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WELDING
44
CHAPTER 5: WELDING 1.
INTRODUCTION
Rosenthal (1946) was the first to determine the temperature distribution during welding of steel, assuming either a point, line or plane heat source moving along the welding axis, neglecting heat losses and variation of physical properties with temperature. His analysis is subject to serious error for temperatures in or near the fusion and heat-affected zones. Recently, it is recognized that fluid motion in weld pools may play an important role in affecting both the heat transfer phenomena and, ultimately, the mechanical properties of the welds produced (for example, Dilawari et al., 1978; Debroy et al., 1980; Oreper and Szekely, 1984). This chapter treats the heat and fluid flow phenomena in the weld pool of both electroslag and arc welding systems and the vaporization rate of laser welding systems.
2.
ELECTROSLAG WELDING
In electroslag welding, the size of the heat-affected zone, the sequence of solidification and the slag metal reactions play a key role in determining the weld properties. The size, shape and nature of the weld pool is a key factor in the production of sound welds. The fluid and heat flow in the weld pool are controlled by both buoyancy-driven natural convection and by an electromagnetic (EM) force field. The buoyancy and EM forces produce vigorous agitation in the molten weld pool. The nature of this agitation is significantly affected by the geometry. When the current field is non-parallel, as in the case with wire electrodes, the EM force field plays a major role, while for parallel current fields, as produced by flat planes, the flow field is driven essentially by buoyancy force. Electroslag welding is used for the welding of thick plates needed in the construction of ships, storage tanks, pressure vessels, bridges, buildings and other such structures. It has been shown to produce relatively defect-free joints at fast deposition rates, not requiring close joint fit-up. The system becomes more efficient with greater plate thickness. However, the ESW process has a major drawback: poor fracture toughness caused by the severe thermal cycle in the heat affected zone which leads to grain coarsening. A schematic of the ESW of two plates is depicted in Fig. 5-1. A consumable wire (or plate) electrode is being fed continuously into a molten slag pool, which is resistively heated by the current passing from the electrode, through the molten slag and the metal pool, to the base plates (1 and 2). Two water cooled copper shoes provide a mold through which a portion of the heat is removed from the system. Figure 5-2 shows a cross-sectional view of the ESW process. The passage of the current through the conducting molten slag and metal phases may generate an EM force field. Additionally, the non-uniformity of the thermal fields within these molten regions gives rise to buoyancy forces. Both the EM and buoyancy fields may generate fluid motion, which in turn could affect the heat transfer rate within the molten regions.
45
;ure 5-1 A schematic of electroslag welding (ESW) process
Figure 5-2 A cross-sectional view of ESW process
46 The governing equations for a melt include: continuity: Δ · ν
=
0
(5-1)
momentum equations:
Ρ ^ = -νΡ + ν · μ ν V + F β
(5-2)
b
EM force vector (F): F = JXB
(5-3)
Maxwell's equations:
f
(5-4)
VXH = J
(5-5)
V· Η=0
(5-6)
VXE =
magnetic flux density vector (B): Β = μ 0Η
(5-7)
Ohm's law: J = g ( Ë + V X B )
E
(5-8)
= -V
(5-9)
heat equation: p C p ^ + V*VT = V*keVT
+
E.J
+ ST
(5-10)
All symbols are listed in the NOMENCLATURE. The molten slag and molten metal pool have separate fluid flow equations coupled by boundary conditions. The electrode, molten slag, metal pool and solidified weld, including the welding plates, have separate EM force field and heat equations which are related through the boundary conditions.
