Modern industries such as marine engineering, vehicle manufacturing, aerospace, and aviation are placing ever-increasing demands on lightweight structural materials. Cast aluminum alloys are highly valued in these fields because of their favorable corrosion resistance, excellent castability, and high specific strength. Among the various forming routes, sand casting is one of the most flexible and economical methods for producing complex-shaped aluminum components. However, the solidification process in sand casting is often accompanied by coarse dendritic structures and a range of metallurgical defects, including shrinkage porosity and macro-segregation. These imperfections significantly reduce the mechanical properties of the final products, such as tensile strength, ductility, and toughness. Therefore, refining the solidification microstructure is a key approach to improving the performance of sand casting parts. In my research, I focused on the design and development of a pouring cup that applies a pulsed magnetic field (PMF) to the melt before it enters the mold cavity, with the aim of improving the microstructure and mechanical behavior of sand casting parts.
For years, researchers have explored various external fields, including ultrasonic vibration, mechanical stirring, and electromagnetic stirring, to control solidification structures. Among these techniques, pulsed magnetic field treatment has drawn considerable attention because it offers non-contact processing, strong instantaneous energy, low energy consumption, and easy control. In particular, PMF can generate a large time derivative of magnetic flux density, often denoted as dB/dt, which induces forced convection and promotes nucleation in the melt. Compared with conventional grain refiners, PMF does not introduce additional impurities into the melt. As a result, the risk of inclusions and gas porosity in the final casting is reduced. In the context of sand casting parts, this advantage is significant because the pouring cup is the first component that contacts the liquid metal, and any treatment performed there can directly influence the quality of the downstream casting. To gain a comprehensive understanding of the process, I combined experimental investigations with numerical simulations using COMSOL Multiphysics. The numerical work allowed me to visualize the magnetic field, flow field, and temperature field inside the pouring cup and to optimize the cup design and electromagnetic parameters before manufacturing and field trials.
In this paper, I present a systematic study covering laboratory simulations, multi-physics numerical modeling, device design, and industrial trials. I first conducted sand-casting-like experiments in the laboratory to investigate the effects of PMF on the solidification structure and mechanical properties of ZL205A aluminum alloy. Then, I developed three-dimensional transient mathematical models for two types of pouring cups: a plug-type pouring cup for large sand molds and a small ring-type pouring cup for small sand molds. Through these models, I analyzed the influence of structural configurations, magnetic field direction, pulse frequency, duty cycle, and coil current on the melt treatment efficiency. Finally, I manufactured the optimized pouring cups, carried out sand casting experiments in an industrial environment, and verified the feasibility of the proposed designs by evaluating the microstructures and mechanical properties of the resulting sand casting parts.
1. Laboratory Simulation of PMF Treatment in Sand Casting Conditions
Before designing the pouring cup, I needed a reliable process window for the electromagnetic parameters. I therefore simulated the sand casting pouring practice in the laboratory using ZL205A alloy, which is a high-strength Al-Cu-Mn system alloy widely employed for structural sand casting parts. The chemical composition of the alloy used in this study is given in Table 1.
| Cu | Mn | Ti | Cd | V | Zr | Si | B | Al |
|---|---|---|---|---|---|---|---|---|
| 4.76 | 0.321 | 0.165 | 0.151 | 0.005 | 0.096 | 0.046 | 0.017 | Bal. |
I melted the alloy in a medium-frequency induction furnace and superheated it to 730 °C with a holding time of 300 s. A graphite mold with dimensions of 45/60 mm inner/outer diameter and 160 mm height was preheated to 300 °C. During the PMF treatment, an induction coil was positioned 10 mm above the melt surface, and the pulsed magnetic field generator was activated according to the parameters listed in Table 2. The treatment lasted for 90 s, after which the melt was cooled to 700 °C, skimmed, and poured into the preheated graphite mold. The entire procedure closely resembled the industrial sand casting pouring sequence, so the obtained data could serve as a basis for the subsequent pouring cup design.
