Strength-Electrical Conductivity Balances of Cu/Martensite Steel Multilayered Sheets with Various Volume Ratios

Abstract

The strength and electrical conductivity balances in Cu/martensite (α′) steel multilayered sheets with various volume ratios were evaluated. Furthermore, the measured tensile properties and electrical conductivity were compared with the values estimated from the rule of mixtures, and the reason for the difference between the measured and estimated values was discussed based on the deformation and fracture behaviors. The multilayered sheets exhibited excellent strength-electrical conductivity balances superior to those of conventional Cu alloys, and their strength and electrical conductivity can be controlled over a wide range by changing the volume fraction of α′ steel. The electrical conductivities of the multilayered sheets with different volume ratio were approximately identical to the estimated values based on the rule of mixtures. However, the ultimate tensile stresses of the multilayered sheets with lower volume fractions of the α′ layer were slightly lower than the estimated values. A significant strain concentration occurred within the α′ steel layer in the multilayered sheet with the lowest volume fraction of α′ steel. Furthermore, cracks and/or voids were formed in the α′ steel layers even during the uniform-deformation stage. Therefore, the early fracture of the α′ steel layers caused lower ultimate tensile stresses in the multilayered sheet with low fractions of the α′ steel than the estimated value based on the rule of mixtures.

Fig. 4 Relationship between σ_UTS and ρ of various Cu alloys,¹³⁾ Cu/Fe multilayered sheets,¹¹⁾ Cu and α′ steel sheets, and present Cu/α′ steel multilayered sheets.

1. Introduction

Cu alloys have been widely adopted as conductive materials owing to their excellent electrical conductivities. Recently, higher electrical conductivity and strength have been simultaneously required for conductive materials because electrical devices are becoming lighter and smaller. However, it is generally known that there is a trade-off relationship between strength and electrical conductivity in Cu alloys.¹⁾ For example, typical precipitation-hardened Cu–Ni–Si, Cu–Be, and Cu–Ti alloys exhibit higher strengths but lower electrical conductivities than pure Cu. As a result, the strength-electrical conductivity balance scarcely improves from that of pure Cu.^2–4) Therefore, various efforts have been made to improve the strength-electrical conductivity balance of the Cu alloys, such as developing new alloy systems and applying severe plastic deformation methods to conventional alloys.^5–7)

Multilayered sheets consisting of dissimilar metals exhibit mechanical or functional properties superior to those of single-layer sheets composed of multilayered sheets.^8–10) Koseki et al. demonstrated that a multilayered steel sheet consisting of Type 420 martensitic steel and Type 304 austenitic steel has a significantly high strength-ductility balance; the product of the ultimate tensile strength and fracture elongation in the multilayered sheets was significantly higher than that in each single-layer sheet.¹¹⁾ Koga et al. reported that pure Cu/pure Fe multilayered sheet with a Cu volume fraction of 60% exhibited excellent temperature dependence of tensile properties; strength increased, similar to the Fe single-phase material, but elongation was maintained, similar to the Cu single-phase material, with decreasing temperature.¹¹⁾ The electrical conductivity of the Cu/Fe multilayered sheet was identical to the value calculated from the volume fraction and electrical conductivity of each layer based on the rule of mixtures. Therefore, it can be said that the electrical conductivity of Cu layers in the multilayered sheet barely deteriorates by stacking with the Fe layers. Furthermore, this strongly suggests that Fe atoms scarcely diffuse into the Cu layer from the adjacent Fe layer during the fabrication of the multilayered sheet or heat treatment processes because the dissolvement of Fe into Cu significantly lowers the electrical conductivity of Cu.¹²⁾ However, the strength-electrical conductivity balance of the Cu/Fe multilayered sheet was lower than that of conventional Cu alloys (see Fig. 4).¹³⁾ This was due to the low strength of pure Fe (130 HV). As the electrical conductivity of the Cu layer scarcely decreased during the fabrication process of the Cu/Fe multilayered sheet, the strength can be improved without decreasing the electrical conductivity by replacing the soft pure Fe layers in the Cu/Fe multilayered sheet with harder carbon steel layers.