47 The system, in general, is three-dimensional in space and becomes time-dependent due to the movement of the various boundaries. Hence, certain simplifications are made in order to make the problem manageable. Computed results are obtained for two-dimensional systems (Dilawari et al., 1978): a rectangular symmetry with a thin plate electrode producing a rectangular weld between two plates of essentially infinite thickness, and a cylindrical symmetry with a wire electrode producing a cylindrical weld. The depth of the metal pool is taken to be 2/3 of the weld width, consistent with the established ESW practice. Simplifications include: (i) constant physical properties, (ii) laminar flow, (iii) direct current, (iv) the weld and electrode surfaces maintained at the melting temperature, (v) negligible effect of interfacial (slag/pool and pool/ingot) movement, and (vi) neglect of the flow induced by the movement of the electrode and molten drops. Tables 5-1 and 5-2 list the property values employed in the computation of the rectangular and cylindrical systems which are of comparable linear dimensions. The total heat input is identical in the two systems. Figures 5-3, 5-4 and 5-5 are computed streamline pattern, velocity vector and isotherms for (a) rectangular and (b) cylindrical systems (Dilawari et al., 1978). It is seen that in the rectangular system, the flow in the slag and the metal phases is driven exclusively by buoyancy forces. In the cylindrical system, however, the EM force field plays a major role in determining the streamline pattern and velocity. The velocities are about 20 times larger in the slag region and 3 times larger in the metal pool of the cylindrical system than those of the rectangular one. A comparison of the temperature profiles in the two systems suggests the much more quiescent nature of the rectangular system. Figure 5-6 illustrates the distribution of the local heat flux at the weld/slag interface. The rapid convective heat transfer by the electromagnetically driven slag flow causes the marked non-uniform distribution of the heat flux in the cylindrical system. The sharp peak appearing in the top right corner of Fig. 5-6 corresponds to the "undercutting phenomena" in welding practice. In contrast, the heat flux distribution in the rectangular system is much more uniform because the circulation in the slag driven by buoyancy force is much weaker. Table 5-1 Physical property values used in computations of rectangular and cylindrical systems Thermal conductivity of slag, 10.5 χ 10"^k J/m-s-K Thermal conductivity of molten metal, 20.9 χ 10"^ J/m-s-K Density of slag at reference temperature, T 0 , 2.75 χ 10"^ kg/m"3 Density of molten metal at reference temperature, T 0 , 7.2 χ 10"^ kg/m"3 Specific heat of electrode, 0.50 kJ/kg-K Specific heat of molten metal or of metal droplet, 0.75 kJ/kg-K Liquidus temperature of electrode or of welding plate material, 1823 Κ Reference temperature, 350 Κ Latent heat of fusion of electrode, 272 kJ/kg 4
Thermal coefficient of cubical expansion for molten slag, 1.0 χ 10" (K)"
1
Thermal coefficient of cubical expansion for molten metal, 1.0 χ 10"^ (Κ)" Electrical conductivity of molten slag, 2.0 χ 10^ mho/m Electrical conductivity of molten metal, 7.14 χ 10"^ mho/m Emissivity of free slag surface, 0.6 Viscosity of molten slag, 1.0 χ 10"^ kg/m-s Viscosity of molten metal, 6.0 χ 10"-^ kg/m-s
1
48 Table 5-2 Numerical values of parameters used in computations of rectangular and cylindrical systems Electrode radium (half thickness), 1.5 χ Weld pool radius (half thickness), 1.5 χ lO'^m Electrode immersion in slag pool, 1.0 χ lO'^m Depth of slag pool, 1.5 χ lO'^m Depth of mcial pool, 3.0 χ lO'^m Magnetic permeability, 1.26 χ 10"^ henry/m 1
Stcphan-Boltzmann constant, 5.73 χ 10" 1 kJ/m^-s-K Electric potential at the immersed surface of the electrode (rectangular system), 14.52 V Breadth of the rectangular electrode, 0.12 m Electric potential at the immersed surface of electrode (cylindrical system), 47.80 V
An important conclusion is derived: the system geometry has a profound influence of whether the EM or buoyancy forces dominate the flow field. For a rectangular geometry such as a plate electrode, the flow field is controlled by buoyancy forces, resulting in a weaker circulation, leading to a better utilization of the energy input. In contrast, when a cylindrical electrode such as a wire is employed to deposit a cylindrical weldment, the EM force dominated, resulting in a very vigorous circulation, leading to a relatively poor thermal efficiency.
ô
as
io
is
DISTANCE FROM CENTER LUE (cm)
3
°-
6
0
5
0
2
RADIAL DISTANCE (cm)
1 0
Figure 5-3 Computed streamline patterns for (a) rectangular system and (b) cylindrical system
49
Figure 5-4 Computed velocity vectors for (a) rectangular system and (b) cylindrical s y s t e m
Figure 5-5 Computed isotherms for (a) rectangular system and (b) cylindrical system
50
Ο
025 O.SO 0.75 LOO DISTANCE F R O M F R E E S L A G SURFACE
1.25 (cm)
1.50
Figure 5-6 Local heat flux distribution at the weld/slag interface, A for rectangular system and Β for cylindrical system
(6)
Figure 5-7 A schematic of TIG welding process, (a) actual case and (b) ideal case
51
3.
