The lightweight demand of the automotive industry has driven the development and application of integrated molding technology for body structural components. In the field of commercial vehicle lightweighting, the mature experience of passenger car body weight reduction cannot be directly transplanted. Especially for core load-bearing components such as commercial vehicle crossbeams, lightweight alloys struggle to meet the strength requirements, while traditional steel materials are constrained by complex multi-step processing. In contrast, ductile iron, leveraging its integrated molding process advantages, combined with superior mechanical properties and adaptability, has become an innovative choice for the lightweight transformation of commercial vehicles. This article focuses on the lightweight design of commercial vehicle crossbeam structural components, achieving the lightweight goal through the integrated design of crossbeam structures. Under the premise of ensuring structural load-bearing performance, the integrated ductile iron casting process is adopted to replace the traditional steel welding structure. Numerical simulation technology is employed to analyze the casting process and optimize the casting parameters. The formation mechanism and control methods of casting defects are systematically studied through this approach.
1. Introduction and Application of Ductile Iron
As an important branch of engineering cast iron, ductile iron exhibits outstanding mechanical properties due to its unique spheroidal graphite microstructure. Compared with traditional cast iron materials, this material possesses excellent tensile strength and toughness, with its comprehensive mechanical performance indicators comparable to carbon steel. In 1943, the International Nickel Company (INCO) conducted experiments on gray cast iron, inspired by the good strengthening and toughening effect of Mg on white cast iron. They successfully obtained ductile iron with excellent comprehensive mechanical properties by adding Ni-Mg or Cu-Mg master alloys containing Mg to the gray cast iron melt. This marked the birth of the method of using Mg as a spheroidizing agent for producing ductile iron. In 1948, H. Morrogh and W.J. Willimas successfully developed ductile iron using Ce as a spheroidizing agent, combined with FeSi and Si-Mn-Zr inoculants. However, due to the high cost of Ce and its weak spheroidizing effect, industrial production of ductile iron could not be realized. Up to now, most industrial ductile iron worldwide uses Mg as the spheroidizing element, while Ce is mainly used to inhibit anti-graphitizing elements and refine molten iron.
The development of ductile iron in China began in 1949, when Professor Wang Zunming of Tsinghua University used various magnesium alloys to trial-produce ductile iron, successfully exploring a series of magnesium alloys that could serve as spheroidizing agents. In the mid-1960s, researchers successfully developed rare earth magnesium ductile iron, which solved the problems of high rejection rates by improving both the spheroidizing agent and the spheroidizing treatment method. Since the 1970s, austempered ductile iron developed in China and Finland achieved a significant improvement in the mechanical properties of ductile iron, further expanding its application scope. Since 2001, China’s ductile iron production has reached the world’s leading level. The production of ductile iron increased from 10.5 million tons in 2011 to 15.3 million tons in 2020, representing a growth of 45.7% in ten years. The proportion of ductile iron in total cast iron production increased from 25.30% to 29.45%. The ratio of ductile iron to gray cast iron increased from 0.53 to 0.68.
2. Materials and Experimental Methods
2.1 Raw Materials and Melting Process
The experimental material used in this research is QT800-5 ductile iron, which represents a high-strength, high-ductility grade widely applied in the automotive industry. The specific chemical composition of the main raw materials is presented in the following table.
| Raw 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 |
The melting process was carried out in a 1000 kg medium-frequency induction furnace. The specific melting procedure was as follows: first, high-purity pig iron, steel scrap, return scrap, and carburizing agent were added to the furnace and melted. When the melting temperature reached 1530-1570 ℃, the melt was held at high temperature to purify the molten iron. Then, electrolytic copper, nickel plate, and ferromolybdenum were added. When the melt temperature reached 1490-1520 ℃, the melt was tapped for nodularization treatment. The nodularizing treatment was performed using the cored wire feeding method with FeSiMg6RE2 alloy as the nodulizing agent, with a particle size range of 10-30 mm and an addition ratio of 1.35%-1.45% of the molten iron mass. The treatment temperature was maintained at 1490-1520 ℃. The inoculation process adopted a three-stage procedure: primary inoculation used Si-Ba inoculant with a particle size of 3-10 mm at 0.3%-0.4% of the iron mass; secondary inoculation was carried out when two-thirds of the molten iron was transferred to the transport ladle, using 0.2%-0.3% of the iron mass; finally, stream inoculation was implemented during pouring with 0.1%-0.2% Si-Ba complex inoculant. The pouring temperature was controlled within 1380-1420 ℃, and the entire pouring process needed to be completed within 10 minutes to avoid nodulizing and inoculation fading.
