2025 Volume 66 Issue 1 Pages 99-106
The effect of Nb content on the mechanical properties and solidification structure of stabilized ferritic stainless cast steel with a basic composition of 18%Cr–0.5%Cu–(0.35–1.1)%Nb–0.035%(C + N) was investigated. Tensile test results for specimens containing 0.35–0.45% Nb showed a tensile strength of 370–380 MPa, 0.2% proof stress of 270–280 MPa and elongation of 7–10%. However, as the Nb content increased to 1.1%, the 0.2% proof stress increased to 360 MPa and the elongation decreased significantly to 1%. In all specimens, Nb(C, N) particles were dispersed in the ferrite matrix, and the particle size became larger with increasing Nb content. Some continuous film-like Nb(C, N) appeared at grain boundaries in the 0.75% Nb sample and increased in the 1.1% Nb sample. These difference in Nb(C, N) distribution should affect the mechanical properties.
The solidification process of the sample containing 0.45% Nb, which was considered as optimal Nb content, was investigated. The results show that δ-ferrite grows as dendrites in the solidification temperature range from 1768 K to 1713 K. The analysis with Scheil’s equation predicted that δ + Nb(C, N) eutectic starts crystalizing at fraction solid of 0.98, but that the δ + Nb(C, N) eutectic does not appear taking into consideration the diffusion of Nb in the solid. The results of the latter analysis considering diffusion in the solid phase were consistent with the solidification structure of 0.45% Nb sample.
This Paper was Originally Published in Japanese in J. JFS 95 (2023) 115–122.
Ferritic stainless steel is extensively used in heat-resistant parts of machines and industrial plants [1]. Ferritic stainless steel exhibits lower heat resistance compared to austenitic stainless steel. However, its lower thermal expansion coefficient leads to less stress and strain under service conditions, which is beneficial for castings with complex shapes. The low alloy cost is another advantage of ferritic stainless steel. In general, stainless cast steel contains a slightly higher amount of C and N than rolled steel products, and their sensitivity to intergranular corrosion is higher [2]. Nb with a higher carbide formation tendency is commonly added to suppress the precipitation of chromium carbonitride at the grain boundary to improve the corrosion resistance of ferritic stainless cast steel [3–8]. On the other hand, as the partition coefficient of Nb to primary δ-ferrite is less than 1, Nb segregates to residual liquid during solidification. An excess Nb content crystallizes eutectic Nb(C, N) at the grain boundary and deteriorates the mechanical properties, particularly ductility. Nb(C, N) also precipitates during cooling in the mold, and its morphology and distribution affect the mechanical properties [9–11]. Higher ductility is desired for stabilized ferritic stainless cast steel to eliminate the risk of crack formation caused by thermal and mechanical strain during welding and finishing processes. Therefore, optimizing the Nb addition is essential for obtaining a suitable solidification structure and good mechanical properties.
In this study, the effect of Nb content on the solidification structure and mechanical properties of JIS SUS430J1L-type steel samples containing Cr: 18, Cu: 0.5, Nb: 0.3–1.1, (C + N): 0.07 (mass%) was investigated. First, the morphology and distribution of Nb(C, N) and its tensile properties were examined in relation to the Nb content. Then, the solidification and precipitation sequences and the redistribution of alloying elements in samples with an appropriate Nb content were analyzed.
The chemical compositions of the samples are listed in Table 1. Appropriately adjusted amounts of raw materials, including ferritic stainless steel scrap, ferro-chrome, ferro-manganese, ferro-silicon, and ferro-niobium were melted at 1953 K in a high-frequency induction furnace and then cast into JIS G 5502 knock-off Kb-type sand molds. The Nb content in the experimental samples varied from 0.34% to 1.1%, which is slightly higher than the values 0.28%–0.80% defined by JIS G 4304 and 4305 for stabilizing elements. The contents of other elements were within the range specified in JIS G 0321.