ARC WELDING
A schematic of a typical TIG (tungsten-inert-gas) welding system is shown in Fig. 5-7. An arc is being struck between an inert electrode and a molten weld pool. The heat generated in the arc plasma jet causes the base metal to melt. The solidification of the molten metal in the weld pool forms the bond between the two pieces that are joined by the process. The key physical phenomena are (Oreper and Szekely, 1984): (i)
The arc striking the surface of the metal pool acts both as a distributed heat flux transmitter to the surface and as the source of a spatially non-uniform electric current passing through the pool.
(ii)
The passage of the electric current through the pool induces an EM force field owing to its interaction with the magnetic field thus generated. This force field gives rise to a recirculation motion in the pool.
(iii) The temperature differences in the pool give rise to a buoyancy field which also produces fluid flow. (iv) Surface tension gradients at the free surface and the shear stress exerted by the imping plasma jet may generate fluid motion. (v)
The convective and conductive heat transfer process in the pool bring about the melting of the work piece through the lateral advancement of the melt-solid interface.
Orper and Szekely (1984) considered phenomena (i) through (iv), ignoring (v) in their formulation. Combining Eqs. (5-1) and (5-2) in the stream function-vorticity form and neglecting the last two terms in the heat equation (5-10), numerical results are obtained using the physical property values in Table 5-3. Table 5-3 Physical properties employed in the computation of a TIG welding process Specific heat of molten and solid metal, 753 J/,kg/,K Thermal conductivity of molten metal, 15.48 W m/,K Thermal conductivity of solid metal, 31.39 W m/,K Thickness of metal plate, 3.1 χ 10"2 m Maximum radius of region of calculation, 4.65 χ 10"^ m Temperature, Κ Initial temperature of metal, 300 Κ Liquidus temperature of metal, 1723 Κ Solidus temperature of metal, 1523 Κ dimcnsionlcss temperature (Τ-T0)J(Ti-T0), 4
1
Coefficient of thermal expansion, 10* K" Latent heat of fusion, 247 kJ/,kg Viscosity of molten metal, 0.006 kg/,m/.s
Magnetic permeability of free space, 1.26 χ 10*6 H/,m Density of molten and solid mcial, 7.2 χ 1(P kg/,m^ Electrical conductivity, 7.14 χ 10^, Ι/Ω-m
52 Three basic operating conditions exist: normal mode type, cathode spot type, and surface tension effect. The normal mode type of operation gives a relatively diffuse current and heatflow distribution, while the cathode spot type provides a more sharply focused input of current and heat. These two types, as defined in Table 5-4, assume zero shear stress at the free surface. Table 5-4 Parameters for characterizing the heat flux and electric-current distribution at the surface Parameter
normal mode
heat flux at the axis of symmetry, m" electric-current distribution, m"
2
scale of electric current density, m" heat-flux distribution, rrf
1
2
2
2.2 χ 1 0
7
3.18 χ 1 0 1.9 x 1 0 1.3 χ 1 0
6 2
Cathode m\ mode 6.135 χ 10
4
1 χ 1()
7
5
5.11 χ 1 0 1.3 χ 1 0
6
2
The third type considers the effect of surface tension whose gradients cannot be accurately assessed at present. Only the results for the normal mode type operation are depicted here, in Figs. 5-8, 5-9 and 5-10.
Figure 5-8 Isotherms for purely conductive heat transfer (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s. The numbers on the curves represent the ratio of actual to characteristic temperatures.
53
Figures 5-8 and 5-9 show the computed temperature profiles respectively for pure conduction and combined buoyancy-electromagnetically driven flow.
Figure 5-9 Isotherms for combined buoyancy and electromagnetically driven flow (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s. The numbers on the curves indicate the ratio of actual to characteristic temperatures.
2
ι
m ( m)
Figure 5-10 Velocity fields for combined buoyancy and electromagnetically driven flow (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s.
