D.H. Lammlein PhD Dissertation - Vanderbilt University

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5m 2 ) is the non-contact area, and 0.246”( 0.0063m) is the non-contact centroid distance. For the wide probe tool this equation gives: 0.71.66 7104.72 2 3 π. 00793 ‐3.003 2 . 0046 2.4e 5. 0063 =1,403 Watts where only the radius of the wide probe tool, 0.118” (0.003m), is substituted from the previous calculation. The slightly increased surface area of the probe and the slightly higher relative velocity of the probe wall result in higher heat input for the fat probed tool at identical parameters. Although torque values for this experiment were not obtained, a secondary estimate of weld thermal input can be obtained via the weld power method by using a torque value taken from a similar experiment under similar conditions. A value of 10.56 N·m is taken from an experiment performed on butted spheres of identical (0.2”) thickness using the same rotary welding apparatus. The tool shoulder from this case was of the same 5/8” diameter but was cupped in geometry to mate with the spherical work. Like the present case this value was taken over a 1000rpm weld, however, the lowest traverse rate recorded during the analog spherical experiment at the 1000rpm rotation rate was 7.8ipm, so an extrapolated value is taken. The probe was 3/16” diameter and threaded and therefore will be compared with the 3/16” diameter probe calculation. Using the weld power method (Pew, Nelson, & Sorensen, 2007): , 10.56 104.7 1106 1287 An upper limit weld power can be calculated using an upper limit weld torque obtained by applying the measured axial load at the very edge of the tool shoulder. For the wide probe modeled case this yields: , 2880.5. 0079 104.7 2394 Watts For the presented models, the Q n and Q w heat rate or power values are used for the narrow probe and wide probe models respectively. Heat was distributed across the weld interface and input into the work material as defined by a user-defined function which 184
applied heat locally in proportion to the local tool velocity. This results in highest energy input at the edge of the shoulder where the local tool velocity is highest. A second user defined function was used to impose the temperature on the work side of the interface onto the adjacent interface cells on the tool side of the interface. This formulation resulted in 5.1% of the total weld power entering the tool in the narrow probe case and 6.6% entering the tool in the wide probe case. The total heat input into cells on both sides of the interface was equal to the values obtained by the Schmidt formula. A rotational velocity of 1000rpm and traverse speed of 5.2ipm were used in the presented models of the narrow probe and wide probe geometries. Additionally, cases were run for the wide probe geometry at all parameters used experimentally in order to provide a numerical estimation of axial force for comparison with experiment. In Fluent, the tool rotational axis is maintained stationary while velocity inlet type boundary conditions are used at the inlet and outlet boundaries to establish material flow past the tool, as when traversing. The outlet boundary condition (BC) can be defined as a velocity inlet with negative magnitude because the flow far from the tool in the work is known to be the traverse rate. A pressure outlet or outflow boundary condition at the outlet would not take advantage of this knowledge and would instead force fluent to arrive at some flow profile. Thermal BC's are defined as 10 W/(m 2·K) for exposed surfaces and 150 W/(m 2·K) for intimate metal to metal contact. The rotating and exposed surface of the tool was defined as 50 W/(m 2·K). Additionally, a wall emissivity of 0.2 was applied to exposed aluminum surfaces and 0.5 to exposed steel surfaces. The FSW environment presents an extreme fluid case for Fluent software. Temperature and velocity gradients are steep in the vicinity of the tool and the tool rotational velocity is high with respect to the traverse speed. Iterating the models starting with the full zero shear viscosity and tool rotational velocity results in divergence errors. In CFD software it is best practice that best guess initial conditions be provided to simplify arrival at a solution. The models presented were initialized using the full experimental traverse speed. From this point, it was found that a solution was best reached in the presented models by successive increases in tool rotational velocity and zero shear viscosity from a small fraction of their desired values to, eventually, their 185
Page 1 and 2:
FRICTION STIR WELDING OF SPHERES, C
Page 3 and 4:
TABLE OF CONTENTS TABLE OF FIGURES
Page 5 and 6:
Appendix ..........................