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The tensile test was conducted using JIS No. 14 samples with a parallel diameter of 8 mm and a gauge length of 44 mm. The tensile speeds were 0.7 mm/min in the elastic range and 5.0 mm/min in the plastic range, respectively, in accordance with JIS Z 2241. A sample for microstructure observation was taken from the gripping portion of the tensile specimen to analyze the relationship between mechanical properties and microstructure observations.
The solidification and subsequent precipitation sequences were examined using the 0.45% Nb sample, which exhibited appropriate tensile properties. Approximately 40 g of a block sample with the composition shown in Table 2 was placed in a mullite tamman tube with an inner diameter of 20 mm and height of 120 mm and heated to 1823 K in a siliconit furnace under an argon atmosphere. The molten sample was cooled at a rate of 0.05 K/s, and the solidification temperature was determined by thermal analysis. To analyze the phase transformation sequences and changes in the microstructure, a series of samples were quenched at 1756 K, 1743 K, 1673 K, 1623 K, and 1523 K by dropping the tamman tube into water. The quenching temperatures were determined based on the equilibrium diagram of the Fe–Nb quasi-binary system calculated using Thermo-Calc [12] (Database: Thermotech, TCFE9), as shown in Fig. 1. The quenched specimens were cut at the center and mirror-finished by wet polishing with water-resistant abrasive paper and buffing with diamond paste and colloidal silica. The microstructure was then observed by scanning electron microscopy (SEM, JEOL: JSM-IT800SHL). In addition, the sample quenched at 1756 K was analyzed using an electron probe micro-analyzer (EPMA, Shimazu: EPMA-8050G) to clarify the redistribution behavior of the solute elements during solidification. Line analysis was performed across the secondary dendrite arms at an acceleration voltage of 15 kV, irradiation current of 100 nA, and beam diameter of 0.1 µm. Repeated point analysis was also performed on a 1 mm square of the quenched specimen at an interval of 100 µm to determine the partition coefficients of solute elements to the primary δ-ferrite. The dendrite arm spacing of the sample was approximately 200 µm; thus, the analyzed area contained several dendrite arms. As the random sampling method [13–15], the partition coefficient of each element was evaluated under the assumption that the concentration distribution obtained by sorting the point analysis results in ascending order of concentration follows Scheil’s equation, a micro-segregation model. Random sampling method provides a reasonable partition coefficient when the number of analysis points is not less than 100 [15].
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Pseudo Fe–Nb binary equilibrium phase diagram calculated by Thermo-calc. The arrows indicate quenching temperatures of 1756 K, 1743 K, 1673 K, 1623 K and 1523 K.
As shown in Fig. 2, the stress–strain curves of the samples in Table 1 significantly varied depending on the Nb content. The abrupt increase in stress at 1% strain can be attributed to changes in the tensile speed. The effects of Nb content on the tensile strength (σB), 0.2% proof stress (σ0.2), and elongation (δ) are illustrated in Fig. 3. The tensile strengths were 370–380 MPa at 0.46% Nb, 338 MPa at 0.75% Nb, and 375 MPa at 1.1% Nb. The 0.2% proof stress was 270–280 MPa at 0.46% Nb, 290 MPa at 0.75% Nb, and 354 MPa at 1.1% Nb. The increase in the 0.2% proof stress with increasing Nb content can be attributed to the solid solution strengthening effect of Nb in the ferrite phase [16], as well as the plastic constraint induced by Nb(C, N) precipitating at grain boundaries. Conversely, the elongation decreased significantly from 10% to 1% as the Nb content increased from 0.46% to 1.1%. The lower mechanical property limits for SUS430J1L rolled steel in JIS G 4304 is 390 MPa for tensile strength, 205 MPa for 0.2% proof stress, and 22% for elongation. Compared with these lower limits, the tensile strength of the test cast samples was similar, but their 0.2% proof stress was significantly higher, and their elongation was substantially lower. This difference from rolled steel can be attributed to the cast microstructure. Therefore, to clarify the microstructure, EPMA analysis was performed on samples cut from the gripping portion of the tensile specimens. Figure 4 shows backscattered electron (BSE) images and characteristic X-ray images of Nb. A white compound phase is distributed in the ferrite