54
Figure 5-10 shows plots of the computed velocity field corresponding to the temperatures shown in Fig. 5-9. Initially, the flow appears to be dominated by buoyancy forces, but subsequently, double circulating loops appear, suggesting that buoyancy forces play an important role at a distance from the axis of symmetry. The rather weaker counterclockwise circulation pattern in the center of the weld pool is attributed to EM forces, which seem to gain strength with the growth of the weld pool; i.e., during the latter stage of the process. The prediction for the shape of the weld pool is presented in Fig. 5-11, corresponding to the normal mode of operation after 8 sec. of welding. The dotted, dashed and solid lines correspond to the pure conduction, buoyancy-driven flow, and combined buoyancy-electromagnetically driven flow, respectively. The profiles predicted on the basis of these three mechanisms are essentially the same.
0
1
2
χ (mm)
3
4
Figure 5-11 Shape of the weld pool under normal mode operation after 8 s of welding. Dotted, dashed and solid lines represent pure conduction, buoyancy driven and combined buoyancy-electromagnetically driven flow, respectively. For the cathode spot mode of operation, the circulation is counterclockwise, not shown. Thus, the flow field is dominated by EM forces, which are generated by the strongly divergent current field. After a certain time, a deep pool may develop, with relatively high circulating velocities. The greater pool depths may be attributed to the higher intensity of the heat input near the axis of symmetry, combined with the counterclockwise circulation which brings hot fluid from the center of the free surface down the vertical axis of the pool. Surface-tension driven flows are found to have a profound influence on weld-pool behavior. When the surface tension gradient is negative, the resulting radial outflow of hot fluid from the center will by produced in a wider but shallower weld pool. For positive surfacetension gradients, which may occur in the presence of impurities, the resultant counterclockwise circulation will result in a deep penetration of the weld pool. In short, surface-tension effects produce high surface velocities and marked variations in the weld-pool shape.
55
4.
LASER WELDING
The vaporization process is particularly important in the laser welding of alloys containing volatile components because of the potential loss of element from the laser melted pools. The principal factors that govern the alloy element vaporization rates are the temperature distribution at the surface of the molten pool and the composition of the melt. The temperature distribution, in turn, is controlled by the rate of absorption of beam energy by the workpiece, convection in the molten region driven by surface tension and buoyancy forces, etc. The rate of absorption of the beam energy is dependent upon the plasma composition which in turn is affected by the temperature distribution at the surface of the molten pool. Figure 5-12 is a schematic diagram of the experimental setup for the determination of weld pool temperature during the laser welding of AISI 202 stainless steel (Khan and Debroy, 1984). A carbon dioxide laser, Coherent Model Everlase 525-1, capable of producing a maximum output power of 575 watts in the continuous wave mode, was used. Samples were placed on a remotely controlled, electrically operated table capable of providing - linear motion. 2 All welding was carried out inside a Plexiglas box by means of a 2.54 χ 10 m diameter, 0.127 focal length Zn-Se lens with an anti-reflective coating. The total rate of alloying element vaporization was evaluated from the measured values of the loss sample weight resulting from element vaporization and the laser-material interaction time. The experiment also yielded plasma concentration and changes in the weld composition during the laser welding of AISI 202 stainless steel. Figure 5-13 depicts the 3measured rate of vaporization as a function of laser 4power at the 3 welding speed of 3.5 χ 10~ m/s and the shielding gas flow rate of 1.0 χ 10' m /s. The increase in the overall vaporization rate with an increase in the laser power is due to two factors. First, the higher is the laser power, the higher is the surface area of the molten zone where vaporization can occur. Secondly, the vaporization is a strong function of temperature. The high vaporization rate at high laser power is due to the fact that the weld pool temperature increases with an increase in laser power.
LASER
BEAM
STATIONARY Q U A R T Z TUBE
\-i—SAMPLE
ο
Figure 5-12 A schematic of experimental setup for laser welding.
200
300
400
500
600
Figure 5-13 A plot of measured vaporization rate versus laser power for AISI 202 stainless steel at a welding 3 speed of 3.5 χ 10" m/s and a shielding gas flow rate of 4 1 χ ΙΟ" m V s
56
5.
REFERENCES
David, S.A. and Vitek, J.M. (eds.) (1990), Recent Trends in Welding Science and Technology, ASM International, Material Park, OH. Khan, P.A.A. and Debroy, T. (1984), "Alloying Element Vaporization and Weld Pool Temperature during Laser Welding of AISI 202 Stainless Steel," Metallurgical Trans. B, Vol. 15B, pp. 641 - 644.
CHAPTER 5: WELDING 1.
