Free Access
Issue
Metall. Res. Technol.
Volume 116, Number 6, 2019
Article Number 623
Number of page(s) 8
DOI https://doi.org/10.1051/metal/2019052
Published online 11 October 2019

© EDP Sciences, 2019

1 Introduction

Quality defects in steel products often originate from the non-metallic inclusions therein, which causes quality problems in the steel production process, for example, slabs will be accompanied by mechanical impact during lifting and placing of crane, and fracture accident occurs due to the existence of internal micro-cracks in the slab, thus resulting in severe economic losses to the steel companies. One of the important factors related to the quality of the slab is the non-metallic inclusions which embedded in slabs. To this end, control of non-metallic inclusions in steel has received growing attention in recent years, especially in the production of high quality slabs for the Advanced High Strength Steel (AHSS). With a good plasticity, MnS in steel can lower the stress concentration between oxides and matrix by wrapping hard oxides in steel to form composite inclusions, which prevents the crack initiation and propagation on oxides effectively. This in turn can enhance the mechanical properties of steel products including ingots, especially the impact toughness.

Al2O3, a common high melting oxide, is quite harmful to the steel performance. If an appropriate process is employed to control the precipitation of MnS on the Al2O3 surface during the solidification of molten steel, an MnS + Al2O3 composite inclusion wrapped by MnS can be obtained. The encapsulation degree of such composite oxides is significantly important for improving the performance of steel products. Previous studies [1,2] have also explored the behavior of Al2O3 inclusion particles in Al-deoxidized titanium alloyed steel during the solidification process, as well as its effect on the production of MnS. Oikawa et al. [2] found that the Al2O3 inclusion in free-cutting steel can be used as heterogeneous nucleation core for the MnS formation. According to a study on the effect of Al2O3 inclusion on the precipitation of MnS during Fe-10% Ni alloy solidification by Ohta and Suito [3,4], the molten steel pushed the Al2O3 inclusions more obviously in the solidification process at an S mass fraction of around 0.01%. Besides, MnS + Al2O3 composite inclusion would be produced when Al2O3 was the heterogeneous nucleation core for MnS. Ma and Cui [5] demonstrated that the encapsulation of soft MnS outside oxides could improve the fatigue properties of steel. The beneficial trait of composite ductile inclusions is that after the sulfide is wrapped on the oxide surface, the stress concentration between the oxide and the matrix is lowered. This prevents the crack initiation and propagation on the oxide, thereby enhancing the fatigue endurance and fracture toughness [68].

Existing researches on the formation rate of MnS + Al2O3 inclusion focus on the aspects of S content, Mn content, Al content and Al2O3 size. However, few have studied the non-negligible effects of molten steel cooling rate during solidification on the precipitation of MnS and the formation of composite inclusions. This study designs a set of water-cooling copper-chill mold test equipment to investigate the effect of cooling rate on the precipitation of MnS + Al2O3 composite inclusion in steel by applying different cooling conditions to the ingot cooling after the rapid cooling of test steel in the early solidification stage. Meanwhile, Thermo-Calc is used to calculate the precipitation trends of MnS and Al2O3 in the high temperature phase. By testing the impact toughness of ingots obtained in different cooling conditions, we study how the composition and morphology of inclusions in the ingots affect the mechanical properties of test steel, especially the fracture toughness.

2 Experimental section

2.1 Experimental equipment

A water-cooling copper mold with controllable cooling rates is designed specifically for experimentation herein, which allows the test steel to solidify at different cooling rates. Figure 1 displays the schematic of the test equipment. In the water-cooling copper mold system, the water-cooling wall is a composite plate comprising a copper plate and a steel plate. While the copper plate is in direct contact with the molten steel on the inner surface, the steel plate is on the outer surface. There is a water channel for circulating cooling water within the copper plate. To measure the heat exchange capacity of copper plate during the experiment, five K-type thermocouples are distributed at different positions inside the plate. Meanwhile, an R-type thermocouple is installed on top of the copper mold for measuring the ingot cooling process during solidification, i.e. the temporal variation of ingot core temperature. The No. 3 K-type thermocouples among the five K-type thermocouples distributed on the mold copper plate is used to measure the temperature change on ingot surface during the test. All the temperature data over the entire test period are converted via a temperature acquisition module and eventually displayed on the connected computer.

