In the field of heavy-duty vehicle manufacturing, the tractor saddle is a critical load-bearing component that connects the tractor and the semi-trailer. During operation, it must withstand severe impact loads from multiple directions, which demands high mechanical strength and sound internal quality. Traditionally, such saddles are produced from forged steel, which offers excellent toughness but comes with high material and energy costs. To reduce cost while maintaining adequate performance, the industry has attempted to replace steel with ductile iron through casting. However, the transition from forged steel to cast iron introduces a series of typical sand casting defects, such as gas holes, cold shut, shrinkage porosity, shrinkage cavity, sand inclusion, and slag inclusion. These defects significantly increase the scrap rate and lead to substantial economic losses. In this study, I focused on optimizing the sand mold casting process for a tractor saddle to minimize these sand casting defects. I designed the gating system, filtering system, and exhaust system based on theoretical analysis, performed numerical simulations using AnyCasting, and verified the optimized process through physical testing. The results demonstrate that the optimized process dramatically reduces sand casting defects and improves the yield from about 60% to more than 95%.
1. Structural Analysis and Process Design
The tractor saddle under investigation has a symmetrical, large thin-wall structure with outer dimensions of 755 mm × 580 mm × 322 mm. The gross weight of the casting is 63 kg, the average wall thickness is 10 mm, and the maximum wall thickness is 30 mm. The geometry was modeled in three dimensions, and the structural design was evaluated for castability. Sharp corners were replaced with rounded fillets, and reinforcing ribs were arranged to avoid excessive hot spots. The material was selected as ductile iron grade QT700-6, which provides a tensile strength of at least 700 MPa, elongation of at least 6%, and a hardness range of 240–290 HBW. The chemical composition of QT700-6 is listed in Table 1.
| C | Si | Mn | S | P | Mg | Ce | Cr | Cu | Mo |
|---|---|---|---|---|---|---|---|---|---|
| 3.5–3.9 | 1.8–2.1 | 0.35–0.5 | 0.006–0.02 | ≤0.05 | 0.035–0.055 | 0.006–0.03 | ≤0.1 | 0.4–0.6 | ≤0.15 |
Since the saddle is a large thin-wall component, I adopted a two-part flask molding with horizontal parting. The most important assembly surface was placed in the lower mold to avoid the upward floatation of inclusions. The linear shrinkage rate of the casting was calculated using the following relation:
$$K = \frac{L_{\text{mold}} – L_{\text{casting}}}{L_{\text{casting}}} \times 100\%$$
For ductile iron with a pearlitic matrix, the free shrinkage is 0.9%–1.1% and the hindered shrinkage is 0.6%–0.8%. Considering the complex core geometry, a shrinkage rate of 1.1% was used in the pattern design. The tolerances followed the standard DCTG 11–13 for sand casting, and the machining allowance was selected as RMAG F–H. A draft angle of 3° was applied to facilitate pattern withdrawal.
One sand core was used for the entire inner cavity. The core was made with cold-box resin sand and fixed with both vertical and horizontal core prints. The vertical upper core print dimensions were 108 mm × 75 mm × 42 mm, while the lower core print was 108 mm × 98 mm × 42 mm. The horizontal core prints were arranged symmetrically on the left and right sides, with rear prints of 64 mm and front prints of 35 mm, both with a taper of 3°. The core design ensured precise positioning and firm locking during mold assembly.
