Thin-walled shell components are widely used in high-power engines for various applications such as military equipment and marine propulsion. However, their casting processes often suffer from poor stability, severe shrinkage porosity and hot tearing defects, leading to low yield and difficulty in meeting production requirements. To address these challenges, I investigated the investment casting process of a complex thin-walled shell casting made of 1Cr20Ni14Si2 heat-resistant stainless steel. The study combined theoretical calculations with numerical simulations to analyze and predict the formation of shrinkage and hot tearing defects. The ultimate goal was to optimize the casting process to reduce costs and improve product quality. This work holds both academic significance and practical value for the manufacturing of similar thin-walled components.
Casting defects such as shrinkage porosity and hot tearing are among the most critical issues in thin-walled castings. Shrinkage porosity arises when liquid metal cannot adequately feed the solidifying regions, especially in areas with complex geometry or varying wall thickness. Hot tearing occurs during the final stage of solidification when thermal stresses exceed the material’s tearing resistance. Both defects are strongly influenced by the gating system design and casting process parameters. Numerical simulation using ProCAST 2018 provides an efficient method to visualize filling, solidification, temperature fields, and stress evolution, enabling process optimization without costly trial-and-error experiments. In this study, I aimed to optimize the gating system and casting parameters to eliminate casting defects in a connecting shell component.
1. Component and Material Analysis
The studied component is a connecting shell used in a high-power engine. Its overall dimensions are 216 mm × 180 mm × 240 mm, featuring an eccentric tee structure connected by thin-walled arcs. The top flange has a maximum diameter of 137 mm, while more than 70% of the surface area has a wall thickness of only 3.5 mm, with gradual transitions between 3.5 mm and 5 mm. The complex geometry, especially the thin walls and varying thickness sections, makes the casting particularly susceptible to casting defects. The flange and the central inclined groove require only minor machining, while the rest must be cast to near-net shape with no defects allowed under non-destructive inspection.
The original design specified GX25CrNiSi20-14 (06Cr25Ni20) stainless steel. However, after comparing mechanical properties and high-temperature performance, I selected 1Cr20Ni14Si2 stainless steel as a suitable alternative. The chemical composition of 1Cr20Ni14Si2 is given in Table 1. Its yield strength is 295 MPa and tensile strength is 590 MPa, both higher than those of 06Cr25Ni20, while elongation and reduction of area remain at 35% and 50%, respectively. The alloy has a liquidus temperature of 1440 °C and a solidus temperature of 1268 °C, resulting in a wide freezing range of 172 °C, which tends to promote the formation of casting defects. Thus, careful process design is essential.
| Element | C | Cr | Ni | Si | Mn | P | S | Fe |
|---|---|---|---|---|---|---|---|---|
| Specification | ≤0.20 | 19–22 | 12–15 | 1.5–2.5 | ≤1.5 | ≤0.035 | ≤0.03 | Bal. |
| Actual | 0.100 | 20.235 | 14.028 | 1.612 | 0.55 | 0.005 | 0.008 | Bal. |
2. Gating System Design
For the investment casting of this thin-walled shell, I adopted a top-gated system with a pouring cup, a sprue, a runner, and two ingates. The dimensions were calculated using the cup capacity feeding method. The crucial thermal node at the wall transition was simplified as a rectangular bar with an equivalent diameter \(D_c \approx 14\) mm. According to the design charts, the weight coefficient \(k\) was determined as 0.92, giving the ingate equivalent diameter \(D_g = k D_c = 13\) mm. The runner was sized to be at least 1.15–1.25 times the ingate diameter, leading to a runner cross-section side length of 50 mm. The pouring cup top diameter was 100 mm, bottom diameter 70 mm, and height 80 mm. Two gating system schemes were proposed: Scheme 1 without padding (feeder pad) at the ingate, and Scheme 2 with padding added to improve feeding and smooth filling. The detailed dimensions are presented in Table 2.
| Scheme | D (cup top) | D1 (cup bottom) | H (cup height) | A (runner side) | L (runner length) | Dg (ingate) | Padding |
|---|---|---|---|---|---|---|---|
| Scheme 1 | 100 | 70 | 80 | 60 | 260 | 22 | No |
| Scheme 2 | 100 | 70 | 80 | 60 | 260 | 22 | Yes |
3. Material Database and Numerical Setup
A critical prerequisite for accurate numerical simulation is the establishment of a reliable material database for 1Cr20Ni14Si2. I used the Scheil solidification model in ProCAST to calculate thermophysical and mechanical properties as functions of temperature. Figures 1 and 2 show the computed density and solid fraction curves, respectively. The density at room temperature is about 7662 kg/m³, decreasing sharply near the liquidus. The solid fraction curve confirms the wide freezing range, which directly influences the formation of casting defects.
