Model prediction of the effect of in-mold electromagnetic stirring on negative segregation under bloom surface

Yu-kun Huo; Li-hua Zhao; Hang-hang An; Min Wang; Chang-dong Zou

doi:10.1007/s12613-019-1906-6

Volume 27 Issue 3

Mar. 2020

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Article Contents

Abstract

1. Introduction

2. Numerical model

3. Model verification

4. Results and discussion

5. Conclusions

Acknowledgements

References

International Journal of Minerals, Metallurgy and Materials > 2020 > 27(3): 319-327. > DOI: 10.1007/s12613-019-1906-6

Yu-kun Huo, Li-hua Zhao, Hang-hang An, Min Wang, and Chang-dong Zou, Model prediction of the effect of in-mold electromagnetic stirring on negative segregation under bloom surface, Int. J. Miner. Metall. Mater., 27(2020), No. 3, pp.319-327. https://dx.doi.org/10.1007/s12613-019-1906-6

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Research Article

Model prediction of the effect of in-mold electromagnetic stirring on negative segregation under bloom surface

Yu-kun Huo^1,2),
Li-hua Zhao^1,2), ,,
Hang-hang An^1,3),
Min Wang¹⁾ and
Chang-dong Zou⁴⁾

Author Affilications

1)
State Key Laboratory of Advanced Metallurgy, University of Science and Technology Beijing, Beijing 100083, China
2)
School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing, Beijing 100083, China
3)
National Engineering Research Center of Flat Rolling Equipment, University of Science and Technology Beijing, Beijing 100083, China
4)
Institute of Research of Iron and Steel (IRIS) of Shasteel of Jiangsu Province, Zhangjiagang 215625, China

Funds: National Natural Science Foundation of China (No. 51774031) and the Foundation of State Key Laboratory of Advanced Metallurgy, University of Science and Technology Beijing, China (No. 41602014). The authors are thankful to Institute of Research of Iron and Steel(IRIS)of Shasteel, China for the support on the field test

Corresponding author:
Li-hua Zhao E-mail: zhaolihua@metall.ustb.edu.cn
Received: 21 April 2019; Revised: 02 September 2019; Accepted: 10 September 2019; Available Online: 28 October 2019

Graphical Abstract

Abstract

Abstract

Aiming at the problem of negative segregation under a bloom surface, a coupling macrosegregation model considering electromagnetic field, flow, heat, and solute transport was established based on the volume average method to study the effect of in-mold electromagnetic stirring (M-EMS) on the negative segregation under the bloom surface. In the model, the influence of dendrite structure on the flow and solute transport was described by the change of permeability. The model was validated by the magnetic induction intensity of M-EMS and carbon segregation experiment. The results show that the solute C in the solidified shell in the turbulent zone of the bloom undergoes two negative segregations, whereby the first is caused by nozzle jet, and the second by the M-EMS. The severities of the negative segregation caused by M-EMS at different currents and frequencies are also different, and the larger the current is, or the smaller the frequency is, the more serious will be the negative segregation. With the M-EMS, the solute C distribution in the liquid phase of the bloom is more uniform, but the mass fraction of C in the liquid phase is higher than that without M-EMS.
- continuous casting,
- in-mold electromagnetic stirring,
- bloom,
- negative segregation,
- numerical simulation

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1. Introduction

The in-mold electromagnetic stirring (M-EMS) is one of the effective technical means to refine the solidification structure, enlarge the equiaxed crystal zone, and improve the center segregation of casting strands [1‒4]; however, the use of M-EMS will also produce a negative segregation band under the bloom surface [5‒8]. This negative segregation band, which is difficult to eliminate in subsequent processes, may adversely affect the product’s hardenability, surface hardness, and mechanical properties. At present, the formation mechanism of the negative segregation band has not been uniformly explained [9‒11]. The application effect of the electromagnetic stirrer is affected by its installation position, current, frequency, and casting parameters [12‒14]. Sometimes, accurately obtaining the influence law of a single factor through industrial tests is difficult. Therefore, the influence of M-EMS parameters on negative segregation under the bloom surface needs to be determined by numerical simulation combined with industrial test.

