HEAT FLOW INTO FRICTION STIR WELDING TOOLS

Transcription

1 HEAT FLOW INTO FRICTION STIR WELDING TOOLS Terry Dickerson, Qing-yu Shi, Hugh R Shercliff Cambridge University Engineering Department, Trumpington Street, Cambridge, CB2 1PZ, UK. ABSTRACT Using experimental welding data and thermal modelling the transient heat losses into FSW tools were calculated. Heat loss into the tool enabled the welding efficiency to be determined. The body of this work covered the welding of two light alloys using solid steel tools. Finally the use of grooves in the tool to impede heat flow was investigated as a way of increasing welding efficiency. The work concludes that using solid tools the steady state heat loss into the tool is about 10% of the total heat generated; for short welds that start with a cold tool this can be higher. Grooves in the tool shank increase the thermal efficiency and stability of the welding process. 1. INTRODUCTION For friction stir welds (FSW) in light alloys most of the heat is transported into the work piece; however a proportion of the heat is conducted into the welding tool which reduces welding efficiency. The welding efficiency (f) is defined here as: Page 1 of 11 f = (P-q)/P [1] Where P is the total mechanical power input to the weld and q is the heat power loss into the welding tool. For a weld with tool rotation speed ω and reaction torque (M), P is defined as: P = ω.m [2] It is assumed that all of the mechanical power is converted into heat and is transferred into and through the weld and tool (no welding flash, acoustic energy loss small, etc). It is convenient to work in weld power (Watts) rather than the heat input (Joules/mm). The aims of this work were to: determine the heat loss into the tool whilst friction stir welding and evaluate f investigate an improvement in the thermal design of the welding tool In particular the above were investigated for the transient phase at the start of welds. The following investigations were carried out and are reported on here:

2 Welding experiments allowed the assessment of the total mechanical power input, P, allowed weld interface and tool temperatures to be determined. Finite element models of the tool were constructed with fixed temperature boundary conditions derived from the welding experiments. Transient analyses were run to simulate the heating of the tool from the start of welding. Predicted temperatures from the models were validated against the experimentally derived tool temperatures. Weld efficiency factors f are calculated. The use of grooves in the tool flank was investigated as a method of improving the thermal stability and efficiency of welding. 2. WELDING EXPERIMENTS A series of welds in two combinations of light alloys was carried out. The welds were instrumented with thermocouples and the tool was driven through a dynamometer which gave dynamic readings of forces and torques. The data from two of these welds were used for this work, the weld identifiers and material combinations are listed in Table 1. The weld panels were (150 60)mm by 3mm thick and joined as butt welds with a weld length of 105mm. The tools can be seen in Fig.1; they had a Ø18mm shoulder with a M6 threaded probe that was adjustable for length. The tools were made from hardened tool steel and were solid apart from a Ø6mm axial hole in which the adjustable probe fitted. The welding parameters used are listed in Table 1, they are not necessarily optimal. 2.1 WELDING DA TA The welding data taken from the dynamometer is very detailed, only selected values derived from this data will be given here. Generally the values are averaged, either over a section of a weld and/or from multiple welds. Table 1 Details of the welds used in the experimental work Identifier Material Combinations Description DMC-A T6 / 2024-T3 dissimilar aluminium alloys Welding Speed 200mm/min Tool Rotation Speed 600rpm DMC-B4 AM60 / AM60 magnesium alloy 400mm/min 700rpm Page 2 of 11