2. Theoretical Basis for Filling and Solidification
To predict and eliminate sand casting defects, it is essential to understand the fluid flow and heat transfer phenomena during mold filling and solidification. The governing equations include the continuity equation, the momentum equations, and the energy equation. For an incompressible Newtonian fluid, the continuity equation is:
$$ \frac{\partial \rho}{\partial t} + \frac{\partial (\rho u)}{\partial x} + \frac{\partial (\rho v)}{\partial y} + \frac{\partial (\rho w)}{\partial z} = 0 $$
The momentum conservation in the x, y, and z directions can be expressed as:
$$ \rho \left( \frac{\partial u}{\partial t} + u \frac{\partial u}{\partial x} + v \frac{\partial u}{\partial y} + w \frac{\partial u}{\partial z} \right) = -\frac{\partial p}{\partial x} + \mu \nabla^2 u + \rho g_x $$
$$ \rho \left( \frac{\partial v}{\partial t} + u \frac{\partial v}{\partial x} + v \frac{\partial v}{\partial y} + w \frac{\partial v}{\partial z} \right) = -\frac{\partial p}{\partial y} + \mu \nabla^2 v + \rho g_y $$
$$ \rho \left( \frac{\partial w}{\partial t} + u \frac{\partial w}{\partial x} + v \frac{\partial w}{\partial y} + w \frac{\partial w}{\partial z} \right) = -\frac{\partial p}{\partial z} + \mu \nabla^2 w + \rho g_z $$
The energy conservation equation, including the latent heat release during solidification, is:
$$ \rho C_p \left( \frac{\partial T}{\partial t} + u \frac{\partial T}{\partial x} + v \frac{\partial T}{\partial y} + w \frac{\partial T}{\partial z} \right) = \frac{\partial}{\partial x} \left( k \frac{\partial T}{\partial x} \right) + \frac{\partial}{\partial y} \left( k \frac{\partial T}{\partial y} \right) + \frac{\partial}{\partial z} \left( k \frac{\partial T}{\partial z} \right) + S_T $$
where \( S_T \) represents the latent heat source term. In numerical simulation, the temperature field and the solid fraction are coupled. The relationship between temperature and solid fraction during solidification can be approximated as:
$$ f_s = \frac{T_L – T}{T_L – T_S} $$
where \( T_L \) is the liquidus temperature and \( T_S \) is the solidus temperature. The temperature-gradient method is often used to predict shrinkage defects. The temperature gradient \( G \) is given by:
$$ G = \sqrt{ \left( \frac{\partial T}{\partial x} \right)^2 + \left( \frac{\partial T}{\partial y} \right)^2 + \left( \frac{\partial T}{\partial z} \right)^2 } $$
When the parameter \( G / \sqrt{R} \) (where \( R \) is the cooling rate) falls below a critical value, shrinkage porosity is likely to occur. In my simulation, I used the built-in defect criterion in AnyCasting to visualize the probability of sand casting defects such as shrinkage porosity and shrinkage cavities.
3. Design of Gating, Filtering, and Exhaust Systems
The gating system directly influences the filling pattern and the occurrence of sand casting defects. I selected a closed gating system with the area ratio:
$$ \sum A_{\text{inner}} : \sum A_{\text{runner}} : \sum A_{\text{sprue}} = 1 : 1.65 : 1.9 $$
To ensure sufficient filling without turbulence, the filling time was calculated using:
$$ t = S \sqrt{G} $$
where \( G \) is the total pouring weight (157 kg) and \( S \) is a coefficient depending on wall thickness. For a wall thickness of 8–15 mm, \( S = 1.85 \). Thus, \( t = 14 \text{ s} \).
For a bottom-gating system, the average pressure head is:
$$ H_P = H_0 – \frac{h}{2} $$
where \( H_0 \) is the height between the inner gate and the pouring basin (30 cm) and \( h \) is the saddle height (32 cm). Therefore \( H_P = 14 \text{ cm} \). The minimum residual pressure head was checked using:
$$ H_M \ge L \tan \alpha $$
with \( L = 195 \text{ mm} \) and \( \alpha = 8^\circ \), giving \( H_M \ge 27 \text{ mm} \). The total inner gate area was determined by the well-known formula:
$$ \sum A_{\text{inner}} = \frac{G}{0.3 \mu t \sqrt{H_P}} $$
Substituting \( G = 157 \text{ kg} \), \( \mu = 0.64 \), \( t = 14 \text{ s} \), and \( H_P = 14 \text{ cm} \), the calculation yields \( \sum A_{\text{inner}} = 15.6 \text{ cm}^2 \).