2.2 Finite Element Theory and Numerical Simulation
The finite element method is a crucial numerical solution technique for continuous system problems. It discretizes a continuous body into a finite number of simple units, where physical quantities at each unit are expressed as nodal function values through interpolation functions. The main steps of the finite element analysis process include discretization, element analysis, assembly of the global stiffness matrix, constraint processing, and post-processing visualization analysis. The finite element method has strong expandability and can be extended to other continuous medium analysis fields by adjusting corresponding physical quantities.
The mold filling and solidification processes were simulated using commercial casting simulation software, which provides comprehensive casting process simulation capabilities. The key simulation parameters used in this study are summarized in the table below.
| Parameter | Value |
|---|---|
| Density | 6459 kg/m³ |
| Liquidus temperature | 1457 K |
| Solidus temperature | 1418 K |
| Latent heat | 210 kJ/kg |
| Specific heat capacity | 0.78 kJ/(kg·K) |
| Heat transfer coefficient (metal-mold) | 700 W/(m²·K) |
| Gravity direction | Z-axis negative |
| Gravity magnitude | 980 cm/s² |
| Mesh size | 2 mm |
3. Structural Analysis and Integrated Design of the Commercial Vehicle Crossbeam
3.1 Static Analysis of the Original Crossbeam Structure
Static analysis is used to calculate displacement, stress, strain, and force of the structure under fixed loads, with the core objective of evaluating the stability and reliability of the structure under static loads without considering inertia and damping effects. The main assumptions in static analysis include: equilibrium state assumption, rigid body assumption, linear elastic assumption, and time independence assumption. Objects satisfying these assumptions are ideal linear elastic bodies, which facilitate the establishment of mathematical models and provide support for the reliability and scientificity of static analysis results.
The original commercial vehicle crossbeam weighed 50.5 kg and consisted of 7 parts: 4 L-shaped steel plates, 2 fixing seats, and 1 U-shaped steel plate. The fixing seats and steel plates were made of different materials. The main technical parameters of the crossbeam are shown in the table below.
| Component | Material | Elastic Modulus (MPa) | Yield Strength (MPa) | Poisson’s Ratio | Mass (kg) | Dimensions (mm) |
|---|---|---|---|---|---|---|
| Fixing seat | QT550 | 172000 | 320 | 0.275 | 9.78×2 | 814×237×600 |
| Steel plate | B510L | 200000 | 355 | 0.3 | (4.46+5.07)×2 | — |
The crossbeam structure was placed in the vehicle frame environment for finite element analysis to avoid distortion of stress analysis results. The frame longitudinal beams and crossbeam steel plates were meshed using elastic shell elements, while the crossbeam casting with its complex shape was meshed using first-order tetrahedral elements. Bolt, rivet connections, and thrust rods were simulated using rigid elements in Hypermesh. The original crossbeam structure was subjected to static analysis under four typical operating conditions: acceleration, turning, torsion, and emergency braking. The analysis results revealed that the maximum equivalent stress of 680.6 MPa occurred under the torsion condition at the threaded hole location, which exceeded the material yield strength threshold of 550 MPa.
3.2 Integrated Structural Design of the Crossbeam
Considering the structural design principles and key factors, the integrated ductile iron crossbeam casting structure was designed based on the original crossbeam. The design considerations included: minimum wall thickness criterion, rib direction flexibility criterion, and favorable stress state criterion. For the material selection, the equivalent stiffness substitution method was used to calculate the integrated crossbeam wall thickness based on the following formula:
$$E_{sl} \cdot t_{sl}^{3} = E_{dl} \cdot t_{dl}^{3} \quad (3-1)$$
where $E_{dl}$ and $E_{sl}$ are the elastic moduli of ductile iron and steel, respectively; $t_{dl}$ and $t_{sl}$ are the thicknesses of ductile iron and steel, respectively. The calculation showed that the minimum thickness of ductile iron was 8.0 mm to ensure sufficient stiffness of the crossbeam. Based on the static analysis results of the original crossbeam, the main structure of the integrated crossbeam was changed from an I-shaped section to a circular shaft section to enhance the torsional resistance and avoid structural instability. Strengthening ribs were optimally designed for the weak structural areas, and support platforms and assembly surfaces were integrated in the central shaft area. The maximum profile dimensions of the integrated crossbeam were 839×403×221 mm.