The steel studied in this experiment is the Fe–C–Mn–Si advanced high-strength steel for automobiles, which has a Si mass fraction of 2%, and an Mn mass fraction of 2%. A 50 kg vacuum induction melting furnace is used to melt the steel. Industrially pure iron (purity 99.8%) is placed in the magnesia crucible. After melting of steel in an air atmosphere, it is subjected to measure the oxygen content before deoxidation. Subsequently, vacuum chamber is evacuated and then filled with high-purity Ar. This procedure is repeated three times to attain the desired vacuum degree (< 50 Pa). Afterwards, deoxidation and alloying of molten steel are performed with Al particle deoxidizer (purity 99%) and alloy materials of graphite, electrolytic manganese, ferrosilicon and FeS. The smelted steel is then poured into the baked water-cooling copper-chill mold, and cooled according to the established scheme. Finally, the samples are processed and collected.

thumbnail Fig. 1

Schematic diagram of experimental system.

2.2 Testing procedures and research methods

The ingot cooling rate of the designed copper-chill mold is controlled by adjusting the water flow rate and pressure. Tests are conducted at the same cooling water pressure and flow rate. Figure 2 presents the water pressure and flow rate used for testing ingots No. 1 and No. 2. As is clear, the water pressure of copper mold’s inlet circuit is approximately 32 psi, where C1 and C2 represent the cooling water circuits on the two sides within copper plate, respectively. Meanwhile, the flow rate of cooling water is approximately 27 GPM. The flow rate and pressure of cooling water used in this study are determined based on the actual cooling water consumed by the mold in the continuous casting production of steelworks. During the experiment, the cooling rate of ingots is presented by monitoring the temperature readings displayed on the computer by the R-type thermocouple that is inserted into the ingot core. When the R-type thermocouple shows a temperature of 1400 °C (solidus temperature of test steel = 1434 °C), the water-cooling copper mold is opened to take out the ingots. Ingot 1 is exposed directly in the air for air cooling, while ingot 2 is placed in a container with thermal insulation sand that completely submerges the ingot. In Figure 3, the variation trend of ingot core temperature (1000 s) over the experimental period is illustrated. During the initial 248 s, the ingots are cooled in the copper mold (MC). Then, the copper mold is opened, and the ingot 1 is directly cooled in the air (AC) to room temperature, while the ingot 2 is cooled in a thermal insulation sand (SC) to room temperature.

After cooling of the two test ingots (ingots 1 and 2) to room temperature, eight impact samples are cut separately at the quarter and center positions of ingot thickness direction, which are processed into standard Charpy specimens 10 × 10 × 55 mm-V (2 mm) in size. These processed specimens are then subjected to impact test on a Charpy impact tester separately at a room temperature (22 °C) and −16 °C. Afterwards, the tested specimens are polished stepwise with waterproof sandpaper, followed by the morphology, size observation and composition analysis of non-metallic inclusions in the steel specimens under a ZEISS-Supra 55 scanning electron microscope.

thumbnail Fig. 2

Pressure and flow rate of cooling water during the test.

thumbnail Fig. 3

Temperature variation of ingot core during continuous cooling.

3 Results and discussion

3.1 Quantity and size of inclusions

For sampling analysis of the ingots in two different cooling regimes, samples are taken at the quarter and center positions along the ingot thickness direction. Then, the samples are subjected to metallographic grinding and polishing. After preparation of specimens, the inclusions therein are counted for number and analyzed for size. Ten identically sized areas (approximately 1.8 mm2) are selected for each specimen at various positions and averaged ultimately. Figure 4 below details the statistical results. Where ingot 1 represents the ingot cooled by air cooling; and ingot 2 represents the ingot cooled by sand cooling. It can be seen in the figure that the total quantity of inclusions at the center of ingot 1 is 44.7% higher than that at the ingot 2 center, which are dominated by small particle inclusions (less than 4 µm in size). The total quantity of inclusions at the quarter position of ingot 2 is 70.2% higher than that at the ingot 1 center, and the inclusions sized larger than 4 µm predominate the quarter position of the ingot 2. At the center position of ingot, inclusions sized between 1 and 4 µm are highest in quantity, while the large-sized inclusions (8 ∼ 20 µm) are minimum in number. Distribution of total inclusion quantity of ingot 2 in the thickness direction is more uniform than the ingot 1 in the thickness direction. In other words, the inclusions at the center and quarter positions of ingot 2 are almost quantitatively equal, while for ingot 1, the number of inclusions at the core is 81.3% greater than that at the quarter position.

thumbnail Fig. 4

Statistical results of inclusions number and size in the sample.