The thermal conductivity, enthalpy, and Newtonian viscosity of the alloy were also computed. The mechanical properties required for hot tearing prediction, such as Young’s modulus, Poisson’s ratio, thermal expansion coefficient, plastic modulus, and yield stress, were defined as temperature-dependent. For example, the plastic modulus was approximated as one-twentieth of Young’s modulus for simplicity, while the yield stress above 400 °C was estimated based on literature data for similar heat-resistant steels.
The simulation setup in ProCAST involved three modules: Mesh, Cast, and Viewer. In the Mesh module, the casting and gating system were discretized with triangular elements. The casting itself used a fine mesh size of 2 mm to accurately resolve thin walls, while the gating system used a coarser mesh of 3 mm. The ceramic shell was generated with a uniform thickness of 8 mm. The interface heat transfer coefficient between casting and shell was set to 900 W/m²·K, representing a typical COINC interface. Air cooling was applied as the boundary condition. The gravity direction was along the Y-axis, and the filling velocity was specified at the top of the pouring cup. The casting was treated as an elasto-plastic material to calculate stress and hot tearing indicators, while the gating system was treated as linear-elastic and the shell as rigid.
4. Initial Process Parameters
Before running simulations, I estimated the initial casting process parameters. The pouring temperature was chosen based on the alloy’s liquidus and the component structure. Since the casting is thin-walled and complex, a higher pouring temperature improves fluidity. The initial value was set to 1560 °C, which is 120 °C above the liquidus, reasonable for this Cr-containing steel whose viscosity is relatively high. The pouring rate was calculated using the formula:
$$ q = \frac{K \cdot l_{max}}{\delta_c} $$
where \(K\) is a coefficient equal to 0.05 for top gating, \(l_{max}\) is the maximum length of the thin-wall region, and \(\delta_c\) is the minimum wall thickness. This yielded \(q = 3.37\) kg/s, which I rounded to 3 kg/s for convenience. The mold preheating temperature was selected based on the casting size and alloy. Since the shell casting weighs about 7 kg and has extensive thin walls, I chose an initial shell temperature of 800 °C, consistent with typical values for steel investment casting.
5. Simulation of the Two Gating Schemes
5.1 Filling behavior
Simulation results for Scheme 1 (without padding) showed that during filling, the molten metal entered the cavity through the ingates but exhibited obvious splash and turbulence at the ends of the runner. This could entrain gas and inclusions, increasing the risk of casting defects. In contrast, Scheme 2 (with padding) produced a smoother filling front, as the padding helped guide the flow and reduce splashing. The total filling times were 4.5 s for Scheme 1 and 4.8 s for Scheme 2. Both schemes achieved a final fill fraction of 98% due to solidification shrinkage, but no mistun occurred.
5.2 Solidification behavior
Figure 3 shows the solidification progress for both schemes. At 25% solidification, the thin walls had already solidified in both cases, but the inclined groove region still contained liquid metal surrounded by partially solidified zones, creating isolated liquid pools that could lead to shrinkage porosity. Scheme 2 exhibited a lower solid fraction near the ingates (0.3) compared with Scheme 1, indicating better feeding capability. The total solidification time was 2436 s for Scheme 1 and 2470 s for Scheme 2. The longer solidification time in Scheme 2 allowed more feeding time and promoted a more favorable temperature gradient, acting to reduce casting defects.
5.3 Prediction of shrinkage porosity
Using ProCAST’s total shrinkage porosity criterion, I calculated the volume and distribution of shrinkage defects. In Scheme 1, the shrinkages were concentrated at the inclined groove and scattered throughout the thin-wall regions. The total shrinkage volume was 0.55 cc, with 0.32 cc at the groove and 0.23 cc in the thin walls. Scheme 2 significantly reduced the shrinkage volume to 0.29 cc (0.22 cc at the groove and 0.07 cc in the thin walls), a reduction of about 47%. The improvement is attributed to the padding, which maintained a liquid channel in the ingate region longer and enabled better feeding of the solidifying casting.
