The pursuit of lightweight commercial vehicles has driven the development of innovative manufacturing processes, particularly in the domain of structural components. While passenger vehicles have benefited from advanced lightweighting technologies, commercial vehicle load-bearing structures such as crossbeams present unique challenges. Traditional steel fabrication involves multi-step forming and welding operations, while lightweight alloys often lack the necessary strength. Ductile iron, with its excellent castability, superior mechanical properties, and cost-effectiveness, emerges as a compelling alternative for integrated structural components. In this study, I focused on the integral design and sand casting process optimization of a commercial vehicle crossbeam, with the goal of achieving weight reduction while maintaining structural integrity. The central technical challenge was to control sand casting defects, including shrinkage cavities and gas porosity, through systematic numerical simulation and experimental validation.
Materials and Experimental Procedures
The material selected for this research was QT800-5 ductile iron, which offers an excellent combination of high strength and moderate ductility. The raw materials included high-purity Q10 pig iron with low sulfur and manganese contents, carbon steel scrap, and returns. A FeSiMg6RE2 rare-earth magnesium nodulizer was employed for spheroidization treatment via the cored wire feeding method, while a Si-Ba inoculant was used for in-ladle inoculation followed by 75SiFe for stream inoculation. The chemical composition of the raw materials is summarized in Table 1.
| Material | C | Si | Mn | P | S | RE | Mg | Ba | Fe |
|---|---|---|---|---|---|---|---|---|---|
| High-purity pig iron | 4.25 | 0.23 | 0.05 | 0.018 | 0.010 | — | — | — | Balance |
| Steel scrap | 0.26 | 0.22 | 0.2 | 0.02 | 0.01 | — | — | — | Balance |
| Nodulizer | — | 41-44 | — | — | — | 2-3 | 6-7 | — | Balance |
| Si-Ba inoculant | — | 60-65 | — | — | — | — | — | 4-6 | Balance |
| 75SiFe | — | 72-80 | — | — | — | — | — | — | Balance |
Melting was conducted in a 1000 kg medium-frequency induction furnace. The charge materials were placed in the furnace in sequence, and the temperature was raised to the range of 1530–1570 °C for superheating and purification. After slag removal, alloying elements such as copper, nickel, and molybdenum were added. The melt was then tapped at 1490–1520 °C for nodularization treatment. The spheroidization process consumed 1.35–1.45% nodulizer through cored wire feeding, followed by a three-stage inoculation procedure. The pouring temperature was controlled in the range of 1380–1420 °C, and the entire pouring operation was completed within 10 minutes to prevent nodulizer fading. The actual pouring was performed on an automated molding line.
Structural Design and Static Analysis
The original crossbeam structure consisted of seven individual steel components: four L-shaped steel plates, two mounting brackets, and one U-shaped plate. The total mass of the assembly was 50.5 kg. Table 2 presents the technical parameters of the original components. The assembly process required multiple forming, machining, and welding steps, which significantly increased production complexity and cost.
| Component | Material | Elastic modulus (MPa) | Yield strength (MPa) | Poisson’s ratio | Mass (kg) |
|---|---|---|---|---|---|
| Mounting bracket | QT550 | 172000 | 320 | 0.275 | 9.78 × 2 |
| Steel plate | B510L | 200000 | 355 | 0.3 | (4.46 + 5.07) × 2 |
To achieve structural integration, I employed an equal-stiffness replacement approach to determine the appropriate wall thickness for the ductile iron design. The governing equation is expressed as:
$$ E_{sl} t_{sl}^3 = E_{dl} t_{dl}^3 $$
where \(E_{sl}\) and \(E_{dl}\) represent the elastic moduli of steel and ductile iron, respectively, and \(t_{sl}\) and \(t_{dl}\) denote their corresponding thicknesses. The calculation yielded a minimum wall thickness of 8.0 mm for the ductile iron component. Given that the torsional loading condition was identified as the most critical from the static analysis of the original structure, the integrated design employed a circular shaft cross-section for the main body to enhance torsional stiffness. The final integrated crossbeam design, with overall dimensions of 839 × 403 × 221 mm, achieved a mass reduction to 36 kg—a 28.7% decrease compared to the original structure.
