Introduction

The diameter of steel cords can be less than 0.2 mm after hot rolling and cold drawing from billets. The breakage of steel during cold drawing and fabrication is mainly caused by nonmetallic inclusions, and is a crucial issue. The effects of inclusions on steel properties are different owing to their different plasticities. Nondeformable inclusions are extremely detrimental to the drawing performance of tire cord steels and can cause fracture[1]; therefore, deformable inclusions are required for such steels.

The control of inclusions in tire cord steels has been reviewed and investigated.[2,3,4,5,6] Considering the deformation behavior of inclusions during hot rolling and cold rolling, several experimental investigations[7,8,9,10] and finite-element calculations[11,12] were reported. There have not been sufficient investigations so far on the deformation of inclusions in tire cord steels, and the factors influencing their deformability are not clearly known yet. The inclusions in tire cord steels can be roughly classified into two categories[13,14]: the SiO2-MnO-Al2O3 system and the SiO2-CaO-Al2O3 system. To achieve good plasticity, the composition of inclusions should be controlled by the regions having low melting points,[5,15] namely, the region around spessartine in the SiO2-MnO-Al2O3 system and that surrounded by anorthite-tridymite-pseudo wollastonite-gehlenite in the SiO2-CaO-Al2O3 system. The melting point is the temperature at which a substance changes from the solid state to the liquid state; this depends on the type of crystals and the associated forces, which are related to the strengths of covalent or ionic bonds. The inclusions exhibit plasticity at high temperatures owing to their softened state even though they are not liquids. Cold drawing of tire cord wires is usually carried out at room temperature, which is significantly lower than the solidus temperature of inclusions. Deformability will no longer have a direct relationship with the melting temperature. It is insufficient to control the inclusion composition only in regions with low melting points if the goal is to improve the deformation of inclusions.

Some researchers have tried to identify other parameters that influence the plasticity of inclusions. Bernard et al.[16] studied the relationship between deformability and viscosity of the SiO2-MnO-Al2O3 and SiO2-CaO-Al2O3 systems at high temperatures. Three composition fields were defined in the ternary diagrams. Near the corner corresponding to the pure SiO2 phase, inclusions were too viscous to be deformed. If the inclusions recrystallized, they could become undeformable, whereas the inclusions with low viscosity were more deformable.[16] Kimura et al.[17] studied the fracture behaviors of different oxide inclusions including alumina, zirconia, zircon, and silica during hot rolling and cold drawing on a laboratory scale. It was reported that the magnitude of fracture of the oxide inclusions was affected by the compressive strengths of the oxides, and could be predicted from the Young’s moduli and mean atomic volumes of the oxides.[17] However, only a few single oxides were investigated; this cannot adequately represent the complex phenomena that occur during the industrial production of tire cord steels.

The characteristic evolution of oxide inclusions in tire cord steels during hot rolling has been investigated in one of our previous papers.[18] The deformabilities of inclusions in tire cord steels during both the industrial hot-rolling process and cold drawing were evaluated, and the influencing factors were discussed.

Experimental Methods

The two experimental heats of tire cord steel produced by the basic oxygen furnace (BOF)–ladle furnace (LF)–continuous casting (CC) route in a previous study[18] were considered in the current study. The experimental conditions including refining slag, steel composition, casting and rolling parameters, and the inclusion characteristics can be found in Reference 18.

Billets of dimension 150 × 150 mm2 for Heat 1 and 240 × 180 mm2 for Heat 2 were hot rolled into wire rods using a four-pass rolling process, and then cold drawn into wires of diameter 0.2 mm. The billets were reheated to 1293 K (1020 °C) before rolling, and the temperature of the rods during rolling was controlled between 1173 K and 1213 K (900 °C and 940 °C). During the hot-rolling process, the reduction ratios (RR) (defined by Eq. [1]) of rolled rods were, respectively, 5.4, 20.3, 88.4, and 947.0 for Heat 1; and 10.3, 50.5, 169.7, and 1818.3 for Heat 2.

$$ \eta = \frac{{A_{\text{billet}} }}{{A_{\text{rod}} }}, $$
(1)

where \( \eta \) is the RR, \( A_{\text{billet}} \) is the cross-sectional area of the billet (mm2), and \( A_{\text{rod}} \) is the cross-sectional area of the rolled rod (mm2).