5.4 Hot tearing tendency
Hot tearing was evaluated using the Hot Tearing Indicator (HTI) calculated by the stress module. The highest HTI values appeared near the ingates and at the junction between the flange and the thin walls (labeled region A). For Scheme 1, the maximum HTI at region A was 0.0145, while for Scheme 2 it was 0.0122, indicating a lower hot tearing risk in Scheme 2. I further monitored the dynamic effective stress and elastic strain at a point in region A (point A for each scheme). At the end of solidification (around 800 s), the effective stress in Scheme 2 was about 34.3 MPa, slightly lower than 35.7 MPa in Scheme 1. The final effective stress after complete cooling was 255 MPa for Scheme 2, below the yield strength of 295 MPa, whereas Scheme 1 reached 264 MPa. Thus, Scheme 2 was selected as the optimal gating system because it reduces both shrinkage porosity and hot tearing tendency.
6. Effect of Casting Parameters on Shrinkage and Hot Tearing
6.1 Pouring temperature
Using the optimized Scheme 2 gating system, I studied the influence of pouring temperature on casting defects. The pouring rate was fixed at 3 kg/s and shell temperature at 800 °C. Seven temperatures from 1530 °C to 1650 °C were tested. Table 3 summarizes the resulting shrinkage volumes. The total shrinkage volume decreased dramatically from 0.86 cc at 1530 °C to only 0.03 cc at 1650 °C, a reduction of 96.5%. Raising the pouring temperature increases the fluidity of the melt and slows the cooling rate, thus prolonging the effective feeding time. The spatial distribution of shrinkage also changed: at lower temperatures, defects were present throughout the thin walls and the inclined groove; as temperature increased, the groove defects disappeared first, and at 1650 °C nearly all shrinkage was eliminated.
| Temperature (°C) | 1530 | 1550 | 1570 | 1590 | 1610 | 1630 | 1650 |
|---|---|---|---|---|---|---|---|
| Shrinkage volume (cc) | 0.86 | 0.39 | 0.36 | 0.25 | 0.12 | 0.19 | 0.03 |
The hot tearing behavior at region A was also examined. As shown in Table 4, the maximum HTI varied between 0.0118 and 0.0135 for temperatures of 1590–1650 °C. Interestingly, the highest HTI occurred at 1630 °C. At the monitoring point, the effective stress at the end of solidification (about 900 s) decreased as temperature increased, from 37.4 MPa at 1590 °C to 31.1 MPa at 1650 °C. However, the final effective stress after full cooling increased with pouring temperature: 266.7 MPa at 1590 °C, 273.6 MPa at 1610 °C, 287.9 MPa at 1630 °C, and 277.3 MPa at 1650 °C. The increase in final stress is related to higher residual stresses from a larger thermal gradient during cooling. Nevertheless, all values remained below the yield strength, so hot cracking was not expected.
| Temperature (°C) | 1590 | 1610 | 1630 | 1650 |
|---|---|---|---|---|
| HTI | 0.0118 | 0.0121 | 0.0135 | 0.0127 |
6.2 Pouring rate
The pouring rate was varied from 2.0 to 5.0 kg/s while fixing the pouring temperature at 1560 °C and shell temperature at 800 °C. The results in Table 5 show that the minimum shrinkage volume occurred at 3.0 kg/s (0.29 cc), while both lower and higher rates produced more shrinkage. At 2.5 kg/s, the volume was 0.57 cc, and at 4.0 kg/s it reached 0.51 cc. This indicates that an excessively slow fill leads to premature solidification and inadequate feeding, while too fast a fill causes turbulence and gas entrapment, both of which contribute to casting defects. The optimum pouring rate corresponds to a balanced filling velocity that avoids these adverse effects.
| Pouring rate (kg/s) | 2.0 | 2.5 | 3.0 | 3.5 | 4.0 | 4.5 | 5.0 |
|---|---|---|---|---|---|---|---|
| Shrinkage volume (cc) | 0.31 | 0.57 | 0.29 | 0.46 | 0.51 | 0.44 | 0.43 |
In terms of hot tearing, the HTI at region A remained almost unchanged at about 0.0114 for all pouring rates. The effective stress evolution at a monitoring point was also nearly identical: at 800 s the stress was about 34.4 MPa, and the final stress converged to 270 MPa with an elastic strain of 0.032. Therefore, the pouring rate primarily affects shrinkage porosity rather than hot tearing for this casting, as long as the rate is within a reasonable range.