| No. | Frequency (Hz) | Duty cycle (%) | Peak current (A) | dB/dt (mT/s) | Melt temperature (°C) | Treatment time (s) | Pouring temperature (°C) |
|---|---|---|---|---|---|---|---|
| 1 | 0 | 0 | 0 | 0 | 730 | 90 | 700 |
| 2 | 20 | 20 | 100 | 3150 | 730 | 90 | 700 |
| 3 | 40 | 20 | 100 | 8000 | 730 | 90 | 700 |
| 4 | 50 | 20 | 100 | 150000 | 730 | 90 | 700 |
After casting, I extracted metallographic samples from both the edge and the center of the ingots, while tensile specimens were machined from the bottom section. The metallographic samples were ground, polished, and etched with Keller’s reagent. I used an optical microscope to observe the microstructure, and the average grain size was measured by the linear intercept method using at least 100 grains per sample. A field-emission scanning electron microscope was used to examine the Al₂Cu phase morphology. Tensile tests were carried out at a crosshead speed of 1 mm/min, and Vickers hardness was measured on an automatic hardness tester.
1.1 Effect of PMF on primary α-Al grains
The as-cast microstructure of the untreated ZL205A alloy was dominated by coarse and irregular rosette-shaped primary α-Al grains. When the pulsed magnetic field was applied, the rosette crystals were effectively refined, and fine globular grains began to appear. The degree of refinement increased with increasing dB/dt up to 8000 mT/s, beyond which coarsening occurred. The edge region always showed finer grains than the center region, which can be attributed to the higher cooling rate at the mold wall. Table 3 summarizes the measured average grain sizes under different conditions.
| dB/dt (mT/s) | Edge grain size (μm) | Center grain size (μm) |
|---|---|---|
| 0 | 71.53 | 80.45 |
| 3150 | 61.31 | 69.67 |
| 8000 | 52.34 | 62.46 |
| 150000 | 61.59 | 70.67 |
At the optimum dB/dt of 8000 mT/s, the edge and center grain sizes decreased by 26.83% and 22.36%, respectively, compared with the untreated sample. The morphology of the Al₂Cu phase also changed from a coarse network into small, fragmented, or nearly spherical particles, particularly in the edge region. This is important because the size and distribution of the Al₂Cu phase directly affect the mechanical properties of sand casting parts.
1.2 Mechanical properties after PMF treatment
Table 4 presents the tensile properties of ZL205A samples processed with different dB/dt values. The yield strength, ultimate tensile strength, and elongation all increased first and then decreased with increasing dB/dt. At 8000 mT/s, the yield strength, tensile strength, and elongation reached 160 MPa, 241 MPa, and 11.5%, respectively, representing increases of 28%, 17.5%, and 91.66% relative to the untreated condition. The hardness measurements in Table 5 show the same trend: the edge hardness increased from 55.3 HV to 62.5 HV, and the center hardness increased from 51.2 HV to 56 HV, corresponding to improvements of 13% and 9.3%.
| dB/dt (mT/s) | Yield strength (MPa) | Tensile strength (MPa) | Elongation (%) |
|---|---|---|---|
| 0 | 125 | 206 | 6.0 |
| 3150 | 138 | 225 | 8.5 |
| 8000 | 160 | 241 | 11.5 |
| 150000 | 145 | 230 | 9.0 |
| dB/dt (mT/s) | Edge hardness (HV) | Center hardness (HV) |
|---|---|---|
| 0 | 55.3 | 51.2 |
| 8000 | 62.5 | 56.0 |
1.3 Mechanism of grain refinement by pulsed magnetic field
The grain refinement observed in the experiments can be explained from both thermodynamic and kinetic viewpoints. When a pulsed magnetic field is applied to the melt, the electromagnetic energy is absorbed by the liquid metal. This energy can lower the critical nucleation activation energy. The modified nucleation activation energy can be expressed as
$$ \Delta G_m^* = \frac{16\pi \sigma_{sl}^3}{3\Delta G_m^2} = \frac{16\pi \sigma_{sl}^3}{3\left[\Delta G_v + \frac{\mu_0 \Delta \chi_{S-L}}{2} |H|^2\right]^2} $$
where \(\mu_0\) is the vacuum permeability, \(\Delta \chi_{S-L}\) is the change in volume magnetic susceptibility between the solid and liquid phases, \(H\) is the magnetic field strength, and \(\Delta G_v\) is the Gibbs free energy difference between the liquid and solid phases. The nucleation rate per unit volume per unit time is given by
$$ I = n_s^* \varepsilon v n e^{-\frac{\Delta G_m^*}{kT}} e^{-\frac{Q}{kT}} $$
where \(k\) is the Boltzmann constant, \(Q\) is the activation energy for diffusion, \(n_s^*\) is the number of atoms on the critical nucleus surface, \(\varepsilon\) is the probability of an atom jumping in a given direction, and \(v\) is the vibration frequency of atoms in the liquid phase. According to the above equation, the electromagnetic energy reduces the critical Gibbs free energy barrier for nucleation, thereby increasing the nucleation rate. In addition, the pulsed magnetic field promotes the migration and collision of atomic clusters in the melt. The average radius of atomic clusters in the liquid can be described by
$$ R(T) = r \left[ \frac{\alpha}{Z_1^2} \left( \frac{1+C_0}{C_0} \exp\left(\frac{\Delta G_r^*}{k}\left(\frac{1}{T} – \frac{1}{T_m}\right)\right) – 1 \right) – 1 \right] $$
where \(T\) is the temperature, \(T_m\) is the melting point, \(r\) is the interatomic distance, \(\alpha\) is a geometric factor, \(Z_1\) is the coordination number of the liquid, \(C_0\) is the relative concentration of activated atoms at the melting point, and \(\Delta G_r^*\) is the activation energy for cluster formation. The PMF energy reduces the size of atomic clusters, increases the number of clusters close to the critical size, and thus increases the nucleation rate. Consequently, the α-Al grains are refined, and the improved structural uniformity enhances the mechanical properties of sand casting parts.
