- Email: [email protected]
A thermomechanical hot channel approach for friction stir welding
Journal of Materials Processing Technology 174 (2006) 190–194
A thermomechanical hot channel approach for friction stir welding S. Mandal a , K. Williamson b,∗ a b
Department of Mechanical Engineering, Old Dominion University, Norfolk, VA 23529, USA Department of Industrial Technology, East Carolina University, Greenville, NC 27858, USA
Received 10 June 2004; received in revised form 25 January 2005; accepted 13 December 2005
Abstract This paper provides a theoretical framework for developing a thermomechanical hot channel (THC) approach for augmenting the friction stir welding (FSW) process. The model follows from the Rosenthal solution for moving point sources where heat input from the tool shoulder acts as a warm source while plasmas or laser heat sources provide higher energy input. The THC approach aims at decreasing tool wear by reducing the demand for frictional heat from the tool shoulder and pin. In the proposed model, the THC processing temperature is approximately 1000 ◦ C, which is close to the temperature required for stir welding steel. © 2006 Elsevier B.V. All rights reserved. Keywords: Friction stir welding; Tool wear; Pre-heating; Thermomechanical hot channel
1. Introduction Friction stir welding (FSW) has emerged as an innovative method for joining low melting temperature alloys like aluminum (Al) and magnesium (Mg). The process combines frictional heating with intense plastic deformation to produce cost efficient joints with better mechanical properties than conventional fusion welding techniques. FSW has already been applied to the construction of aluminum structures in both the transportation [1] and aerospace [1] industries, and it is being thoroughly investigated for application to steel structures in the shipbuilding industry. Although most of the research on FSW has been on aluminum, investigations for steel are still relatively new. Laboratory trials carried out on low carbon steel by Thomas [1] from TWI showed that the heat required to induce plastic deformation similar to aluminum is much higher. FSW for steels require the tool to spin at much faster speeds resulting in a far higher tool wear rates for steel joints compared to softer aluminum joints. This factor is highly pronounced for long welds. Thomas [1,5] and Lienert et al. [6] reported significant tool wear during the friction stir welding of steel. In their experiments, Thomas minimized tool wear by using pre-drilled holes, less than the diameter of the tool pin as a means to reduce the high tool wear
during plunging. In their research, Lienert et al. [6] suggested that increased wear during plunging may be due to the high load spikes resulting from the greater flow stress of the cold workpiece. Besides pre-drilling holes to minimize wear, allowing the tool to spin in one position for a sufficient time plasticizes additional workpiece material and creates additional heat input through plastic work. These ongoing experiments reflect significant effort towards the development of more robust FSW techniques for steel that offsets problems with tool wear. Frequent replacement of worn out tools is expensive and it carries additional cost through delays and reduced production rate. Also, the presence of inclusions from worn out tools, in the weld would significantly reduce the quality of the joint. Given the significance of tool wear during the friction stir welding of steel and other harder materials, this study investigates thermomechanical hot channels as an approach for taking FSW beyond the limits of aluminum and other softer materials. Here, the goal is to combine FSW with pre-heating sources that create a thermomechanical hot channel (THC) ahead of the FSW tool. The idea is to pre-heat the workpiece, and reduce the amount of frictional heat and subsequently the tool wear rate (Fig. 1). 2. Theory
∗
Corresponding author. Tel.: +1 252 328 9722; fax: +1 252 328 1618. E-mail address: [email protected] (K. Williamson).