Steel samples were taken from the billets and the hot-rolled rods, and examined after grinding and polishing using an automatic scanning electron microscope (SEM) coupled with energy dispersive X-ray spectroscopy (EDS), named Aspex, to analyze inclusion behavior during hot rolling. The Aspex apparatus has been described in Reference 19. During Aspex detection, 16 uniformly distributed chords that started from the geometric center and moved to the edge of the inclusion were used to measure the minimum, maximum, and average diameters, as well as the areas of the inclusions. The aspect ratio of an inclusion is defined as the ratio of the maximum and minimum diameters of the inclusion. The accelerating voltage of Aspex was set at 20 kV, and a magnification of 250 was used during scanning. The accuracy of detection decreased as the size of the inclusions decreased; therefore, only inclusions with maximum diameters exceeding 3 μm were detected in order to improve the speed and accuracy of detection.

The detection positions of billets and rods after first and second rolling were along the longitudinal sections and located at a quarter of the thickness or diameter, while those of wire rods after the third and fourth hot-rolling passes were along the longitudinal sections and located at the centers of the cross sections. The polishing planes were parallel to the rolling direction. For a better comparison between the billet and rolled rods, all the inclusion diameters used were the equivalent circle diameters (D e), defined as the diameter of a circle having the same area as the inclusion.

Only oxide inclusions—defined as those in which the mole fraction of MnS was less than 10 pct—were taken into account in the current study. In Heat 1 refined by slag with a basicity of 2.1, the inclusions in the billet were usually located in regions having low melting points in the SiO2-CaO-Al2O3 and SiO2-MnO-Al2O3 systems. In Heat 2 refined by slag with a basicity of 0.8, the oxide inclusions in the billet were divided into several types, including the SiO2-CaO-Al2O3-type oxides having low melting points and the high melting point SiO2-MnO-Al2O3-type inclusions (having Al2O3 mass fraction less than 10 pct and high concentrations of SiO2).

The cold-drawn wires were too thin for directly detecting the inclusions, and thus only those on the fractured cross sections of wires were analyzed using an SEM equipped with EDS. Due to the lower fracture ratio of wires cold drawn from Heat 2, only the fractured wires produced during the drawing process of Heat 1 were analyzed.

Oxide Inclusions in Steel

Oxide Inclusions During Hot Rolling

There was no difference in the composition distribution of inclusions in the first hot-rolled rod relative to that in the billet.[18] The morphologies of large oxide inclusions after first hot rolling are shown in Figure 1. A few large inclusions in the Al2O3-SiO2-CaO system were elongated along the rolling direction during the hot-rolling process. Most of the elongated inclusions were not fractured during first rolling owing to the small rolling reduction. The second image in Figure 1(b) shows the morphology of a nonuniform inclusion in Heat 2 after hot rolling. The deformation of the SiO2-MnO phase in zone 1 (shown in the figure) was good, while the SiO2 phase in zone 2 was nondeformable. It was estimated from the third inclusion in Figure 1(b) that the deformability of SiO2-type inclusions was lower during hot rolling.

Fig. 1
figure 1

Morphologies of large inclusions in the first-rolled rods of (a) Heat 1 and (b) Heat 2

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Different RRs were observed in the rod obtained from the billet after each rolling due to the difference in billet dimensions between the two heats. The RR of the billet was assumed to be 1. Figure 2 shows the number density, defined as the number of detected oxide inclusions per unit scanned area under different RRs for the hot-rolling process. The number density of oxide inclusions was lower in hot-rolled rods than that in the billet. There were two possible reasons for this phenomenon: (i) only inclusions with diameters exceeding 3 μm were considered, and the inclusions crushed to particles smaller than 3 μm could not be detected; (ii) more MnS might have precipitated on the surfaces of oxides during hot rolling, which would also cause a decrease in the number of oxide inclusions, since oxide inclusions were defined as those with mole fractions of MnS smaller than 10 pct. In the first three rolling passes, the number density of oxide inclusions in Heat 1 was small, suggesting that the inclusions in Heat 1 were easily crushed to small particles.