INTRODUCTION
Rosenthal (1946) was the first to determine the temperature distribution during welding of steel, assuming either a point, line or plane heat source moving along the welding axis, neglecting heat losses and variation of physical properties with temperature. His analysis is subject to serious error for temperatures in or near the fusion and heat-affected zones. Recently, it is recognized that fluid motion in weld pools may play an important role in affecting both the heat transfer phenomena and, ultimately, the mechanical properties of the welds produced (for example, Dilawari et al., 1978; Debroy et al., 1980; Oreper and Szekely, 1984). This chapter treats the heat and fluid flow phenomena in the weld pool of both electroslag and arc welding systems and the vaporization rate of laser welding systems.
2.
ELECTROSLAG WELDING
In electroslag welding, the size of the heat-affected zone, the sequence of solidification and the slag metal reactions play a key role in determining the weld properties. The size, shape and nature of the weld pool is a key factor in the production of sound welds. The fluid and heat flow in the weld pool are controlled by both buoyancy-driven natural convection and by an electromagnetic (EM) force field. The buoyancy and EM forces produce vigorous agitation in the molten weld pool. The nature of this agitation is significantly affected by the geometry. When the current field is non-parallel, as in the case with wire electrodes, the EM force field plays a major role, while for parallel current fields, as produced by flat planes, the flow field is driven essentially by buoyancy force. Electroslag welding is used for the welding of thick plates needed in the construction of ships, storage tanks, pressure vessels, bridges, buildings and other such structures. It has been shown to produce relatively defect-free joints at fast deposition rates, not requiring close joint fit-up. The system becomes more efficient with greater plate thickness. However, the ESW process has a major drawback: poor fracture toughness caused by the severe thermal cycle in the heat affected zone which leads to grain coarsening. A schematic of the ESW of two plates is depicted in Fig. 5-1. A consumable wire (or plate) electrode is being fed continuously into a molten slag pool, which is resistively heated by the current passing from the electrode, through the molten slag and the metal pool, to the base plates (1 and 2). Two water cooled copper shoes provide a mold through which a portion of the heat is removed from the system. Figure 5-2 shows a cross-sectional view of the ESW process. The passage of the current through the conducting molten slag and metal phases may generate an EM force field. Additionally, the non-uniformity of the thermal fields within these molten regions gives rise to buoyancy forces. Both the EM and buoyancy fields may generate fluid motion, which in turn could affect the heat transfer rate within the molten regions.
45
;ure 5-1 A schematic of electroslag welding (ESW) process
Figure 5-2 A cross-sectional view of ESW process
46 The governing equations for a melt include: continuity: Δ · ν
=
0
(5-1)
momentum equations:
Ρ ^ = -νΡ + ν · μ ν V + F β
(5-2)
b
EM force vector (F): F = JXB
(5-3)
Maxwell's equations:
f
(5-4)
VXH = J
(5-5)
V· Η=0
(5-6)
VXE =
magnetic flux density vector (B): Β = μ 0Η
(5-7)
Ohm's law: J = g ( Ë + V X B )
E
(5-8)
= -V
(5-9)
heat equation: p C p ^ + V*VT = V*keVT
+
E.J
+ ST
(5-10)
All symbols are listed in the NOMENCLATURE. The molten slag and molten metal pool have separate fluid flow equations coupled by boundary conditions. The electrode, molten slag, metal pool and solidified weld, including the welding plates, have separate EM force field and heat equations which are related through the boundary conditions.