0924-0136/$ – see front matter © 2006 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2005.12.012
The analytical model presented in this paper is a twodimensional model based on Rosenthal’s model of thermal
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
191
workpiece, which is fairly accurate as long the shoulder stays in contact with the workpiece. The heat generated is given by, 2π µFωr dα = 2πµFωr (2) Q= 0
where F is the downward force applied by the tool on the workpiece, ω the angular velocity of rotation of tool, µ the coefficient of friction between the tool and the workpiece and r is the radius of the tool shoulder. The temperature rise due to stir tool can therefore be calculated from the equation, 2πµFωr −v(ξ+R)/2α Fωr −v(ξ+R)/2α e e = (3) 2πkR kR Kannatey-Asibu [2] shows that the temperature rise due to dual pre-heating source displaced from the origin is given by, T =
Fig. 1. Schematic of FSW coupled to THC.
distribution [7] during welding. Three point heat sources are assumed to be moving on a semi-infinite medium. Here, we consider distributed point sources; however, solutions are available for line source configurations. To simplify the problem, heat transfer by radiation and convection on the surface of the medium is neglected.
The Rosenthal model for temperature distribution of a singlepoint moving source is given by, Q −v(ξ+R)/2α e 2πkR
P1 −v(ξ +Ra )/2α e 2πkRa
(4)
where θ 1 the temperature rise due to pre-heating source closest to the stir tool, P1 the heat input from the first pre-heating source, ξ = ξ − d1 , d1 being the separation distance between the stir tool and the first pre-heating source and Ra = (ξ )2 + y + z2 . Similarly, the rise in temperature due to the second preheating source is given by,
2.1. Calculating the temperature distribution during the welding process
T = T0 +
θ1 =
(1)
where Q is the rate of heat supply by the friction stir tool, k the thermal conductivity of the surface, α the thermal diffusivity, ξ = x − vt (moving coordinate system), T0 the initial temperature of the workpiece, which is considered as 25 ◦ C, T the temperature rise due to minor heat source alone, t the time, v the linear speed of the stir tool (welding speed) and R = ξ 2 + y2 + z2 . Mathematically, R is the radius of a sphere whose value tends to zero as the total heat delivered to the plate reaches the maximum. In the present case, the temperature rise due to the minor heat source, T is considered as the heat generated by the stir tool positioned at the origin of the moving coordinate system.
θ2 =
P e−v(ξ +Rb )/2α 2πkRb
(5)
where θ 2 the temperature rise due to second pre-heating source. ξ = ξ − d2 , d2 being the distance between the stir tool and the
second pre-heating source and Rb = (ξ )2 + y2 + z2 . From Eqs. (1)–(3), it can be concluded that the temperature rise due to the heat generated by the stir tool and the two preheating sources can be given by, 1 Q −v(ξ+R)/2α P −v(ξ +Ra )/2α Θ= e + e 2πk R Ra P −v(ξ +Rb )/2α (6) + e Rb
This equation could be generalized to give the temperature rise due to n number of pre-heating sources. ⎤ ⎡ n √ 1 ⎣ Q −v(ξ+R)/2α Pi 2 2 2 (7) e Θ= + e−(v(ξ−di )+ (ξ−di ) +y +z )/2α ⎦ 2πk R 2 (ξ − di ) + y2 + z2 i=1
The heat generated by the stir tool is mainly through friction between the workpiece and the shoulder of the rotating tool. To a lesser extent there is also a contribution from the tool pin, which depends on the thickness of the plate and the pin length [4]. However, a two-dimensional model, such as this one neglects the effect of thickness. Hence, only the heat generated from the tool shoulder is considered. Feng et al. [3] and Song and Kovacevic [4] calculated the frictional heat input from the tool shoulder. In this case it is assumed that the heat is generated by a line contact between the rim of the tool shoulder and the