6.3 Shell preheating temperature
Shell preheating temperatures were tested from 750 °C to 1050 °C while keeping pouring temperature at 1560 °C and pouring rate at 3 kg/s. Table 6 lists the shrinkage volumes. At 750 °C, the volume was 0.71 cc. Increasing the shell temperature to 950 °C reduced the volume to 0.05 cc, and at 1050 °C the shrinkage volume was only 0.01 cc, a 98% reduction compared to 750 °C. A hotter shell reduces the cooling rate of the casting, providing more time for melt to feed the solidifying zones. The distribution gradually shifted from the inclined groove to the thin-wall regions, and at 1050 °C defects were almost absent.
| Shell temperature (°C) | 750 | 800 | 850 | 900 | 950 | 1000 | 1050 |
|---|---|---|---|---|---|---|---|
| Shrinkage volume (cc) | 0.71 | 0.29 | 0.24 | 0.22 | 0.05 | 0.10 | 0.01 |
The hot tearing behavior is summarized in Table 7. The HTI remained nearly constant at 0.0114–0.0115 for shell temperatures of 750–1050 °C. However, the effective stress at the end of solidification (around 800 s) decreased notably with increasing shell temperature: 37.5 MPa at 750 °C, 32.5 MPa at 850 °C, 28.1 MPa at 950 °C, and 24.1 MPa at 1050 °C. The final effective stress after complete cooling was approximately 268 MPa for all cases, with a final elastic strain of 0.039. Thus, raising the shell temperature reduces the stress level during the critical final stage of solidification, which helps lower the risk of hot tearing, even though the final stress remains similar.
| Shell temperature (°C) | 750 | 850 | 950 | 1050 |
|---|---|---|---|---|
| HTI | 0.0115 | 0.0114 | 0.0115 | 0.0114 |
7. Determination of Optimal Process Parameters
Based on the single-factor studies, I designed four combined schemes to further refine the casting process. The pouring rate was fixed at 3 kg/s, while pouring temperature was set to either 1610 °C or 1650 °C, and shell temperature was set to either 950 °C or 1050 °C. The four combinations are listed in Table 8. The shrinkage volumes predicted by simulation are also shown. Scheme 4 (1650 °C, 3 kg/s, 1050 °C) yielded zero shrinkage volume, while Scheme 1 gave 0.13 cc, Scheme 2 gave 0.02 cc, and Scheme 3 gave 0.06 cc. Thus, Scheme 4 was the best in terms of eliminating shrinkage porosity.
| Scheme | Pouring temp (°C) | Pouring rate (kg/s) | Shell temp (°C) | Shrinkage volume (cc) |
|---|---|---|---|---|
| 1 | 1610 | 3 | 950 | 0.13 |
| 2 | 1610 | 3 | 1050 | 0.02 |
| 3 | 1650 | 3 | 950 | 0.06 |
| 4 | 1650 | 3 | 1050 | 0 |
For hot tearing, Scheme 3 had the highest HTI of 0.014, while Schemes 1, 2, and 4 gave 0.012. Monitoring point analysis (point 131649) showed that at 800 s, the effective stresses were 23.2, 21.1, 22.5, and 20.2 MPa for Schemes 1 to 4, respectively. The final effective stresses were 257, 259, 291, and 259 MPa. Scheme 4 again provided a stable final stress well below the yield strength. Consequently, I selected Scheme 4 as the optimal casting process: pouring temperature 1650 °C, pouring rate 3 kg/s, and shell preheating temperature 1050 °C.
8. Experimental Verification
8.1 Wax pattern and shell fabrication
To validate the numerical results, I performed actual investment casting trials. Wax patterns were injected using a semi-automatic wax injection machine. The injection pressure was 3.5 MPa, wax temperature 61 °C, injection time 25 s, and holding time 55 s. After cooling and de-molding, the patterns were inspected for surface integrity and then assembled onto the pre-designed runner system. The ceramic shell was built using colloidal silica as the binder and zircon/mullite as refractory materials. A total of five layers were applied, with the sequence: face coat (zircon slurry, 120-mesh zircon sand), second coat (60-mesh mullite), third to fifth coats (30-mesh mullite), and a sealing coat. Each layer was dried under controlled humidity. After drying, the wax was removed in a steam autoclave at 180 °C for 17 minutes, followed by sintering of the shell at 1170 °C for 60 minutes plus 30 minutes of holding.