3.3 Static Analysis of the Integrated Crossbeam
The integrated crossbeam was also placed in the vehicle frame environment for finite element analysis. The stress analysis was performed under the same four operating conditions as the original crossbeam: acceleration, turning, torsion, and emergency braking. The static analysis results showed that the maximum equivalent stress of the integrated crossbeam was 558.4 MPa, also located at the threaded hole under the torsion condition. The safety factor was calculated as follows:
$$\text{Safety Factor} = \frac{\sigma_{yield}}{\sigma_{max}} = \frac{800 \text{ MPa}}{558.4 \text{ MPa}} = 1.4 \quad (3-2)$$
The mass of the integrated crossbeam was reduced from 50.5 kg to 36 kg, a decrease of 28.7%, satisfying the strength requirements of commercial vehicles. The comparison of the static analysis results for both crossbeam structures is summarized in the following table.
| Operating Condition | Original Crossbeam Maximum Stress (MPa) | Integrated Crossbeam Maximum Stress (MPa) | Location |
|---|---|---|---|
| Acceleration | 85.1 | 69.3 | Threaded hole |
| Turning | 396.1 | 311.1 | Threaded hole |
| Torsion | 680.6 | 558.4 | Threaded hole |
| Emergency braking | 95.4 | 95.4 | Threaded hole |
4. Casting Process Design and Numerical Simulation
4.1 Casting Process Design
The selection of the parting surface is critical for ensuring product forming quality and reducing defects. The maximum outer contour of the crossbeam was selected as the parting surface. For the sand core design, the internal cavity structure of the crossbeam casting was relatively complex, requiring the use of sand cores. Self-hardening resin sand was selected as the sand core material. Since the sand core is usually surrounded by molten metal inside the casting, it must have low moisture absorption and low gas production, while ensuring good exhaust capability. A complete sand core exhaust system was established to prevent gas-related casting defects.
The pouring temperature and pouring speed were initially determined based on structural characteristics and empirical formulas. The pouring time was calculated using the following empirical formula:
$$t = f\left(\sqrt{G} + \sqrt[3]{0.2WG}\right) \quad (4-1)$$
where $f$ is the material coefficient (generally 0.6-0.8), $G$ is the casting mass, and $W$ is the casting wall thickness. The calculated pouring time was 7.27 s, and the pouring speed was set at 11 cm/s. Considering the large wall thickness variation of the integrated crossbeam casting (8.0 mm to 55.4 mm), the pouring temperature range was initially determined to be 1380-1420 ℃. For the casting shrinkage rate, since QT800-5 is a pearlitic ductile iron with shrinkage obstruction predominance, the casting shrinkage rate was taken as 1.2%.
Due to the complex structure of the integrated crossbeam, a closed gating system was adopted to ensure complete filling. The cross-sectional area ratio of the gating system components was set as follows:
$$F_{sprue} : F_{runner} : F_{ingate} = 2.50 : 1.25 : 1 \quad (4-2)$$
Four ingates were designed to ensure uniform filling of the large-span crossbeam and satisfy the sequential solidification principle. The ingate cross-sectional area was calculated using the flow rate method:
$$A_{g} = \frac{V}{v_{g} \times t} \quad (4-3)$$
where $A_{g}$ is the ingate cross-sectional area (mm²), $V$ is the volume of molten metal passing through the ingate (mm³), $v_{g}$ is the velocity of molten metal (m/s), and $t$ is the mold filling time (s). The calculated ingate cross-sectional area was 1200 mm², the runner area was 1500 mm², and the sprue area was 3000 mm². The layout employed a two-cavity mold with two identical crossbeam castings per mold, and the box size was 1100×900×350 mm.