matrix in all samples. This compound phase was identified as carbonitride Nb(C, N) because the characteristic X-ray image showed a significantly high concentration of Nb and higher C and N concentrations. The volume fractions of Nb(C, N) were 0.42% for the 0.46% Nb sample, 0.67% for the 0.75% Nb sample, and 1.3% for the 1.1% Nb sample; thus, the amount of Nb(C, N) increased as the Nb content increased. Furthermore, the distribution of Nb(C, N) varied with the Nb content; Nb(C, N) was uniformly distributed as fine particles with a diameter of a few µm in the 0.46% Nb sample, whereas Nb(C, N) was coarsely distributed at the grain boundaries in the 1.1% Nb sample. To reveal the morphology and distribution of Nb(C, N), the 1.1% Nb sample was deep-etched using vilella reagent (1 g of picric acid, 5 ml hydrochloric acid, 100 ml ethanol), and SEM observations were performed. As shown in Fig. 5(a), dendritic coarse Nb(C, N) particles were continuously formed at the grain boundaries. In addition, the non-equilibrium δ + Nb(C, N) eutectic phase, as shown in Fig. 5(b), crystallized with a small volume fraction of approximately 0.05%. Therefore, the decrease in elongation and increase in 0.2% proof stress with increasing Nb content can be attributed to the solid solution strengthening of Nb and the plastic constraint by the continuous distribution of Nb(C, N) at the grain boundaries. Because neither Cr carbides nor Cr nitrides formed, the matrix phase was stabilized by Nb in all samples. Therefore, the optimal Nb content for this cast steel was determined to be 0.45%, considering both its grain boundary corrosion resistance and mechanical properties.
Nominal stress–strain curves of Sample 3 (Nb: 0.46%), Sample 8 (Nb: 0.75%) and Sample 9 (Nb: 1.1%).
Relationship between mechanical properties and Nb content of specimens.
BSE images and characteristic X-ray images of Nb in Sample 3 (Nb: 0.46%), Sample 8 (Nb: 0.75%) and Sample 9 (Nb: 1.1%). (online color)
BSE images of (a) Nb(C, N) precipitated at grain boundary and (b) eutectic Nb(C, N) in deep etched sample 9 (Nb: 1.1%).
The solidification and subsequent solid-state transformation of the 0.45% Nb sample were investigated. The thermal analysis curve of the sample cooled at 0.05 K/s from the molten state is displayed in Fig. 6. A clear recalescence appeared on the cooling curve, revealing that the temperature of the molten metal increased to 1768 K upon nucleation and that rapid crystal growth started at 1760 K. The cooling curve indicates that this cast steel solidifies into a single-phase, δ-ferrite, and the solidification temperature range of 1768 K–1713 K.
Thermal analysis curve of experimental alloy with 0.45%Nb.
Figure 7 depicts the microstructures of samples quenched at different temperatures during thermal analysis. The fine black spherical inclusions distributed in each sample represent SiO2-based oxides formed as the deoxidation reaction product. In the samples quenched at (a) 1756 K and (b) 1743 K, the δ-ferrite crystallized into a dendritic shape, and fine δ + Nb(C, N) eutectic was observed in the dendrite arms. This δ + Nb(C, N) eutectic is attributed to the rapid cooling of the melt remaining among the dendrite arms. The microstructures of the samples quenched at (c) 1673 K and (d) 1623 K consisted of single-phase δ-ferrite. In contrast, fine Nb(C, N) particles were distributed in the δ-ferrite phase at (e) 1523 K and (f) room temperature. These Nb(C, N) particles precipitated during cooling after solidification. The Nb(C, N) particles surrounded most SiO2 particles, suggesting a crystallographic similarity between Nb(C, N) and SiO2. Because the flaky and plate-like development of Nb(C, N) particles also indicates a specific crystallographic relation between Nb(C, N) and the matrix δ-ferrite phase, we analyzed the crystal orientations of both phases using electron backscatter diffraction. As indicated by the arrows in the polar point diagram in Fig. 8, the orientation relationship is {001}Nb(C, N)//{001}δ and ⟨001⟩Nb(C, N)//⟨110⟩δ. NaCl-type precipitates and body-centered cubic ferrite matrix have a well-known Baker–Nutting orientation relationship, ({001}NaCl//{001}α and ⟨001⟩NaCl//⟨110⟩α) [17,18]. Accordingly, flaky Nb(C, N) particles should grow with a good matching orientation relationship with the δ-ferrite phase. Thus, Nb(C, N) particles began to precipitate at a temperature between 1623 K and 1523 K, and the number and size increased with cooling. This result is consistent with the equilibrium diagram of the Fe–Nb quasi-binary system of this cast steel shown in Fig. 1.