Wu et al. [1] found that increasing the torque of M-EMS is beneficial to refine the solidification structure of the billet and expand the equiaxed crystal zone by examination of interdendritic corrosion. Using numerical simulation, Cho et al. [15] analyzed the magnetic field, flow field, temperature distribution, and inclusions trajectory of a billet with M-EMS, and plant trials were carried out. The strand central porosity and surface defects were significantly improved after the M-EMS parameters were optimized. Jiang and Zhu [8] reported that a higher stirring current of M-EMS will not continue to increase the equiaxed crystal ratio of the billet but will result in a more serious negative segregation of the billet, and the effect of equiaxed grains sedimentation and the liquid thermosolutal flow resulted in a negative segregation zone near the billet center and positive segregation at the center. Moreover, Oh and Chang [16] conducted a series of plant tests as well as relevant laboratory experiments to study the influence of various EMS modes on reducing macrosegregation. The combination stirring mode was found to be effective for macrosegregation reduction.

In recent years, many studies [7‒8,17‒19] have adopted numerical simulation methods to explore the macrosegregation behavior in strands, but only few studies have investigated the effect of M-EMS on negative segregation under the bloom surface by numerical simulation. Numerical simulation methods can reveal the formation process of negative segregation more intuitively and clearly. In the present study, aiming at the problem of internal segregation of 300 mm × 390 mm section of a steel mill, a three-dimensional mathematical model is established to study the influence of the M-EMS parameters on the negative segregation under the surface of the bloom. The location of the negative segregation band is determined more accurately, and the function rule of M-EMS is reflected quantitatively, which provides significant guidance for improving the internal quality of casting strands.

2. Numerical model

Based on the commercial software FLUENT of finite volume method, a coupling model considering the electromagnetic field, flow, heat, and solute transport is established, and the simulation results are verified by an industrial test. An LX82A high carbon steel is simulated; the stirrer center is 0.59 m from the meniscus, and the submerged entry nozzle structure has an upward inclination of 15° with four side hole nozzles. Fig. 1(a) shows the model and mesh division of the electromagnetic stirrer. The magnetic field is simulated using Maxwell software and then added to the FLUENT momentum source item by interpolation. Fig. 1(b) shows the model and mesh division of the bloom mold zone. The model performs proper grid-intensive treatment on the nozzle wall and solidification zone. The calculation domain of this model is set to 1.5 m from the meniscus since there the M-EMS no longer affects the macrosegregation behavior. The model only considers the segregation of C elements in the bloom. The process parameters and physical properties of the steel for simulation are presented in Table 1.

Fig. 1. Model and mesh division: (a) M-EMS; (b) bloom.

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Table 1. Process parameters and physical properties of the steel for simulation

Bloom section / mm²	300 × 390
Effective length of mold / mm	720
Pull velocity / (m∙min⁻¹)	0.6
Density / (kg∙m⁻³)	6956
Viscosity / (Pa∙s)	0.0048
Latent heat / (J∙kg⁻¹)	270000
Specific heat / (J∙kg⁻¹∙K⁻¹)	0.2043T + 406.1
Solid phase thermal conductivity / (W∙m⁻¹∙K⁻¹)	25
liquid phase thermal conductivity / (W∙m⁻¹∙K⁻¹)	33
Initial C content / wt%	0.82
Thermal expansion coefficient / K⁻¹	1.2 × 10⁻⁵
Solute expansion coefficient	0.011
Solute C liquidus slope	−8300
Equilibrium partition coefficient	0.34
Pure iron melting temperature / K	1808

| Show Table

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2.1 Basic assumption

In the process of simulating the electromagnetic field, flow, heat, and solute transport of the turbulent region of the bloom, the following assumptions are made to the coupling model:

(1) Molten steel is taken as an incompressible Newtonian fluid, and the density conforms to Boussinesq approximation.

(2) The mushy zone is regarded as a porous medium, and the flow in the mushy zone conforms to Darcy’s law; the influence of dendrite structure on flow and solute transport is expressed by permeability.

(3) The effects of other solute elements on C segregation are ignored.

(4) The influence of molten steel flow on electromagnetic field is ignored.

2.2 Governing equation

(1) Continuity equation:

$\frac{{\partial \left( {\rho {{{u}}_i}} \right)}}{{\partial {{{x}}_i}}} = 0$

(1)

where $\rho$ is the mixed fluid density, kg/m³; ${{{u}}_i}$ is the component of fluid velocity in $i$ direction, m/s; ${{{x}}_i}$ is the length in $i$ direction, m.