3.2 Morphology and composition of inclusions

Morphological observation and composition analysis of the inclusions in specimens are performed by Scanning Electron Microscopy (SEM) and Energy Dispersive Spectroscopy (EDS). In Figure 5, the morphologies of typical inclusions observed from the present test samples are displayed. Figure 5a depicts the Al2O3 inclusions with irregular shapes, whereas Figure 5b depicts the MnS inclusions in the test steel that are in an irregular spindle shape. Composite inclusions with MnS wrapping Al2O3 are also observable from the electron microscope, as shown in Figure 5c. While the encapsulated inclusions are either elliptical, square or round, the centrally wrapped Al2O3 inclusions are generally spherical. In Figure 6, the SEM-EDS analysis results of typical inclusions in the specimens are presented. From the EDS spectra, the distribution of components in the inclusions is clearly visible, according to which the type and composition of the inclusions can be identified.

A 30 mm2 area is selected each at the quarter and core positions of ingots 1 and 2 in the thickness direction for the composition and type analysis of non-metallic inclusions. The quantitative density of non-metallic inclusions in the test steel is defined as: (1)

Where, NA is the quantitative density of inclusions; N is the total number of inclusions; A is the area of the tested sample, mm2.

According to the results of inclusion type analysis, the inclusions occupying the largest proportion in ingots 1 and 2 are MnS; followed by the MnS + Al2O3 composite inclusions; while the oxide inclusions, such as Al2O3 and SiO2, come in third place. The proportion of the first two inclusion types exceeds 80%. Figure 7 illustrates the relationship between number density and size for various types of inclusions at the quarter and core positions of specimens (ingots 1 and 2) in the thickness direction. From the perspective of inclusion types, for ingot 1, all types of inclusions are quantitatively larger at the core position than at the quarter, as is clear in Figures 7a and 7b, especially the number of MnS inclusions. As for ingot 2, it can be seen in Figures 7c and 7d that the MnS + Al2O3 composite inclusions are markedly more in number than the oxide inclusions, especially at the ingot core position. Contrastively, in ingot 1, oxide inclusions are more in number than the MnS + Al2O3 inclusions.

thumbnail Fig. 5

Morphology of typical inclusions in the test ingots. a: oxides; b: sulfides; c: composite inclusions.

thumbnail Fig. 6

SEM-Mapping result of inclusions in the test ingots.

thumbnail Fig. 7

Statistics of inclusions quantitative density and size in the test ingots. a: quarter point of ingot 1; b: center of ingot 1; c: quarter point of ingot 2; d: center of ingot 2.

3.3 Thermodynamic calculation of inclusions precipitation

Thermo-Calc software is used to calculate the formation trends of MnS and Al2O3 inclusions with a system capacity of 1 mol in our test steel that has a matrix composition of Fe-0.19wt%C-2.0wt%Mn-2.0wt%Si. In Figure 8, the influence of temperature changes on the precipitation amounts of MnS and Al2O3 is illustrated in the test steel containing 0.003 wt% S. As can be seen in the figure, the binding force between Al and O elements is quite large, and Al2O3 is already precipitated from the molten steel at temperatures higher than 1600 °C. Meanwhile, the calculations demonstrate that MnS begins to be produced when the temperature is between 1100–1405 °C. At this time, if the diameter of precipitated Al2O3 satisfies the nucleation conditions of MnS + Al2O3 inclusions, MnS can nucleate and grow centering around Al2O3, thereby forming wrapped MnS + Al2O3 composite inclusions.

Nakae and Wu [9] proposed a wetting interface energy model to calculate the trapping/pushing behavior of Al2O3 particles, in order explain the formation mechanism of MnS + Al2O3 composite inclusions in steel. Their results showed that the contact angle of Al2O3 particles at the solid–liquid interface is 137°, and that the Al2O3 particles could be pushed to the liquid phase region during solidification. When the Al2O3 particles in molten steel were pushed quickly to the front of solidification liquid phase, the Mn and S elements were enriched in the liquid phase front during the molten steel solidification. MnS inclusions began to precipitate when the actual concentration product of [Mn][S] was greater than the dissolved equilibrium concentration product [10]. At this point, Al2O3 was pushed to the liquid solidification front by the action of liquid phase, which could exactly be used as the nucleation core of MnS to reduce the nucleation energy of its precipitation, thereby producing MnS + Al2O3 composite inclusions. A theoretical calculation study on the critical capturing/pushing velocity of Al2O3 particles at the solid–liquid interface and the radius of particles by Stefanescu et al. [11] found that the Al2O3 particles were pushed to the solidification front region when the pushing rate was greater than the solidification rate of molten steel, which was conducive to forming the MnS + Al2O3 inclusions. However, their research was confined to the theoretical calculation of Al2O3 particle capture by MnS at the solidification phase interface only, which failed to investigate the effect of cooling rate on the MnS + Al2O3 inclusion formation after pushing of Al2O3 to the solidification front. According to the present experimental data, slow cooling is helpful for the nucleation and growth of MnS on the Al2O3 particles. This also well explains why the ingot 2 in Figure 7 yields more MnS + Al2O3 composite inclusions than the ingot 1 when cooled slowly after 1400 °C. This phenomenon is especially prominent at the ingot core position.