BSE images of specimens quenched at (a) 1756 K, (b) 1743 K, (c) 1673 K, (d) 1623 K, (e) 1523 K and (f) furnace cooled to R.T.
IPF map and stereographic projection of flaky Nb(C, N) and delta ferrite phase. (online color)
Each solute element segregated during solidification, and the micro-segregation of these elements influenced the amount of δ + Nb(C, N) eutectic formed at the final solidification stage and the distribution and morphology of Nb(C, N) precipitated after solidification. Therefore, considering the secondary dendrite arm as a segregation unit of solute elements [15], we investigated the behaviors of Cr, Nb, C, and N, which are the main elements influencing the mechanical properties and corrosion resistance of this cast steel.
3.3.1 Partition coefficients and redistribution of solute elements during solidificationEPMA line analysis was performed across the secondary dendrite arms of the sample quenched at 1756 K. The results are presented in Fig. 9. All alloying elements exhibited a normal segregation tendency, with lower concentrations at the center and higher concentrations toward the boundaries of the dendrite arms. Random sampling method [13–15] was performed to determine the concentration changes of Cr and Nb with increasing solid fraction. In this method, the area fraction of solid-phase fs was evaluated as the ratio of the cumulative number of analyzed points to the total number of measurement points, referring to the analysis by Yamamoto et al. [15]. Figure 10 plots the analyzed Cr and Nb concentrations in relation to the solid fraction. The concentrations of both elements increased as the solid fraction (fs) increased. The low Cr content at solid fraction fs = 0.1 can be attributed to Nb(C, N) precipitating in the analyzed region during rapid cooling, as indicated by the significantly high Nb concentration, exceeding 15% at these points.
Distributions of alloy elements in secondary dendrite arm in specimen quenched at 1756 K. (online color)
Cr and Nb distribution measured with Random sampling method and estimated by Scheil’s equation for the sample quenched at 1756 K.
Assuming that the Cr and Nb distributions in the solid phase followed Scheil’s equation (eq. (1)), which was derived under the condition of complete mixing in the liquid phase and no solute diffusion in the solid phase, we obtained eq. (2) by taking the logarithm of both sides of eq. (1).
| \begin{equation} C_{s} = kC_{0}(1 - f_{s})^{(k - 1)} \end{equation} | (1) |
| \begin{equation} \mathit{ln}\frac{C_{s}}{C_{0}} = \mathit{ln}\,k_{1} + (k_{2} - 1)\mathit{ln}\,f_{L} \end{equation} | (2) |
In eqs. (1) and (2), k: the partition coefficient; Cs: the solute concentration of the solid phase at the freezing front (mass%); fL: the liquid fraction; C0: the initial concentration (mass%).
By plotting ln fL on the horizontal axis and ln(Cs/C0) on the vertical axis and applying linear approximation using the least-squares method, we evaluated the partition coefficients k1 and k2 from the intercept and slope of the straight line. The least-squares method was applied in the high linearity range of the solid fraction fs = 0.2–0.7 [15], and the initial concentration C0 was used as the average concentration at all measurement points. The average values (k1 + k2)/2 obtained for Cr and Nb are listed in Table 3. Table 3 also lists the partition coefficients of Cr, Nb, C, and N calculated using Thermo-Calc. The experimental partition coefficients of Cr and Nb obtained using the random sampling method were comparable to those estimated by Thermo-Calc.
