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Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys
Journal Pre-proof Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys Duanzhi Duan, Changsheng Li, Jianjun Ding, Yangpeng Liu, Lin Sun, Qijing Lin, Zhuangde Jiang PII:
S0272-8842(20)30414-4
DOI:
https://doi.org/10.1016/j.ceramint.2020.02.092
Reference:
CERI 24315
To appear in:
Ceramics International
Received Date: 11 December 2019 Revised Date:
4 February 2020
Accepted Date: 11 February 2020
Please cite this article as: D. Duan, C. Li, J. Ding, Y. Liu, L. Sun, Q. Lin, Z. Jiang, Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys, Ceramics International (2020), doi: https://doi.org/10.1016/j.ceramint.2020.02.092. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2020 Published by Elsevier Ltd.
Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys Duanzhi Duan, Changsheng Li*, Jianjun Ding, Yangpeng Liu, Lin Sun, Qijing Lin, Zhuangde Jiang State Key Laboratory for Manufacturing Systems Engineering, School of Mechanical Engineering, Xi’an Jiaotong University, Xi’an 710049, China
Abstract Novel multi-layer brazed diamond segments were fabricated using NiCr–x(CuCe) composite alloys. Differential scanning calorimetry curves of the composite alloys were measured and analysed. The microstructures of the alloy segments and surface topographies of the brazed diamond segments were characterised. Performance tests of the alloy segments and brazed diamond segments were performed. The undercooling degree of the Ni–Cr alloy in the composite alloy increased with the Cu–Ce alloy addition, which led to coarse NiCu-rich regions and Ni3Si phases. A brazed diamond segment with a 5% Cu–Ce alloy addition exhibited the highest wear resistance and machining performance and the best surface morphology after a wear test. An excessive Cu–Ce alloy addition led to a rapid decrease in wear resistance of the brazed diamond segment owing to the large number of coarse NiCu-rich phases falling off from the composite alloy. The mechanism of the reduction in thermal damage to diamonds by the Cu–Ce alloy is elucidated. Initially the Cu–Ce particles melted and mainly Ni atoms diffused into the Cu–Ce liquid, thereby leading to the formation of NiCu-rich regions and Ce2Ni7 and CeNi2
— — — — *Corresponding author. E-mail address: [email protected] (C.S. Li). 1
phases, which in turn promoted the diffusion. The melting temperature of the Ni–Cr composite alloy was significantly reduced by the addition of the Cu–Ce alloy. Keywords: Ni–Cr alloy; microstructure; differential scanning calorimetry; Cu–Ce alloy; performance
1. Introduction Artificial synthetic diamond grains have been extensively used for the fabrication of grinding tools to process hard, fragile, and other hard-to-machine materials (e.g., jade, marble, engineering ceramics, glasses, and metal-matrix composites) [1]. Most of the existing diamond grit tools are fabricated by electroplating and hot-press sintering. The defects of the diamond tools (e.g., low grain protrusion and low bond strength between grains and metal matrix) make them more susceptible to a grain premature pull-out. Single-layer brazed diamond tools have been suggested as alternatives to overcome the above defects. In brazed diamond tools, the interfacial bonding strength between diamond grits and the bonding alloy is improved by active brazing with “work” element (e.g., Ti and Cr) addition into the brazing alloy owing to the formation of a chemical bond at the interface. However, these brazed diamond tools have only one layer of diamonds and short lifetimes [2]. Multi-layer brazed diamond tools are considered promising because of the high interfacial bonding strengths and long lifetimes. However, no extensive studies on multi-layer brazed diamond tools have been reported [3, 4].