47 The system, in general, is three-dimensional in space and becomes time-dependent due to the movement of the various boundaries. Hence, certain simplifications are made in order to make the problem manageable. Computed results are obtained for two-dimensional systems (Dilawari et al., 1978): a rectangular symmetry with a thin plate electrode producing a rectangular weld between two plates of essentially infinite thickness, and a cylindrical symmetry with a wire electrode producing a cylindrical weld. The depth of the metal pool is taken to be 2/3 of the weld width, consistent with the established ESW practice. Simplifications include: (i) constant physical properties, (ii) laminar flow, (iii) direct current, (iv) the weld and electrode surfaces maintained at the melting temperature, (v) negligible effect of interfacial (slag/pool and pool/ingot) movement, and (vi) neglect of the flow induced by the movement of the electrode and molten drops. Tables 5-1 and 5-2 list the property values employed in the computation of the rectangular and cylindrical systems which are of comparable linear dimensions. The total heat input is identical in the two systems. Figures 5-3, 5-4 and 5-5 are computed streamline pattern, velocity vector and isotherms for (a) rectangular and (b) cylindrical systems (Dilawari et al., 1978). It is seen that in the rectangular system, the flow in the slag and the metal phases is driven exclusively by buoyancy forces. In the cylindrical system, however, the EM force field plays a major role in determining the streamline pattern and velocity. The velocities are about 20 times larger in the slag region and 3 times larger in the metal pool of the cylindrical system than those of the rectangular one. A comparison of the temperature profiles in the two systems suggests the much more quiescent nature of the rectangular system. Figure 5-6 illustrates the distribution of the local heat flux at the weld/slag interface. The rapid convective heat transfer by the electromagnetically driven slag flow causes the marked non-uniform distribution of the heat flux in the cylindrical system. The sharp peak appearing in the top right corner of Fig. 5-6 corresponds to the "undercutting phenomena" in welding practice. In contrast, the heat flux distribution in the rectangular system is much more uniform because the circulation in the slag driven by buoyancy force is much weaker. Table 5-1 Physical property values used in computations of rectangular and cylindrical systems Thermal conductivity of slag, 10.5 χ 10"^k J/m-s-K Thermal conductivity of molten metal, 20.9 χ 10"^ J/m-s-K Density of slag at reference temperature, T 0 , 2.75 χ 10"^ kg/m"3 Density of molten metal at reference temperature, T 0 , 7.2 χ 10"^ kg/m"3 Specific heat of electrode, 0.50 kJ/kg-K Specific heat of molten metal or of metal droplet, 0.75 kJ/kg-K Liquidus temperature of electrode or of welding plate material, 1823 Κ Reference temperature, 350 Κ Latent heat of fusion of electrode, 272 kJ/kg 4
Thermal coefficient of cubical expansion for molten slag, 1.0 χ 10" (K)"
1
Thermal coefficient of cubical expansion for molten metal, 1.0 χ 10"^ (Κ)" Electrical conductivity of molten slag, 2.0 χ 10^ mho/m Electrical conductivity of molten metal, 7.14 χ 10"^ mho/m Emissivity of free slag surface, 0.6 Viscosity of molten slag, 1.0 χ 10"^ kg/m-s Viscosity of molten metal, 6.0 χ 10"-^ kg/m-s
1
48 Table 5-2 Numerical values of parameters used in computations of rectangular and cylindrical systems Electrode radium (half thickness), 1.5 χ Weld pool radius (half thickness), 1.5 χ lO'^m Electrode immersion in slag pool, 1.0 χ lO'^m Depth of slag pool, 1.5 χ lO'^m Depth of mcial pool, 3.0 χ lO'^m Magnetic permeability, 1.26 χ 10"^ henry/m 1
Stcphan-Boltzmann constant, 5.73 χ 10" 1 kJ/m^-s-K Electric potential at the immersed surface of the electrode (rectangular system), 14.52 V Breadth of the rectangular electrode, 0.12 m Electric potential at the immersed surface of electrode (cylindrical system), 47.80 V
An important conclusion is derived: the system geometry has a profound influence of whether the EM or buoyancy forces dominate the flow field. For a rectangular geometry such as a plate electrode, the flow field is controlled by buoyancy forces, resulting in a weaker circulation, leading to a better utilization of the energy input. In contrast, when a cylindrical electrode such as a wire is employed to deposit a cylindrical weldment, the EM force dominated, resulting in a very vigorous circulation, leading to a relatively poor thermal efficiency.
ô
as
io
is
DISTANCE FROM CENTER LUE (cm)
3
°-
6
0
5
0
2
RADIAL DISTANCE (cm)
1 0
Figure 5-3 Computed streamline patterns for (a) rectangular system and (b) cylindrical system
49
Figure 5-4 Computed velocity vectors for (a) rectangular system and (b) cylindrical s y s t e m
Figure 5-5 Computed isotherms for (a) rectangular system and (b) cylindrical system
50
Ο
025 O.SO 0.75 LOO DISTANCE F R O M F R E E S L A G SURFACE
1.25 (cm)
1.50
Figure 5-6 Local heat flux distribution at the weld/slag interface, A for rectangular system and Β for cylindrical system
(6)
Figure 5-7 A schematic of TIG welding process, (a) actual case and (b) ideal case
51
3.