where n is the number of pre-heating sources, Pi the pre-heating sources for i = 1–n and di gives the distance of the ith pre-heating source from the stir tool. 2.2. Plotting the analytical model for single heat source To show the benefits of augmenting the FSW process with a THC, we will compare the temperature distribution of FSW with a single-point moving heat source to FSW coupled to a THC with multiple point sources moving at the same velocity as the FSW
192
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
tool. In order to generate contour plots of temperature distribution in case of a single-point moving source, the following values were assumed to arrive at a solution for the equation. The downward force applied on the workpiece by the stir tool, F = 100 N; co-efficient of friction between tool and workpiece, µ = 0.6; radius of tool shoulder, r = 9.525 mm; thermal conductivity of the surface, k = 0.028 W/mm K (the workpiece is assumed as mild steel); density of mild steel, ρ = 7.8 × 10−6 kg/mm3 ; specific heat of mild steel, c = 0.75 kJ/kg K; thermal diffusivity of mild steel, α = k/(ρc) = 4.786 mm2 /s. The plots have a range of −80 to +50 mm along the x-axis (the length of the specimen) and −40 to +40 mm along the y-axis (the width of the specimen). The rotation speed of the stir tool and the welding speed were varied to study the effects they had on the temperature distribution on the workpiece. In the first case the welding speed was fixed at 4 mm/s and the rotation speed of the tool was evaluated at 400 and 3000 rpm. For the next case, the welding speed was reduced to 1 mm/s with the tool rotation speeds being 400 and 3000 rpm. From Figs. 2–5 it can be seen that for FSW without a THC, the heat generated is concentrated over a very small region. For FSW at 400 rpm, the temperature 5 mm ahead of the tool is close to room temperature and increasing the speed to 3000 rpm does not result in any significant decrease in the temperature gradient. Here, the increase in the spindle speed had a negligible impact on the heat distribution around the tool. This is partly due to the low thermal diffusivity of mild steel as compared to aluminum where a significant amount of heat is transferred to the front of the tool during FSW. The negligible change in temperatures ahead of the tool for mild steel suggests that the tool consistently encounters regions of the workpiece that is at room temperature thereby increasing the amount of energy required and subsequently tool wear. For steels, these results suggest that increasing spindle speed only increases wear without any significant benefit to the process. Here, decreasing the welding speed from 4 to 1 mm/s does improve the temperature distribution as
Fig. 3. Temperature distribution for single-point moving source at 3000 rpm and 4 mm/s.
Fig. 4. Temperature distribution for single-point moving source at 400 rpm and 1 mm/s.
shown in Figs. 4 and 5. For a spinning speed of 400 rpm and weld speed of 1 mm/s, the temperature 5 mm ahead of the tool is around 100 ◦ C, and for 3000 rpm it is over 400 ◦ C. This shows that at a welding speed of 1 mm/s the stir tool requires to do lesser amount of work to continue the welding process, thus reducing the wear on itself. However, this comes at the cost of a drastic cut down in the welding speed from 4 to 1 mm/s, which is obviously a costly affair considering the decrease in production rate. 3. Results for multiple point heat sources
Fig. 2. Temperature distribution for single-point moving source at 400 rpm and 4 mm/s.
For FSW combined with THC, the welding speed and tool rotation speed are fixed at 4 mm/s and 400 rpm, respectively. The other parameters used to calculate temperature distribution in a normal FSW are also retained. In this case pre-heating sources
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
Fig. 5. Temperature distribution for single-point moving source at 3000 rpm and 1 mm/s.
