basis of an analytic model describing the hardness variation across the weld. ACKNOWLEDGMENTS This work was part of the Brite-Euram project ‘‘JoinDMC’’ with project number GRD1-1999-10551. The extensive help of the other partners is acknowledged, in particular Mr. F. Palm at EADS for the production of the welds. The authors thank the staff of the German Aerospace Centre (DLR) for their assistance with the collection of the instrumented weld data. P.J.W. acknowledges a Royal Society-Wolfson Merit Award. REFERENCES 1. W.M. Thomas, E.D. Nicholas, J.C. Needham, M.G. Murch, P. Temple-Smith, and C.J. Dawes: Friction Stir Butt Welding. 1991, International Patent No. PCT/GB92/02203. 2. T. Nagasawa, M. Otsuka, T. Yokota, and T. Ueki: in Magnesium Technology 2000, Minerals and Metals Society, 2000, pp. 383-87. 3. M.C. Juhas, G.B. Viswanathan, and H.L. Fraser: in 2nd International Symposium on FSW (CD-ROM), Gothenburg, Sweden, 2000, TWI. 4. A.P. Reynolds, W. Tang, T. Gnaupel-Herold, and H. Prask: Scripta Mater., 2003, vol. 48 (9), pp. 1289-94. 5. W.M. Thomas, P.L. Threadgill, and E.D. Nicholas: Sci. Technol. Weld. Joining, 1999, vol. 4 (6), p. 365. METALLURGICAL AND MATERIALS TRANSACTIONS A
6. E.D. Nicholas: in ICAA-6: 6th International Conference on Aluminium Alloys, Toyohashi, Japan, 1998, Japan Institute of Light Metals. 7. T. Hashimoto, S. Jyogan, K. Nakata, Y.G. Kim, and M. Ushio: in 1st International Symposium on FSW (CD-ROM), Thousand Oaks, CA, 1999, TWI. 8. A.P. Reynolds, W.D. Lockwood, and T.U. Seidel: Mater. Sci. Forum, 2000, vol. 331–337, pp. 1719-24. 9. M.J. Russel and H.R. Shercliff: in 1st International Symposium on FSW (CD-ROM), Thousand Oaks, CA, 1999. 10. L. Luijendijk: J. Mater. Process. Technol., 2000, vol. 103, pp. 29-35. 11. Y. Li, L.E. Murr, and J.C. McClure: Scripta Met, 1999, vol. 40 (9), pp. 1041-46. 12. H. Larsson, L. Karlsson, S. Stoltz, and E.L. Bergqvist: in 2nd International Symposium on FSW (CD-ROM), Gothenburg, Sweden, 2000, TWI. 13. S. Tanaka and M. Kumagai: in 3rd International Symposium on FSW (CD-ROM), Kobe, Japan, 2001, TWI. 14. J.H. Ouyang and R. Kovacevic: J. Mater. Eng. Perform., 2002, vol. 11 (1), pp. 51-63. 15. M. Peel, A. Steuwer, and P.J. Withers: Metall. Mater. Trans. A, 2006, vol. 37A, pp. 2195–206. 16. R. Johnson: in 3rd International Symposium on FSW (CD-ROM), Kobe, Japan, 2002, TWI. 17. T.U. Seidel and A.P. Reynolds: Sci. Technol. Weld. Joining, 2003, vol. 8 (3), pp. 175-83. 18. P. Colegrove and H.R. Shercliff: Sci. Technol. Weld. Joining, 2003, vol. 8 (5), pp. 360-88. 19. K. Lindner, Z. Khandkar, J. Khan, W. Tang, and A.P. Reynolds: in 4th International Symposium on FSW (CD-ROM), Utah, 2003, TWI. 20. Q. Shi, T. Dickerson, and H.R. Shercliff: in 4th International Symposium on FSW (CD-ROM), Utah, 2003, TWI. 21. Y.J. Chao and X. Qi: in 1st International Symposium on FSW (CDROM), Thousand Oaks, CA, 1999, TWI.
METALLURGICAL AND MATERIALS TRANSACTIONS A
22. P. Colegrove, M. Painter, D. Graham, and T. Miller: in 2nd International Symposium on FSW (CD-ROM), Gothenburg, Sweden, 2000. 23. T. Dickerson, Q. Shi, and H.R. Shercliff: in 4th International Symposium on FSW (CD-ROM), Utah, 2003, TWI. 24. H. Schmidt, J. Hattel, and J. Wert: Modell. Simul. Mater. Sci. Eng., 2004, vol. 12 (1), pp. 143-57. 25. W.M. Rohsenow and J.P. Hartnett: Handbook of Heat Transfer. McGraw-Hill, New York, London, 1973. 26. K.C. Mills: Recommended Values of Thermophysical Properties for Selected Commercial Alloys. ASM International, 2001. 27. A. Federation: The Properties of Aluminium and its Alloys, 6th ed. Aluminium Federation, London, 1968. 28. E.A. Brandes and G.B. Brook, eds.: Smithells Metal Reference Book. Butterworth-Heinemann, Oxford, 1992. 29. Y.J. Chao, X. Qi, and W. Tang: J. Manuf. Sci. Eng., 2003, vol. 125 (1), pp. 138-45. 30. W. Blum, Q. Zhu, R. Merkel, and H.J. McQueen: Mater. Sci. Eng. A, 1996, vol. 205 (1–2), pp. 23-30. 31. L.D. Oosterkamp, A. Ivankovic, and G. Venizelos: Mater. Sci. Eng. A, 2000, vol. 278 (1–2), pp. 225-35. 32. J.R. Cho, W.B. Bae, W.J. Hwang, and P. Hartley: J. Mater. Process. Technol., 2001, vol. 118 (1–3), pp. 356-61. 33. S. Spigarelli, E. Evangelista, and H.J. McQueen: Scripta Mater., 2003, vol. 49 (2), pp. 179-83. 34. I.M. Hutchings: Tribology: Friction and Wear of Engineering Materials. E. Arnold, 1992. 35. G.E. Dieter: Mechanical Metallurgy. 3rd ed. McGraw-Hill, New York, 1986. 36. M. Peel, A. Steuwer, M. Preuss, and P.J. Withers: Acta Mater., 2003, vol. 51 (16), pp. 4791-801. 37. J. Hagstrom and R. Sandstrom: Sci. Technol. Weld. Joining, 1997, vol. 2 (5), pp. 199-208.
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