Finite element analysis was conducted to evaluate the structural performance of both the original and integrated crossbeam designs. Four representative loading conditions were simulated: acceleration, turning, torsion, and emergency braking. For the integrated design, a three-dimensional tetrahedral mesh with 3,309,713 elements and 133,147 nodes was generated for the casting, while shell elements were used for the frame components. The maximum equivalent stresses for each condition are compared in Table 3.
| Loading condition | Original crossbeam (MPa) | Integrated crossbeam (MPa) |
|---|---|---|
| Acceleration | 85.1 | 69.3 |
| Turning | 396.1 | 311.1 |
| Torsion | 680.6 | 558.4 |
| Emergency braking | 95.4 | 95.4 |
The analysis revealed that the maximum stress under all conditions occurred at the threaded holes due to stress concentration. For the integrated crossbeam, the peak stress of 558.4 MPa occurred under torsional loading, which was below the allowable stress of 800 MPa for QT800-5 ductile iron. The corresponding safety factor was 1.4, satisfying the strength requirements for commercial vehicle applications.
Casting Process Design
The casting process design for the integral crossbeam required careful consideration of several key aspects: parting surface selection, sand core design, and gating system layout. The parting surface was positioned at the maximum cross-sectional contour of the casting to facilitate mold release and minimize the projected area on the parting plane. Self-hardening resin sand was selected for the core material, and a comprehensive core venting system was designed to facilitate the escape of gases generated during pouring.
The casting is characterized by significant wall thickness variations (8.0 mm to 55.4 mm), which presents challenges in achieving uniform solidification. Ductile iron exhibits a mushy solidification mode, characterized by a wide freezing range and a tendency for isolated liquid pools to form. To address these challenges, I employed a closed gating system with a cross-sectional area ratio of F直:F横:F内 = 2.50:1.25:1. Four ingates were designed to achieve uniform filling across the long-span casting. The ingate area was calculated using the flow-rate method:
$$ A_g = \frac{V}{v_g \times t} $$
where \(A_g\) is the total ingate cross-sectional area, \(V\) is the volume of molten metal passing through the ingates, \(v_g\) is the flow velocity, and \(t\) is the mold filling time. The calculated ingate area was 1200 mm², resulting in a runner area of 1500 mm² and a sprue area of 3000 mm². The pouring time was determined from the empirical relationship:
$$ t = f\left(\sqrt{G} + \sqrt{0.2WG^3}\right) $$
where \(f\) is the material coefficient (0.6–0.8), \(G\) is the casting mass, and \(W\) is the average wall thickness. For an initial pouring temperature of 1400 °C and pouring speed of 11 cm/s, the calculated filling time was 7.27 seconds. The casting shrinkage allowance for this pearlitic ductile iron was set at 1.2%.
Numerical Simulation of the Casting Process
The finite element method was employed to simulate both the mold filling and solidification processes. A mesh size of 2 mm was ultimately selected, yielding a total of 1.2266795 × 10⁷ elements. The physical properties of QT800-5 used in the simulation are listed in Table 4, including temperature-dependent density, thermal conductivity, and specific heat. The heat transfer coefficient between the molten metal and the sand mold was set to 700 W/m²·K, with gravity acting in the negative Z direction.
| Property | Value |
|---|---|
| Density (kg/m³) | 6459 |
| Liquidus temperature (K) | 1457 |
| Solidus temperature (K) | 1418 |
| Latent heat (kJ/kg) | 210 |
| Specific heat capacity (kJ/kg·K) | 0.78 |
Initial Gating System Evaluation
The initial gating system design featured a four-ingate configuration. Simulation results at a pouring temperature of 1400 °C and pouring speed of 11 cm/s indicated that the mold filling process was relatively stable, with complete filling achieved in 7.27 seconds. However, solidification analysis revealed insufficient feeding capability at the two ends of the casting. The solidification time at the thick-walled end regions reached 436 seconds, creating isolated liquid pools that could not be compensated during the final solidification stage. These regions exhibited a high risk of sand casting defects, particularly shrinkage cavities and porosity.
