Numerical simulation on direct chilled continuous casting process for

Numerical simulation on direct chilled continuous casting process for preparing clad slab of aluminium alloys L. Wu, T. M. Wang*, J. Li, H. Chen, J. B. Sun and Z. Q. Cao The numerical simulation is used to investigate the influence of processing parameters on direct chilled continuous casting process for preparing clad slab of Al–1Mn and Al–10Si alloys. In order to get a clear bonding interface avoiding unexpected mixing, a special dividing plate is set between molten clad materials. The effects of cooling intensity and cooled part height of dividing plate, casting speed and pouring temperature on the temperature field, liquid fraction, and solidification shell of clad slab have been studied in detail. A practical experiment has been done based on the simulation studies. It was found that a clad slab with excellent metallurgical bonding can be obtained when the cooling water flow of dividing plate is 250 L h21, the cooled part height of dividing plate is 20 mm, the casting speed is 80 mm min21 and the pouring temperatures of Al–1Mn and Al–10Si alloys are 710 and 670uC respectively. Keywords: Numerical simulation, Continuous casting, Clad slab, Aluminium alloy

Introduction As a new type of material, clad metals are widely used in many fields such as aerospace, petroleum, chemical, automobile and shipbuilding industries and have become one of research focuses in material field because they possess excellent physical, chemical and mechanical properties that conventional monolithic alloys do not have.1,2 These metals are conventionally manufactured by roll bonding,3 explosive welding,4 diffusion bonding,5 extrusion cladding,6 and spray deposition technique.7 The casting method to prepare clad metals is more efficient and economical than other processes. Therefore, this method has drawn great attention in recent years. Wagstaff et al.8 developed a technology referred to as fusion technology that produced clad material by direct chill casting. Jiang et al.9 prepared three-layer composite ingot of 4045/3004/4045 aluminium alloys by direct chill semi-continuous casting process. Fu et al.10 prepared aluminium alloy circular clad ingots by an innovative direct chill casting process. In recent works, we have proposed a direct chill continuous casting method with a special dividing plate to prepare clad ingot of Al–1Mn (wt-%) alloy and Al– 10Si (wt-%) alloy. The Al–1Mn alloy has excellent corrosion resistance and low strength, while the Al–10Si alloy has high strength, weldability and poor corrosion resistance. Therefore, the Al–1Mn and Al–10Si clad ingot can combine their advantages.2 However, the direct chilled continuous casting process for preparing School of Materials Science and Engineering, Dalian University of Technology, Dalian 116024, China *Correspondence author, email [email protected]

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ß 2014 W. S. Maney & Son Ltd. Received 23 April 2013; accepted 5 January 2014 DOI 10.1179/1743133614Y.0000000104

clad slab is complicated because of undetermined processing parameters. In recent years, decreasing computational costs and increasing power of commercial modelling packages is making it easier to apply numerical modelling as an additional tool to understand the continuous casting process. Many scholars have researched the heat transfer, fluid flow and solidification phenomenon of casting process by numerical simulation.11–18 Zhang et al.14 developed a comprehensive mathematical model to describe the interaction of the multiple physics fields during the conventional DC casting process of 7XXX aluminium alloys. Feng et al.17 did numerical simulation of the temperature field in high speed steel composite roll during continuous pouring process for cladding. Sun et al.18 numerically studied the influence of casting temperature and casting speed on temperature field during the fabrication of clad slab by semi-continuous casting. In order to optimise the processing parameters of the proposed method, case study by simulation is done to investigate the effects of pouring temperature, casting speed, cooling intensity and cooled part height of the dividing plate on the direct chilled continuous casting process in this paper.

Experimental Figure 1 shows the sketch of direct chilled continuous casting process for preparing clad slab of aluminium alloys. The right hand part of Fig. 1 is the structural diagram of the dividing plate with different cooling conditions in the two sides. The dividing plate was placed into the mould between molten clad materials and the cooling water flow in it was controllable. The

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Numerical simulation on direct chilled continuous casting process

1, liquid alloys control device; 2, Al–1Mn alloy melt; 3, Al–1Mn alloy solidified shell; 4, dividing plate; 5, Al–10Si alloy melt; 6, sprues; 7, mould; 8, secondary cooling water; 9, semi-solid Al–10Si alloy; 10, interface; 11, bimetal slab; 12, starter block; 13, heat insulation layer; 14, stainless steel; 15, cooling water 1 Schematic illustration of continuous casting of bimetal slab

cooled part of the dividing plate acted as a mould wall. The Al–1Mn melt was poured into the left section of the mould first. Once the starter block was drawn, the Al– 10Si melt was fed to the right section of the mould to contact with the semi-solid surface of the Al–1Mn alloy below the dividing plate. The technological parameters were adjusted to maintain an excellent metallurgical bond at the interface during the steady state period of the continuous casting.