(2) Momentum conservation equation:

$\rho \frac{{\partial {{{u}}_i}{{{u}}_j}}}{{\partial {{{x}}_j}}} = \frac{\partial }{{\partial {{{x}}_j}}}\left[ {{\mu _{{\rm{eff}}}}\left( {\frac{{\partial {{{u}}_i}}}{{\partial {{{x}}_j}}} + \frac{{\partial {{{u}}_j}}}{{\partial {{{x}}_i}}}} \right)} \right] - \frac{{\partial {{P}}}}{{\partial {{{x}}_i}}} + \rho {{g}} + {{{F}}_{\rm{B}}} + {{{F}}_{\rm{E}}} + {{{S}}_{\rm{p}}}$

(2)

where ${{{u}}_j}$ is the component of fluid velocity in $j$ direction, m/s; ${{{x}}_j}$ is the length in $j$ direction, m; ${\mu _{{\rm{eff}}}}$ is the effective viscosity, ${{P}}$ is the pressure, ${{g}}$ is the acceleration of gravity, ${{{F}}_{\rm{B}}}$ is the buoyancy, including thermal buoyancy and solute buoyancy, ${{{F}}_{\rm{E}}}$ is the electromagnetic force, and ${{{S}}_{\rm{p}}}$ is the Darcy attenuation term.

(3) Low Reynolds number k‒ε turbulence equation:

$\rho \frac{{\partial k{{{u}}_i}}}{{\partial {{{x}}_i}}} = \frac{\partial }{{\partial {{{x}}_i}}}\left( {{\mu _{{\rm{eff}}}}\frac{{\partial k}}{{\partial {\sigma _k}}}} \right) + {{G}} - \rho \varepsilon + \rho {{D}} + {{{S}}_k}$

(3)

$\rho \frac{{\partial {{{u}}_j}\varepsilon }}{{\partial {{{x}}_j}}} = \frac{\partial }{{\partial {{{x}}_j}}}\left( {\frac{{{\mu _{{\rm{eff}}}}}}{{{\sigma _\varepsilon }}}\frac{{\partial \varepsilon }}{{\partial {{{x}}_j}}}} \right) + {C_1}{f_1}\frac{\varepsilon }{k}{{G}} - {C_2}{f_2}\frac{\varepsilon }{k}\rho \varepsilon + \rho {{E}} + {{{S}}_\varepsilon }$

(4)

where k and ε are the turbulent kinetic energy and dissipation rate, respectively. The empirical constant $C_1$ is 1.44, $C_2$ is 1.92, ${\sigma _\varepsilon }$ is 1.3, and ${\sigma _k}$ is 1.0, ${f_1}$ is 1.0, ${f_2}$ is 1.0, and the other coefficients in Eqs. (3) and (4) can be found elsewhere [7].

(4) Energy equation:

$\rho {c_{\rm{P}}}{{{u}}_i}\frac{{\partial H}}{{\partial {{{x}}_i}}} = \frac{\partial }{{\partial {{{x}}_i}}}\left( {{k_{{\rm{eff}}}}\frac{{\partial H}}{{\partial {{{x}}_i}}}} \right)$

(5)

where ${k_{{\rm{eff}}}}$ is the effective thermal conductivity, which is the sum of the thermal conductivities of laminar flow ${k_{\rm{l}}}$ and turbulent flow ${k_{\rm{t}}}$ , ${c_{\rm{P}}}$ is the specific heat, and H is the total heat enthalpy. H can be split into the sensible enthalpy ( $h$ ) and the latent enthalpy ( $\Delta H$ ), which is given by

$\left\{ \begin{aligned} & H = h + \Delta H \\ & h = {h_{{\rm{ref}}}} + \mathop \smallint \limits_{{T_{{\rm{ref}}}}}^T {c_{\rm{P}}}{\rm{d}} T \\ & \Delta H = {f_{\rm{l}}}{L_{\rm{h}}} \end{aligned} \right.$

(6)

where ${h_{{\rm{ref}}}}$ is the enthalpy at the reference temperature ${T_{{\rm{ref}}}}$ , ${L_{\rm{h}}}$ is the latent heat of molten steel solidification, and ${f_{\rm{l}}}$ is liquid fraction. The liquid phase temperature ( ${T_{\rm{l}}}$ ) and solid phase temperature ( ${T_{\rm{s}}}$ ) of the molten steel is calculated by the following formula:

${T_{\rm{l}}} = {T_{{\rm{pure}}}} + \mathop \sum \limits_i {m_i}{C_{{\rm{l}},i}}$

(7)

${T_{\rm{s}}} = {T_{{\rm{pure}}}} + \mathop \sum \limits_i {m_i}{C_{{\rm{l}},i}}/{k_i}$

(8)

where ${T_{{\rm{pure}}}}$ is the pure iron melting temperature, K; ${m_i}$ and ${k_i}$ are the liquid phase line slope and equilibrium distribution coefficient for solute $i$ , respectively; and ${C_{{\rm{l}},i}}$ is the local liquid phase mass fraction of $i$ element.

(5) Solute transfer equation:

$\rho {{{u}}_i}\frac{{\partial c}}{{\partial {{{x}}_i}}} = \frac{\partial }{{\partial {{{x}}_i}}}\left( {\rho {D_{{\rm{eff}}}}\frac{{\partial c}}{{\partial {{{x}}_i}}}} \right) + {{{S}}_{\rm{C}}}$

(9)

where $c$ is the mass fraction of solute C, ${D_{{\rm{eff}}}}$ is the effective diffusion coefficient, the value of which is the sum of laminar flow ${D_{\rm{l}}}$ and turbulent diffusion coefficient ${D_{\rm{t}}}$ ; ${{{S}}_{\rm{C}}}$ is the source item of the conservation equation of solute transport, and it is composed of the diffusion source item ${{{S}}_{{\rm{c}},{\rm{dif}}}}$ and the convection diffusion source item ${{{S}}_{{\rm{c}},{\rm{con}}}}$ caused by Fick’s Law. The expressions for ${{{S}}_{{\rm{c}},{\rm{dif}}}}$ and ${{{S}}_{{\rm{c}},{\rm{con}}}}$ are given as

${{{S}}_{{\rm{c}},{\rm{dif}}}} = \frac{\partial }{{\partial {{{x}}_i}}}\left[ {\rho {f_{\rm{s}}}{D_{\rm{s}}}\frac{{\partial \left( {{C_{\rm{s}}} - C} \right)}}{{\partial {{{x}}_i}}}} \right] + \frac{\partial }{{\partial {{{x}}_i}}}\left[ {\rho {f_{\rm{l}}}{D_{{\rm{eff}}}}\frac{{\partial \left( {{C_{\rm{l}}} - C} \right)}}{{\partial {{{x}}_i}}}} \right]$

(10)

${{{S}}_{{\rm{c}},{\rm{con}}}} = \frac{\partial }{{\partial {{{x}}_i}}}\left[ {\rho \left( {{{{u}}_i} - {{{u}}_{i,{\rm{s}}}}} \right)\left( {{C_{\rm{l}}} - C} \right)} \right]$

(11)

where ${f_{\rm{s}}}$ is the solid fraction, ${D_{\rm{s}}}$ is the solute diffusive coefficient in the solid phase, ${C_{\rm{s}}}$ and ${C_{\rm{l}}}$ are the solute concentrations in the solid and liquid phase, respectively, $C$ is the locally averaged concentration of solute, ${{{u}}_{i,{\rm{s}}}}$ is the velocity of solid phase in $i$ direction.

(6) Equation for solving electromagnetic force.

Maxwell’s equations:

$\left\{ \begin{aligned} & \nabla \times {{{H}}_{\rm{m}}} = {{J}} \\ & \nabla \times {{E}} = - \frac{{\partial {{B}}}}{{\partial t}} \\ & \nabla \cdot {{B}} = 0 \\ & {{{B}} = \mu {{{H}}_{\rm{m}}}} \\ & {{{J}} = \sigma {{E}}} \end{aligned} \right.$

(12)

where ${{{H}}_{\rm{m}}}$ is the magnetic field strength, A/m; ${{J}}$ is the current density, A/m²; ${{E}}$ is the electric field strength, V/m; ${{B}}$ is the magnetic induction intensity, T; $t$ is the time, s; $\mu$ is the magnetic conductivity, H/m; and $\sigma$ is the conductivity, S/m. The electromagnetic stirring process is calculated using uniform electromagnetic force:

${ F} = \frac{1}{2}{\rm{Re}}\left( {{{J}} \times {{{B}}^ * }} \right)$

(13)

where Re represents the real part of the plural, and ${{{B}}^ * }$ is a conjugate complex number.