3 2.2 WELD TEMP ERATURES The welds had thermocouples inset into the panel surfaces such that the tool shoulder traversed directly over them. These thermocouple junctions were originally located 4mm from the centre-line of the welds with one on the retreating side and one on the advancing side. The thermocouples therefore travelled through the intensely stirred region near the intersection of the shoulder and the probe. This enabled the near interface temperature to be recorded. Because of their location, the thermocouples were prone to movement and damage during welding. Figure 2 shows the welds and the resulting positions of the thermocouples. Two of the thermocouples gave valid readings throughout the welding. In addition channel 5 (Ch5) of DMC-A5 gave meaningful results up to the time that the probe passed its original position. This broken thermocouple carried on recording after the probe had passed by and despite this the temperature results look quite believable. However, because of the significant displacements of the thermocouple wires, see Fig.2(a), results after the probe passed the original thermocouple location are not used. Figure 3 shows the temperature data. Ch4 of weld DMC-A5, on the retreating side of the weld, moved approximately 5mm during welding. This has the effect of compressing the position readings under the tool and shifting the readings after the tool had passed along the weld distance axis by -5mm. 200 C 260 C 240 C 300 C (a) Figure 1 Two welding tools showing temper colours and approximate temperatures: (a) DMC A5 (Aluminium Alloys) and (b) DMC-B4 (Magnesium Alloy). (b) Page 3 of 11

4 Advancing Side + Ch5 Ch4 + 10mm (a) Retreating Side (b) Advancing Side + Ch5 + Ch4 10mm Retreating Side Figure 2 The twelds showing the thermocouples embedded after welding: (a) DMC-A5 (b) DMC-B4. On the left is a photograph of the weld cap and on the right an X-ray image. The marks the thermocouple position before welding and the + is after welding. The dotted lines represent the nominal path through which the thermocouples travel relative to the tool. Temperature / [ C] DMC-A5 Ch4 DMC-A5 Ch5 DMC-B4 Ch4 DMC-B4 Ch5 Passage of shoulder over original position of thermocouples Tool position along Weld / [mm] Figure 3 Welding interface temperature measurements for DMC-A5 and DMC-B4 as a function of tool position during welding, zero is at the plunge position. Page 4 of 11

5 2.3 WELD INTER FACE TEMPERATURE ESTIMATE The tilt angle on the tool means that the trailing edge of the tool is plunged into the panels and the leading edge is raised above the panels. This means that the thermocouples will be below the weld interface when the tool first covers them. As the tool moves over the thermocouples they will move closer to the weld interface. These effects will tend to lower the recorded temperature at the tool leading edge when compared to the interface temperature. Visual examination of the finished welds confirmed that the remaining thermocouples were either at or close to the weld surface. Therefore at the trailing edge of the tool the thermocouples should have recorded the interface temperatures. In Fig.2, before the tool probe reaches the thermocouples, the temperatures for the dissimilar material weld are almost identical. The joint-line thermally insulates the weld panels and the thermal properties of the two materials are similar, this suggests that the heat generation is approximately symmetric about the joint-line (despite the incomplete filling of the shoulder, see Fig.2). Because the tool rotates at relatively high speed, the best estimate at this stage is that the temperature of the tool working face is uniform and at the peak temperature recorded by the thermocouples. This assumption was validated using temperature data taken from the tools. 2.4 TOOL TEMPE RATURES Prior to welding, the steel tools were polished to a bright finish. During welding heating of the tools caused them to discolour. These tempering colours are used as a guide to the temperature profile of the tools and validation of the modelling. Figures 1(a) and (b) show photographs of the welding tools used for the aluminium alloy welds and the magnesium alloy welds respectively. The estimated temperatures are shown in the figures. The colours and hence temperatures are subject to interpretation and hence error; the value underlined is less certain than those without. 3. THERMAL MODELLING OF THE TOOLS The welds were relatively short and the tools did not have time to reach steady-state. Therefore, transient thermal models of the tool were produced using ABAQUS/Standard. The same mesh was used for modelling the tool used for each material combination. Different boundary conditions allowed the tool to be simulated for the different welds. Page 5 of 11