k2 (3) e where Cm is a function of the turbulent Reynolds number which is a constant value; k and e are the turbulence kinetic energy and its rate of dissipation respectively, and obtained from the following transport equations: Turbulent kinetic energy equation LðrkÞ Lðrui kÞ L m Lk z zG{re (4) ~ mz t Lt Lxi Lxi sk Lxi

Simulation model

where

mt ~rCm

Assumptions In order to simplify the calculation, some assumptions are made as follows: (i) the molten aluminium alloys behave as an incompressible fluid (ii) ignore the interface reaction between Al–1Mn and Al–10Si alloys (iii) there is no diffusion in solid phase (iv) ignore the effect of liquid surface fluctuations in the flow.

sk51?0 Dissipation rate of turbulent kinetic energy LðreÞ Lðrui eÞ L m Le e e2 z (5) zc1 G{c2 r ~ mz t Lt Lxi Lxi k se Lxi k where c151?44, c251?92, cm50?09, sk51?0, se51?3

Energy equation

Continuity equation

Momentum equation Lðrui Þ L rui uj Lp L Lui z ~rgi { z meff Lt Lxi Lxi Lxi Lxi

Lui Lmi Lmj z Lxj Lxj Lxi

Governing equations for thermal field

Governing equations for flow field L(rui ) ~0 Lxi

G~ut

(1)

(2)

where r is the density of liquid melt, ui and uj are time average velocity, P is the pressure, meff is the effective viscosity coefficient meff5mlzmt, ml is laminar viscosity, mt is turbulent viscosity, gi is gravitational acceleration. The standard k–e model, which is a semi-empirical model, is used to model transport of turbulence kinetic energy k and its dissipation rate e. The first form of this model is

LðrH Þ z+ðr~ vH Þ~+ðk+T Þ Lt

(6)

The enthalpy of the material ÐH is computed as the T sum of the sensible enthalpy h~ Tref cp dT and the latent heat DH~fL DHf in the energy equation as 8 TwTL > ðT Tref : L S 0 TvTS where ~ v is the fluid velocity, k is the thermal conductivity, Tref is the reference temperature, Cp is specific heat at constant pressure, DHf is the latent heat of the material, fL is the liquid fraction, and TL and TS are liquidus and solidus temperatures respectively. The


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Table 1 Parameters of material properties Parameters

Al–1Mn

Al–10Si

Density/kg m23 Viscosity/Pa s Specific heat/J kg21 K21 Thermal conductivity/W m21 K21 Latent heat of fusion/kJ kg21 Solidus temperature/K Liquidus temperature/K

2750 0.0013 893 193 386 916 927

2710 0.0013 864 155 481.5 830 846.5

symmetry plane as shown in Fig. 3b. When flow field and thermal field are solved base on FLUENT, the boundary conditions are described as follows: Inlet and outlet boundary

The velocity boundary condition is used in the region. Free surface boundary

2 1/2 symmetrical computational grid

The boundary condition is treated as the static adiabatic wall.

latent heat is released uniformly between the liquidus and solidus temperatures. In other words, the enthalpy changes linearly over the solidification range. The release of latent heat within the mushy region is properly incorporated in the energy equation as a source term.

Mould cooling boundary

The thermal boundary condition is treated as Cauchytype boundary condition, and the heat transfer coefficient between the mould and the metal is assumed to vary with the solid fraction and is written according to equation (8)

Physical properties and boundary conditions In this paper, the commercial software FLUENT was used to solve the flow field and thermal field. For the centrosymmetric distribution of aluminium alloys in width direction of the mould, a 1/2 symmetry model was used to reduce the calculation amount, as shown in Fig. 2. The size of the mould and watercooled divider are 15061606500 and 150615655 mm respectively. The model contains 17364 volume elements. The physical properties of Al–1Mn and Al–10Si alloys are given in Table 1. Figure 3a and b shows the front view and top view of the clad slab geometry model respectively. The position of the symmetry plane is shown in Fig. 2 and the sprue plane passing through the sprues is parallel to the

h~hcontact ð1{fS Þzhgap fS

where hcontact is intended to reflect good thermal contact between the mould, hgap reflects poor thermal contact associated with the gap formed when the metal is solidified and fS is the solid fraction and is computed as fS51–fL. Second cooling boundary

In terms of the thermal boundary condition, the secondary cooling boundary is divided into two zones, that is, the air cooled zone and the water cooled zone. The heat transfer coefficient in air cooled zone is given according to equation (9)

3 a front view and b top view of clad slab geometry model

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a symmetry plane, 4000 W m22 K21; b sprue plane, 4000 W m22 K21; c symmetry plane, 5000 W m22 K21; d sprue plane, 5000 W m22 K21 4 Influence of cooling intensity of dividing plate on temperature field of clad slab

hair ~seT 3

(9)

where s is Stefan–Boltzmann constant, e is blackness of the clad slab surface and T is temperature of the clad slab surface. Based on the empirical model of Weckman and Niessen,19 the heat transfer coefficient in water cooled zone is given as follow h~ 8 : > :{1:67|105 z704T Q1=3 ðTw2000 CÞ (10) where T- is the average temperature of water, DT is the temperature difference between ingot surface and cooling water, DTx is the difference between the ingot surface temperature and the water saturation temperature (at this temperature, 372?8 K, the cooling water will boil off the surface of the clad slab); Q is water flowrate.