To investigate the effect of the number of inner gates on the filling behavior and final quality, I designed three variants with 4, 6, and 8 inner gates. The dimensions of each gating system are summarized in Table 2.
| Variant | Number of inner gates | Single inner gate area (cm²) | Total inner gate area (cm²) | Runner area (cm²) | Sprue area (cm²) |
|---|---|---|---|---|---|
| A | 4 | 3.9 | 15.6 | 25.74 | 29.64 |
| B | 6 | 2.6 | 15.6 | 25.74 | 29.64 |
| C | 8 | 1.95 | 15.6 | 25.74 | 29.64 |
Each inner gate was designed as a square cross-section. The side lengths of the inner gates were approximately 2.0 cm for the 4-gate variant, 1.6 cm for the 6-gate variant, and 1.4 cm for the 8-gate variant. The runner was a single square channel with a side length of 5 cm, and the sprue was circular with a diameter of 6 cm.
For the filtering system, I used a combination of a choke at the sprue exit, a ceramic foam filter (90×90×22 mm, 10 ppi) placed in the runner, and a slag pocket at the end of the runner. This multi-stage filtering approach effectively removes inclusions and prevents the formation of slag-related sand casting defects. The exhaust system was designed separately for the mold cavity and the sand core. The total exhaust area was calculated as:
$$ S = (1.5 \sim 4) \times \frac{G}{t \cdot \rho \cdot \mu \cdot h_p} $$
With the values used in this study, the required cavity exhaust area was 11 cm². I placed one large vent (8 cm²) at the end of the runner and two small vents (1.5 cm² each) at the highest points of the cavity. For the sand core, multiple vent channels were formed along the core prints to ensure that gases generated from the core binder could escape smoothly.
4. Numerical Simulation of the Initial Designs
I performed numerical simulations using AnyCasting software. The geometry was imported as an STL file, and the mesh was generated with different cell sizes for the core, runner, and cavity to balance accuracy and computational cost. The core mesh volume was 2.2 mm³, the runner mesh volume was 4.5 mm³, and the cavity mesh volume was 9.1 mm³. The initial conditions were set as follows: pouring temperature 1395°C, gravity 9.8 m/s² in the +x direction, and filling time 14 s. The thermal properties of FCD 700 were entered from the material database.
Simulation results for the 4-gate variant showed that the liquid metal entered the cavity with a high velocity and significant fluctuations. The maximum velocity reached 179 mm/s while the minimum was around 0.7 mm/s, indicating turbulent flow and potential air entrapment. These conditions are unfavorable because they can lead to gas porosity, one of the most common sand casting defects. The 8-gate variant produced a smoother filling in the early and middle stages, but at 95% fill some regions still exhibited somewhat unstable flow. The 6-gate variant gave the most stable velocity field throughout filling, with a small and uniform velocity difference and no obvious turbulence. This behavior is essential to prevent gas and sand inclusion defects.
The solidification simulation was used to examine the temperature field and the formation of isolated liquid pools. For the 4-gate system, isolated liquid regions appeared at a solidification fraction of about 66%, and at 85% solidification there were still several dispersed pools. The 8-gate system showed elongated isolated regions near the cavity edges, which are prone to shrinkage porosity. In contrast, the 6-gate system exhibited only small, localized isolated pools that disappeared gradually, resulting in a more uniform directional solidification. The defect probability parameter was then evaluated using the built-in criteria of AnyCasting. The 4-gate variant displayed many defect-prone spots on the upper surface and in the main cavity, while the 8-gate variant showed numerous clusters of potential shrinkage defects on both sides of the casting. The 6-gate variant had only a few scattered spots, one of which was located near the locking hole, but overall it presented the lowest probability of shrinkage defects. Therefore, I selected the 6-gate gating system as the optimal initial design.
| Variant | Filling stability | Isolated liquid regions | Defect probability | Feasibility |
|---|---|---|---|---|
| 4 gates | Large velocity fluctuation, turbulence | Multiple dispersed pools | High | Poor |
| 6 gates | Smooth, small velocity gradients | Small and localized | Low | Good |
| 8 gates | Good in early stages, unstable later | Long slender pools | Moderate to high | Moderate |
5. Orthogonal Experiment and Process Optimization
Although the 6-gate gating system significantly reduced sand casting defects, some minor defects remained. To further minimize the porosity and improve productivity, I employed an orthogonal experimental design. Four factors were studied: pouring temperature, filling pressure, pouring velocity, and liquid metal weight. Each factor was set at three levels, as shown in Table 4.