are placed ahead of the tool. In a previous section, Eq. (7) was developed to calculate temperature distribution on a workpiece when n number of pre-heating sources is added to a normal FSW process. For the purpose of the current evaluation, two pre-heating sources are used, i.e. n = 2. The distances between the pre-heating source and stir tool (d1 ) and that between the two pre-heating sources (d2 ) are altered to generate different plots. The distance between the first pre-heating source and the stir tool is varied from 10 to 15 mm and the distance between the two pre-heating sources is assumed as half the distance between the stir tool and the first pre-heating source. The distances assumed here are arbitrary. They could be varied depending upon the workpiece material and the pre-heating temperature required to stir weld them successfully. For friction stir welding process with pre-heating sources, the heat input is concentrated in a hot channel ahead of the stir tool. This means that stirring would occur in a warmer region than in typical stir welding. Potentially, the energy demand for plastic work by the tool is considerably reduced with consequences for reduced wear as well. From Figs. 6–8, it can be seen that the temperature of the workpiece around 5 mm ahead of the tool is approximately 1100 ◦ C for d1 = 7.5 mm, 900 ◦ C for d1 = 10 mm and around 500 ◦ C for d1 = 15 mm. In the first two cases that is d1 = 7.5 and 10 mm, the predicted temperatures are around the temperature reported by Thomas [1,5] in his experiments on stir welding steel. This means that if the thermomechanical hot channel ahead of the stir tool were at temperatures as predicted by this model, then less work is required by the stir tool to raise the temperature of workpiece and reduce flow stresses as required for stir welding. On the contrary, in case of the single moving source, even at 3000 rpm, the temperature 5 mm ahead of the tool was little over 50 ◦ C at a similar welding speed as the present case, i.e. 4 mm/s. The reduction of welding speed from 4 to 1 mm/s increased the temperature to around 500 ◦ C. However, this means that weld production rate is reduced four times. Also, at a rotational speed of 3000 rpm, the wear rate in
193
Fig. 6. Temperature distribution for multi-point moving source of heat with d1 = 7.5 mm and d2 = 3.75 mm.
Fig. 7. Temperature distribution for multi-point moving source of heat with d1 = 10 mm and d2 = 5 mm.
Fig. 8. Temperature distribution for multi-point moving source of heat with d1 = 15 mm and d2 = 7.5 mm.
194
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
the tool would be much higher than in case of 400 rpm used in multiple point moving sources. This model thus demonstrates that the method of a thermomechanical hot channel ahead of the stir tool is a viable option for reducing tool wear and also increasing welding speeds. 4. Conclusions The concept of a thermomechanical hot channel for preheating during FSW of high hardness materials like steel is analyzed with respect to its effects on tool life improvement. The results indicate that the presence of pre-heating sources ahead of the stir tool, significantly reduce the temperature gradient, as compared to a conventional FSW. This would lead to lesser amount of work done by the tool, which translates to lesser tool wear and longer tool life. In summary it should be noted that the results indicated in Figs. 6–8 are based only on theoretical parameters. Actual parameters would depend on experimental data for the heat sources used to develop the THCs. For example, it is possible that any changes to the heat input from the pre-heating sources maybe offset by corresponding changes to the heat input from friction via the rotational speed of the tool. Design optimization techniques could provide good estimates for the distances between the sources and the optimal rotational and linear speeds
that could be used for stir welding harder materials like steel to produce longer welds with a consistent weld quality. Acknowledgement This work was supported by the National Science Foundation’s Division of Design, Manufacturing & Industrial Innovation under Award #0343646. References [1] W. Thomas, Friction stir welding of ferrous materials: a feasibility study, in: Proceedings of First International Symposium on Friction Stir Welding, Thousand Oaks, California, 1999. [2] E. Kannatey-Asibu Jr., Thermal aspects of split-beam laser welding concept, J. Eng. Mater. Technol. 113 (1991) 215–221. [3] Z. Feng, J.E. Gould, T.J. Lienert, A heat flow model for friction stir welding of steel, in: Proceedings of Hot Deformation of Aluminum Alloys, 1998, pp. 149–158. [4] M. Song, R. Kovacevic, Thermal modeling of friction stir welding in a moving coordinate system and its validation, Int. J. Mach. Tools Manuf. 43 (2003) 605–615. [5] W. Thomas, Feasibility of friction stir welding steel, Sci. Technol. Weld. Join. 4 (1999) 365–372. [6] T.J. Lienert, W.L. Stellwag Jr., B.B. Grimmett, R.W. Warke, Friction stir welding studies on mild steel, Suppl. Weld. J. (2003) 1s–9s. [7] D. Rosenthal, Mathematical theory of heat distribution during welding and cutting, Weld. J. (1941) 20–25.