6 3.1 GEOMETRY A ND MESH Only the tool and its holders were meshed; axisymmetric elements were used. Geometrically the model consisted of three parts: The tool, the tool holder and the machine-tool toolholder. These parts were thermally connected by an interface conductance. Figure 4 shows a picture of the geometry and the mesh used. 3.2 MATERIAL PROPERTIES The analyses only warranted the use of approximate properties; these were obtained for tool steel from handbooks 1,2 and are listed in Table 2 C L Machine Tool Holder Tool Holder X X X X 3.3 BOUNDARY CONDITIONS FSW Tool The thermal loading was applied by giving the tool probe volume and (a) (b) shoulder surface a constant Figure 4 Axisymmetric finite element model of the welding tool and its holders. temperature equal to the maximum temperature recorded by DMC-A5 Ch4 and DMC-B4 Ch4. This constant Table 2 Properties for Tool Steel Property Value, units Density 7800 kg m -2 interface temperature was assumed Specific Heat Capacity 500 J kg -1 k -1 because, after an initial peak, the total Conductivity 45 W.m -2 K -1 mechanical heat input to the weld was approximately constant. Actually the heat input showed a small (~5%) decline from the start to the finish of the welds which suggesting the interface temperature was increasing, As thermocouple data indicated the welds were in thermal steady-state this reduction in the power required was associated with heating of the tool. As this transient heat flow in to the tool is the subject of this work the assumption of constant temperature gives a second order error and this was considered acceptable. Page 6 of 11

7 The boundary conditions are listed in Table 3; the total welding time included ramping of the temperature over a 5 second period to represent the plunge, the dwell and welding. Table 3 Boundary conditions applied to the finite element models Description Symbol in Value for Fig 4(b) DMC-A5 DMC-B4 Fixed temperature at machine tool chuck X 30 C Contact conductivity between tool holders O 2800 W m -2 K -1 Fixed Temperature at tool/weld interface 475 C 380 C Heating (weld) time, including dwell N/A 41.5 s 21.8 s Convection (heat transfer coefficient) 100 W m -2 K -1 Free Surfaces Radiation (emissivity) MODELLING RESULTS The predicted temperatures at the end of the weld simulations are shown as contour plots in Fig.5(a) and (b); these may be directly compared to the indicated temperatures in Fig.1(a) and (b). In Fig.6 the model and observed temperatures are compared and they show remarkably good agreement which suggests the boundary condition assumptions were valid. Distance Along Tool Axis / [mm] DMC-A5 - Model DMC-B4 - Model DMC-A5 - Observed DMC-B4 - Observed (a) (b) Figure 5 Temperature predictions for the welding tools; the 2D axisymmetric meshes have been expanded into 3D. (a) used for weld DMC A5 (aluminium alloys) and (b) used for weld DMC-B4 (magnesium alloy) Temperature / [ C] Figure 6 Tool surface-temperature predictions compared to the temperature observations (from the tempering colours). Page 7 of 11

8 The constant 1400 temperature assumption and the heating of the tool meant that the power loss into the tool DMC-A5 DMC-B4 changed as the 200 welding progressed; 0 / / Steady State the variations for the Time / [s] two welds are shown (a) (b) 1400 in Figs.7. In the DMC-A5 figures the graphs 1200 DMC-B4 show a peak during 1000 the plunge/dwell 800 period. The power loss reduces as the welding progresses 200 because the tool is 0 / / Steady heating-up. Figure State Weld Distance / [mm] 7(b) also shows the Figure 7 Predicted heat flow into the tool as a function of: power losses into the (a) Welding time; zero time is at the start of the traverse tool during the (b) Tool position along the weld. welding are almost identical for both welds; this is a coincidence and a consequence of the shorter dwell time and faster welding speed for DMC-B4. The models were run to steady state (>1000s) and these results are also shown on the right in Figs.7. Power Loss into Tool, q / [W] Power Loss into Tool, q / [W] 4. WELDING EFFICIENCIES The weld efficiencies (f) were calculated at two positions of the two types of weld: At an arbitary position near the middle of the welds between 60-70mm (f M ). The measured value of P and predicted values of q at this distance are used to calculate the efficiencies. This value is typical for short welds where the tool has not been preheated. Page 8 of 11