4.2 Initial Gating System and Defect Analysis
The initial gating system scheme was established with 4 risers. Numerical simulation was conducted with a pouring temperature of 1400 ℃, pouring speed of 10 cm/s, and mold temperature of 20 ℃. The filling process analysis showed that at 1.06 s, the molten iron filled the sprue and runner with minimal temperature loss. At 3.31 s, the molten iron was divided into four streams through the four branch ingates, uniformly spreading across the bottom of the mold cavity. At 5.04 s, the iron liquid filled to the middle of the crossbeam with stable liquid level and relatively fast overall filling speed. When filling reached 99% at 7.27 s, the casting was basically filled without obvious flow-related casting defects. However, the solidification analysis revealed significant issues. The solidification time of each region indicated that the complex thick-walled areas at both ends of the casting had the longest solidification time. The liquid phase distribution before complete solidification showed closed liquid regions formed at the thick-walled parts of the casting. These closed regions, characterized by complex geometric features and large cross-section thickness, typically became the final solidification zones. If adequate molten metal replenishment was not achieved, shrinkage cavities and shrinkage porosity casting defects would be highly prone to form. The initial scheme produced a shrinkage porosity volume of 751.16 mm³.
4.3 Gating System Optimization
To ensure adequate feeding during solidification and eliminate the solidification collapse zone, the riser configuration was optimized. The number of risers was increased from 4 to 8 (4×Φ70 mm×120 mm), and they were strategically positioned at both ends of the crossbeam. Under identical pouring temperature and speed conditions, the optimized gating system was simulated. The filling process of the optimized scheme yielded a complete fill time of 7.27 s with stable filling behavior and no cold shut or misrun casting defects. The temperature gradient uniformity was improved, with the central thin-wall area showing faster heat dissipation. The solidification simulation showed that the solidification time at both ends of the casting was significantly reduced from 436 s to 307.67 s. More importantly, the risers at both ends provided flowing molten metal during solidification, eliminating the isolated molten pools observed in the original scheme. The final solidification zones shifted to the gating system. The shrinkage porosity volume decreased to 685.85 mm³, representing an 8.69% reduction compared to the initial scheme. The shrinkage defects were primarily distributed in the risers, which fulfilled the intended function of collecting gas and inclusions while minimizing the impact of casting defects on the mechanical performance of critical areas.
4.4 Effect of Pouring Temperature on Casting Quality
Using a benchmark condition of 1400 ℃ pouring temperature, 11 cm/s pouring speed, and 20 ℃ mold temperature, additional simulations were conducted at pouring temperatures of 1380 ℃ and 1420 ℃ to systematically investigate the influence of temperature on filling behavior, solidification characteristics, and casting defect formation. The filling process at 1380 ℃ was essentially similar to that at 1400 ℃, but the flow rate was slower in local areas, particularly in complex regions. At 4.38 s, the molten iron had not yet filled the areas near the gate at both ends. The overall temperature after filling was lower than that at 1400 ℃, with no observed misrun or cold shut casting defects. At 1420 ℃, the molten iron filled faster with more vortex phenomena observed. The overall temperature was higher than that at 1400 ℃, but the temperature field distribution pattern remained basically the same.
The simulation results of gas entrainment at different pouring temperatures are shown in the following table.
| Pouring Temperature (℃) | Gas Entrainment Volume (mm³) |
|---|---|
| 1380 | 1700.58 |
| 1400 | 1682.44 |
| 1420 | 1697.82 |
The gas entrainment distribution showed a non-linear pattern with increasing pouring temperature. When the temperature increased from 1380 ℃ to 1400 ℃, the gas entrainment decreased because higher temperature improved the fluidity of molten iron, leading to more stable filling and reduced turbulence-induced gas entrapment. However, when the temperature further increased to 1420 ℃, the gas entrainment increased, attributed to the reduced surface tension of the molten iron and enhanced gas evolution at higher temperatures. The gas entrainment was more severe at the complex end regions of the casting, where multi-directional molten iron flow caused intense turbulence, and at the central region where two streams of molten iron collided.