2.3 Boundary conditions

(1) Free surface.

The free surface of the mold is set as an adiabatic wall, and the other variables are zero gradient boundary conditions.

(2) Inlet and outlet of the model.

The inlet and outlet of the model are all velocity boundary conditions. The outlet velocity is set to be constant and is equal to the pull velocity. The inlet velocity is calculated according to the quality conservation calculation of the inlet-outlet cross-sectional area, the molten steel temperature is the pouring temperature, and the initial content of C is 0.82wt%. To describe the turbulence phenomenon, a semi-empirical formula is used to determine the turbulence kinetic energy $k$ and the dissipation rate $\varepsilon$ :

$k = 0.01{{u}}_{{\rm{in}}}^{\rm{2}}$

(14)

$\varepsilon = k_{}^{{\rm{1}}{\rm{.5}}}/D$

(15)

where ${{u}}_{{\rm{in}}}^{}$ is the inlet velocity, m/s; D is the inlet diameter, m.

(3) Heat transfer conditions on strand surface.

(a) Mold zone.

The distribution of the heat flux density of the instantaneous mold along the direction of pull and the average heat flux density are calculated by the following formulas:

${q_{\rm{s}}} = 2680000 - b\sqrt {L/v}$

(16)

$b = \frac{{1.5 \times \left( {2680000 - \bar q} \right)}}{{\sqrt { {{L_{\rm{m}}}/v}} }}$

(17)

$\bar q{\rm{ = }}\frac{{{C_{\rm{W}}} \cdot m \cdot \Delta T}}{{{{{S}}_{{\rm{eff}}}}}}$

(18)

where ${q_{\rm{s}}}$ is the heat flux, W/m²; L is the distance from the meniscus, m; b is an unknown coefficient, v is the casting speed, m/s; ${L_{\rm{m}}}$ is the effective length of the mold, m; $\bar q$ is average heat flux of mold, ${C_{\rm{w}}}$ is the specific heat of water, J/(kg∙°C); $m$ is the mold water flow, kg/s; $\Delta T$ is the difference between the inlet and outlet temperatures of water, °C; and ${{{S}}_{{\rm{eff}}}}$ is the effective cooling zone of the mold wall, m².

(b) Secondary cooling zone.

${q_{\rm{s}}} = {h_{\rm{c}}}\left( {{T_{\rm{b}}} - {T_{\rm{w}}}} \right)$

(19)

where ${h_{\rm{c}}}$ is the heat transfer coefficient between the billet and the cooling water, W/(m²·°C); ${T_{\rm{b}}}$ is the surface temperature of the strand, °C; and ${T_{\rm{w}}}$ is the temperature of cooling water, °C.

3. Model verification

3.1 Magnetic induction intensity distribution in mold and model verification

KANETEC TM-701 Gaussian meter was used to measure the magnetic induction intensity value of the axial center of the electromagnetic stirrer when it was empty in mold and the magnetic induction intensity of the stirrer center under different currents. The magnetic induction intensity was the largest at the stirrer center (Fig. 2(a)). Because the stirrer installation position was in the lower part of the mold, the free surface was weakened by the magnetic induction intensity interference. Fig. 2(b) shows the cross-sectional tangential velocity distribution at the stirrer center position; it can be found that the rotational velocity of the molten steel increased with the increase of the stirring current. Fig. 3(a) compares the center axial magnetic induction intensity simulation results of the mold with the field measurement results, where the corresponding magnetic stirring parameters are 750 A and 1.5 Hz. The simulation results basically agree with the measured values; the central axial magnetic induction intensity of the mold change law for the stirring center was large, and the two sides were small. Fig. 3(b) illustrates the changes of the magnetic induction intensity at the stirrer center position at different currents. The intensity increases linearly with the increase of the current, and the slope is about 0.065 mT/A.

Fig. 2. Simulated magnetic induction intensity distribution (a) and velocity distribution of stirrer center position (b).

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Fig. 3. Simulation result verification: (a) magnetic induction intensity verification; (b) relationship between magnetic induction intensity and current at stirrer center.