The strengthening effect can also be interpreted through the Hall-Petch relationship:
$$ \sigma_b = \sigma_0 + k d^{-1/2} $$
where \(\sigma_b\) is the tensile strength, \(d\) is the average grain size, and \(k\) is a material constant. Finer grains increase the strength because grain boundaries hinder dislocation motion. The experimental results confirmed that the PMF treatment improved the strength and hardness of ZL205A alloy, which is beneficial for sand casting parts requiring high load-bearing capacity.
2. Multi-Physics Mathematical Model for the Pouring Cup
After establishing the beneficial effect of PMF treatment from laboratory experiments, I proceeded to design two categories of pouring cups for sand casting: a plug-type pouring cup for large molds and a small ring-type pouring cup for small molds. Numerical simulation was essential to determine the cup geometry and electromagnetic parameters. I used COMSOL Multiphysics to construct three-dimensional finite element models that coupled the electromagnetic field, fluid flow, and heat transfer. The models allowed me to visualize the distribution of magnetic flux density, Lorentz force, velocity, and temperature inside the molten pool.
2.1 Material properties
Table 6 summarizes the magnetic and physical properties used in the models. The ZL205A melt was assigned to the plug-type pouring cup, while the ZL114A melt was assigned to the small pouring cup. The B-H curve of the iron core was derived from a magnetization curve, and the specific heat and thermal conductivity of both alloys were treated as temperature-dependent functions.
| Property | ZL114A melt | ZL205A melt | Air | Iron core | Coil |
|---|---|---|---|---|---|
| Relative permeability | 1 | 1 | 1 | B-H curve | 1 |
| Relative permittivity | 1 | 1 | 1 | 1 | 1 |
| Electrical conductivity (S/m) | 8.77×10⁶ | 3.71×10⁶ | 0 | 3.18×10⁷ | 1.72×10⁶ |
| Property | Value |
|---|---|
| Density (kg/m³) | 2420 |
| Specific heat (J/(g·K)) | Temperature-dependent |
| Thermal conductivity (W/(m·K)) | Temperature-dependent |
| Thermal expansion coefficient (1/K) | 2.3×10⁻⁵ |
| Dynamic viscosity (Pa·s) | 0.0013 |
| Property | Value |
|---|---|
| Density (kg/m³) | 2360 |
| Specific heat (J/(g·K)) | Temperature-dependent |
| Thermal conductivity (W/(m·K)) | Temperature-dependent |
| Thermal expansion coefficient (1/°C) | 2.7×10⁻⁵ |
| Dynamic viscosity (Pa·s) | 0.0011 |
2.2 Model assumptions and governing equations
Several reasonable assumptions were introduced in the numerical model: the melts were incompressible Newtonian fluids; thermal buoyancy was handled using the Boussinesq approximation; the free surface of the melt was assumed to be flat; the magnetic Reynolds number was much smaller than unity, so the influence of fluid flow on the electromagnetic field was negligible; displacement currents were neglected; and the solidification shrinkage was not included in the model. Under these assumptions, the Maxwell equations can be written as
$$ \nabla \times \vec{H} = \vec{J} + \frac{\partial \vec{D}}{\partial t} $$
where \(\vec{H}\) is the magnetic field intensity, \(\vec{J}\) is the current density, and \(\vec{D}\) is the electric flux density. In the present low-frequency problem, the displacement current term can be neglected, so
$$ \nabla \times \vec{H} = \vec{J} $$