| Level | A: Pouring temperature (°C) | B: Filling pressure (Pa) | C: Pouring velocity (m/s) | D: Metal weight (kg) |
|---|---|---|---|---|
| 1 | 1395 | 1.0 | 0.34 | 157 |
| 2 | 1385 | 0.9 | 0.30 | 155 |
| 3 | 1405 | 1.1 | 0.40 | 159 |
Using the standard \( L_9(3^4) \) orthogonal table, I conducted nine simulation runs. The responses were the total porosity volume and the solidification time, as listed in Table 5.
| Trial | A | B | C | D | Porosity volume (cm³) | Solidification time (h) |
|---|---|---|---|---|---|---|
| 1 | 1 | 1 | 1 | 1 | 0.986 | 0.5575 |
| 2 | 1 | 2 | 2 | 2 | 0.654 | 0.5581 |
| 3 | 1 | 3 | 3 | 3 | 0.473 | 0.5542 |
| 4 | 2 | 1 | 2 | 3 | 0.813 | 0.5458 |
| 5 | 2 | 2 | 3 | 1 | 0.634 | 0.5411 |
| 6 | 2 | 3 | 1 | 2 | 0.746 | 0.5417 |
| 7 | 3 | 1 | 3 | 2 | 0.682 | 0.5708 |
| 8 | 3 | 2 | 1 | 3 | 0.914 | 0.5728 |
| 9 | 3 | 3 | 2 | 1 | 0.753 | 0.5739 |
The range analysis was carried out separately for each response. The calculated mean values \( K_{ij} \) and ranges \( R \) are summarized in Table 6.
| Response | Statistic | A | B | C | D |
|---|---|---|---|---|---|
| Porosity volume (cm³) | K1 | 0.704 | 0.827 | 0.882 | 0.791 |
| K2 | 0.731 | 0.734 | 0.740 | 0.694 | |
| K3 | 0.783 | 0.657 | 0.596 | 0.733 | |
| R | 0.079 | 0.170 | 0.286 | 0.097 | |
| Solidification time (h) | K1 | 0.5566 | 0.5580 | 0.5573 | 0.5575 |
| K2 | 0.5429 | 0.5573 | 0.5593 | 0.5569 | |
| K3 | 0.5725 | 0.5566 | 0.5554 | 0.5576 | |
| R | 0.0296 | 0.0014 | 0.0039 | 0.0007 |
From the range analysis, for porosity volume, the order of influence was \( C > B > D > A \), indicating that pouring velocity had the greatest effect, followed by filling pressure, metal weight, and pouring temperature. For solidification time, the order was \( A > C > B > D \). Since porosity must be minimized first, and solidification time is a secondary criterion, I obtained a preliminary optimum of \( A_1B_3C_3D_2 \) from the porosity results and \( A_2B_3C_3D_2 \) from the solidification time results. To resolve the conflict, a matrix analysis was performed to combine the two indicators. The final comprehensive optimum was determined as:
$$ A_2 \, B_3 \, C_3 \, D_2 $$
i.e., pouring temperature 1385°C, filling pressure 1.1 Pa, pouring velocity 0.40 m/s, and liquid metal weight 155 kg.
The optimized parameters were then used in a confirmation simulation. The velocity field remained smooth throughout the entire filling process, with no visible splashing or flow separation. The solidification sequence was directional: thin sections solidified first, followed by the thicker main body, and only a small isolated liquid region appeared near the final solidification zone. The predicted porosity volume decreased to 0.401 cm³, which was lower than any value in the original orthogonal matrix. The defect probability map showed no shrinkage cavities or significant shrinkage porosity in the critical regions; the small line defect near the locking hole disappeared completely. This simulation confirmed that the optimized process effectively suppresses sand casting defects.