thumbnail Fig. 8

Calculation of precipitation temperature of MnS and Al2O3 in Fe-0.19C-2.0Mn-2.0Si Steel.

3.4 Analysis of mechanical properties

In Table 1, the impact toughness test results of processed specimens on a Charpy impact tester are detailed at room temperature and −16 °C. According to the test results, the impact energy of ingot 2 is greater than the ingot 1 at both the quarter and center positions, especially at the center position. At room temperature, ingot 2 exhibits impact energies 15% and 25% greater than ingot 1 at the quarter and center positions, respectively. At a low temperature (−16 °C), the impact energy of ingot 2 is 21% and 28% greater than ingot 1 at the quarter and center positions, respectively.

From the test results, it is clear that the inclusion composition and type of specimens are somewhat influential to the steel mechanical properties. This necessitates the study on the effects of single Al2O3 and composite inclusions on the mechanical properties of the test steel. Given the differences in elastic modulus and linear expansion coefficient between the inclusions and the matrix, tessellated stress was generated between the two during cooling process, which was actually one cause of the stress concentration and microcrack initiation in the material matrix [12]. Raghupathy et al. [12] and Brooksbank and Andrew [13] studied in detail the stresses between inclusions and matrix, who put forward a model for computing the radial stress of inclusions in steel. By assuming the inclusions in steel to be ideally spherical, a computational model shown in Figure 9 is built. The formula for calculating the radial tessellated stress of inclusions in steel is obtained as formula (2) [13]. (2)

where, αm is linear expansion coefficient of matrix and α is linear expansion coefficient of composite inclusion; νm is Poisson’s ratio of matrix and ν is Poisson’s ratio of composite inclusion; Em is Young’s modulus of matrix and E is Young’s modulus of composite inclusion, is temperature difference between room temperature and temperature when pearlite transformation is completed, and d is volume fraction of inclusions in matrix. The parameters used in the calculation are shown in Table 2 [13,14].

For the MnS + Al2O3 composite inclusions, the Young’s modulus and Poisson’s ratio used in the calculation are weighted averages based on the volume fractions of MnS and Al2O3. The thermal expansion coefficient of the composite inclusions is , where αMnS denotes the thermal expansion coefficient of MnS; and is the thermal expansion coefficient of Al2O3. When R1 = R2, there is only one inclusion type, i.e. Al2O3 inclusions, and the model is calculated according to Figure 9b. The tessellated stress of composite inclusions in steel and the size relationship between particles Al2O3 and MnS are derived based on the calculation model in equation (2), as shown in Figure 10. It is clear in the figure on the right that when R2 / R1 > 3, the tessellated stress of composite inclusions in steel approaches 0 MPa, suggesting that the smaller the size of Al2O3 particles in the composite inclusions, the less the tessellated stress of composite inclusions in steel. When R2 / R1 < 2, the tessellated stress of composite inclusions begins to increase rapidly. And when R2 / R1 = 1, the inclusions are single Al2O3 particles, at which point the inclusion tessellated stress in steel reaches a maximum of 1182 MPa. This shows that the single Al2O3 particles have a considerably greater tessellated stress than the composite inclusions in steel.

SEM observation of ingot 1 formed in this experiment finds the occurrence of crack sources around Al2O3 inclusions, while no crack source was noted around the MnS + Al2O3 composite inclusions, as shown in Figure 11. For this phenomenon, it can be explained that, during the external force loading of specimens, the tessellated stress field around the Al2O3 inclusions overlaps with the specimens’ principal stress field, thus reducing the diffusion resistance of crack front around the inclusions. On the other hand, according to the foregoing calculations, the dislocation slip of matrix material causes vacancy accumulation around the inclusions after the tessellated stress exceeds the yield strength, which promotes the formation of cleavage fracture cracks [15,16]. Thus, the Al2O3 inclusions easily become crack initiation points in the crack propagation process. Many researchers have also obtained similar conclusions by using the the theoretical analysis of Brooksbank for their subjects, such as Turkdogan [17] and Pan et al. [15]. The calculation demonstrates that compared with the single Al2O3 inclusions in the test steel, the stress concentration around the MnS + Al2O3 composite inclusions decreases obviously. The impact test results also verify that the higher the content of composite inclusions, the greater the impact energy and the better the fracture toughness of the test steel. Hence, to obtain a test steel cast structure with excellent mechanical properties, MnS + Al2O3 inclusions should be formed therein to replace the single Al2O3 inclusions.