Results and discussion The direct chilled continuous casting process has been simulated to understand the distribution of thermal and flow fields. The distribution of liquid fraction near the dividing plate is determined by the distribution of temperature field. The status of solidification, feature

of the sumps and thickness of solidification shell can be also obtained, which have important effects on the metallurgical bonding of the clad slab.

Effects of cooling intensity of dividing plate The influence of cooling intensity of dividing plate hd on the temperature field of clad slab is shown in Fig. 4. Cooling intensity is controlled by adjusting the cooling water flow in dividing plate. The cooling intensity is 4000 and 5000 W m22 K21 when the cooling water flow is 180 and 250 L h21 respectively. The temperature at sprue plane is higher than that at symmetry plane at the same vertical height because of the high temperature melt poured continuously from the gates. The cooling intensity increase results in a decrease on temperature around the dividing plate. Figure 5 shows the influence of cooling intensity of dividing plate on the liquid fraction of clad slab. The thickness of solidification shell at the sprue plane is smaller than that at the symmetry plane and the depth of the sump at the sprue plane is deeper than that at the symmetry plane. While the shell width of Al–1Mn alloy inside the mould is two times the size of Al–10Si alloy, their depth of the sump is basically the same. This is because Al–1Mn melt is cooled by both the mould and the cooled part of dividing plate but Al–10Si melt is cooled by the mould only. Compared with the Fig. 5b and d, it is found that the sump of Al–1Mn melt

a symmetry plane, 4000 W m22 K21; b sprue plane, 4000 W m22 K21; c symmetry plane, 5000 W m22 K21; d sprue plane, 5000 W m22 K21 5 Influence of cooling intensity of dividing plate on liquid fraction of clad slab


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a symmetry plane, 20 mm; b sprue plane, 20 mm; c symmetry plane, 30 mm; d sprue plane, 30 mm 6 Influence of cooled part height of dividing plate on temperature field of clad slab

a symmetry plane, 20 mm; b sprue plane, 20 mm; c symmetry plane, 30 mm; d sprue plane, 30 mm 7 Influence of cooled part height of dividing plate on liquid fraction of clad slab

Al–1Mn alloy at the horizontal section beneath the dividing plate. The thickness of solidification shell is defined as the horizontal distance between the dividing plate bottom and the liquid fraction contour line of 0?1, as shown in Fig. 5a. In details, by increasing the cooling intensity from 4000 to 5000 W m22 K21, the thickness of Al–1Mn solidification shell formed on the dividing plate is increased by 4 mm at the sprue plane as well as that at the symmetry plane. That is to say, increasing the cooling intensity has the same effect on the solidification shell thickness of Al–1Mn alloy at the different position along the dividing plate.

becomes shallower and flatter when the cooling intensity increases. The thickness of Al–1Mn solidification shell formed on the dividing plate becomes larger with the cooling intensity increases. When the Al–1Mn solidification shell leaves the dividing plate and contacts with the Al–10Si melt, it remelts partly and the liquid fraction increases. Actually, the increase of liquid fraction near the interface will improve the diffusion rate of alloy elements, which can enhance higher metallurgical bond strength at the interface. Figure 10a shows the influence of cooling intensity of dividing plate on the solidification shell thickness of

a symmetry plane, 70 mm min21; b sprue plane, 70 mm min21; c symmetry plane, 80 mm min21; b sprue plane, 80 mm min21 8 Influence of casting speed on temperature field of clad slab

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a symmetry plane, 70 mm min21; b sprue plane, 70 mm min21; c symmetry plane, 80 mm min21; d sprue plane, 80 mm min21 9 Influence of casting speed on liquid fraction of clad slab