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3.2 Change of carbon content in bloom and model verification

Drilling sampling was used to obtain a sample cross section with a thickness of 20 mm on the bloom. Then a drill of 4 mm in diameter was used to drill holes at 15 mm and 38 mm in the inner and outer arc center line directions of the sample (Fig. 4(a)). A carbon-sulfur meter was used to detect the mass fraction of C in the drill cuttings. Fig. 4(b) shows that the simulation result agrees well with the measured results of the C segregation in the direction of the center line of the bloom wide surface under 750 A and 1.5 Hz. Table 2 compares the measured value with the simulated value of the C content at the bloom drilling position. Compared with the simulation results, the mass fraction of C near the side of the inner arc is higher, and that near the outer arc side is lower; the deviation is less than 2%, which shows that the prediction of this model is accurate and reliable.

Table 2. Comparison of measured and simulated values of C content in drilling position

Drilling position	Measured / wt%	Simulated / wt%
15 mm from the inner arc surface	0.792	0.786
38 mm from the inner arc surface	0.851	0.834
15 mm from the outer arc surface	0.771	0.786
38 mm from the outer arc surface	0.824	0.833

| Show Table

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Fig. 4. Drilling position schematic diagram (a) and verification of C segregation results at the turbulent zone outlet (b).

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4. Results and discussion

4.1 Effect of M-EMS on C content distribution in turbulent zone

Figs. 5(a) and 5(b) are images of contours of the C element distribution on the center symmetrical plane (Y = 0 m) of the narrow surface of the bloom without and with M-EMS. It can be seen that irrespective of M-EMS application, the position of the initial shell of the bloom presented a negative segregation band, which is caused by the impact of the nozzle jet on the mold wall. Fig. 6 is an enlarged image of the dotted line box in Fig. 5(a), which shows that the position on the mold wall of the impacting nozzle jet was about 0.04 m away from the meniscus, followed by the formation of upper and lower vortexes. Solute C was easily enriched between the upper vortex and the mold wall to form a positive segregation; the mass fraction of C here was about 1.00wt%. When the shell solidified to the impact zone of the nozzle jet, continuously washed by the flow, the solute C in the mushy zone began to form a negative segregation. The negative segregation began to form at 1 mm from the bloom surface, and the degree of negative segregation at 5.6 mm under the surface was maximized.

Fig. 5. (a, b) Solute C distribution and (c, d) flow field profile in turbulent zone: (a, c) without M-EMS; (b, d) with M-EMS.

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Fig. 6. Flow field and solute C distribution of nozzle jet zone.

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Figs. 5(c) and 5(d) show images of the flow field contours in the turbulent zone without and with M-EMS. When M-EMS was not applied, the molten steel moved downward. When M-EMS was applied, a pair of vortexes were produced in the lower part of the stirrer center position, which led to the difference of solute distribution in the bloom. Fig. 7 shows the change of solute C on the central axis of the bloom without and with M-EMS. When M-EMS was not applied, with the casting in progress, the solute C enriched in the mushy zone diffused to the bloom center; therefore, the mass fraction of C in the liquid phase will change in gradient. When M-EMS was applied, however, under the action of tangential rotation and vortexes, the solute C in the liquid phase tended to be more homogeneous.

Fig. 7. Change of solute C on the central axis of the bloom.

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4.2 Effect of M-EMS on macrosegregation in model outlet

Fig. 8 shows a C segregation image of contours of the model outlet without and with M-EMS. It can be found that due to the impact of the nozzle jet, a significant negative segregation band appeared at the position of the initial shell at the center of the wide and narrow surface of the bloom. Due to the fact that turbulence flow and electromagnetic force have little effect on the corner of bloom, the negative segregation degree of the solidification shell in this position was remarkably weaker. In addition, it can also be found that the casting strand experienced two negative segregations with M-EMS, and the solute distribution in the liquid phase was more uniform than that without M-EMS.

Fig. 8. C segregation of the turbulent zone outlet (a) without and (b) with M-EMS.