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The most commonly used active brazed alloys are based on the Ni–Cr and Cu– Sn–Ti alloys [5–14]. The Ni–Cr alloy has become the prevalent material in the development of single-layer brazed diamond grit tools owing to its high hardness and good abrasion resistance [15, 16]. However, possible thermal damage attributed to the Ni–Cr alloy, including brazing residual stresses near the interfaces between grains and the bonding alloy and corrosion pits on grain surfaces, have been reported [16–18]. The thermal damage limits the potentially superior machining performance of the brazed diamond tools [19, 20]. In this study, to decrease or even avoid thermal damage, multi-layer brazed diamond segments were fabricated using the Ni–Cr alloy and four NiCr–x(CuCe) composite alloys (x = 2–20 wt.%). Differential scanning calorimetry (DSC) curves of the Ni–Cr composite alloys were measured upon heating and cooling and analysed. The microstructures of the alloy segments and surface topographies of the brazed diamond segments were studied. In addition, a wear test was performed on the brazed diamond segments. The mechanism of the reduction in thermal damage of the brazed diamonds was analysed.
2. Materials and methods A self-made NiCrBSi brazing alloy powder containing 10 wt.% of chromium, Cu–20-wt.%-Ce alloy powder (Changsha Tianjiu Metal Materials Co. Ltd., China), artificial synthetic diamond grits (35/40 US mesh), and cylindrical 45-steel substrate with a diameter of 16 mm was used. The size of the Cu–Ce powder alloy was smaller 3
than 74 µm. The compositions of the different brazing alloys used in the experiment are listed in Table Ⅰ. Table Ⅰ Compositions of five brazing alloys (wt.%).
Brazing alloy
Ni–Cr alloy powder
Cu–Ce alloy powder
Ni–Cr alloy
100
0
No. 1 Ni–Cr composite alloy
98
2
No. 2 Ni–Cr composite alloy
95
5
No. 3 Ni–Cr composite alloy
90
10
No. 4 Ni–Cr composite alloy
80
20
Diamond grits were ultrasonically irrigated in acetone. The brazing alloy or mixture of the brazing alloy and diamond grits was filled into a graphite mould cavity. The graphite moulds were placed into a furnace and were subjected to a brazing heat treatment at a peak temperature of 1050 ℃ with a hold up of 10 min. Alloy segments and diamond segments were obtained upon furnace cooling, which are denoted according to the brazing alloy used in the experiment. For example, the alloy segments and diamond segments based on the Ni–Cr alloy are denoted as “conventional”. The alloy segments and diamond segments based on No. 1 composite alloy are denoted as No. 1. Figure 1 shows the multi-layer brazed diamond segments obtained through the above brazing process.
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Fig. 1 Multi-layer brazed diamond segments.
The metallographic process included mechanical grinding, disc polishing, and chemical etching by a solution of HCl, H2SO4, and CuSO4. The microstructures of the alloy segments after the metallographic process and morphologies of the diamond segments were observed using tungsten-filament scanning electron microscopy (SEM, SU3500, Hitachi, Japan) with energy-dispersive spectrometry (EDS). The phase compositions of the composite alloy segments were analysed by X‐ ray diffraction (XRD, D8 ADVANCE, Bruker, Germany). The microhardness of the alloy segments after fine polishing was measured using an HXD-1000TMC/LCD Vickers microhardness tester by following the American Society for Testing and Materials (ASTM) E384-11ɛ1 standard. A force of 500 gf (5 N) was loaded with a dwell time of 10 s. Each presented microhardness is the mean value of nine indentations. The wear resistance of the multi-layer brazed diamond segments were evaluated based on their wear losses. A lower wear loss implies a higher wear resistance. The wear losses of the diamond segments were measured using an MPX-2A friction and 5
wear tester and same machining parameter setup. The load was 100 N, the rotation speed was 800 r/min, and the test time was 30 min. A resin-bonded silicon carbide wheel with a diameter of 200 mm and thickness of 20 mm was used as the counterpart material. The machining performance of the segments was evaluated using their wear ratios (ratio of the weight loss of removed counterpart material to the weight loss of brazed diamond segment wear). The wear losses (∆m, g) of the segments and counterpart material were determined by weighing them on an analytical balance with a measurement accuracy of ±0.001 g before and after the wear test. The wear test for each brazed diamond segment was repeated three times to obtain valid and statistically significant values.