Fig. 2
figure 2

Number densities of oxide inclusions under different reduction ratios (RRs)

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The variations of mean D e of oxide inclusions under different RRs during hot rolling are shown in Figure 3. The error bars in this figure and the following figures represent 95 pct confidence intervals of the mean values. The mean D es of oxide inclusions in the two heats were similar and did not change when the RR was smaller than 100; however the values decreased when the RR became higher than 1000. The mean value for Heat 1 decreased to 1.3 μm. Larger fluctuations were observed for the D e of inclusions in Heat 1, which was due to the lower amount, uneven elongation, and crushing of inclusions in Heat 1.

Fig. 3
figure 3

Mean equivalent circular diameters of oxide inclusions under different RRs

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Inclusions Causing Fracture of Wires During Cold Drawing

The fracture ratios of wires after Heat 1 caused by inclusions during cold drawing were 60 to 80 km once, while that of Heat 2 was less than 120 km once. Due to the low fracture ratio of Heat 2, only the inclusions present at the fracture surfaces of wires from Heat 1 were collected, and have been summarized in Table I. The solidus and liquidus temperatures were calculated using FactsageTM 7.0, a commercial software used for thermodynamic calculations. The morphologies of two of the inclusions are shown in Figure 4. Most of the inclusions causing fracture were located near the centers of wires and had sizes ranging from 9.7 to 25.3 μm, which were significantly larger than the average size of 2.8 μm for the inclusions in the slab.

Table I Compositions of the Inclusions Causing the Fracture of Wires During Cold Drawing
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Fig. 4
figure 4

Inclusions on the cross sections of fractured cold-drawn wire surfaces

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A normalized composition of inclusions is presented in Figure 5, which suggests that some of the inclusion compositions were similar to the average composition of inclusions in the slab. There were some inclusions with high Al2O3 or CaO content, which were considered as exogenous due to, for instance, slag entrainment or refractory erosion. Another reason for inclusions with high Al2O3 content was the limited detection area in the billet. Inclusions with average Al2O3 content of 20 pct were detected on an area of 66 mm2 in the billet of Heat 1. Each inclusion in Table I appeared once, after wire drawing for more than 60 km. The calculated liquidus temperatures of the inclusions ranged from 1654 K to 2397 K (1381 °C to 2124 °C), and the solidus temperatures from 1490 K to 1654 K (1217 °C to 1381 °C), indicating that some inclusions had low melting points.

Fig. 5
figure 5

Compositional distribution of inclusions on the fracture of wires of Heat 1

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Deformability of Oxide Inclusions in Steel

Effect of RR on the Aspect Ratio of Oxide Inclusions After Hot Rolling

Figure 6 shows the aspect ratios of oxide inclusions at different RRs. The aspect ratios were larger in rolled rods than those in the billet owing to the deformation of inclusions. The aspect ratio decreased with the increasing RR during hot rolling, which was attributed to the fracture of inclusions. The aspect ratio after first rolling was used to characterize the high-temperature deformability of inclusions. The aspect ratio of oxide inclusions of Heat 1 was greater, indicating that the inclusions in Heat 1 had better deformability during hot rolling, which was in turn attributed to the different natures of oxide inclusions.

Fig. 6
figure 6

Aspect ratios of oxide inclusions at different RRs

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Relationship Between High-Temperature Deformability and Melting Point of Oxide Inclusions

The relationship between the aspect ratio and the ratio of oxide inclusions (MgO + Al2O3)/(SiO2 + MnO) after the first hot-rolling process is shown in Figure 7. The aspect ratio increased initially and then decreased with the increasing concentration ratio of (MgO + Al2O3)/(SiO2 + MnO). The aspect ratios of oxide inclusions for concentration ratios increased from 0.35 to 0.55 for Heat 1 and from 0.15 to 0.35 for Heat 2.

Fig. 7
figure 7

Effect of the (MgO + Al2O3)/(SiO2 + MnO) ratio on the aspect ratio of oxide inclusions after first hot rolling

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The relationship between the aspect ratio and melting point of oxide inclusions in the first-rolled rod of Heat 2 is presented in Figure 8. Both the solidus and liquidus temperatures of inclusions calculated by FactsageTM 7.0 were considered. The aspect ratio was inversely proportional to the melting temperature of the inclusion. The temperature during first hot rolling was high (1213 K to 1293 K or 940 °C to 1020 °C), giving rise to softening and facilitating deformation of the oxide inclusions.