In this study, multilayered Cu/martensitic (α′) steel sheets with three different volumes of α′ steel were fabricated, and their strength-electrical conductivity balances were systematically investigated. Furthermore, the experimentally obtained tensile properties and electrical conductivity were compared with the estimates from the rule of mixtures, and the reasons for the difference between the experimental and estimated values were discussed.

2. Experimental Procedure

2.1 Materials

Oxygen-free copper with less than 0.04 mass%O and carbon steel (SCM435: Fe–0.35%C–0.75%Mn–0.30Mo, in mass%) sheets were used as starting materials in this study. The Cu and carbon steel sheets were stacked so that the volume fractions of the carbon steel were 30%, 60% and 80%. Then, the stacked sheets were subjected to hot rolling to bond them. The fabricated multilayered Cu/carbon steel sheets were solution-treated at 1063 K for 0.6 ks and then quenched into cold water. The carbon steel layers in the multilayered sheets transformed into full martensite structure after the quenching. Hereinafter, the Cu/α′ steel multilayered sheets with 80%, 60%, and 30% volume fractions of the α′ steel layer are referred to as 80%α′, 60%α′, and 30%α′ sheets, respectively.

2.2 Microstructural observation

Microstructural observations were conducted mainly along the transverse direction (TD) of the sheets using field-emission scanning electron microscopy (SEM; JEOL JSM-7900F) under an acceleration voltage of 15 kV. Observations were performed in each layer individually because it was difficult to finish the Cu and α′ layers simultaneously. The specimens for the microstructural observation of the Cu layers were prepared by electropolishing after mechanical polishing. Electropolishing was carried out at 5 V and 243 K for 180 s using a 30% Nital solution (methanol:nitric acid = 7:3 (volume ratio)), and SUS304 stainless steel was used as the cathode. The specimen for microstructural observation of the α′ layers was finished using a colloidal silica suspension.

2.3 Tensile test

Specimens for tensile tests with a gauge dimension of 10 mm length, 1.5 mm width and 2 mm thickness were cut so as the loading direction was parallel to the rolling direction (RD). Tensile tests were conducted at room temperature and at an initial strain rate of 10⁻³ s⁻¹ using an Instron-type universal testing machine.

2.4 Measurement of electrical resistivity

Specimen pieces with a length of 10 mm, width of 2 mm and thickness of 1 mm were cut so that the longitudinal direction was parallel to the RD. The electrical resistivity of the multilayered sheets was measured using a direct-current four-terminal sensing method at 293 K. The electrical conductivity (%IACS) was obtained by dividing the measured electrical resistivity by that of pure copper (17.241 nΩm).¹³⁾

2.5 Digital image correlation analysis

Digital image correlation analyses were performed using a subset of 41 pixels and a step of 3 pixels. VIC-2D software (Correlated Solutions Inc.) was used for analysis. The SEM images for DIC analysis were obtained before and after the tensile test.

3. Result

3.1 Microstructure

Figures 1 are SEM images showing Cu layers and α′ steel layers in the (a) and (d) 80%α′, (b) and (e) 60%α′, and (c) and (f) 30%α′ sheets, respectively. Table 1 lists the measured area fraction of the α′ steel layer, the layer thickness, and the Vickers hardness of the Cu and α′ steel layers. The Cu and α′ steel layer thicknesses were varied as listed in Table 1. Variations in the thickness of the layers in a dissimilar material are often observed in multilayered sheets because of the difference in work-hardenabilities between each layer.^14–16) In all sheets, the Cu layers consisted of coarse grains with annealing twins. The grain sizes were approximately 30 µm in the 80%α′ and 60%α′ sheets and 60 µm in the 30%α′ sheet. Since the grain growth is restricted by adjacent dissimilar layers in multilayered sheets,¹⁷⁾ it is reasonable that the 30%α′ sheet with the largest layer thickness has the largest grain size in the three. On the other hand, the Vickers hardness of the Cu layer was approximately identical among the sheets. Therefore, the effect of grain refinement ranging from 60 µm to 30 µm on the Vickers hardness is not significant. In the α′ steel layers, a typical lath martensitic microstructure was observed in all sheets (Figs. 1(d)–(f)). Moreover, the Vickers hardness of the α′ steel layers in all sheets was approximately identical to that of the α′ steel single-layer sheet, which is consistent with a previous study.¹⁸⁾

Fig. 1

SEM images of (a), (b) and (c) Cu and (d), (e) and (f) α′ steel layers in (a), (d) 80%α′, (b), (e) 60%α′ and (c), (f) 30%α′ sheets.