The solidification time increased with increasing pouring temperature: 307.67 s at 1380 ℃, 326.96 s at 1400 ℃, and 346.33 s at 1420 ℃. The solidification sequence was basically identical at all temperatures. The thin-wall central shaft solidified first, followed by other thin-wall regions, with the complex thick-wall regions and gating system solidifying last. Isolated liquid zones formed in complex structural regions before complete solidification, and these zones were prone to shrinkage casting defects. The shrinkage porosity volumes at different pouring temperatures are summarized below.
| Pouring Temperature (℃) | Shrinkage Porosity Volume (mm³) |
|---|---|
| 1380 | 667.92 |
| 1400 | 685.85 |
| 1420 | 702.11 |
These results demonstrate a clear trend: with increasing pouring temperature, the shrinkage porosity volume increased consistently. Higher pouring temperatures improved fluidity but simultaneously aggravated solidification shrinkage effects and increased temperature gradients, ultimately leading to a higher tendency for shrinkage casting defects.
4.5 Effect of Pouring Speed on Casting Quality
The influence of pouring speed was investigated by comparing simulations at 10 cm/s, 11 cm/s, and 12 cm/s. At a lower pouring speed of 10 cm/s, the flow stability was effectively improved during the filling process. The overall temperature field after filling was similar to that at 11 cm/s, but the temperature at thin-wall regions was lower. At 12 cm/s, the molten iron flow velocity was significantly increased, and the filling time was reduced to 6.36 s. Higher flow velocity caused splashing and local vortex phenomena due to the angles between ingates and cavity walls, particularly pronounced at the higher pouring speed. However, the temperature field after filling exhibited better uniformity with a minimum temperature of 1286 ℃.
The gas entrainment distribution patterns were consistent across different pouring speeds, concentrated at the complex end regions, the central impact zone, and the terminal solidification regions. The quantitative gas entrainment results are shown below.
| Pouring Speed (cm/s) | Gas Entrainment Volume (mm³) |
|---|---|
| 10 | 1655.77 |
| 11 | 1682.44 |
| 12 | 1700.39 |
The gas entrainment showed a monotonically increasing trend with increasing pouring speed. At higher pouring speeds, the kinetic energy of the molten iron increased, generating high shear forces at the cavity contact surfaces, leading to boundary layer separation and vortex formation that entrapped more gas into the melt. Additionally, the shorter rising time for bubbles at faster filling rates caused some bubbles to be captured between dendrites or in closed liquid regions, forming dispersed surface defects.
The solidification time varied with pouring speed: 305.69 s at 10 cm/s, 307.67 s at 11 cm/s, and 348.38 s at 12 cm/s. The differences in solidification time primarily originated from the filling stage, where faster pouring speeds resulted in more uniform temperature distribution and slower heat loss. The shrinkage porosity volumes at different pouring speeds are presented below.
| Pouring Speed (cm/s) | Shrinkage Porosity Volume (mm³) |
|---|---|
| 10 | 696.92 |
| 11 | 685.85 |
| 12 | 680.85 |
Data analysis revealed that when the pouring speed increased from 10 cm/s to 12 cm/s, the shrinkage porosity volume decreased. Faster pouring speeds enabled rapid filling with reduced heat loss, yielding a more uniform filling temperature field and consequently a lower tendency for shrinkage defects.
5. Orthogonal Experiment and Process Parameter Optimization
5.1 Orthogonal Experiment Design
To comprehensively analyze the influence of various process parameters on casting quality, an L9(3²) orthogonal experiment was designed with two factors and three levels. The factors were pouring temperature (1380 ℃, 1400 ℃, and 1420 ℃) and pouring speed (10 cm/s, 11 cm/s, and 12 cm/s), with gas entrainment volume and shrinkage porosity volume as the evaluation indicators. The orthogonal experiment table is shown below.
| Experiment No. | Pouring Speed A (cm/s) | Pouring Temperature B (℃) | Gas Entrainment Volume (mm³) | Shrinkage Porosity Volume (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 |
5.2 Analysis of Variance (ANOVA)
The analysis of variance method was applied to decompose the main effects of multiple factors and multiple levels from the orthogonal experiment results. By comparing the F values of each factor, the process parameters with significant effects on gas entrainment volume and shrinkage porosity volume could be identified, while statistically insignificant interference factors could be eliminated. The ANOVA results for gas entrainment volume are shown below.