6. Experimental Verification
After the simulation-based optimization, I produced a batch of tractor saddles using the optimized casting process. The external surface of the castings was smooth and free of visible defects such as cracks, cold shuts, or sand inclusion. Ultrasonic inspection was performed on 14 randomly sampled products from two batches. No unacceptable internal discontinuities were detected, indicating that the optimized gating and pouring scheme successfully reduced sand casting defects to a negligible level.
Metallographic examination was conducted after etching with nital solution. The microstructure was evaluated for graphite type, graphite size, pearlite content, and nodularity. The test results are summarized in Table 7.
| Sample | Elongation (%) | Yield strength (MPa) | Tensile strength (MPa) | Hardness (HBW) | Graphite size (grade) | Pearlite content (%) | Nodularity |
|---|---|---|---|---|---|---|---|
| 01 | 10.0 | 424 | 739 | 245 | 6 | 80 | VI |
| 02 | 10.3 | 412 | 725 | 247 | 6 | 81 | VI |
| 03 | 10.7 | 426 | 747 | 248 | 6 | 80 | VI |
| 04 | 10.0 | 411 | 749 | 240 | 6 | 75 | VI |
| 05 | 10.2 | 436 | 728 | 246 | 6 | 80 | VI |
| 06 | 10.1 | 432 | 735 | 249 | 6 | 81 | VI |
| 07 | 10.3 | 416 | 736 | 247 | 6 | 81 | VI |
| 08 | 10.6 | 414 | 745 | 245 | 6 | 80 | VI |
| 09 | 10.5 | 433 | 742 | 247 | 6 | 82 | VI |
| 10 | 10.2 | 421 | 728 | 249 | 6 | 82 | VI |
| 11 | 10.0 | 424 | 726 | 246 | 6 | 81 | VI |
| 12 | 10.1 | 410 | 731 | 248 | 6 | 80 | VI |
| 13 | 10.1 | 433 | 744 | 246 | 6 | 83 | VI |
| 14 | 10.5 | 424 | 742 | 247 | 6 | 82 | VI |
All samples exceeded the specified elongation of 6% and tensile strength of 700 MPa. The hardness values were within the 240–290 HBW range. The graphite was uniformly distributed spherical graphite with a size grade of 6, and the pearlite content was 80% or above except for one sample that had 75% but still met the requirement. No flake graphite was observed. The nodularity was consistently grade VI. In production, the overall qualification rate rose from approximately 60% under the original process to above 95% with the optimized process, confirming that the combination of a closed 6-gate gating system, a properly designed exhaust and filtering system, and the carefully selected process parameters effectively eliminated the majority of sand casting defects.
7. Conclusion
In this work, I developed an optimized sand mold casting process for a tractor saddle that was originally a forged steel component. By replacing steel with ductile iron, a significant cost reduction was achieved without compromising mechanical integrity. The main conclusions are as follows:
- The structural design and process parameters were determined, including the material QT700-6, a horizontal two-part mold, a linear shrinkage rate of 1.1%, and a draft angle of 3°.
- A closed gating system with a ratio of 1:1.65:1.9 for the inner gate, runner, and sprue areas was designed. Among the three variants, the 6-gate system produced the most stable filling, the smallest isolated liquid regions, and the lowest probability of shrinkage defects.
- The filtering and exhaust systems were improved by combining a choke, a ceramic foam filter, and a slag pocket, and by adding properly sized vents for both the cavity and the core.
- Orthogonal experimental optimization revealed that pouring velocity is the most critical factor affecting porosity, and the optimal combination was pouring temperature 1385°C, filling pressure 1.1 Pa, pouring velocity 0.40 m/s, and liquid metal weight 155 kg.
- Experimental verification on production castings confirmed that the optimized process reduces sand casting defects such as shrinkage porosity, shrinkage cavity, gas holes, and sand inclusion. The qualification rate increased from 60% to over 95%, and all mechanical and metallographic requirements were satisfied.
This study demonstrates that numerical simulation combined with orthogonal experiments is an effective approach for eliminating sand casting defects in large thin-wall ductile iron castings. The methodology can be readily extended to other similar structural castings.