Abstract
In order to obtain dense microstructure, superior performance, and prevent crack formation in investment casting couplers, an orthogonal test was designed to determine the optimal parameters for steel casting. The optimal parameters were identified as a pouring temperature of 1,570°C, mold shell preheating temperature of 425°C, and pouring time of 29 seconds. Casting simulations were conducted to analyze the influence of the temperature field and stress field on the formation of shrinkage porosities and cracks under optimal parameters. Additionally, the KGT model was employed to simulate the microstructure of key casting parts. X-ray inspection of the trial-produced couplers revealed no internal cracks or pores, and the microstructures of various parts were consistent with the simulation results. After heat treatment, the average tensile strength of the couplers reached 1,020 MPa. The results indicate that the casting parameters are reasonable after orthogonal optimization, resulting in fine grains and performance exceeding railway standard requirements.
Introduction
Locomotive couplers are critical components for connecting vehicles and transmitting traction force in railway freight cars. They require excellent comprehensive properties to prevent internal porosities, cracks, and significant microstructural defects. Researchers have conducted extensive studies on casting processes, casting stresses, and related defects, leading to significant improvements in crack reduction and product qualification rates. This paper focuses on the precision casting process of locomotive couplers, aiming to address the issues of crack and shrinkage defect formation during production.
Based on ProCAST numerical simulations, orthogonal tests were designed to optimize the pouring temperature, mold shell temperature, and pouring time, which are key parameters influencing casting quality. The optimal parameters for high-quality mold filling and solidification were determined. Temperature field, stress field, and microstructure simulations were conducted to validate the optimal casting parameters obtained from the orthogonal tests. Finally, precision-cast couplers were trial-produced to provide a reference for related product manufacturing.
Simulation Model and Orthogonal Test
1.1 Simulation Model and Initial Conditions
The locomotive coupler has a hollow internal structure with significant wall thickness variations. Its overall dimensions are 594 mm × 370 mm × 350 mm, and its mass is approximately 70 kg. A vertical top-pouring, two-sided diversion gating system was adopted, with a feeder set at the top of the coupler head. The coupler casting and its gating system .
The coupler was produced using E-grade cast steel ZG25MnCrNiMo. The chemical composition of the steel, determined by furnace-front testing, is shown in Table 6. The steel exhibits excellent wear resistance, low-temperature impact resistance, and high strength. The mixed model in the thermodynamic database of ProCAST software was used to calculate the thermal conductivity, density, enthalpy, and other physical properties of the cast steel at different temperatures. The initial conditions for the filling and solidification simulations are listed in Table 1.
Table 1: Initial conditions for filling solidification simulation
| Parameter | Value |
|---|---|
| Pouring temperature (T1) | 1,550°C |
| Mold shell preheating temperature (T2) | 400°C |
| Ambient temperature (T0) | 20°C |
| Pouring time (t) | 15 s |
| Heat transfer coefficient between mold shell and casting (h1) | 500 W·m-1 |
| Heat transfer coefficient between mold shell and air (h2) | 10 W·m-1 |
| Mesh size (pouring system/riser) | 8 mm |
| Mesh size (casting body) | 5 mm |
Table 6: Chemical composition of ZG25MnCrNiMo cast steel
| Element | C | Si | Mn | Cr | Ni | Mo | P | S |
|---|---|---|---|---|---|---|---|---|
| Content (%) | 0.26 | 0.45 | 1.4 | 0.55 | 0.45 | 0.25 | ≤0.35 | ≤0.35 |
1.2 Casting Stress Calculation Model
Thermal stress simulations were performed using the thermo-elastoplastic model. Since hot crack defects primarily occur in the mushy zone, stress analysis was focused on this region. The elastic-plastic constitutive model, specifically the classical bilinear hardening model, was employed to describe the physical properties of the E-grade steel:
sigma = begin{cases} E_1 \varepsilon & (\varepsilon \leq \varepsilon_s) \\ sigma_{0.2} + E_2 (\varepsilon – \varepsilon_s) & (\varepsilon > \varepsilon_s) end{cases}
where σ is the stress (MPa), σ0.2 is the yield strength (MPa), ε is the strain, εs is the yield strain, and E1 and E2 are the elastic moduli (MPa) of the material.