A thermomechanical hot channel approach for friction stir welding S. Mandal a , K. Williamson b,∗ a b
Department of Mechanical Engineering, Old Dominion University, Norfolk, VA 23529, USA Department of Industrial Technology, East Carolina University, Greenville, NC 27858, USA
Received 10 June 2004; received in revised form 25 January 2005; accepted 13 December 2005
Abstract This paper provides a theoretical framework for developing a thermomechanical hot channel (THC) approach for augmenting the friction stir welding (FSW) process. The model follows from the Rosenthal solution for moving point sources where heat input from the tool shoulder acts as a warm source while plasmas or laser heat sources provide higher energy input. The THC approach aims at decreasing tool wear by reducing the demand for frictional heat from the tool shoulder and pin. In the proposed model, the THC processing temperature is approximately 1000 ◦ C, which is close to the temperature required for stir welding steel. © 2006 Elsevier B.V. All rights reserved. Keywords: Friction stir welding; Tool wear; Pre-heating; Thermomechanical hot channel
1. Introduction Friction stir welding (FSW) has emerged as an innovative method for joining low melting temperature alloys like aluminum (Al) and magnesium (Mg). The process combines frictional heating with intense plastic deformation to produce cost efficient joints with better mechanical properties than conventional fusion welding techniques. FSW has already been applied to the construction of aluminum structures in both the transportation [1] and aerospace [1] industries, and it is being thoroughly investigated for application to steel structures in the shipbuilding industry. Although most of the research on FSW has been on aluminum, investigations for steel are still relatively new. Laboratory trials carried out on low carbon steel by Thomas [1] from TWI showed that the heat required to induce plastic deformation similar to aluminum is much higher. FSW for steels require the tool to spin at much faster speeds resulting in a far higher tool wear rates for steel joints compared to softer aluminum joints. This factor is highly pronounced for long welds. Thomas [1,5] and Lienert et al. [6] reported significant tool wear during the friction stir welding of steel. In their experiments, Thomas minimized tool wear by using pre-drilled holes, less than the diameter of the tool pin as a means to reduce the high tool wear
during plunging. In their research, Lienert et al. [6] suggested that increased wear during plunging may be due to the high load spikes resulting from the greater flow stress of the cold workpiece. Besides pre-drilling holes to minimize wear, allowing the tool to spin in one position for a sufficient time plasticizes additional workpiece material and creates additional heat input through plastic work. These ongoing experiments reflect significant effort towards the development of more robust FSW techniques for steel that offsets problems with tool wear. Frequent replacement of worn out tools is expensive and it carries additional cost through delays and reduced production rate. Also, the presence of inclusions from worn out tools, in the weld would significantly reduce the quality of the joint. Given the significance of tool wear during the friction stir welding of steel and other harder materials, this study investigates thermomechanical hot channels as an approach for taking FSW beyond the limits of aluminum and other softer materials. Here, the goal is to combine FSW with pre-heating sources that create a thermomechanical hot channel (THC) ahead of the FSW tool. The idea is to pre-heat the workpiece, and reduce the amount of frictional heat and subsequently the tool wear rate (Fig. 1). 2. Theory
∗
Corresponding author. Tel.: +1 252 328 9722; fax: +1 252 328 1618. E-mail address: [email protected] (K. Williamson).
0924-0136/$ – see front matter © 2006 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2005.12.012
The analytical model presented in this paper is a twodimensional model based on Rosenthal’s model of thermal
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
191
workpiece, which is fairly accurate as long the shoulder stays in contact with the workpiece. The heat generated is given by, 2π µFωr dα = 2πµFωr (2) Q= 0
where F is the downward force applied by the tool on the workpiece, ω the angular velocity of rotation of tool, µ the coefficient of friction between the tool and the workpiece and r is the radius of the tool shoulder. The temperature rise due to stir tool can therefore be calculated from the equation, 2πµFωr −v(ξ+R)/2α Fωr −v(ξ+R)/2α e e = (3) 2πkR kR Kannatey-Asibu [2] shows that the temperature rise due to dual pre-heating source displaced from the origin is given by, T =
Fig. 1. Schematic of FSW coupled to THC.
distribution [7] during welding. Three point heat sources are assumed to be moving on a semi-infinite medium. Here, we consider distributed point sources; however, solutions are available for line source configurations. To simplify the problem, heat transfer by radiation and convection on the surface of the medium is neglected.