9 The steady state condition (f SS ). Thermocouple and dynometer results (not shown here) indicate the weld panels reach thermal equilibrium within the first 30mm of the start of the weld. Therefore the average mechanical heat input during the last 10mm of the welds is used as the steady state values of P. The values of q are taken from the models run to the steady state condition. The weld efficiency results are shown in Table 4. Table 4 Heat flows during friction stir welding and the welding efficiencies. Weld Identifier Mechanical Heat Input. P [W] Between 60-70mm Heat Loss Into Tool, q [W] f M Mechanical Heat Input. P [W] At steady state Heat loss into tool, q [W] DMC-A DMC-B f SS 5. DISCUSSION The temperatures and heat flows into welding tools have been modelled. Constant temperature boundary conditions were used to drive the heat flow into the weld tool models; these boundary conditions were based on experimental observations. The resulting temperature profiles of the models were compared to experimental observations; these were made from tempering colours that were difficult to interpret and so prone to error. However, the temperature predictions and observations showed remarkably good agreement, which indicated that the model results for the heat flow are credible. Heat flow into the tool was used to calculate the thermal efficiency factor f, which determines the proportion of heat flow into the weld. The values of f M and f SS for the welds are shown in Table 4; as would be expected the steady state values are higher than for the short welds. These efficiency values can be used when interpreting the welding power data derived from torque and rotation speed data. Figure 8 Temperature predictions for the welding tool with grooves. The model is otherwise identical to Fig.4(a). Page 9 of 11

10 As previously 1400 mentioned and 1200 DMC-A5 DMC-A5: Grooves in Tool indicated in Fig.7, the heat flow into the tool had not reached steady state during the short welds made for this work. This is 0 / / Steady attributed to the State Time / [s] thermal capacity of Figure 9 Predicted heat flow into the tool as a function of welding the tool and its time for tools without and with grooves. holders. Some designs of tool have grooves in the shank just above the shoulder presumably to increase the thermal resistance and lower the thermal mass of the tool. The effect of such grooves was investigated for the tool used on DMC-A5; all other conditions were the same as previously used. The geometry and temperature results can be seen in Fig.8; comparing with Fig.5(a) it can be seen that the temperature gradient along the tool axis direction is much greater. The comparison of the heat flows for the tools with and without the grooves is shown in Fig.9. As can be seen from the figure, the tool with the grooves not only has a lower heat loss through the tool but also the heat loss stabilizes much more quickly. The tool with the grooves is almost at its working temperature when the welding traverse starts. Putting grooves into the tools therefore is advantageous from the point of view of weld efficiency and process consistency. Another advantage is that the need for cooling the machine tool bearings will be reduced. Of course the tool will need to be properly designed to prevent failure and this will limit the depth and shape of the grooves and hence the gains in efficiency. 6. CONCLUSIONS Power Loss into Tool, q / [W] For friction stir welding of light alloys using a steel tool: using a solid tool the steady state welding efficiency is about 0.9, ie 10% of the heat generated is lost through the tool. For short welds that start with a cold tool the welding efficiency may be 0.8 or lower. Tools with grooves cut into the shank should be used to decrease the heat loss through the tool and increase the process stability. The work here suggests the efficiency can be raised to 0.94, even for short welds. Page 10 of 11

11 7. ACKNOWLEDGEMENTS: This work has been supported by the European Community under the Competitive and Sustainable Growth Programme ( ). Project name: Joining Dissimilar Materials and Composites by Friction Stir Welding. Project No.: GRD Contract No.: G5RD- CT The authors wish to thank Mr. Frank Palm at EADS (Ottobrunn, Germany) and technical staff at DLR (Cologne, Germany) for their help with the experiments. 8. REFERENCES 1 Brandes E A and Brook G B (eds); Smithells Metals Reference Book, 7th Edition, Tennent R M (ed); Science Data Book, Oliver and Boyd, Page 11 of 11