Gating System Optimization
To enhance the feeding capability and mitigate the risk of sand casting defects, I optimized the gating system by increasing the number of risers from four to eight, each measuring 70 mm in diameter and 120 mm in height, positioned strategically at both ends of the crossbeam. The optimized scheme was evaluated under identical process conditions. The solidification time at the critical end regions was significantly reduced from 436 s to 307.67 s, and the isolated liquid pools were effectively eliminated. The quantitative comparison of shrinkage defects between the initial and optimized schemes is presented in Table 5.
| Scheme | Shrinkage volume (mm³) | Defect reduction (%) |
|---|---|---|
| Initial gating system | 751.16 | — |
| Optimized gating system | 685.85 | 8.69 |
The optimized gating system effectively redirected most shrinkage defects to the risers, which serve their intended purpose of collecting the final liquid metal. This improvement was crucial in minimizing sand casting defects within the actual load-bearing structure of the crossbeam.
Effect of Process Parameters on Casting Quality
Following the gating system optimization, I systematically investigated the influence of two critical process parameters—pouring temperature and pouring speed—on the formation of sand casting defects. The base case was defined as a pouring temperature of 1400 °C, pouring speed of 11 cm/s, and mold temperature of 20 °C.
Effect of Pouring Temperature
Three pouring temperatures were evaluated: 1380 °C, 1400 °C, and 1420 °C. The mold filling process was stable across all temperatures, with no observable cold shuts or misruns. However, the gas entrainment (卷气) varied notably. As the pouring temperature increased from 1380 °C to 1400 °C, the gas entrainment volume decreased from 1700.58 mm³ to 1682.44 mm³ due to improved fluidity of the molten metal. A further increase to 1420 °C resulted in higher gas entrainment (1697.82 mm³), attributed to the reduced surface tension and increased gas evolution at higher temperatures. Regarding solidification, the shrinkage porosity volume showed a monotonic increasing trend with temperature: 667.92 mm³ at 1380 °C, 685.85 mm³ at 1400 °C, and 702.11 mm³ at 1420 °C. Higher pouring temperatures increase the total superheat that must be dissipated, enlarging the temperature gradients and promoting the formation of isolated liquid pools, which deteriorate the feeding efficiency and exacerbate sand casting defects.
Effect of Pouring Speed
Pouring speeds of 10 cm/s, 11 cm/s, and 12 cm/s were evaluated. At the lowest speed, the filling process was the most stable, resulting in the lowest gas entrainment volume of 1655.77 mm³. However, the flow velocity considerably influenced the filling stability—at the highest speed, significant splashing and localized eddies were observed, increasing the gas entrainment to 1700.39 mm³. In contrast, the shrinkage porosity volume exhibited a decreasing trend with increasing pouring speed: 696.92 mm³ at 10 cm/s, 685.85 mm³ at 11 cm/s, and 680.85 mm³ at 12 cm/s. Faster pouring reduces in-mold heat losses, resulting in more uniform temperature distribution and improved feeding efficiency, thereby diminishing shrinkage-related sand casting defects.
Orthogonal Experiments and Variance Analysis
To comprehensively evaluate the coupled effects of the two process parameters, I designed a two-factor, three-level orthogonal experiment matrix. The nine experimental runs and their corresponding results are summarized in Table 6.