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Figure 10 shows the Cr and Nb concentrations in the solid phase, as obtained by substituting the experimental partition coefficients into Scheil’s equation (eq. (1)). The calculated Cr and Nb contents significantly exceed the measured contents when the solid fraction fs exceeds 0.9. This can be attributed to Scheil’s equation, which infinitely concentrates solute elements as the solid fraction fs approaches 1, whereas the diffusion of solutes becomes non-negligible under steeper solute concentration gradients [15].
Therefore, the effect of diffusion in solids was analyzed in terms of the redistribution of Nb. Nb was selected because of its smaller partition coefficient and higher segregation tendency. Brody–Flemings [19], Clyne–Kurz [20], and Ohnaka [21] proposed theories and simulation models for evaluating the effect of solid-phase diffusion during solidification. Among them, the Ohnaka model provides more accurate analysis [22], providing eqs. (3)–(7) for the solute concentration CL of the liquid phase at a solid fraction fs on the premise of complete liquid mixing.
| \begin{equation} C_{L} = C_{0}(1 - \varGamma f_{s})^{\frac{k - 1}{\Gamma}} \end{equation} | (3) |
| \begin{equation} \varGamma = 1 - \frac{\beta k}{1 + \beta} \end{equation} | (4) |
| \begin{equation} \beta = \frac{2D_{S}\theta}{S^{2}} \end{equation} | (5) |
| \begin{equation} D_{s} = D_{0}\exp\left(-\frac{Q}{RT}\right) \end{equation} | (6) |
| \begin{equation} S = \frac{\lambda_{2}}{2} \end{equation} | (7) |
Here, θ: the solidification time (s); Ds: the diffusion coefficient of the solute in the solid phase (m2/s); λ2: the secondary dendrite arm spacing (µm).
The behavior of solute elements in the solid phase during solidification was analyzed in one dimension using the forward difference method [23–26]. Figure 11 shows a schematic of the analysis. When the half-region of the secondary dendrite arm spacing was divided into 100 segments, the solid–liquid interface at the solid fraction fs = I was located between elements (I) and (I + 1). The solute concentration in the liquid phase of element (I + 1) is expressed by Ohnaka’s equation (eq. (3)). The solute concentration in the solid phase of element (I) is Cs* = kCL. The solute concentrations in each solid-phase element at this time are represented as Cs(1)–Cs(I). The time Δt required for the solid–liquid interface to advance by one element is obtained by dividing the solidification time θ by the number of divisions N = 100. The diffusion in solid during Δt is calculated using the diffusion equation, which produces eqs. (8)–(10). The data presented in Table 4 were used for the analysis.
| \begin{equation} \text{Element i} = 1\quad C_{S}'(1) = C_{S}(1) + \frac{D_{S}\Delta t}{\Delta x^{2}}\{C_{S}(2) - C_{S}(1)\} \end{equation} | (8) |
| \begin{align} &\text{Element i} = \text{2–($\text{I}-1$)}\\ & C_{S}'(i) = C_{S}(i) + \frac{D_{S}\Delta t}{\Delta x^{2}}\{C_{S}(i + 1) - 2C_{S}(i) + C_{S}(i - 1)\} \end{align} | (9) |
| \begin{align} &\text{Element i} = \text{I}\\ & C_{S}'(I) = C_{S}(I) + \frac{D_{S}\Delta t}{\Delta x^{2}}\{C_{S}(I - 1) - C_{S}(I)\} \end{align} | (10) |
Analysis model for redistribution of alloying element during solidification.
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Here, CS(i): the solute concentration in the solid phase of element i at time t; $C'_{S}(i)$: the solute concentration in the solid phase of element i at time t + Δt; Δx: the width of the element (the value is obtained by dividing the analysis target region S by the number of divisions N = 100).
The Nb distribution calculated using this method is shown in Fig. 12, along with those obtained using the random sampling method and Scheil’s equation. Regarding the Nb distribution in the solid fraction range of fs = 0.2–0.7 used to determine the partition coefficient, Ohnaka’s equation produced slightly higher Cs values than Scheil’s equation; however, the difference was negligibly small. The experimental values obtained using the random sampling method were slightly higher; however, they exhibited a similar relationship to fs. At the late stage of solidification around fs = 0.9, Ohnaka’s equation produced lower Cs values than Scheil’s equation, approaching the measured values. These results demonstrate the significance of diffusion within the solid phase at the final solidification stage.