Table 1

Impact energy of the test steel at different temperatures.

thumbnail Fig. 9

Scheme used in calculation of Schematic diagram of tessellated stress between inclusion and matrix. a: Al2O3 + MnS composite inclusion; b: Al2O3 inclusion.

Table 2

Thermal expansion coefficient and Young’s modulus of inclusions and matrix.

thumbnail Fig. 10

Relationship between the tessellated stress and inclusion size.

thumbnail Fig. 11

Photograph of microcracks around Al2O3 inclusion in ingot 1.

4 Conclusions

This research work is a basic laboratory test item, the purpose of current study is to find out a method to improve the quality of the test steel slab, i.e. to control the internal cracking of the slab. Studies in this paper have shown that the fracture toughness of the test steel can be improved by slow cooling in the high temperature section. This research work is extremely important for the quality control of the test steel slab in our future study. The main conclusions are summarized as follows:

  • water-cooling copper-chill mold is used to solidify the molten steel quickly in the early stage of solidification. The experimental results show that slow cooling treatment near the solidus temperature of test steel facilitates the nucleation and growth of MnS on the Al2O3 particles. This in turn helps form more MnS + Al2O3 composite inclusions;

  • MnS + Al2O3 inclusions from ingot 2 can reduce the inlay stress between inclusions and matrix, and improve the impact toughness of test steel, especially the low temperature toughness. At −16 °C, the impact energy of ingot 2 at the center position is 28% higher than the corresponding value of ingot 1;

  • calculations reveal that the tessellated stress of composite inclusions in test steel approaches 0 MPa when Al2O3 around by the encapsulating MnS, at which point the inclusions exhibit good fracture toughness. When the inclusion is single Al2O3, the tessellated stress in the test steel reaches 1182 MPa. After the tessellated stress exceeds the yield strength, the dislocation slip of matrix material leads to vacancy accumulation around the inclusions. As this promotes the formation of cleavage fracture cracks, the mechanical properties of test steel are degraded considerably.

Acknowledgments

The authors gratefully acknowledge support provided by China Scholarship Council, Grant number 201606080085.

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Cite this article as: Zhanfang Wu, Zhenyu Liu, Shengtao Qiu, Xiangyang Li, Effect of composition and morphology of non-metallic inclusions on fracture toughness in as-cast AHSS, Metall. Res. Technol. 116, 623 (2019)

All Tables

Table 1

Impact energy of the test steel at different temperatures.

Table 2

Thermal expansion coefficient and Young’s modulus of inclusions and matrix.

All Figures

thumbnail Fig. 1

Schematic diagram of experimental system.

In the text
thumbnail Fig. 2

Pressure and flow rate of cooling water during the test.

In the text
thumbnail Fig. 3

Temperature variation of ingot core during continuous cooling.

In the text
thumbnail Fig. 4

Statistical results of inclusions number and size in the sample.

In the text
thumbnail Fig. 5

Morphology of typical inclusions in the test ingots. a: oxides; b: sulfides; c: composite inclusions.

In the text
thumbnail Fig. 6

SEM-Mapping result of inclusions in the test ingots.

In the text
thumbnail Fig. 7

Statistics of inclusions quantitative density and size in the test ingots. a: quarter point of ingot 1; b: center of ingot 1; c: quarter point of ingot 2; d: center of ingot 2.

In the text
thumbnail Fig. 8

Calculation of precipitation temperature of MnS and Al2O3 in Fe-0.19C-2.0Mn-2.0Si Steel.

In the text
thumbnail Fig. 9

Scheme used in calculation of Schematic diagram of tessellated stress between inclusion and matrix. a: Al2O3 + MnS composite inclusion; b: Al2O3 inclusion.

In the text
thumbnail Fig. 10

Relationship between the tessellated stress and inclusion size.

In the text
thumbnail Fig. 11

Photograph of microcracks around Al2O3 inclusion in ingot 1.

In the text

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