ARC WELDING
A schematic of a typical TIG (tungsten-inert-gas) welding system is shown in Fig. 5-7. An arc is being struck between an inert electrode and a molten weld pool. The heat generated in the arc plasma jet causes the base metal to melt. The solidification of the molten metal in the weld pool forms the bond between the two pieces that are joined by the process. The key physical phenomena are (Oreper and Szekely, 1984): (i)
The arc striking the surface of the metal pool acts both as a distributed heat flux transmitter to the surface and as the source of a spatially non-uniform electric current passing through the pool.
(ii)
The passage of the electric current through the pool induces an EM force field owing to its interaction with the magnetic field thus generated. This force field gives rise to a recirculation motion in the pool.
(iii) The temperature differences in the pool give rise to a buoyancy field which also produces fluid flow. (iv) Surface tension gradients at the free surface and the shear stress exerted by the imping plasma jet may generate fluid motion. (v)
The convective and conductive heat transfer process in the pool bring about the melting of the work piece through the lateral advancement of the melt-solid interface.
Orper and Szekely (1984) considered phenomena (i) through (iv), ignoring (v) in their formulation. Combining Eqs. (5-1) and (5-2) in the stream function-vorticity form and neglecting the last two terms in the heat equation (5-10), numerical results are obtained using the physical property values in Table 5-3. Table 5-3 Physical properties employed in the computation of a TIG welding process Specific heat of molten and solid metal, 753 J/,kg/,K Thermal conductivity of molten metal, 15.48 W m/,K Thermal conductivity of solid metal, 31.39 W m/,K Thickness of metal plate, 3.1 χ 10"2 m Maximum radius of region of calculation, 4.65 χ 10"^ m Temperature, Κ Initial temperature of metal, 300 Κ Liquidus temperature of metal, 1723 Κ Solidus temperature of metal, 1523 Κ dimcnsionlcss temperature (Τ-T0)J(Ti-T0), 4
1
Coefficient of thermal expansion, 10* K" Latent heat of fusion, 247 kJ/,kg Viscosity of molten metal, 0.006 kg/,m/.s
Magnetic permeability of free space, 1.26 χ 10*6 H/,m Density of molten and solid mcial, 7.2 χ 1(P kg/,m^ Electrical conductivity, 7.14 χ 10^, Ι/Ω-m
52 Three basic operating conditions exist: normal mode type, cathode spot type, and surface tension effect. The normal mode type of operation gives a relatively diffuse current and heatflow distribution, while the cathode spot type provides a more sharply focused input of current and heat. These two types, as defined in Table 5-4, assume zero shear stress at the free surface. Table 5-4 Parameters for characterizing the heat flux and electric-current distribution at the surface Parameter
normal mode
heat flux at the axis of symmetry, m" electric-current distribution, m"
2
scale of electric current density, m" heat-flux distribution, rrf
1
2
2
2.2 χ 1 0
7
3.18 χ 1 0 1.9 x 1 0 1.3 χ 1 0
6 2
Cathode m\ mode 6.135 χ 10
4
1 χ 1()
7
5
5.11 χ 1 0 1.3 χ 1 0
6
2
The third type considers the effect of surface tension whose gradients cannot be accurately assessed at present. Only the results for the normal mode type operation are depicted here, in Figs. 5-8, 5-9 and 5-10.
Figure 5-8 Isotherms for purely conductive heat transfer (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s. The numbers on the curves represent the ratio of actual to characteristic temperatures.
53
Figures 5-8 and 5-9 show the computed temperature profiles respectively for pure conduction and combined buoyancy-electromagnetically driven flow.
Figure 5-9 Isotherms for combined buoyancy and electromagnetically driven flow (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s. The numbers on the curves indicate the ratio of actual to characteristic temperatures.
2
ι
m ( m)
Figure 5-10 Velocity fields for combined buoyancy and electromagnetically driven flow (normal mode), at (a) 2.5 s, (b) 5 s, and (c) 8 s.