Table 1 Area fraction of α′ steel, average layer thickness and Vickers hardness in each layer of Cu/α′ steel multilayered sheets.

3.2 Tensile properties and electrical conductivity

Figure 2 shows the nominal stress-nominal strain curves of the pure Cu, α′ steel sheets and Cu/α′ steel multilayered sheets with different volume fractions of α′ steel. Table 2 lists the average values of the ultimate tensile stress (σ_UTS), uniform strain (ε_uniform), total strain (ε_total) and electrical conductivity (ρ), obtained from the tensile tests and direct current four-terminal sensing measurements. Calculated strength-elongation (σ_UTS × ε_total) and strength-electrical conductivity (σ_UTS × ρ) balances are also shown in Table 2. The σ_UTS of the Cu/α′ steel multilayered sheets were lower than that of the α′ steel sheet but significantly higher than that of the Cu sheet. σ_UTS decreased with decreasing the volume fraction of α′ steel in the multilayered sheets. Interestingly, ε_total of the multilayered sheets also decreased with decreasing the volume fraction of α′ steel, and the 60% and 30%α′ sheets exhibited a lower ε_total than that of α′ steel sheets, despite the high volume fraction of Cu with high ε_total. As a result, the σ_UTS × ε_total decreased with decreasing the volume fraction of α′ steel, and the 30%α′ sheet exhibited the lowest σ_UTS × ε_total of approximately a tenth of that in the α′ steel sheet. The 80%α′ sheet had an intermediate σ_UTS × ε_total between those of the Cu and α′ steel sheets. Several studies have reported that multilayered sheets have mechanical properties superior to those of single-layer sheets.^8,9) Unfortunately, in the present Cu/α′ steel multilayered sheets, the strength–elongation balance could not be improved by forming a multilayered structure.

Fig. 2

Nominal stress-nominal strain curves in pure copper, α′ steel and Cu/α′ steel multilayered sheets.

Table 2 Measured ultimate tensile stress σ_UTS, uniform strain ε_uniform, total strain ε_total and electrical resistivity ρ of pure copper, α′ steel and Cu/α′ steel multilayered sheets. Calculated strength-elongation σ_UTS × ε_total and strength-electrical conductivity σ_UTS × ρ balances are also shown.

Figure 3 shows the true stress and work-hardening rate as a function of the true strain in the α′ steel and Cu/α′ steel multilayered sheets. The work-hardening rate decreased with decrease in the volume fraction of the α′ steel. All specimens satisfied the plastic instability condition; the true stress became equal to the work-hardening rate, suggesting that early fractures did not occur in the specimens in the macroscopic view. In the 80%α′ sheet, ε_uniform was slightly smaller than the α′ steel, but ε_total was slightly larger. Thus, the necking strain (ε_local = ε_total − ε_uniform) was improved by the multilayered structure. However, in the 60%α′ and 30%α′ sheets, both ε_uniform and ε_local were lower than those in the α′ steel sheets, resulting in low ε_total. Consequently, these results suggest that the Cu layer does not contribute to the improvement of elongation in the Cu/α′ steel multilayered sheet, and it causes lower values of σ_UTS × ε_total.

Fig. 3

True stress-true strain and work hardening curves of α′ steel and the Cu/α′ steel multilayered sheets with different volume fraction of α′ steel.

The ρ value of the Cu sheet was approximately 100%IACS, and the α′ steel sheet had 6%IACS, both of which were consistent with a previous study.^13,19) The values of ρ of the Cu/α′ steel multilayered sheets were of the opposite order of the volume fraction of α′ steel. σ_UTS × ρ values of the multilayered sheets were significantly higher than that in the single-layered sheets. Therefore, it can be concluded that strength-electrical conductivity balance can be improved by forming the multilayered structure with layers of high conductivity and strength.