The magnetic flux density is related to the magnetic field intensity by
$$ \vec{B} = \mu \vec{H} $$
where \(\mu\) is the magnetic permeability. The current density is governed by Ohm’s law:
$$ \vec{J} = \sigma \vec{E} $$
where \(\sigma\) is the electrical conductivity and \(\vec{E}\) is the electric field intensity. The Lorentz force and electromagnetic heat generated by the interaction between the induced current and the magnetic field are expressed as
$$ \vec{F} = \vec{J} \times \vec{B} $$
$$ Q = \vec{J} \cdot \vec{E} $$
The continuity, momentum, and energy equations are given by
$$ \frac{\partial \rho}{\partial t} + \nabla \cdot (\rho \vec{U}) = 0 $$
$$ \frac{\partial(\rho \vec{U})}{\partial t} + \nabla \cdot (\rho \vec{U}\vec{U}) = \nabla \cdot (\mu_{eff} \nabla \vec{U}) – \nabla P + S_m $$
$$ \frac{\partial(\rho T)}{\partial t} + \nabla \cdot (\rho \vec{U} T) = \nabla \cdot \left(\frac{k}{c_p} \nabla T\right) + S_{th} $$
where \(\rho\) is the melt density, \(\vec{U}\) is the velocity vector, \(P\) is the pressure, \(T\) is the temperature, \(k\) is the thermal conductivity, and \(c_p\) is the specific heat. The source term \(S_m\) includes the electromagnetic force and thermal buoyancy, while \(S_{th}\) includes Joule heating. The thermal buoyancy is defined as
$$ S_i = \rho g \beta (T – T_0) $$
where \(g\) is the gravitational acceleration, \(\beta\) is the thermal expansion coefficient, and \(T_0\) is a reference temperature. Since the pulsed magnetic field varies with time, the time-averaged Lorentz force over one pulse period is
$$ \vec{F}_m = \frac{1}{T_p} \int_0^{T_p} f \, dt $$
To model the turbulent flow in the melt, I adopted the standard \(k\)-\(\varepsilon\) model. The turbulent viscosity is
$$ \mu_t = \rho C_\mu \frac{k^2}{\varepsilon} $$
where \(k\) is the turbulent kinetic energy and \(\varepsilon\) is its dissipation rate. The transport equations for \(k\) and \(\varepsilon\) are
$$ \frac{\partial(\rho k)}{\partial t} + \nabla \cdot (\rho \vec{U} k) = \nabla \cdot \left[\left(\mu + \frac{\mu_t}{\sigma_k}\right)\nabla k\right] + G_k – \rho \varepsilon + S_k $$
$$ \frac{\partial(\rho \varepsilon)}{\partial t} + \nabla \cdot (\rho \vec{U} \varepsilon) = \nabla \cdot \left[\left(\mu + \frac{\mu_t}{\sigma_\varepsilon}\right) \nabla \varepsilon\right] + C_1 \frac{\varepsilon}{k} G_k – C_2 \rho \frac{\varepsilon^2}{k} + S_\varepsilon $$
where \(G_k\) represents the generation of turbulent kinetic energy due to mean velocity gradients:
$$ G_k = \mu_T \left[ \nabla \vec{U} \cdot \left(\nabla \vec{U} + (\nabla \vec{U})^T\right) \right] $$
The model constants are listed in Table 9.
| C₁ | C₂ | σk | σε | Cμ |
|---|---|---|---|---|
| 1.44 | 1.92 | 1.0 | 1.3 | 0.09 |
2.3 Boundary conditions and mesh
For the plug-type pouring cup, the melt initially had a zero velocity and a uniform temperature of 730 °C. The ambient temperature was set to 25 °C. The outer boundary of the air domain was treated as the far-field where the magnetic flux density vanished. The melt sidewall and bottom were assumed to be thermally insulated, while the free surface was exposed to air with a heat transfer coefficient of 25 W/(m²·°C). The periodic pulsed current was applied to the coil region as an interpolation function. The boundary conditions are summarized in Table 10.