| Trial no. | Pouring speed (cm/s) | Pouring temp. (°C) | Gas entrainment (mm³) | Shrinkage (mm³) |
|---|---|---|---|---|
| 1 | 10 | 1380 | 1677.19 | 679.18 |
| 2 | 10 | 1400 | 1655.77 | 696.92 |
| 3 | 10 | 1420 | 1668.84 | 713.20 |
| 4 | 11 | 1400 | 1682.44 | 685.85 |
| 5 | 11 | 1420 | 1697.82 | 702.11 |
| 6 | 11 | 1380 | 1700.58 | 667.92 |
| 7 | 12 | 1420 | 1704.32 | 697.20 |
| 8 | 12 | 1400 | 1700.39 | 680.85 |
| 9 | 12 | 1380 | 1721.51 | 663.00 |
Analysis of variance (ANOVA) was applied to assess the statistical significance of each factor on the two quality indices. Table 7 presents the ANOVA results for gas entrainment volume.
| Source | Sum of squares | df | Mean square | F-value | p-value |
|---|---|---|---|---|---|
| Pouring speed | 2643.000 | 2 | 1321.500 | 24.266 | 0.006 |
| Pouring temp. | 567.608 | 2 | 283.804 | 5.211 | 0.077 |
| Error | 108.919 | 2 | 54.459 | — | — |
The ANOVA results demonstrate that pouring speed has a statistically significant effect on gas entrainment (p = 0.006 < 0.05), whereas the effect of pouring temperature is not significant (p = 0.077 > 0.05). Conversely, for shrinkage porosity, the analysis (Table 8) reveals that pouring temperature is the dominant significant factor (p = 0.017 < 0.05), while pouring speed does not exert a statistically significant influence (p = 0.167 > 0.05).
| Source | Sum of squares | df | Mean square | F-value | p-value |
|---|---|---|---|---|---|
| Pouring speed | 406.121 | 2 | 203.061 | 2.847 | 0.167 |
| Pouring temp. | 1747.142 | 2 | 873.571 | 12.247 | 0.017 |
| Error | 142.659 | 2 | 71.330 | — | — |
Based on the factor-level means, the optimal parameter combination was identified as a pouring speed of 11 cm/s and a pouring temperature of 1380 °C. The selected parameters minimize gas entrainment while concurrently reducing shrinkage porosity. A confirmation simulation using these optimized parameters demonstrated excellent filling stability and refined solidification control, with the shrinkage porosity volume reduced to 667.92 mm³—an 11.74% improvement over the initial process scheme.
Production Trial and Characterization
To validate the optimized casting process, integrated crossbeam castings were produced using the final process parameters: pouring temperature of 1380 °C, pouring speed of 11 cm/s, and mold temperature of 20 °C. The chemical composition of the produced castings is shown in Table 9. The carbon equivalent and residual magnesium content met the specification requirements, while the phosphorus and sulfur concentrations were sufficiently low to prevent the formation of phosphide eutectics and sulfide inclusions.
| Element | C | Si | Mn | P | S | Mg | Ni | Mo | Cu |
|---|---|---|---|---|---|---|---|---|---|
| wt.% | 3.61 | 2.67 | 0.105 | 0.007 | 0.007 | 0.043 | 0.438 | 0.151 | 0.803 |
Metallographic specimens were extracted from three representative locations: Position 1 (near the ingate), Position 2 (the thin-walled shaft region), and Position 3 (far from the ingate). X-ray inspection of the critical load-bearing regions revealed no internal defects such as gas holes, cracks, or shrinkage porosity, confirming the effectiveness of the optimized process in controlling sand casting defects. The metallographic analysis, summarized in Table 10, revealed a predominantly pearlitic microstructure with ≥90% pearlite content in all samples. The graphite nodularity consistently reached grade 2, with graphite size distributions in the 6–7 range. Position 2, which experienced the fastest cooling rate, exhibited the smallest graphite nodules and the highest nodule count (252 nodules/mm²), while Position 1 near the ingate displayed larger graphite and a lower nodule count (180 nodules/mm²).