The Rosenthal model for temperature distribution of a singlepoint moving source is given by, Q −v(ξ+R)/2α e 2πkR
P1 −v(ξ +Ra )/2α e 2πkRa
(4)
where θ 1 the temperature rise due to pre-heating source closest to the stir tool, P1 the heat input from the first pre-heating source, ξ = ξ − d1 , d1 being the separation distance between the stir tool and the first pre-heating source and Ra = (ξ )2 + y + z2 . Similarly, the rise in temperature due to the second preheating source is given by,
2.1. Calculating the temperature distribution during the welding process
T = T0 +
θ1 =
(1)
where Q is the rate of heat supply by the friction stir tool, k the thermal conductivity of the surface, α the thermal diffusivity, ξ = x − vt (moving coordinate system), T0 the initial temperature of the workpiece, which is considered as 25 ◦ C, T the temperature rise due to minor heat source alone, t the time, v the linear speed of the stir tool (welding speed) and R = ξ 2 + y2 + z2 . Mathematically, R is the radius of a sphere whose value tends to zero as the total heat delivered to the plate reaches the maximum. In the present case, the temperature rise due to the minor heat source, T is considered as the heat generated by the stir tool positioned at the origin of the moving coordinate system.
θ2 =
P e−v(ξ +Rb )/2α 2πkRb
(5)
where θ 2 the temperature rise due to second pre-heating source. ξ = ξ − d2 , d2 being the distance between the stir tool and the
second pre-heating source and Rb = (ξ )2 + y2 + z2 . From Eqs. (1)–(3), it can be concluded that the temperature rise due to the heat generated by the stir tool and the two preheating sources can be given by, 1 Q −v(ξ+R)/2α P −v(ξ +Ra )/2α Θ= e + e 2πk R Ra P −v(ξ +Rb )/2α (6) + e Rb
This equation could be generalized to give the temperature rise due to n number of pre-heating sources. ⎤ ⎡ n √ 1 ⎣ Q −v(ξ+R)/2α Pi 2 2 2 (7) e Θ= + e−(v(ξ−di )+ (ξ−di ) +y +z )/2α ⎦ 2πk R 2 (ξ − di ) + y2 + z2 i=1
The heat generated by the stir tool is mainly through friction between the workpiece and the shoulder of the rotating tool. To a lesser extent there is also a contribution from the tool pin, which depends on the thickness of the plate and the pin length [4]. However, a two-dimensional model, such as this one neglects the effect of thickness. Hence, only the heat generated from the tool shoulder is considered. Feng et al. [3] and Song and Kovacevic [4] calculated the frictional heat input from the tool shoulder. In this case it is assumed that the heat is generated by a line contact between the rim of the tool shoulder and the
where n is the number of pre-heating sources, Pi the pre-heating sources for i = 1–n and di gives the distance of the ith pre-heating source from the stir tool. 2.2. Plotting the analytical model for single heat source To show the benefits of augmenting the FSW process with a THC, we will compare the temperature distribution of FSW with a single-point moving heat source to FSW coupled to a THC with multiple point sources moving at the same velocity as the FSW
192
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