| Physical field | Location | Boundary condition |
|---|---|---|
| Magnetic field | Air outer surface | Far-field, B = 0 |
| Flow field | Free surface | Initial velocity = 0 |
| Flow field | Melt sidewall | No-slip wall |
| Flow field | Melt bottom | No-slip wall |
| Temperature field | Free surface | λ ∂T/∂n = h(T − T_env), h = 25 W/(m²·°C) |
| Temperature field | Melt sidewall | Adiabatic |
| Temperature field | Melt bottom | Adiabatic |
I used structured hexahedral meshes for the melt, iron core, and coil regions, while unstructured tetrahedral meshes were applied to the air and far-field regions. Boundary layer meshes were generated near the melt walls to resolve the strong gradients of velocity and temperature. The mesh density was refined until the variation in the predicted peak magnetic flux density was less than 2%. The computing was performed on an E5-2680v4 server.
3. Numerical Simulation Results and Design Optimization
Using the established models, I systematically compared the performance of different pouring cup structures, magnetic field directions, and electromagnetic parameters. The design indicators included the dB/dt value, the volume fraction of the melt exposed to a magnetic flux density in the range of 10–131 mT, and the uniformity of the flow and temperature distributions. Based on the laboratory experiments, the target dB/dt range was set as 8000–15000 mT/s, and the effective magnetic treatment volume was required to be greater than 25% of the total melt volume.
3.1 Comparison of different structural schemes for the plug-type pouring cup
I investigated three structural configurations for the plug-type pouring cup: single-ended, inner-annular, and symmetric. The melt volume in the cup was 1.5×10⁸ mm³. The coil current was set at 200 A with a frequency of 20 Hz and a duty cycle of 50%. Table 11 summarizes the magnetic treatment volumes at 10–131 mT and the corresponding dB/dt values for the three schemes.
| Structure | Treatment volume (mm³) | Volume fraction (%) | dB/dt (mT/s) |
|---|---|---|---|
| Single-ended | 2.68×10⁷ | 17.87 | 10480 |
| Inner-annular | 4.37×10⁷ | 29.13 | 10480 |
| Symmetric | 3.05×10⁷ | 20.33 | 10480 |
The single-ended scheme concentrated the magnetic flux near one end of the coil, leaving a large portion of the melt untreated. The symmetric scheme produced two hemispherical high-field regions on both sides of the melt, but the central zone still had low field intensity. The inner-annular scheme, in which the coil surrounded the annular melt volume, provided the largest effective treatment volume and satisfied the 25% requirement. Therefore, I selected the inner-annular structure as the basis for the plug-type pouring cup design.
3.2 Influence of magnetic field direction
Within the inner-annular structure, I further compared the same-direction and opposite-direction coil connections. The vector plots showed that when the two coil groups were connected in the opposite direction, the magnetic flux distribution formed a vortex-like pattern inside the melt, which is more favorable for driving convection. The treatment volumes are compared in Table 12.
| Field direction | Treatment volume (mm³) | Volume fraction (%) |
|---|---|---|
| Same direction | 3.13×10⁷ | 20.89 |
| Opposite direction | 4.37×10⁷ | 29.13 |
The opposite-direction connection increased the effective treatment volume by about 39.3% compared with the same-direction connection. I therefore adopted the opposite-direction inner-annular coil configuration in the final design.
3.3 Effect of frequency and duty cycle
I evaluated nine combinations of frequency (20, 40, 50 Hz) and duty cycle (20%, 40%, 50%). For a given frequency, increasing the duty cycle expanded the magnetic field coverage in the melt, although the rate of expansion gradually slowed down. For a given duty cycle, increasing the frequency reduced the magnetic penetration depth due to a stronger skin effect. Table 13 summarizes the dB/dt and treatment volume for each combination.
| Frequency (Hz) | Duty cycle (%) | dB/dt (mT/s) | Treatment volume (mm³) | Volume fraction (%) |
|---|---|---|---|---|
| 20 | 20 | 18400 | 3.26×10⁷ | 21.73 |
| 20 | 40 | 11800 | 4.05×10⁷ | 27.00 |
| 20 | 50 | 10480 | 4.37×10⁷ | 29.13 |
| 40 | 20 | 24200 | 2.47×10⁷ | 16.47 |
| 40 | 40 | 15800 | 3.10×10⁷ | 20.67 |
| 40 | 50 | 13900 | 3.38×10⁷ | 22.53 |
| 50 | 20 | 27600 | 2.15×10⁷ | 14.33 |
| 50 | 40 | 17200 | 2.82×10⁷ | 18.80 |
| 50 | 50 | 15100 | 3.07×10⁷ | 20.47 |
It is clear that the combination of lower frequency and higher duty cycle produces a lower dB/dt and a larger treatment volume. The parameter set of 20 Hz and 50% duty cycle met the target dB/dt range and delivered the largest effective treatment volume. Hence, I used 20 Hz and 50% duty cycle as the operating conditions for the subsequent simulations and experiments.