| Source | Sum of Squares | df | Mean Square | F | p |
|---|---|---|---|---|---|
| Pouring Speed | 2643.000 | 2 | 1321.500 | 24.266 | 0.006 |
| Pouring Temperature | 567.608 | 2 | 283.804 | 5.211 | 0.077 |
| Error | 108.919 | 2 | 54.459 | — | — |
The F value for pouring speed was 24.266 with p = 0.006 < 0.05, indicating that pouring speed had a statistically significant effect on gas entrainment volume. In contrast, pouring temperature had no significant effect on gas entrainment (p = 0.077 > 0.05). The range analysis showed that the range of pouring speed was 124.42, while the range of pouring temperature was only 60.68, further confirming that pouring speed was the dominant factor affecting gas entrainment. The ANOVA results for shrinkage porosity volume are presented below.
| Source | Sum of Squares | df | Mean Square | F | p |
|---|---|---|---|---|---|
| Pouring Speed | 406.121 | 2 | 203.061 | 2.847 | 0.167 |
| Pouring Temperature | 1747.142 | 2 | 873.571 | 12.247 | 0.017 |
| Error | 142.659 | 2 | 71.330 | — | — |
For shrinkage porosity volume, the F value for pouring temperature was 12.247 with p = 0.017 < 0.05, indicating a statistically significant effect. Pouring speed did not show a significant effect on shrinkage porosity (p = 0.167 > 0.05). The range analysis confirmed this finding, with the range of pouring temperature (102.45) being significantly larger than that of pouring speed (48.29). The factor-level effects on the responses are summarized below.
| Factor | Level | Gas Entrainment K (mm³) | Shrinkage Porosity K (mm³) |
|---|---|---|---|
| Pouring Speed A | A1 (10 cm/s) | 5001.80 | 2089.34 |
| A2 (11 cm/s) | 5080.84 | 2055.88 | |
| A3 (12 cm/s) | 5126.22 | 2041.05 | |
| Pouring Temperature B | B1 (1380 ℃) | 5099.28 | 2010.10 |
| B2 (1400 ℃) | 5038.60 | 2063.62 | |
| B3 (1420 ℃) | 5070.98 | 2112.55 |
Based on the ANOVA results and factor-level analysis, the optimal combination of process parameters was determined by balancing gas entrainment and shrinkage porosity. Pouring speed significantly influenced gas entrainment, with lower speeds favoring lower gas entrainment. Pouring temperature significantly influenced shrinkage porosity, with lower temperatures favoring reduced shrinkage defects. Considering both indicators comprehensively, the optimal process parameters were established as: pouring speed 11 cm/s, pouring temperature 1380 ℃, and mold temperature 20 ℃. This combination achieved a good balance between minimizing gas entrainment and minimizing shrinkage casting defects.
The optimized parameter combination corresponds to experiment No. 6 in the orthogonal array, which yielded a shrinkage porosity volume of 667.92 mm³. Compared to the initial process scheme, this optimized result represents an 11.74% reduction in casting defects. The verification simulation using these optimal parameters showed that the mold filling process was smooth and continuous, with uniform filling distribution across all regions of the casting. The solidification process strictly followed the sequential solidification principle: the thin-wall central region solidified first at 40.72 s, followed by the two ends and areas with large contact surfaces with the sand mold, and finally the complex thick-wall regions and gating system. The solidification end time was 307.67 s, and the shrinkage porosity volume was effectively controlled.
6. Production Trial and Experimental Verification
6.1 Casting Production Conditions
Based on the numerical simulation optimization results, the production trial of the integrated ductile iron crossbeam was carried out using the optimized process scheme. The optimized gating system featured a composite design with four ingates (total cross-sectional area of 1200 mm²), a single runner, and strict control of the cross-sectional area ratio (Fsprue:Frunner:Fingate = 2.50:1.25:1). The pouring temperature was maintained at 1380 ℃ with a pouring speed of 10 cm/s and mold temperature of 20 ℃. After complete cooling of the casting, the subsequent processing included shakeout, cleaning, cutting off the gating system, grinding burrs, and surface treatment through shot blasting.