| Position | Nodularity grade | Graphite size grade | Pearlite content (%) | Carbide (%) | Phosphide eutectic |
|---|---|---|---|---|---|
| 1 (near ingate) | 2 | 6 | ≥90 | 1 | None |
| 2 (thin-walled) | 2 | 7 | ≥95 | 1 | None |
| 3 (far from ingate) | 2 | 6 | ≥90 | 1 | None |
Tensile tests were performed on specimens extracted from the three positions, and the results are presented in Table 11. All specimens exceeded the QT800-5 specification requirements, with ultimate tensile strengths above 854 MPa, yield strengths above 546 MPa, and elongations above 6.0%. The thin-walled position (Position 2) exhibited the highest strength values (UTS 891 MPa, YS 562 MPa) but the lowest elongation (6.0%), consistent with the finer graphite structure resulting from faster cooling. The near-ingate position (Position 1) demonstrated the best ductility (7.6% elongation) because of its slower cooling rate and coarser graphite. The hardness profile across the different positions, shown in Table 12, consistently exceeded 270 HV with minimal variation, indicating uniform mechanical properties throughout the casting.
| Position | Yield strength (MPa) | Ultimate tensile strength (MPa) | Elongation (%) |
|---|---|---|---|
| 1 | 546 | 875 | 7.6 |
| 2 | 562 | 891 | 6.0 |
| 3 | 552 | 854 | 6.5 |
| Position | Measurement 1 | Measurement 2 | Measurement 3 | Average |
|---|---|---|---|---|
| 1 | 280.8 | 285.8 | 285.8 | 284.1 |
| 2 | 272.2 | 293.8 | 279.7 | 281.9 |
| 3 | 297.3 | 280.8 | 279.7 | 285.9 |
Fracture surface analysis of the tensile specimens was performed using scanning electron microscopy. The fracture morphology exhibited typical brittle features characteristic of high-strength ductile iron—namely, cleavage facets and river patterns. As the distance from the ingate increased, the cleavage facet size grew and the river patterns became more pronounced, confirming the reduced ductility observed at positions farther from the ingate. Energy-dispersive X-ray spectroscopy (EDS) analysis identified oxide inclusions, primarily calcium oxide (CaO), within the matrix. The sharp, angular morphology of these inclusions acts as severe stress concentrators, making them significantly more detrimental to mechanical properties than the spheroidal graphite particles, which do not form cavities as readily.
Conclusion
In this research, I systematically designed and optimized the sand casting process for an integral ductile iron crossbeam for commercial vehicles. The key findings and contributions are summarized as follows:
(1) The integrated structural design reduced the crossbeam mass from 50.5 kg to 36 kg, a reduction of 28.7%, while maintaining structural integrity. Finite element analysis confirmed that the maximum equivalent stress of 558.4 MPa under torsional loading satisfies the safety factor of 1.4 for commercial vehicle applications.
(2) The gating system was optimized by increasing the riser count from four to eight, strategically positioned at the beam ends. This modification reduced the shrinkage porosity volume from 751.16 mm³ to 685.85 mm³, an 8.69% reduction, effectively controlling sand casting defects in critical load-bearing regions.
(3) Orthogonal experiments combined with variance analysis revealed that pouring speed significantly influences gas entrainment (p = 0.006), while pouring temperature significantly influences shrinkage porosity (p = 0.017). The optimal parameter combination was determined as a pouring speed of 11 cm/s, a pouring temperature of 1380 °C, and a mold temperature of 20 °C.
(4) Production trial castings exhibited stable chemical composition, uniform microstructure with ≥90% pearlite content, and consistently high nodularity (grade 2). The mechanical properties exceeded the QT800-5 specification requirements: tensile strength ≥ 854 MPa, yield strength ≥ 546 MPa, elongation ≥ 6.0%, and hardness ≥ 270 HV.
(5) The fracture surface analysis revealed brittle fracture characteristics with cleavage facets and river patterns. Oxide inclusions, particularly calcium oxide, were identified as more detrimental to mechanical properties than the spheroidal graphite, highlighting the importance of melt cleanliness in minimizing sand casting defects.
This research demonstrates the combined effectiveness of structural integration, numerical simulation, and statistical optimization in developing a reliable sand casting process for high-performance ductile iron components. The established methodology provides a valuable framework for the lightweight design and quality control of commercial vehicle structural castings.