3.4 Effect of current intensity on the melt
With the frequency fixed at 20 Hz and the duty cycle at 50%, I studied three current levels: 100 A, 150 A, and 200 A. The increase in current intensity enhanced the magnetic flux density and expanded the high-field region. Table 14 lists the dB/dt, treatment volume, and the maximum velocity at a representative point in the melt.
| Current (A) | dB/dt (mT/s) | Treatment volume fraction (%) | Velocity at point A (m/s) |
|---|---|---|---|
| 100 | 5256 | 16.41 | 0.031 |
| 150 | 7888 | 23.77 | 0.041 |
| 200 | 10480 | 29.13 | 0.049 |
The velocity at the representative point increased by 32.26% when the current increased from 100 A to 150 A, and by 58.06% when the current increased from 100 A to 200 A. In the flow field, the Lorentz force drove the melt from the sidewalls toward the coil center, creating a complex multi-vortex pattern. At 100 A, the flow was concentrated near the coil, and a large stagnation zone appeared in the central region. At 200 A, the forced convection became strong enough to reduce the dead zone and improve the overall flow uniformity. The temperature distributions under different current levels indicated that PMF treatment reduced the horizontal and vertical temperature gradients in the melt. At 200 A, the isotherms became more uniform, and the cooling rate was accelerated. Therefore, I selected 200 A as the optimum coil current.
3.5 Simulation of the small pouring cup
For small sand molds, the maximum melt volume was 6.28×10⁶ mm³. I designed a ring-type pouring cup in which the coil was wound annularly around the melt container. Using the same optimum electromagnetic parameters of 20 Hz, 50% duty cycle, and 200 A, the simulation predicted a dB/dt of 8960 mT/s. The distribution of magnetic flux density inside the melt showed a typical saddle shape, with higher values near the coil and lower values in the central region. Nevertheless, the entire melt remained within the effective magnetic field range, indicating that the small pouring cup could treat the full melt volume. The flow field exhibited four main vortices, with a maximum velocity of about 0.05 m/s in the annular zone. The temperature field was divided into a lower-temperature upper region and a higher-temperature lower region, demonstrating that the pulsed magnetic field promoted heat exchange between the melt and the environment.
3.6 Model verification
To validate the numerical model, I measured the actual current waveform using a Hall sensor and an oscilloscope. The measured current waveform was then loaded into the model, and the predicted magnetic flux density was compared with the measured value at the same location. The measured peak value was 98 mT, while the simulated value was 92 mT, corresponding to a difference of 6.52%. The acceptable agreement confirmed that the model was reliable for subsequent design purposes. This verification gave me confidence in predicting the performance of the pouring cups when processing sand casting parts.
4. Manufacturing and Field Application of the Pouring Cups
After completing the numerical optimization, I manufactured two types of pouring cups: a large plug-type pouring cup and a small ring-type pouring cup. Both devices were designed to apply a pulsed magnetic field to the molten metal in the pouring cup before the start of pouring, thereby improving the solidification structure and mechanical properties of the resulting sand casting parts.
4.1 Plug-type pouring cup for large sand molds
The plug-type pouring cup was designed for large sand molds with an overall size of 834 mm × 700 mm × 590 mm and a maximum melt capacity of 1.6×10⁸ mm³. The cup body was I-shaped with a funnel-shaped bottom to reduce the impact of the melt during pouring. The magnetic field generator consisted of an iron core and a coil of 120 turns. The iron core had dimensions of 80 mm × 180 mm × 140 mm. The coil was made of a conductor with a cross-sectional area of 70 mm². A heat-insulating plate was placed inside the cup body to protect the coil from the high-temperature melt, and a cooling fan was installed to further dissipate heat. The entire cup was suspended by lifting lugs and could be easily moved by an overhead crane. Figure 1 shows the schematic assembly of the plug-type pouring cup and the actual device.