6.2 Chemical Composition Analysis
The chemical composition of the integrated crossbeam was analyzed three times, and the average values are presented in the table below.
| Element | C | Si | Mn | P | S | Mg | Ni | Mo | Cu |
|---|---|---|---|---|---|---|---|---|---|
| Content (%) | 3.61 | 2.67 | 0.105 | 0.007 | 0.007 | 0.043 | 0.438 | 0.151 | 0.803 |
The chemical composition of the trial-produced integrated crossbeam casting was stable, with carbon equivalent and residual magnesium content meeting the requirements. The P and S contents were sufficiently low, effectively avoiding the formation of phosphide eutectics and sulfides, which can be harmful to the mechanical properties and promote casting defects.
6.3 Microstructure Analysis
Since the integrated crossbeam has a large horizontal span, the solidification conditions differ significantly at various positions relative to the ingate. Different regions of the casting exhibit different cooling rate gradients, leading to potential variations in microstructure and mechanical properties. Therefore, samples were taken from three representative positions: position 1 (near the ingate), position 2 (thin-wall central shaft), and position 3 (far from the ingate).
The graphite morphology at position 1, sampled from the near-riser region with slow cooling, exhibited the largest average graphite ball size and the lowest graphite ball count of 180 balls/mm². Position 2, located in the thin-wall central region with the fastest cooling rate, showed the smallest average graphite ball size with a count of 252 balls/mm². Position 3, far from the riser, displayed uniform graphite sizes with a ball count of 212 balls/mm², intermediate between positions 1 and 2. The metallographic evaluation results are summarized below.
| Sample No. | Spheroidization Grade | Graphite Size Grade | Pearlite Content (%) | Carbide (%) | Phosphide Eutectic |
|---|---|---|---|---|---|
| 1 | 2 | 6 | ≥90 | 1 | — |
| 2 | 2 | 7 | ≥95 | 1 | — |
| 3 | 2 | 6 | ≥90 | 1 | — |
All three samples achieved grade 2 spheroidization (spheroidization rate ≥ 90%), with graphite ball diameters conforming to grade 6-7 characteristics. The matrix structure consisted of black spheroidal graphite, white ferrite, and dark gray pearlite, with pearlite content exceeding 90%. X-ray flaw detection of both the central thin-wall region and the end regions revealed no gas holes, cracks, or shrinkage casting defects, confirming the integrity of critical load-bearing areas of the crossbeam.
6.4 Mechanical Properties
Tensile tests were performed on specimens taken from the three positions of the casting. The results are presented in the table below.
| Sample No. | Yield Strength Rp0.2 (MPa) | Tensile Strength Rm (MPa) | Elongation A (%) |
|---|---|---|---|
| 1 | 546 | 875 | 7.6 |
| 2 | 562 | 891 | 6.0 |
| 3 | 552 | 854 | 6.5 |
Position 1, located in the thick-walled near-riser region with the slowest cooling rate, exhibited the best toughness with a yield strength of 546 MPa, tensile strength of 875 MPa, and elongation of 7.6%. Position 2, the thin-wall central shaft with the fastest cooling rate, demonstrated the highest strength with a yield strength of 562 MPa and tensile strength of 891 MPa, but the lowest elongation of 6.0%. Position 3, far from the riser with complex geometry and thick walls, was the last to solidify and had the highest tendency for gas entrainment and shrinkage casting defects, showing the lowest strength with a tensile strength of 854 MPa.
The hardness values at different positions of the crossbeam body were measured and are summarized below.
| Sample No. | Hardness Measurement 1 (HV) | Hardness Measurement 2 (HV) | Hardness Measurement 3 (HV) | Average Hardness (HV) |
|---|---|---|---|---|
| 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 |
The hardness values at all positions of the crossbeam were similar, exceeding 270 HV, which meets the application requirements. The relatively small hardness variations indicate adequate feeding and minimal shrinkage casting defects under the current process scheme. The tensile strength values exceeded 800 MPa and the elongation exceeded 5% at all test positions, meeting the requirements of the QT800-5 grade and confirming that the optimized casting process can achieve the required mechanical properties for the integrated crossbeam.