Nb distributions in solid evaluated by Random sampling method, Scheil’s and Ohnaka’s equation.
Cr, Nb, C, and N, with partition coefficients less than 1, concentrate in the residual liquid during solidification, causing micro-segregation in the dendritic δ-ferrite phase, and occasionally, eutectic crystallization. Therefore, we evaluated the changes in the Nb, C, and N contents in the liquid phase during solidification using the partition coefficients calculated by Thermo-Calc and Scheil’s equation for Nb and the equilibrium solidification equation (eq. (11)) for C and N.
| \begin{equation} C_{L} = \frac{C_{0}}{f_{s}k + (1 - f_{s})} \end{equation} | (11) |
Here, CL: the C and N concentration in the liquid phase (mass%); k: the equilibrium partition coefficient of C and N calculated using Thermo-Calc, C0: the initial concentration of C and N (mass%).
Figure 13 shows the liquidus surface diagram of an Fe–Nb–C/N pseudo-ternary system with 18% Cr, as calculated using Thermo-Calc. Because Nb(C, N) can use various C/N ratios [27], the calculated concentration changes of Nb and C + N in the liquid phase are depicted in Fig. 13. For the 0.45% Nb sample, both the Nb and C + N concentrations increased during crystallization of primary δ-ferrite, and these concentrations reached the δ + Nb(C, N) eutectic line at fs = 0.98. This result indicates that 2% of δ + Nb(C, N) eutectic crystallizes at the end of solidification. Figure 13 shows similar residual liquid composition changes for the 0.75% and 1.1% Nb samples, predicting their δ + Nb(C, N) eutectic amounts as 5% and 7%, respectively. However, the 0.75% and 1.1% Nb samples were quenched immediately after solidification and exhibited no eutectics. The few δ + Nb(C, N) eutectics in the experimental samples can be attributed to Nb diffusion in the solid phase.
Changes in Nb, C and N concentration in residual liquid during solidification estimated by Scheil’s equation, traced on liquidus surface diagram for Fe–18%Cr–Nb–(C + N) alloy.
Therefore, we calculated the changes in the Nb content of the residual liquid using Ohnaka’s equation. The results are presented in Fig. 14. For the 0.45% Nb sample, the liquid composition did not reach δ + Nb(C, N) even at the solid fraction fs = 1, indicating that the concentration of Nb in the liquid phase was suppressed by diffusion in δ-ferrite during solidification, and there was no δ + Nb(C, N) eutectic crystallization. This result agrees with the microstructure observations shown in Fig. 7. Conversely, a small amount of fine δ + Nb(C, N) eutectic appeared between the dendrite arms of the specimens quenched at 1756 K and 1743 K, as shown in Fig. 6, probably due to the suppression of the diffusion of Nb, C, and N during quenching. Figure 14 also shows the results of Ohnaka’s equation evaluation for the 0.75% and 1.1% Nb samples with eutectic amounts of 1% and 4%, respectively. Although these values are significantly less than Scheil’s equation predictions, they are greater than the measured values (0.75% Nb: 0% and 1.1% Nb: 0.02%). This result can be explained by the distribution of residual liquid; although the calculations assume that the dendrites are plate-like that is, one-dimensional diffusion, the actual residual liquid is distributed in small pools among dendrite arms at the end of solidification, and the diffusion occurs three-dimensionally.
Changes in Nb, C and N concentration in residual liquid during solidification estimated by Ohnaka’s equation, traced on liquidus surface diagram for Fe–18%Cr–Nb–(C + N) alloy.
The effect of Nb content on the mechanical properties and solidification mechanism of stabilized ferritic stainless cast steel with a basic composition of Cr: 18, Cu: 0.5, Nb: 0.3–1.2, (C + N): 0.07 (mass%) was investigated, and the following conclusions were obtained.