Figure 4 shows the relationship between σ_UTS and ρ for various conventional Cu alloys,¹³⁾ Cu/Fe multilayered sheets,¹¹⁾ Cu and α′ steel sheets, and the present Cu/α′ steel multilayered sheets. Notably, σ_UTS × ρ of all Cu/α′ steel multilayered sheets were considerably higher than those of the Cu alloys and single-layer sheets. The value of 5.0 × 10⁴ MPa·%IACS of 60%α′ sheet was the highest in the present multilayered sheets and considerably higher than 3.0 × 10⁴ MPa·%IACS of Corson alloys, which have approximately same ρ. The 80%α′ sheet has higher σ_UTS and σ_UTS × ρ values than the high-strength Cu–Be and Cu–Ti alloys. The strength and electrical conductivity in Cu/α′ steel multilayered sheets can be controlled over a wide range by tuning the volume ratio of each layer. In Section 4, the effects of volume ratio on the strength and electrical conductivity in the Cu/α′ multilayered sheet will be quantitatively discussed.

Fig. 4

Relationship between σ_UTS and ρ of various Cu alloys,¹³⁾ Cu/Fe multilayered sheets,¹¹⁾ Cu and α′ steel sheets, and present Cu/α′ steel multilayered sheets.

3.3 Deformation and fracture behaviors

Figure 5 shows the ε_xx distributions of the (a) 80%α′ and (b) 30%α′ sheets when the applied stress reached their σ_UTS (nominal strains of 0.03 and 0.012, respectively). The dotted white lines indicate the interface between Cu and α′ steel layers. The color bar denotes the strain, and the maximum and minimum values are twice the average strain and 0, respectively. In 80%α′ sheet (Fig. 5(a)), the strain tended to be uniformly distributed independent of layers, although some portions exhibited a twice higher strain than the average one. Whereas, in the 30%α′ sheet (Fig. 5(b)), the strain was inhomogeneously distributed and the regions with high strains over seven times higher than the average strain existed in the α′ steel layer, as indicated by arrows. Since the strain distribution becomes rather inhomogeneous as deformation progresses,^20–22) a larger strain concentration but a lower applied strain in the 30%α′ sheet compared with those in the 60% sheet in Fig. 5 can state that an extremely inhomogeneous strain distribution was developed in the 30%α′ sheet. Generally, in the dual-phase steel consisting of soft ferrite and hard α′ phases, the strain is preferentially introduced into the soft ferrite phase.^23,24) However, in a ferrite/α′ steel multilayered sheet, it has been reported that strain concentration into hard α′ layers occurs owing to a layered structure parallel to the tensile axis.²⁵⁾ It has also been revealed that the strain concentrations into the ferrite/α′ multilayered steel occur preferentially in the thin α′ layer region. Thus, layer thickness should be one of the keys to the strain concentration into the hard layers in multilayered sheets. In the 80%α′ sheet, the thickness of α′ steel layers was nearly three times larger than that in 30%α′ sheet, and it must suppress the strain concentration in a hard α′ steel layer and give a homogeneous strain distribution. The high strain region in the 30%α′ sheet tended to be continuously distributed along approximately 45 degrees from the tensile axis, which is also consistent with previous studies.^23,26–28)

Fig. 5

ε_xx strain distribution at nominal strain of 0.03 and 0.012 in (a) 80%α′ and (b) 30%α′ sheets, respectively.

Figure 6 shows the SEM images of (a) 80%α′ and (b) 30%α′ sheets tensile-deformed to failure. In the 80%α′ sheet, coarse cracks or voids were detected in the α′ steel layers, as indicated by the arrows (Fig. 6(a)). Coarse voids were also observed in the α′ steel layers near the fracture surface in the 30%α′ sheet, as indicated by the arrows (Fig. 6(b)). In contrast, no cracks were detected in the Cu layers of either sheet. Thus, it can be concluded that cracks and voids were nucleated in the α′ steel layers and triggered the final fracture of the Cu/α′ multilayered sheets, regardless of the volume fraction of each layer. The low elongation of the α′ steel sheet (Fig. 2) and the high strain concentration in the α′ steel layers (Fig. 5(b)) reasonably explain the preferential crack and void nucleation in the α′ steel layer (Fig. 6).

Fig. 6

SEM images of the cross-sections of the fracture surfaces of (a) 80%α′ and (b) 30%α′ sheets.