1.3 Microstructure Simulation Model
A combination of the cellular automata and finite element method (CAFE) was used for microstructure simulation. This model includes the continuous nucleation Gauss distribution model and the dendrite tip growth dynamics model, known as the KGT (Kurz-Givoanola-Trivedi) model. The mathematical expressions for these models are as follows:
fracdnd(ΔT)=2π⋅ΔTσnmaxexp(−2(ΔTσ)2(ΔT−ΔTmax)2)
where n is the grain density, ΔT is the undercooling, ΔTmax is the average nucleation undercooling, ΔTσ is the standard deviation of nucleation undercooling, and nmax is the maximum nucleation density obtained by integrating from 0 to infinity.
The kinetic parameters and nucleation Gaussian distribution parameters of ZG25MnCrNiMo are shown in Table 2.
Table 2: Dynamics parameters and nucleation Gaussian distribution parameters of ZG25MnCrNiMo cast steel
| Element | C | Cr | Mn | Mo | Ni |
|---|---|---|---|---|---|
| ns,max (m^-3) | 1×10^9 | 1×10^9 | 1×10^9 | 1×10^9 | 1×10^9 |
| nv,max (m^-3) | 1×10^11 | 1×10^11 | 1×10^11 | 1×10^11 | 1×10^11 |
| ΔTs,max (K) | 8 | 8 | 8 | 8 | 8 |
| ΔTv,max (K) | 8 | 8 | 8 | 8 | 8 |
| ΔTs,σ (K) | 2 | 2 | 2 | 2 | 2 |
| ΔTv,σ (K) | 2 | 2 | 2 | 2 | 2 |
| m | -83.0242 | -1.83704 | -5.17283 | -2.61329 | -3.87222 |
| k | 0.166396 | 0.907287 | 0.737152 | 0.781625 | 0.803188 |
| Γ (×10^-7 K·m) | 3 | 3 | 3 | 3 | 3 |
| D1 (×10^-9 m^2·s^-1) | 3 | 3 | 3 | 3 | 3 |
1.4 Orthogonal Test Design
To optimize the volume of shrinkage porosities and shrinkages (VSP) and the stress values at three nodes in the casting’s stress concentration areas, an orthogonal test was designed. Smaller VSP and σ values indicate a lower tendency for casting defects. The three most critical factors affecting quality objectives were selected as design variables: pouring temperature (A), mold shell temperature (B), and pouring time (C). The initial conditions for these factors were based on empirical values and material properties, with ranges determined as follows: pouring temperature (1,530-1,590°C), mold shell preheating temperature (350-500°C), and pouring time (28-34 s). Four levels were selected for each factor, and an L16(4^3) orthogonal test was designed using the Taguchi method. The factors and levels of the orthogonal test are shown in Table 3, and the orthogonal test design and results are shown in Table 4.
Table 3: Factor level of orthogonal test
| Level | A/°C | B/°C | C/s |
|---|---|---|---|
| 1 | 1,530 | 350 | 28 |
| 2 | 1,550 | 400 | 30 |
| 3 | 1,570 | 450 | 32 |
| 4 | 1,590 | 500 | 34 |
Table 4: Design and results of Taguchi orthogonal experiment
| Seq | A/°C | B/°C | C/s | VSP/cm^3 | (S/N)/dB | σ/MPa | (S/N)/dB |
|---|---|---|---|---|---|---|---|
| 1 | 1,530 | 350 | 28 | 1.048 | -0.407 | 346.124 | -51.641 |
| 2 | 1,530 | 400 | 30 | 0.887 | 0.257 | 344.063 | -51.762 |
| 3 | 1,530 | 450 | 32 | 0.910 | 0.436 | 345.167 | -51.881 |
| 4 | 1,530 | 500 | 34 | 1.082 | 0.128 | 347.598 | -51.999 |
| 5 | 1,550 | 350 | 30 | 0.894 | 0.284 | 344.913 | -51.952 |
| 6 | 1,550 | 400 | 32 | 1.021 | -0.587 | 347.017 | -51.959 |
| 7 | 1,550 | 450 | 34 | 1.054 | 0.104 | 346.577 | -51.998 |
| 8 | 1,550 | 500 | 28 | 1.089 | -0.011 | 343.366 | -52.056 |
| 9 | 1,570 | 350 | 32 | 1.108 | -0.118 | 353.024 | -52.032 |
| 10 | 1,570 | 400 | 34 | 0.990 | -0.098 | 349.998 | -52.040 |
| 11 | 1,570 | 450 | 28 | 1.124 | -0.190 | 352.604 | -52.064 |
| 12 | 1,570 | 500 | 30 | 1.252 | -0.367 | 352.572 | -52.104 |
| 13 | 1,590 | 350 | 34 | 0.897 | -0.279 | 345.059 | -52.095 |
| 14 | 1,590 | 400 | 28 | 1.125 | -0.337 | 352.233 | -52.101 |
| 15 | 1,590 | 450 | 30 | 1.074 | -0.356 | 347.722 | -52.122 |
| 16 | 1,590 | 500 | 32 | 0.900 | -0.287 | 345.979 | -52.156 |
1.5 Orthogonal Test Results and Analysis
The orthogonal test results indicate that the three factors have different degrees of influence on the volume of shrinkage porosities and shrinkages (VSP) and stress values at critical parts. The variance P-values show that the pouring temperature (A) has a significant impact on both VSP and stress values, while the mold shell preheating temperature (B) and pouring time (C) have minimal influence on stress values. The range analysis results show that the factors’ influence on VSP is in the order of A > C > B. According to the Taguchi method, the optimal parameter combination for minimizing VSP is A3B4C1: pouring temperature of 1,570°C, mold shell preheating temperature of 500°C, and pouring time of 28 s. For stress values, the optimal parameter combination is A3B1C2: pouring temperature of 1,570°C, mold shell preheating temperature of 350°C, and pouring time of 30 s.