6.5 Fracture Morphology and EDS Analysis
The fracture morphology of the tensile specimens was examined using scanning electron microscopy. The fracture surface exhibited numerous cleavage facets and continuous river patterns around the graphite in the matrix structure. The cleavage steps originated from crack propagation along specific crystallographic planes, while the river patterns were generated from the direction adjustment of cleavage cracks when crossing grain boundaries with different orientations. These features indicated the relatively low toughness of the ductile iron material. Comparison of the cleavage facet sizes and river pattern extents across the three specimen positions revealed that the farther from the ingate, the larger the cleavage facets became, indicating inferior plasticity at positions farther from the ingate.
Energy dispersive spectroscopy analysis was performed on the fracture surfaces to determine the chemical composition of inclusions and graphite nodules. At a location featuring a large diamond-shaped phase, the EDS elemental mapping revealed that this phase was mainly composed of Ca and O, with contents of 7.8% and 5.7%, respectively, accompanied by Fe segregation around it. This was identified as an oxidation inclusion defect, primarily CaO. The sharp morphology of the CaO phase cuts through the matrix, creating stress concentration during tensile loading, initiating cracks that ultimately lead to fracture. The visible voids in the micrograph originated from this mechanism. The analysis of graphite nodules showed that the selected region consisted of a high carbon content with low Fe and Si levels distributed mainly around the periphery. Graphite nodules, as brittle phases, cause some harm to the matrix, with many cracks propagating along the graphite-matrix interface. However, compared to sharp brittle inclusions such as CaO, graphite balls have a much smaller detrimental effect on the mechanical properties of the casting.
7. Conclusion and Outlook
This research systematically investigated the integrated casting process for a commercial vehicle crossbeam made of ductile iron, integrating structural design, numerical simulation, process optimization, and experimental verification. The main conclusions are summarized as follows:
(1) The original crossbeam structure was redesigned into an integrated structure, reducing the mass from 50.5 kg to 36 kg, a decrease of 28.7%. Static analysis under four operating conditions showed that the maximum equivalent stress of the integrated crossbeam was 558.4 MPa with a safety factor of 1.4, satisfying the strength requirements of commercial vehicles.
(2) The initial gating system was optimized by increasing the number of risers from 4 to 8 (4×Φ70 mm×120 mm) and positioning them at both ends of the crossbeam. This optimized scheme reduced the shrinkage casting defects from 751.16 mm³ to 685.85 mm³, an 8.69% reduction. The distribution of casting defects was also improved, with the majority of shrinkage collecting in the risers rather than in critical load-bearing regions of the casting.
(3) The orthogonal experiment analysis identified the optimal process parameters: pouring speed of 11 cm/s, pouring temperature of 1380 ℃, and mold temperature of 20 ℃. ANOVA revealed that pouring speed significantly affected gas entrainment volume, while pouring temperature significantly affected shrinkage porosity volume. The optimized process parameters achieved a shrinkage porosity volume of 667.92 mm³, an 11.74% reduction compared to the initial scheme.
(4) The production trial using the optimized process confirmed the reliability of the designed casting process. The microstructure examination showed uniform dispersion of spheroidal graphite with grade 2 spheroidization and pearlite content exceeding 90%. The mechanical property tests confirmed that the integrated crossbeam achieved tensile strength ≥ 854 MPa, yield strength ≥ 546 MPa, elongation ≥ 6.0%, and hardness exceeding 270 HV, meeting all requirements of the QT800-5 grade.
(5) Fracture morphology analysis revealed that the material exhibits typical brittle fracture characteristics with cleavage facets and river patterns. EDS analysis identified CaO oxidation inclusions as harmful casting defects, whose sharp morphology causes more significant damage to the mechanical properties than graphite nodules.
Under the current process conditions, the solidification process of large complex ductile iron castings cannot achieve full synchronization. The optimization of such complex ductile iron castings should focus on establishing directional feeding channels and achieving sequential solidification control through a well-designed gating system, thereby ensuring the density of the casting and avoiding shrinkage casting defects. Future research directions include fatigue modeling combined with vehicle testing to study the service performance of castings, and further enhancing the performance of ductile iron through heat treatment processes such as austempering to expand its application potential in automotive structural components.