Figure 7 shows the fracture surfaces of the (a) pure Cu, (b) α′ steel, (c) 80%α′, (d) 60%α′ and (e) 30%α′ sheets. Coarse dimples were observed on the fracture surface of the pure Cu sheet (Fig. 7(a)), which is a typical ductile fracture surface. The α′ steel sheet exhibited a flat fracture surface without dimples, suggesting that brittle fracture occurred (Fig. 7(b)). The fracture surfaces of the α′ layer in the 80%α′ sheet consisted of numerous dimples (Fig. 7(c)). On the other hand, dimpled and flat fracture surfaces coexisted on the α′ layer in the 30%α′ and 60%α′ sheets, as shown in Fig. 7(d) and (e). Therefore, the fracture behavior of the α′ steel layers changed as their thickness decreased. As shown in Fig. 5(b), extremely high strains were introduced into the α′ steel layers in the 30%α′ sheet, which resulted in high stress concentrations and brittle fractures in the α′ steel layers. In the 30%α′ sheet, a flat fracture surface was detected in the Cu layers, even though the pure Cu single sheet exhibited a ductile fracture. Interestingly, facets were also observed near the Cu/α′ steel interfaces in the 60%α′ and 80%α′ sheets, as indicated by the black arrows in Figs. 7(c) and (d). Similar results were also observed in the ductile ferrite layers in a ferrite/α′ steel multilayered sheet, and it was suggested that the fracture of adjacent α′ steel layers induced stress concentration and led to brittle fracture of the ferrite layers.²⁵⁾ As mentioned above, in the Cu/α′ steel multilayered sheets, large cracks were formed only in the α′ steel layers, similar to a previous study.²⁵⁾ Therefore, the brittle fracture of the Cu layers should be induced by the fracture of the α′ steel layers.

Fig. 7

Fracture surfaces of (a) pure copper and (b) α′ steel sheets, and (c) 80%α′, (d) 60%α′ and (e) 30%α′ sheets.

4. Discussion

It was demonstrated that σ_UTS and ρ of multilayered sheets satisfied the rule of mixtures;^11,29) σ_UTS and ρ can be estimated from the following equations:

  

\begin{equation} \sigma_{\text{UTS}}^{\text{Cu/${\alpha'}$steel}} = \sigma_{\text{UTS}}^{\text{Cu}} \times V^{\text{Cu}} + \sigma_{\text{UTS}}^{\text{${\alpha'}$steel}} \times V^{\text{${\alpha'}$steel}} \end{equation} (1)

  

\begin{equation} \rho^{\text{Cu/${\alpha'}$steel}} = \rho^{\text{Cu}} \times V^{\text{Cu}} + \rho^{\text{${\alpha'}$steel}} \times V^{\text{${\alpha'}$steel}} \end{equation} (2)

The area fractions measured using SEM (Table 1) were used as the volume fractions of the layers (V). Figure 8(a) shows the measured σ_UTS values and estimates from eq. (1) as a function of V^{α′ steel}. In the 80%α′ sheet, the estimated σ_UTS was almost identical to the measured one; thus, the rule of mixtures was valid, similar to previous studies.²⁹⁾ However, the measured values of σ_UTS were slightly lower than the estimated values for the 60%α′ and 30%α′ sheets. Because the Vickers hardness in the α′ steel layer in each Cu/α′ steel multilayered sheet was the same as that of the α′ sheet (Table 1), the strength of the α′ steel layers in the 60%α′ and 30%α′ sheets should not be responsible for the differences between the measured and estimated σ_UTS values. Figure 9 shows a cross-sectional SEM image of the 30%α′ sheet deformed to 0.01 nominal strain within the uniform deformation stage. As indicated by the arrows in Fig. 9, coarse cracks can be observed even in the uniform deformation stage. Therefore, the fractured α′ steel layers did not bear the applied stress, resulting in the measured σ_UTS being lower than those estimated for the 60%α′ and 30%α′ sheets. An extremely high strain, which is over seven times larger than the average strain, was introduced into the α′ steel layer in the 30%α′ sheet (Fig. 5) and caused early fracture of the α′ steel layers even during the uniform deformation stage. Inhomogeneous strain distributions and early fracture of the α′ steel layers were not observed in the 80%α′ sheet. Therefore, it can be said that almost all the α′ steel layers can bear the applied stress, and thus, the measured σ_UTS is nearly identical to the estimated one. In the Cu/α′ steel multilayered sheets, the rule of mixtures is almost valid, except for the case in which early fracture of the α′ steel layer occurs during uniform deformation. The thin α′ steel layers caused strain concentration in the 60%α′ and 30%α′ sheets, as discussed in Section 3.3. Thus, σ_UTS and σ_UTS × ρ can be improved by thickening the α′ steel layers.