Fig. 8
figure 8

Relationship between the aspect ratio and melting point of oxide inclusions in the first-rolled rod of Heat 2

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The cold drawing of wires was carried out at room temperature, which was significantly lower than the melting temperatures of the inclusions. Some of the inclusions causing wire fracture during cold drawing had low melting points. It was insufficient to evaluate the deformability of inclusions during cold drawing or cold rolling based only on their melting points.

Evaluation of Oxide Deformability at Low Temperature

Hardness of a crystal is the ability to resist plastic deformation due to hydrostatic compression, tensile loading, and shear.[20] Vickers hardnesses and melting temperatures of a few solids are listed in Table II, and drawn together in Figure 9, which shows that Vickers hardness increases with the increasing melting temperatures of the solids.

Table II Vickers Hardness Values and Melting Temperatures of Solids
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Fig. 9
figure 9

Relationship between Vickers hardness and melting temperature of the solids listed in Table II

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Hardness is a macroscopic concept that is governed by intrinsic (bond strength, cohesive energy, and crystal structure) and extrinsic properties, i.e., defects, stress fields, and morphology. The hardness values of most inclusion compounds cannot be easily and precisely measured. Several categories of microscopic models for hardness evaluation have been proposed[21,22,23,24,25,26] based on first-principle calculations.[22,24,25,26] The calculation process is complex, and there are restrictions on the compositions and kinds of materials. Based on Ashby, the strength of a ceramic is proportional to the elastic modulus, and is generally limited by fracture rather than by plastic flow.[27] Jiang et al. found from statistical analysis that Young’s modulus exhibited good linear correlation with Vickers hardness at low temperature.[28]

Young’s modulus refers to the ratio of stress to strain in the elastic deformation range of a material. It is a physical parameter characterizing the bonding force between atoms in a crystal. In the process of inclusion deformation, elastic deformation occurs first followed by plastic deformation when the applied stress exceeds the elastic limit. The deformation of oxides is related to both their Young’s moduli and hardness values. If a solid is found nondeformable after rolling, it is assumed that elastic deformation has occured during the rolling process. Young’s modulus has been used as a substitute for hardness to evaluate the low-temperature deformabilities of inclusions based on the linear relationship between Young’s modulus and hardness.

According to Ashizuka,[29] Young’s modulus and the mean atomic volume of oxides have the following relationship:

$$ E = kV^{g} , $$
(2)

where E is the Young’s modulus (GPa), V is the mean atomic volume (10−6 m3/mol), which can be calculated using Eq. [3],[17] and k and g are coefficients.

$$ V = \frac{M}{\rho \cdot n}, $$
(3)

where M is the molecular weight (kg/mol), ρ is the density (kg/m3), and n is the number of atoms that compose the oxide.

The relationship between lgE and lgV for inclusions of the MgO-Al2O3-SiO2-CaO system is shown in Figure 10, where the Young’s modulus results at room temperature were measured or summarized by others,[17,29,30,31] and the mean atomic volumes were calculated based on Eq. [3]. From the fitting analysis, Young’s modulus and mean atomic volume have the following relationship at low temperature:

Fig. 10
figure 10

Relationship between Young’s modulus and mean atomic volume for the oxides in the MgO-Al2O3-SiO2-CaO system

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$$ E = 39811V^{ - 2.9} $$
(4)

The value of the coefficient of g was similar to Ashizuka’s value of − 2.8.[29] Using Eq. [4], the low-temperature Young’s modulus of the oxide inclusions was calculated.

The relationship between Young’s modulus and liquidus temperature of the oxide inclusions for the two heats are shown in Figure 11. The Young’s modulus did not change until the liquidus temperature reached ~ 1700 K (1427 °C), and then it decreased with the increasing liquidus temperature.

Fig. 11
figure 11

Relationship between Young’s modulus and liquidus temperature of oxide inclusions from (a) Heat 1 and (b) Heat 2

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The relationship between the aspect ratio of oxide inclusions after first hot rolling and the low-temperature Young’s modulus is shown in Figure 12. There is no relationship between high-temperature deformation and low-temperature Young’s modulus of the inclusions in Heat 1 owing to the limited detection of such inclusions. For Heat 2, the aspect ratio of inclusions changed slightly with the increasing Young’s modulus up to 90 GPa and then increased. Due to the small number of inclusions with high Young’s modulus, the confidence interval increased as well. Deformation was inversely proportional to Young’s modulus, but the observed trends were different. The low-temperature Young’s modulus could not be used to predict deformations at high temperatures. One possible reason was the temperature dependence of Young’s modulus. Wachtman et al. reported that the Young’s modulus of some simple oxides would decrease with the increasing temperature owing to internal friction and grain-boundary slipping,[32] as shown in Figure 13. Density decreased with the increasing temperature, resulting in the increase of mean atomic volume according to Eq. [3], and leading to reduction in Young’s modulus based on Eq. [4].