54
Figure 5-10 shows plots of the computed velocity field corresponding to the temperatures shown in Fig. 5-9. Initially, the flow appears to be dominated by buoyancy forces, but subsequently, double circulating loops appear, suggesting that buoyancy forces play an important role at a distance from the axis of symmetry. The rather weaker counterclockwise circulation pattern in the center of the weld pool is attributed to EM forces, which seem to gain strength with the growth of the weld pool; i.e., during the latter stage of the process. The prediction for the shape of the weld pool is presented in Fig. 5-11, corresponding to the normal mode of operation after 8 sec. of welding. The dotted, dashed and solid lines correspond to the pure conduction, buoyancy-driven flow, and combined buoyancy-electromagnetically driven flow, respectively. The profiles predicted on the basis of these three mechanisms are essentially the same.
0
1
2
χ (mm)
3
4
Figure 5-11 Shape of the weld pool under normal mode operation after 8 s of welding. Dotted, dashed and solid lines represent pure conduction, buoyancy driven and combined buoyancy-electromagnetically driven flow, respectively. For the cathode spot mode of operation, the circulation is counterclockwise, not shown. Thus, the flow field is dominated by EM forces, which are generated by the strongly divergent current field. After a certain time, a deep pool may develop, with relatively high circulating velocities. The greater pool depths may be attributed to the higher intensity of the heat input near the axis of symmetry, combined with the counterclockwise circulation which brings hot fluid from the center of the free surface down the vertical axis of the pool. Surface-tension driven flows are found to have a profound influence on weld-pool behavior. When the surface tension gradient is negative, the resulting radial outflow of hot fluid from the center will by produced in a wider but shallower weld pool. For positive surfacetension gradients, which may occur in the presence of impurities, the resultant counterclockwise circulation will result in a deep penetration of the weld pool. In short, surface-tension effects produce high surface velocities and marked variations in the weld-pool shape.
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4.
LASER WELDING
The vaporization process is particularly important in the laser welding of alloys containing volatile components because of the potential loss of element from the laser melted pools. The principal factors that govern the alloy element vaporization rates are the temperature distribution at the surface of the molten pool and the composition of the melt. The temperature distribution, in turn, is controlled by the rate of absorption of beam energy by the workpiece, convection in the molten region driven by surface tension and buoyancy forces, etc. The rate of absorption of the beam energy is dependent upon the plasma composition which in turn is affected by the temperature distribution at the surface of the molten pool. Figure 5-12 is a schematic diagram of the experimental setup for the determination of weld pool temperature during the laser welding of AISI 202 stainless steel (Khan and Debroy, 1984). A carbon dioxide laser, Coherent Model Everlase 525-1, capable of producing a maximum output power of 575 watts in the continuous wave mode, was used. Samples were placed on a remotely controlled, electrically operated table capable of providing - linear motion. 2 All welding was carried out inside a Plexiglas box by means of a 2.54 χ 10 m diameter, 0.127 focal length Zn-Se lens with an anti-reflective coating. The total rate of alloying element vaporization was evaluated from the measured values of the loss sample weight resulting from element vaporization and the laser-material interaction time. The experiment also yielded plasma concentration and changes in the weld composition during the laser welding of AISI 202 stainless steel. Figure 5-13 depicts the 3measured rate of vaporization as a function of laser 4power at the 3 welding speed of 3.5 χ 10~ m/s and the shielding gas flow rate of 1.0 χ 10' m /s. The increase in the overall vaporization rate with an increase in the laser power is due to two factors. First, the higher is the laser power, the higher is the surface area of the molten zone where vaporization can occur. Secondly, the vaporization is a strong function of temperature. The high vaporization rate at high laser power is due to the fact that the weld pool temperature increases with an increase in laser power.
LASER
BEAM
STATIONARY Q U A R T Z TUBE
\-i—SAMPLE
ο
Figure 5-12 A schematic of experimental setup for laser welding.
200
300
400
500
600
Figure 5-13 A plot of measured vaporization rate versus laser power for AISI 202 stainless steel at a welding 3 speed of 3.5 χ 10" m/s and a shielding gas flow rate of 4 1 χ ΙΟ" m V s
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5.
REFERENCES
David, S.A. and Vitek, J.M. (eds.) (1990), Recent Trends in Welding Science and Technology, ASM International, Material Park, OH. Khan, P.A.A. and Debroy, T. (1984), "Alloying Element Vaporization and Weld Pool Temperature during Laser Welding of AISI 202 Stainless Steel," Metallurgical Trans. B, Vol. 15B, pp. 641 - 644.