tool. In order to generate contour plots of temperature distribution in case of a single-point moving source, the following values were assumed to arrive at a solution for the equation. The downward force applied on the workpiece by the stir tool, F = 100 N; co-efficient of friction between tool and workpiece, µ = 0.6; radius of tool shoulder, r = 9.525 mm; thermal conductivity of the surface, k = 0.028 W/mm K (the workpiece is assumed as mild steel); density of mild steel, ρ = 7.8 × 10−6 kg/mm3 ; specific heat of mild steel, c = 0.75 kJ/kg K; thermal diffusivity of mild steel, α = k/(ρc) = 4.786 mm2 /s. The plots have a range of −80 to +50 mm along the x-axis (the length of the specimen) and −40 to +40 mm along the y-axis (the width of the specimen). The rotation speed of the stir tool and the welding speed were varied to study the effects they had on the temperature distribution on the workpiece. In the first case the welding speed was fixed at 4 mm/s and the rotation speed of the tool was evaluated at 400 and 3000 rpm. For the next case, the welding speed was reduced to 1 mm/s with the tool rotation speeds being 400 and 3000 rpm. From Figs. 2–5 it can be seen that for FSW without a THC, the heat generated is concentrated over a very small region. For FSW at 400 rpm, the temperature 5 mm ahead of the tool is close to room temperature and increasing the speed to 3000 rpm does not result in any significant decrease in the temperature gradient. Here, the increase in the spindle speed had a negligible impact on the heat distribution around the tool. This is partly due to the low thermal diffusivity of mild steel as compared to aluminum where a significant amount of heat is transferred to the front of the tool during FSW. The negligible change in temperatures ahead of the tool for mild steel suggests that the tool consistently encounters regions of the workpiece that is at room temperature thereby increasing the amount of energy required and subsequently tool wear. For steels, these results suggest that increasing spindle speed only increases wear without any significant benefit to the process. Here, decreasing the welding speed from 4 to 1 mm/s does improve the temperature distribution as
Fig. 3. Temperature distribution for single-point moving source at 3000 rpm and 4 mm/s.
Fig. 4. Temperature distribution for single-point moving source at 400 rpm and 1 mm/s.
shown in Figs. 4 and 5. For a spinning speed of 400 rpm and weld speed of 1 mm/s, the temperature 5 mm ahead of the tool is around 100 ◦ C, and for 3000 rpm it is over 400 ◦ C. This shows that at a welding speed of 1 mm/s the stir tool requires to do lesser amount of work to continue the welding process, thus reducing the wear on itself. However, this comes at the cost of a drastic cut down in the welding speed from 4 to 1 mm/s, which is obviously a costly affair considering the decrease in production rate. 3. Results for multiple point heat sources
Fig. 2. Temperature distribution for single-point moving source at 400 rpm and 4 mm/s.
For FSW combined with THC, the welding speed and tool rotation speed are fixed at 4 mm/s and 400 rpm, respectively. The other parameters used to calculate temperature distribution in a normal FSW are also retained. In this case pre-heating sources
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
Fig. 5. Temperature distribution for single-point moving source at 3000 rpm and 1 mm/s.