Effects of cooled part height of dividing plate Figure 6 shows the influence of cooled part height of the dividing plate Hd on the temperature field of clad slab. It is indicated that the surface temperature of the Al–1Mn solidification shell formed on the dividing plate decreases obviously when the cooled part height of the dividing plate is increased from 20 to 30 mm. This is because the contact time of Al–1Mn melt and cooled part of dividing plate increases when the continuous casting process is proceeding. The thickness of Al–1Mn

solidification shell formed on the dividing plate becomes larger with the surface temperature decreases, as shown in Fig. 7. Actually, thick solidification shell reduces the diffusion rate of alloy elements, which will result in lower metallurgical bond strength at the interface. Figure 10b shows the influence of cooled part height of the dividing plate on the solidification shell thickness of Al–1Mn alloy at the horizontal section beneath the dividing plate. In details, by increasing the cooled part height from 20 to 30 mm, the thickness of Al–1Mn

10 Influence of a cooling intensity and b cooled part height of the dividing plate, c casting speed and d pouring temperature on solidification shell thickness of Al–1Mn alloy at horizontal section beneath dividing plate


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12 Tensile fractured specimen

dividing plate is decreased by 4 mm at the sprue plane and 6 mm at the symmetry plane respectively. That is to say, improving the casting speed has more obvious effect on the solidification shell thickness of Al–1Mn alloy at the symmetry plane than that at the sprue plane and reduces the thickness difference of solidification shell between sprue and symmetry plane, which results in more uniform distribution of solidification shell thickness of Al–1Mn melt along the dividing plate and benefits the metallurgical bond. The effects of pouring temperature were also researched in this paper. The results show that improving the Al–1Mn pouring temperature TAl–1Mn has little effect on the solidification shell thickness of Al–1Mn alloy at the different position along the dividing plate, as given in Fig. 10d.

Experimental results According to the previous calculated results, this work decides to use the experimental parameters for preparing the clad slab as follows: the cooling water flow of dividing plate is 250 L h21, the cooled part height of the dividing plate is 20 mm, the casting speed is 80 mm min21, the pouring temperature of Al–1Mn alloy is 710uC and the pouring temperature of Al–10Si is 670uC. Figure 11a and b shows the macrostructure of clad slab on cross-section and the microstructure at the interface zone respectively. The tensile strength of the clad slab is 112 MPa and the fractured position is located in the Al–1Mn alloy side, as seen in Fig. 12, indicating the strength of the interfacial region is higher than that of Al–1Mn alloy. It is indicated that the prepared clad slab has excellent metallurgical bonding and straight interface without any discontinuities.

11 a macrostructure of clad slab on cross-section and b microstructure at interface zone

solidification shell formed on the dividing plate is increased by 5?4 mm at the sprue plane and 17?5 mm at the symmetry plane respectively. That is to say, increasing the cooled part height has more obvious effect on the thickness of Al–1Mn alloy at the symmetry plane than that at the sprue plane and increases the thickness difference of solidification shell between sprue and symmetry plane, which results in more non-uniform distribution of solidification shell thickness of Al–1Mn melt along the dividing plate and does not benefit the metallurgical bond.

Conclusions Three dimensional simulation model coupling thermal and flow fields, and the corresponding experiment have been applied to study the influence of technological parameters on direct chilled continuous casting process for preparing clad slab of aluminum alloys. The following points can be concluded. 1. Changing the cooling intensity and cooled part height of the dividing plate and casting speed has obvious effects on the temperature field and solidification shell of clad slab, while changing the pouring temperature has little effects on it. Increasing the cooled part height of dividing plate results in more non-uniform distribution of solidification shell. Improving the cooling intensity of dividing plate has little effect on the uniformity of solidification shell distribution. Higher casting speed results in more uniform distribution of solidification shell. 2. The prepared clad slabs had excellent metallurgical bonding when the cooling water flow of dividing plate is 250 L h21, the cooled part height of the dividing plate is 20 mm, the casting speed is 80 mm min21, the pouring

Effects of casting speed and pouring temperature Figures 8 and 9 show the influence of casting speed vc on the temperature field and liquid fraction of clad slab respectively. It is found that when the casting speed is increased, the surface temperature of the Al–1Mn solidification shell formed on the dividing plate rises, the depth of the melt sump becomes deeper and the thickness of Al–1Mn solidification shell becomes larger. Actually, a higher surface temperature and thinner solidification shell of Al–1Mn melt benefit the metallurgical bond, but excessively high casting speed will result in mixed flow of Al–1Mn and Al–10Si melt and failure of the metallurgical bond. Figure 10c shows the influence of casting speed on the solidification shell thickness of Al–1Mn alloy at the horizontal section beneath the dividing plate. In details, by increasing the casting speed from 70 to 80 mm min21, the thickness of Al–1Mn solidification shell formed on the

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temperature of Al–1Mn alloy is 710uC and the pouring temperature of Al–10Si is 670uC.

Acknowledgements The authors thank the National Natural Science Foundation of China (nos. 51071035 and 51274054), the Key grant Project of Chinese Ministry of Education (no. 313011), and Liaoning BaiQianWan Talents Program.

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