8.2 Melting and pouring
The 1Cr20Ni14Si2 alloy was melted in a 50 kg medium-frequency induction furnace. The melt composition was verified by optical emission spectroscopy. Two sets of trials were conducted: the first used the initial process parameters (1560 °C, 3 kg/s, 800 °C), and the second used the optimized parameters (1650 °C, 3 kg/s, 1050 °C). After pouring, the castings were allowed to cool completely, then the shells were removed by vibration. The gating system was cut off, and the castings were subjected to grinding, shot blasting, pickling, and passivation to obtain the final surface finish.
8.3 Inspection results
Visual inspection revealed that both castings were completely filled with no misruns or surface defects. However, the casting produced with the initial parameters showed some surface blemishes and scale, which were attributed to the lower pouring and shell temperatures. The optimized casting displayed a cleaner and smoother surface, indicating better filling and solidification conditions.
To evaluate internal quality, industrial computed tomography (CT) was performed using an ICT-450 system with a voltage of 430 kV, current of 1.5 μA, focus size 0.4 mm, and voxel size 0.083 mm. The CT images for the initial parameters revealed two significant shrinkage areas at the inclined groove, with maximum diameters of 19.20 mm and 20.34 mm, respectively. The positions matched remarkably well with the simulated shrinkage locations, confirming the accuracy of the numerical model. No cracks were found at region A. In contrast, the CT images for the optimized parameters showed no discernible shrinkage or cracks in any critical region, including the groove and thin walls. The casting passed the quality inspection standard, demonstrating that the optimized process successfully eliminated casting defects.
9. Discussion
The numerical simulations provided deep insight into the formation mechanism of casting defects in thin-walled shell castings. The shrinkage defects in the inclined groove originated from the formation of isolated liquid pools due to non-uniform cooling between thick and thin sections. The gating system with padding proved effective in maintaining a liquid feeding path and prolonging solidification time, which reduced shrinkage by nearly half. Increasing pouring temperature and shell temperature further reduced shrinkage by enhancing melt fluidity and lowering the cooling rate. The pouring rate had an optimum value, as both extremes promoted turbulence or premature solidification. The hot tearing tendency, as indicated by HTI, was more sensitive to pouring temperature and shell temperature than to pouring rate. Higher pouring temperature increased final residual stress, while higher shell temperature reduced stress during the critical solidification stage. The optimal process balanced these effects to keep the maximum stress below the yield strength at all times.
One limitation of this study is that the yield stress data above 400 °C were approximated from literature and linear interpolation. More accurate high-temperature mechanical properties would further improve the reliability of hot tearing predictions. Nevertheless, the successful experimental verification suggests that the current simulation approach is sufficiently accurate for industrial process optimization.
10. Conclusions
- The cup capacity feeding method was successfully applied to design two gating schemes for a 1Cr20Ni14Si2 thin-walled shell casting. Scheme 2 with padding provided smoother filling and better feeding, reducing shrinkage porosity by 47% and lowering the HTI at the critical region from 0.0145 to 0.0122.
- Pouring temperature had a significant influence on casting defects: increasing from 1530 °C to 1650 °C reduced the shrinkage volume by 96.5%. Hot tearing tendency was minimized at intermediate temperatures, but all final stresses remained below the yield strength.
- The pouring rate showed an optimal value of 3 kg/s during this investigation; both lower and higher rates increased shrinkage defects, but the hot tearing behavior was almost unaffected.
- Increasing shell preheating temperature from 750 °C to 1050 °C reduced shrinkage porosity by 98%. The effective stress at the final solidification stage decreased with increasing shell temperature, while the final stress remained stable at about 268 MPa.
- The optimized casting process was determined as pouring temperature 1650 °C, pouring rate 3 kg/s, and shell preheating temperature 1050 °C. Industrial CT inspection of trial castings confirmed that this process produced defect-free castings, verifying the accuracy of the numerical simulations.
This work demonstrates that numerical simulation combined with careful parameter optimization is a powerful approach to eliminate casting defects in thin-walled shell castings. The methodology can be extended to other complex investment cast components, helping to reduce development costs and improve product reliability.