Page 7 and 8:
Figure 16: Left(FE streamline), Mid
Page 9 and 10:
Figure 51: A thermal camera image t
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for a similar conventional weld (3/
Page 13 and 14:
Figure 102: Tapered retraction proc
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setting are indicated. Note that st
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Unsupported welds of this kind with
Page 19 and 20:
Figure 157: A thermal camera image
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ottom sample is a tungsten inert ga
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N - newton (force unit) n - visco-p
Page 25 and 26:
INTRODUCTION Friction Stir Welding,
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educed conduction lengths. In small
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CHAPTER I LITERATURE REVIEW: COMPUT
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frictional condition at the interfa
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contact. The contact area can inclu
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For horizontal sections, a heat inp
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(1.11) where r p is the probe (pin)
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This efficiency term, η pd , is kn
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The heat input on the tool surface
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The added complication of these met
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Figure 5: Lateral cross sections of
Page 47 and 48:
Flow stress in aluminum alloys is d
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Figure 8: Three flow fields a) rota
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and post-weld cool-down. This model
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Williams et al. [60] use a 2D-axisy
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Figure 16: Left(FE streamline), Mid
Page 57 and 58:
Figure 18: Temperature contour over
Page 59 and 60:
Figure 22: Temperature distribution
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Figure 24: Lateral velocity contour
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Figure 26: Experimental(white) vers
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Figure 28: Top view schematic of tu
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[5] Taylor RE, Groot H, Goerz T, Fe
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[41] Heinz B., Skrotzki B., Eggler
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CHAPTER II COMPUTATIONAL ANALYSIS O
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Figure 30: Blind t-joint setup with
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Figure 33: Process forces and later
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Figure 35: Experimental Welds, Axia
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Figure 37: FEA boundary conditions
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A 4500 N axial force is applied eve
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axial force with lateral offset dur
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Figure 42: Contour plot of von Mise
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Figure 44: Deformation contour for
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Figure 46: Quarter-plate model for
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Figure 49: Contour plot of deflecti
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where P is the weld power (W), Q is
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Figure 52: Fluent CFD model zones.
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Figure 55: Thermal contour of weld
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Figure 59: Thermal contour of weld
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Figure 61: Velocity vectors for 0.0
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Figure 65: Contour of velocity magn
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CHAPTER III THE APPLICATION OF SHOU
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Figure 66: A shoulder-less, conical
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Figure 68: Typical axial(Z-axis) fo
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Figure 71: 90° conical tool macros
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Figure 75: Run 3, 80° conical tool
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Figure 79: (90º tool) Ultimate ten
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Figure 83: (90º tool) Lateral forc
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Figure 87: (80º tool) Lateral forc
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The Eulerian, finite volume, CFD so
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Figure 92 shows the increasing elem
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Figure 91: CFD model geometry consi
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Figure 95: Lateral cross-section of
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Figure 96: Attempts a probe tapered
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CHAPTER IV THE FRICTION STIR WELDIN
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applications for its high strength
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Figure 97: Experimental weld sample
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Figure 99: Split protrusion type de
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with a cupped recess. It is however
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of tapered retraction (i.e. move th
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The adjustments in vertical positio
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manners the assigned depth was main
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Figure 108: Lateral macrosections f
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Figure 110: (Supported, cupped tool
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Page 153 and 154:
Page 155 and 156:
apparent strength of 26% parent whi
Page 157 and 158: with a 100º conical tool (0.025”
Page 159 and 160: Figure 119: Macrosection view of an
Page 161 and 162: and tooling. Torque is highly depen
Page 163 and 164: locally in proportion to the local
Page 165 and 166: Figure 126: (threaded tool) Contour
Page 167 and 168: Figure 128: (conical tool) Geometry
Page 169 and 170: Figure 131: (threaded tool) Lateral
Page 171 and 172: The results presented here show tha
Page 173 and 174: Figure 136:(from left: Tapered retr
Page 175 and 176: [8] Metallurgical analysis of ablat
Page 177 and 178: CHAPTER V FRICTION STIR WELDING OF
Page 179 and 180: Collaboration between Advanced Meta
Page 181 and 182: eccentricity, the method of interio
Page 183 and 184: gas tungsten arc welding of small d
Page 185 and 186: Full penetration welds of 4.2” (1
Page 187 and 188: Figure 143: The curvature of the pi
Page 189 and 190: Figure 144: Experimental axial forc
Page 191 and 192: Together, a fine wall thickness tol
Page 193 and 194: Figure 147: Axial force history for
Page 195 and 196: The weld macrosections show complet
Page 197 and 198: Figure 154: Macrosections of welds
Page 199 and 200: Figure 157: A thermal camera image
Page 201 and 202: Figure 159: Average tool shank temp
Page 203 and 204: 471,146 faces. The meshes are fine
Page 205 and 206: Figure 162: A closeup view of mesh
Page 207: 4 2 3 4 5 16 " 1.745 2 0.24
Page 211 and 212: Figure 164: Modeled temperature con
Page 217 and 218: Figure 172: Model pathlines for the
Page 219 and 220: Figure 174: Model pathlines for the
Page 221 and 222: later speed setting, the tool welds
Page 223 and 224: Figure 178: Zoom view of a macro ta
Page 225 and 226: Figure 180: The superficial appeara
Page 227 and 228: Figure 182: The rotary welding appa
Page 229 and 230: Figure 185: A diagram illustrating
Page 231 and 232: CONCLUSION AND FUTURE WORK Conclusi
Page 233 and 234: new FSW process environment is reso
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D.H. Lammlein PhD Dissertation - Vanderbilt University

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