are placed ahead of the tool. In a previous section, Eq. (7) was developed to calculate temperature distribution on a workpiece when n number of pre-heating sources is added to a normal FSW process. For the purpose of the current evaluation, two pre-heating sources are used, i.e. n = 2. The distances between the pre-heating source and stir tool (d1 ) and that between the two pre-heating sources (d2 ) are altered to generate different plots. The distance between the first pre-heating source and the stir tool is varied from 10 to 15 mm and the distance between the two pre-heating sources is assumed as half the distance between the stir tool and the first pre-heating source. The distances assumed here are arbitrary. They could be varied depending upon the workpiece material and the pre-heating temperature required to stir weld them successfully. For friction stir welding process with pre-heating sources, the heat input is concentrated in a hot channel ahead of the stir tool. This means that stirring would occur in a warmer region than in typical stir welding. Potentially, the energy demand for plastic work by the tool is considerably reduced with consequences for reduced wear as well. From Figs. 6–8, it can be seen that the temperature of the workpiece around 5 mm ahead of the tool is approximately 1100 ◦ C for d1 = 7.5 mm, 900 ◦ C for d1 = 10 mm and around 500 ◦ C for d1 = 15 mm. In the first two cases that is d1 = 7.5 and 10 mm, the predicted temperatures are around the temperature reported by Thomas [1,5] in his experiments on stir welding steel. This means that if the thermomechanical hot channel ahead of the stir tool were at temperatures as predicted by this model, then less work is required by the stir tool to raise the temperature of workpiece and reduce flow stresses as required for stir welding. On the contrary, in case of the single moving source, even at 3000 rpm, the temperature 5 mm ahead of the tool was little over 50 ◦ C at a similar welding speed as the present case, i.e. 4 mm/s. The reduction of welding speed from 4 to 1 mm/s increased the temperature to around 500 ◦ C. However, this means that weld production rate is reduced four times. Also, at a rotational speed of 3000 rpm, the wear rate in
193
Fig. 6. Temperature distribution for multi-point moving source of heat with d1 = 7.5 mm and d2 = 3.75 mm.
Fig. 7. Temperature distribution for multi-point moving source of heat with d1 = 10 mm and d2 = 5 mm.
Fig. 8. Temperature distribution for multi-point moving source of heat with d1 = 15 mm and d2 = 7.5 mm.
194
S. Mandal, K. Williamson / Journal of Materials Processing Technology 174 (2006) 190–194
the tool would be much higher than in case of 400 rpm used in multiple point moving sources. This model thus demonstrates that the method of a thermomechanical hot channel ahead of the stir tool is a viable option for reducing tool wear and also increasing welding speeds. 4. Conclusions The concept of a thermomechanical hot channel for preheating during FSW of high hardness materials like steel is analyzed with respect to its effects on tool life improvement. The results indicate that the presence of pre-heating sources ahead of the stir tool, significantly reduce the temperature gradient, as compared to a conventional FSW. This would lead to lesser amount of work done by the tool, which translates to lesser tool wear and longer tool life. In summary it should be noted that the results indicated in Figs. 6–8 are based only on theoretical parameters. Actual parameters would depend on experimental data for the heat sources used to develop the THCs. For example, it is possible that any changes to the heat input from the pre-heating sources maybe offset by corresponding changes to the heat input from friction via the rotational speed of the tool. Design optimization techniques could provide good estimates for the distances between the sources and the optimal rotational and linear speeds
that could be used for stir welding harder materials like steel to produce longer welds with a consistent weld quality. Acknowledgement This work was supported by the National Science Foundation’s Division of Design, Manufacturing & Industrial Innovation under Award #0343646. References [1] W. Thomas, Friction stir welding of ferrous materials: a feasibility study, in: Proceedings of First International Symposium on Friction Stir Welding, Thousand Oaks, California, 1999. [2] E. Kannatey-Asibu Jr., Thermal aspects of split-beam laser welding concept, J. Eng. Mater. Technol. 113 (1991) 215–221. [3] Z. Feng, J.E. Gould, T.J. Lienert, A heat flow model for friction stir welding of steel, in: Proceedings of Hot Deformation of Aluminum Alloys, 1998, pp. 149–158. [4] M. Song, R. Kovacevic, Thermal modeling of friction stir welding in a moving coordinate system and its validation, Int. J. Mach. Tools Manuf. 43 (2003) 605–615. [5] W. Thomas, Feasibility of friction stir welding steel, Sci. Technol. Weld. Join. 4 (1999) 365–372. [6] T.J. Lienert, W.L. Stellwag Jr., B.B. Grimmett, R.W. Warke, Friction stir welding studies on mild steel, Suppl. Weld. J. (2003) 1s–9s. [7] D. Rosenthal, Mathematical theory of heat distribution during welding and cutting, Weld. J. (1941) 20–25.