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Fig. 9(a) shows the results of the solute C distribution in the direction of the center line of the wide surface at the model outlet without and with M-EMS at different stirring frequencies. It can be seen that whether M-EMS is applied or not, the mass fraction of C in the shell undergoes a “V” shape change, which is caused by the impact of the nozzle jet; the solute changed sharply from positive segregation (about 0.84wt%) on the bloom surface to negative segregation (minimum value of about 0.66wt%), and the C content recovered rapidly after leaving the impact zone of nozzle jet. The velocity of the impact position near the wall of the mold was measured as 0.1199 m/s without M-EMS and 0.1222 m/s with M-EMS (Fig. 10); the small velocity increment caused the slight effect of M-EMS on C segregation. At the solid phase interface of the stirrer center position, with the decrease of the stirring frequency, from 2.0, 1.5 to 1.0 Hz, the mass fractions of C in this position were 0.817wt%, 0.814wt%, and 0.812wt%, respectively. This is because the lower the frequency, the greater the stirring intensity, and the content of C in the mushy zone will less. When the shell continued to solidify downward out of the stirrer center, the solute C underwent negative segregation again. With the decrease of the stirring frequency, the negative segregation minimum values were 0.773wt%, 0.768wt%, 0.763wt%, respectively, and the position was 17.8 mm from the bloom surface. Then with continued solidification, the C content began to pick up again; this time the mass fraction of C in the liquid phase was about 0.840wt% with M-EMS and 0.834wt% without M-EMS. Fig. 9(b) shows the effect of different currents on the macrosegregation at turbulent zone outlet; the larger the current (450, 600, and 750 A), the more serious the degree of negative segregation caused by M-EMS (0.779wt%, 0.773wt%, and 0.768wt%, respectively).

Fig. 9. Solute C distribution of the wide surface at turbulent zone outlet: (a) different frequencies; (b) different currents.

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Fig. 10. Effect of the nozzle on the velocity of the impact zone near the wall of the mold.

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From Fig. 9(a), it can be found that without M-EMS, the solute C near the center position of the M-EMS also underwent a slight negative segregation again. The analysis of the surface temperature change of bloom (Fig. 11) shows that from the position of the stirrer center to the mold outlet, due to the decrease in the cooling strength, the surface temperature of the bloom began to pick up, the solidification shell grew slower, while the mushy zone was still washed by the molten steel; therefore, a negative segregation was formed. This shows that the formation of negative segregation is related to the flow rate and solidification rate.

Fig. 11. Temperature change of the bloom surface center.

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5. Conclusions

(1) Based on the finite volume method and the appropriate boundary conditions, a three-dimensional coupling model is established. The simulation results of macrosegregation are verified by a drilling sampling experiment, and the deviation is less than 2%, which shows that the model prediction is reliable.

(2) The formation of negative segregation is related to the flow rate and solidification rate of molten steel. For a bloom with four side hole nozzles, the solute C in the solidification shell undergoes two negative segregations: the first is caused by the nozzle jet, while the second is caused by M-EMS. The degree of negative segregation caused by the nozzle jet is generally more serious than M-EMS.

(3) The larger the current or the smaller the frequency, the more serious the negative segregation caused by M-EMS. The change of magnetic stirring parameters will not affect the position and width of the two negative segregation bands in the bloom.

(4) The distribution of solute C in the liquid phase is more uniform with M-EMS, but the mass fraction of C in the liquid phase is higher than that without M-EMS, which may aggravate the central positive segregation degree of the bloom to a certain extent.

Acknowledgements

This work was financially supported by the National Natural Science Foundation of China (No. 51774031) and the Foundation of State Key Laboratory of Advanced Metallurgy, University of Science and Technology Beijing, China (No. 41602014). The authors are thankful to Institute of Research of Iron and Steel(IRIS)of Shasteel, China for the support on the field test.

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Model prediction of the effect of in-mold electromagnetic stirring on negative segregation under bloom surface

Abstract

1. Introduction

2. Numerical model

2.1 Basic assumption

2.2 Governing equation

2.3 Boundary conditions

3. Model verification

3.1 Magnetic induction intensity distribution in mold and model verification

3.2 Change of carbon content in bloom and model verification

4. Results and discussion

4.1 Effect of M-EMS on C content distribution in turbulent zone

4.2 Effect of M-EMS on macrosegregation in model outlet

5. Conclusions

Acknowledgements

References

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Model prediction of the effect of in-mold electromagnetic stirring on negative segregation under bloom surface

Abstract

1. Introduction

2. Numerical model

2.1 Basic assumption

2.2 Governing equation

2.3 Boundary conditions

3. Model verification

3.1 Magnetic induction intensity distribution in mold and model verification

3.2 Change of carbon content in bloom and model verification

4. Results and discussion

4.1 Effect of M-EMS on C content distribution in turbulent zone

4.2 Effect of M-EMS on macrosegregation in model outlet

5. Conclusions

Acknowledgements

References

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