The best parameter combination obtained from the Taguchi orthogonal test is the average of the optimal parameters for VSP and stress values, resulting in a pouring temperature of 1,570°C, mold shell preheating temperature of 425°C, and pouring time of 29 s. These optimized parameters were used as the initial conditions for the filling and solidification simulations in ProCAST to analyze the temperature field, shrinkage porosities, stress, and microstructure during the precision casting process of the couplers.
Simulation Analysis and Experiments
2.1 Simulation Verification of Orthogonal Test Results
2.1.1 Temperature Field Simulation Analysis
The temperature field distribution at the end of the solidification process. As the solidification proceeds, the cooling rate gradually slows down. The solidification sequence starts from the middle and lower base of the casting, then extends to the surroundings, and finally reaches the shoulders and gate areas. This solidification sequence is consistent with the wall thickness and gating system structure of the casting. Due to significant variations in wall thickness, the gating system structure was designed to control the solidification sequence, reducing the temperature gradient within a certain range. Large cross-sectional changes occur at the connection points between the coupler head, coupler tongue, and shoulders, so gates were set on both sides of the shoulders to facilitate feeding and avoid the formation of shrinkage porosities and shrinkages.
Due to the reasonable pouring temperature and mold shell preheating temperature, the shrinkage porosities and shrinkages after solidification are located only at the top of the coupler head and the abdomen of the uncoupling cylinder, with a small volume of 0.879 cm^3. The critical parts of the casting show no shrinkage porosities or shrinkages.
2.1.2 Stress Field Simulation Analysis
Casting stress is the direct cause of hot cracks. Five nodes were selected at locations prone to fatigue during coupler operation to analyze the casting solidification process and stress. Node 69754 inside the coupler tongue starts to solidify first. At t = 3,500 s, its solidification fraction converges with Nodes 29772, 51539, and 59684, and the solidification rate decreases. Node 59684 at the bottom of the coupler solidifies relatively slowly. Due to asynchronous solidification and contraction, the maximum stress reaches 470 MPa at t = 6,000 s. Node 35369 at the shoulder gate basically completes solidification at t = 5,000 s, influenced by the nearby thick runner. This node solidifies the slowest but experiences relatively low stress close to 200 MPa due to its simple structure and free contraction. According to the analysis of the thermal cracking index of the casting, areas with sudden cross-sectional changes in the coupler, such as the junction between the shoulder and the uncoupling cylinder and the junction between the base and the gating system, have a higher tendency towards hot cracking. This is primarily because these regions are relatively thick and have heat sources (gating system), leading to hindered solidification shrinkage and subsequent thermal stress. However, throughout the casting process, the stress values at critical nodes of the coupler remain significantly lower than the material’s tensile strength, suggesting a low probability of hot cracking in the casting.
In summary, the stress field simulation provides valuable insights into the solidification behavior and stress evolution of the investment casting coupler. The analysis highlights the areas prone to stress concentration and potential hot cracking, guiding the optimization of casting parameters to ensure the production of high-quality, defect-free couplers. By carefully selecting and analyzing the stress at critical nodes, the study demonstrates the effectiveness of numerical simulation in predicting and mitigating casting defects, thereby enhancing the reliability and performance of the final product.