| Parameter | Without PMF | With PMF |
|---|---|---|
| dB/dt (mT/s) | 0 | 10480 |
| Melt temperature at treatment (°C) | 730 | 730 |
| Pouring temperature (°C) | 663 | 663 |
| Treatment time (s) | 240 | 240 |
I conducted field experiments using ZL205A and ZL114A alloys. The melt was heated to 730 °C, and the sprue was closed with a plug. The molten metal was poured into the pouring cup, and the power supply was switched on. The melt was treated for 240 s. When the temperature dropped to 663 °C, a portion of the melt was extracted and poured into a test bar mold for quality assessment. The plug was then removed, and the remaining melt was poured into the sand mold.
The ZL205A samples were examined in both the as-cast and T5 heat-treated conditions. The microstructure comparison showed that the coarse rosette-shaped primary α-Al grains were significantly refined after PMF treatment. The average grain size decreased from 62.27 μm in the untreated casting to 50.64 μm after treatment, which represents a refinement of 18.7%. The mechanical properties of the as-cast test bars are listed in Table 16, and the T5 heat-treated properties are listed in Table 17.
| Condition | Tensile strength (MPa) | Elongation (%) | Hardness (HV) |
|---|---|---|---|
| Without PMF | 309 | 7.0 | 54 |
| With PMF | 349 | 12.5 | 70 |
| Alloy | Condition | Tensile strength (MPa) | Elongation (%) | Hardness |
|---|---|---|---|---|
| ZL205A | Without PMF | 236 | 3.5 | 116 HBW |
| With PMF | 350 | 10.0 | 125 HBW | |
| ZL114A | Without PMF | 299 | 3.5 | 96 HBW |
| With PMF | 344 | 3.5 | 101 HBW |
For the ZL205A castings, the T5 heat-treated tensile strength increased by 48.31%, the elongation increased by 185.71%, and the hardness increased by 7.76% after PMF treatment. For the ZL114A castings, the tensile strength and hardness increased by 15.05% and 5.21%, respectively. These results clearly demonstrate that the pulsed magnetic field applied in the pouring cup can refine the solidification structure and improve the mechanical performance of sand casting parts.
4.2 Small pouring cup for small sand molds
For small sand molds, conventional steel hand ladles are often used to carry the melt from the furnace to the mold. However, ordinary carbon steel ladles can disturb the magnetic field distribution. I therefore selected a 304 stainless steel hand ladle with an inner diameter of 220 mm, a depth of 200 mm, and a rod length of 1500 mm. The small pouring cup was designed as a ring-shaped structure with an outer diameter of 404 mm and a height of 220 mm. The coil was wound annularly with a copper conductor of 4 mm × 8 mm cross-section, with 6 layers and 20 turns per layer, giving a total of 120 turns. A heat-insulating layer and outer shell were added to protect the coil. The small pouring cup was placed on the sand mold, and the hand ladle was inserted into it.
In the first stage, I conducted test-bar experiments to screen the optimal electromagnetic parameters. The ZL114A melt was heated to 730 °C and poured into the hand ladle, which was then placed into the small pouring cup. PMF treatment was applied for 90 s, and the melt was poured into a test-bar mold at 700 °C. The parameters are listed in Table 18.
| No. | dB/dt (mT/s) | Treatment temperature (°C) | Pouring temperature (°C) | Mold type |
|---|---|---|---|---|
| 1 | 0 | 730 | 700 | Test bar |
| 2 | 8960 | 730 | 700 | Test bar |
| 3 | 17600 | 730 | 700 | Test bar |
| 4 | 18400 | 730 | 700 | Test bar |
| 5 | 23000 | 730 | 700 | Test bar |
The microstructure analysis of the ZL114A test bars revealed that the primary α-Al phase was significantly refined after PMF treatment. The grain size data are presented in Table 19. The optimal condition was 8960 mT/s, which reduced the grain size from 77.66 μm to 60.89 μm, a reduction of 21.59%. The as-cast mechanical properties of the test bars are given in Table 20.