Fig. 8

Comparisons of measured and estimated (a) ultimate tensile stress and (b) electrical conductivity.

Fig. 9

Cross-sectional SEM image showing the formation of voids and cracks in the 30%α′ sheet tensile-deformed to a nominal strain of 0.01.

Figure 8(b) shows the measured and estimated ρ values as a function of the volume fraction of α′ steel. The estimated ρ value was approximately identical to the measured value. Thus, ρ in the Cu/α′ steel multilayered sheets satisfies the rule of mixtures independent of the volume fraction of each layer. It is well known that the dissolution of Fe atoms in Cu causes a significant deterioration of ρ in Cu alloys.¹²⁾ The diffusion distance of Fe atoms into the Cu layers during the solution treatment at 1063 K for 0.6 ks, calculated from the interdiffusion coefficient,³⁰⁾ was approximately 0.8 µm. Because the diffusion distance of Fe is much narrower than the thickness of the Cu layers (Table 1), it is strongly suggested that the values of ρ in the Cu layers are not harmed by the heat-treatment condition.

The value of σ_UTS × ρ can be improved by a multilayered structure formed using a combination of high-strength and high-electrical conductivity materials, and its properties can be predicted from the rule of mixtures unless early fracture occurs in the hard layer and atoms in the hard layers diffuse into the high-conductivity layers. Consequently, multilayered sheets with various strengths and electrical conductivities can be easily designed based on the rule of mixtures.

5. Conclusion

Multilayered sheets composed of pure Cu and martensitic (α′) steel with various volume ratios were fabricated, and their strength-electrical conductivity balances were evaluated. Furthermore, the measured tensile properties and electrical conductivity were compared with those estimated from the rule of mixtures. Based on the obtained results, the reasons for the differences between the measured and estimated values were discussed. The main findings are summarized as follows:

1. (1)    Cu and α′ steel layers in the fabricated multilayered sheets had coarse grain and lath martensite microstructures, respectively. The layers were continuously aligned along the rolling direction without interruptions or large bends, whereas the layer thickness varied with the volume ratio of the multilayered sheet.
2. (2)    The ultimate tensile stresses of the multilayered sheets increased with increasing volume fraction of α′ steel. However, the total strain of the multilayered sheets was nearly identical to or slightly lower than that of the α′ single sheet. As a result, the strength-elongation balance of the multilayered sheets was lower than that of Cu and α′ steel.
3. (3)    The electrical conductivity of the multilayered sheets increased with an increase in the volume fraction of Cu. The strength-electrical conductivity balance in the Cu/α′ steel multilayered sheets was significantly higher than that in Cu, α′ steel and commercial Cu alloys.
4. (4)    Large cracks or voids were formed in the α′ steel layers of the multilayered sheets regardless of the volume ratio. Therefore, fractures of the α′ steel layers could lead to fatal fractures of the multilayered sheets.
5. (5)    The measured ultimate tensile stress and electrical conductivity of the multilayered sheets were nearly identical to those estimated based on the rule of mixtures. However, in the case of early fracture occurring in the hard layer owing to the extremely high strain concentration, the measured ultimate tensile stress was lower than the estimated value.

Acknowledgement

The authors are grateful to TAKEFU SPECIAL STEEL CO., LTD. for providing the Cu/carbon steel multilayered sheet. This work was supported by the Japan Copper and Brass Association, The Japan Institute of Metals and Materials, JKA and its promotion funds from KEIRIN RACE, and the Grant-in-Aid for Scientific Research (KAKENHI) Grant No. 20K14605 and 23K17815.

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