Fig. 12
figure 12

Relationship between the aspect ratio after first hot rolling and the low-temperature Young’s moduli of oxide inclusions

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Fig. 13
figure 13

Temperature dependence of Young’s modulus

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Young’s moduli of the inclusions causing wire fracture were also calculated and are listed in Table I. All values were higher than 120 GPa (some were as high as 280 GPa), which was significantly larger than the average Young’s modulus (105 GPa) of inclusions in the slab. Inclusions with high Young’s moduli had higher strength, and were difficult to deform or break during cold working.

The low-temperature Young’s moduli of the oxides—for both the Al2O3-SiO2-CaO and the Al2O3-SiO2-MnO systems—were predicted, and are shown in Figure 14. On the basis of the hypothesis that the deformability of oxides during cold drawing was inversely proportional to the Young’s modulus, in order to reduce the fracture of tire cord steel wires during cold drawing, it is proposed that inclusions be controlled to the dark blue regions where the SiO2 content was high and the Al2O3 content was extremely low. Experimental evidence is required, and will be carried out in future studies. The figure indicates that Al2O3 was detrimental to the deformabilities of oxides in tire cord steels during cold drawing due to its high Young’s modulus. This is why Al is strictly prohibited in the steelmaking process.

Fig. 14
figure 14

Young’s modulus distribution of (a) Al2O3-SiO2-CaO and (b) Al2O3-SiO2-MnO systems

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As reported previously,[18] most inclusions in the billet of Heat 2 were located in the region of small Young’s modulus in spite of having high melting points. This explains the low fracture ratio during cold drawing of wires from Heat 2.

Future Study

Due to differences in composition and structure, the changes in Young’s modulus of complex oxides with temperature are expected to be different from those of simple oxides. There is a lack of data in the literature for high-temperature Young’s moduli of complex oxide inclusions, and the changes in their Young’s moduli with temperature are not clearly described. The determination of the relationship between Young’s modulus of complex oxides and temperature forms the scope of a future study.

The relationship between the deformation of oxide inclusions during hot rolling and low-temperature Young’s modulus of inclusions was different from what was expected. The temperature dependence of Young’s modulus and the different deformation mechanisms at different temperatures are assumed to be the possible reasons. The underlying reason should be studied further.

The deformation of inclusions was only experimentally investigated during the hot-rolling process. No experimental confirmation has been established for the relationship between deformability and low-temperature Young’s modulus of inclusions. The deformability of inclusions during the cold rolling or cold drawing processes of steel should be experimentally studied.

Conclusions

The behavior of oxide inclusions in tire cord steels during hot rolling was analyzed, and the factors influencing the deformability of inclusions at high and low temperatures were discussed. The main conclusions are summarized below.

  1. 1.

    With the increasing RR during hot rolling, the mean D e of the oxide inclusions changed slightly for RR less than 100, but decreased for RR larger than 1000. The aspect ratio of the oxide inclusions decreased, probably because of fracture. The aspect ratio after the first hot-rolling process was used to characterize the high-temperature deformability of inclusions.

  2. 2.

    The aspect ratio of inclusions increased initially and then decreased with the increasing (MgO + Al2O3)/(SiO2 + MnO) ratio. The deformation of oxide inclusions at high temperatures increased with the decreasing melting temperature.

  3. 3.

    Young’s modulus was used to evaluate the low-temperature deformability of inclusions. An empirical formula to calculate the Young’s moduli of oxides using the mean atomic volume was fitted: \( E = 39811V^{ - 2.9} \).

  4. 4.

    Young’s moduli of the inclusions causing wire fracture were significantly greater than the average modulus value. To reduce fractures in tire cord steel wires during cold drawing, it is proposed that inclusions be controlled to those with high SiO2 content and extremely low Al2O3 content. Such a control was possible because the deformability of oxides during cold drawing was inversely proportional to their Young’s moduli. Further experimental confirmation is required in this regard.