| dB/dt (mT/s) | Grain size (μm) |
|---|---|
| 0 | 77.66 |
| 8960 | 60.89 |
| 17600 | 66.91 |
| 18400 | 65.31 |
| 23000 | 61.57 |
| dB/dt (mT/s) | Tensile strength (MPa) | Elongation (%) | Hardness (HV) |
|---|---|---|---|
| 0 | 199.75 | 5.12 | 68.44 |
| 8960 | 240.00 | 9.00 | 77.14 |
| 17600 | 222.50 | 6.80 | 72.35 |
| 18400 | 228.00 | 7.20 | 73.80 |
| 23000 | 236.50 | 8.10 | 75.62 |
Based on these results, I selected dB/dt values of 8960 mT/s and 23000 mT/s for the subsequent sand mold casting experiments. The parameters are listed in Table 21. The ZL114A melt was treated in the small pouring cup and poured into a box-type sand mold. Samples were cut from the castings, subjected to T5 heat treatment, and tested for mechanical properties. The results are summarized in Table 22.
| No. | dB/dt (mT/s) | Treatment temperature (°C) | Pouring temperature (°C) | Mold type |
|---|---|---|---|---|
| 1 | 0 | 730 | 700 | Sand mold |
| 2 | 8960 | 730 | 700 | Sand mold |
| 3 | 23000 | 730 | 700 | Sand mold |
| dB/dt (mT/s) | Tensile strength (MPa) | Elongation (%) | Hardness (HV) |
|---|---|---|---|
| 0 | 200.0 | 6.52 | 67.16 |
| 8960 | 262.0 | 9.02 | 81.68 |
| 23000 | 254.5 | 7.55 | 78.56 |
Compared with the untreated condition, the tensile strength after T5 heat treatment increased by 31% at 8960 mT/s and by 27.25% at 23000 mT/s. The elongation increased by 38.34% at 8960 mT/s and by 15.8% at 23000 mT/s. The hardness increased by 21.62% at 8960 mT/s and by 16.97% at 23000 mT/s. These results confirm that the small ring-type pouring cup can effectively improve the performance of small sand casting parts without adding any grain refiner to the melt.
5. Conclusion
In this research, I designed and developed two types of pulsed magnetic field pouring cups for sand casting, combining laboratory experiments, numerical simulation, and industrial trials. The main conclusions are as follows:
First, laboratory sand-casting simulations on ZL205A alloy demonstrated that a pulsed magnetic field with a dB/dt of 8000 mT/s significantly refined the primary α-Al grains, reducing the average grain size by up to 26.83%. The yield strength, tensile strength, and elongation increased by 28%, 17.5%, and 91.66%, respectively, while the hardness increased by about 13% at the edge and 9.3% at the center. The pulsed magnetic field lowers the nucleation energy barrier and increases the nucleation rate, thereby refining the as-cast structure and improving the mechanical behavior of sand casting parts.
Second, the multi-physics numerical model built in COMSOL Multiphysics enabled a comprehensive analysis of the magnetic field, flow field, and temperature field inside the pouring cup. For the plug-type pouring cup, the inner-annular structure with opposite-direction coil connection provided a larger effective magnetic treatment volume of 4.37×10⁷ mm³, corresponding to 29.13% of the total melt, with a dB/dt of 10480 mT/s. For the small pouring cup, the ring-type design produced a dB/dt of 8960 mT/s, and the entire melt volume of 6.28×10⁶ mm³ was effectively treated. The simulations also showed that lower frequency and higher duty cycle favor magnetic penetration and treatment uniformity, and that increasing the coil current enhances forced convection and reduces temperature gradients in the melt.
Third, the manufactured plug-type pouring cup and small pouring cup were successfully applied in industrial sand casting experiments. In the plug-type cup, the tensile strength of ZL205A castings in the T5 heat-treated condition increased by 48.31% after PMF treatment, while the elongation increased by 185.71%. For ZL114A castings in the plug-type cup, the tensile strength increased by 15.05%. In the small pouring cup, the optimal dB/dt of 8960 mT/s improved the tensile strength, elongation, and hardness of ZL114A castings in the T5 heat-treated condition by 31%, 38.34%, and 21.62%, respectively. These findings validate the feasibility of applying pulsed magnetic field treatment in the pouring cup as an effective method to improve the quality and reliability of sand casting parts.
In summary, the combination of non-contact pulsed magnetic field treatment with optimized pouring cup design offers a promising route for grain refinement and mechanical property enhancement in sand casting. The numerical simulation methodology and experimental procedures developed in this work can be extended to other alloy systems and casting geometries, providing a practical tool for the digital design and optimization of high-performance sand casting parts.