3. Experimental results and discussion 3.1. Melting and solidification of the Ni–Cr composite alloy
To identify their elementary reactions during the heating and cooling, the Ni–Cr composite alloys were studied by DSC (STA 449F3, Netzsch, Germany) under argon atmosphere. The temperature measurement range was 20 to 1050 ℃ with an accuracy of ±0.1 ℃. The heating and cooling rates were 10 ℃/min. Figure 2 shows the DSC curves of the four Ni–Cr composite alloys and Ni–Cr alloy upon heating and cooling. The vertices in the DSC curves correspond to phase transitions during the heating and cooling. The valleys in the heating curves indicate strong melting transitions, whereas the peaks in the cooling curves indicate intense
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solidification transitions. It is difficult to precisely obtain the onset temperatures for the vertices, and thus the peak temperatures were used.
(a) No. 1 Ni–Cr composite alloy
(b) No. 2 Ni–Cr composite alloy
(c) No. 3 Ni–Cr composite alloy
(d) No. 4 Ni–Cr composite alloy
(e) Ni–Cr alloy
Fig. 2 DSC results during the heating and cooling of the Ni–Cr composite alloys.
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As shown in Fig. 2, several valleys are observed during the heating, attributed to the melting heat flows originating from the Cu–Ce alloy and Ni–Cr alloy powders. The Ce–Cu phase diagram [21] shows that the liquidus temperature of the Cu– 20-wt.%-Ce alloy used in our study is approximately 860 ℃ (1133.15 K). Thus, when the Ni–Cr composite alloy powders were heated, initially the Cu–Ce alloy powders melted, and then the Ni–Cr alloy particles melted. As shown in Figs. 2(a)–(d), initially the Cu–Ce alloy powders melted and generated the first endothermic vertex. The Ni– Cr alloy particles then melted and generated the second and third vertices. The endothermal valley decreased with the Cu–Ce addition, mainly because when the temperature was up to 860 °C, initially the Cu–Ce alloy melted and liquefied to spread along the Ni–Cr alloy, thereby improving the wettability of the Ni–Cr alloy and decreasing the melting point of the Ni–Cr composite alloy. Furthermore, one endothermal valley is observed in Fig. 2(e) with the highest heat-flow compared to those in Figs. 2(a)–(d). One endothermal valley is observed in Fig. 2(a) with a considerably higher heat-flow than those in Figs. 2(b)–(d), which implies that this valley corresponds to the melting of the Ni–Cr alloy with the highest weight fraction (Table Ⅰ). During the cooling, initially a Ni-rich phase froze and generated the first exothermal vertex. With the decrease in temperature, the Cu–Ni solid solution started to freeze and generated the second vertex, which explains the two vertices in the DSC
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curves in Figs. 2(a)–(d). The values in Table Ⅱ are the peak temperatures for the vertices attributed to the Ni–Cr alloy during the heating and cooling. Table Ⅱ Comparison of the peak temperatures (°C) and undercooling values (∆T) in the DSC curves for
the Ni–Cr alloy in the composite alloys.
Sample
Heating Tst (℃)
Cooling Tir (℃)
Undercooling (Tst – Tir) (℃)
No. 1 Ni–Cr composite alloy
978.7
966.7
12
No. 2 Ni–Cr composite alloy
975.7
948.7
27
No. 3 Ni–Cr composite alloy
972.7
927.7
45
No. 4 Ni–Cr composite alloy
972.7
921.7
51
As shown in Table Ⅱ, the exothermal peak upon the cooling attributed to the Ni-rich phase is observed at a lower temperature than that of the endothermic peak upon the heating attributed to the Ni–Cr alloy particles. The degree of undercooling was calculated as the difference between the melting temperature upon the heating and solidification temperature upon the cooling, which is associated with the difficulties in the nucleation of a solid phase in a liquid state [22, 23]. Remarkably, the degrees of undercooling were 12, 27, 45, and 51 ℃ for No. 1, No. 2, No. 3, and No. 4 composite alloys, respectively. Notably, the Cu–Ce alloy addition could lead to an increase in degree of undercooling. A low undercooling can promote the nucleation and enhance the nucleus during liquid brazing alloy solidification and thus lead to fine structures [24–26]. Thus, the Ni-rich phases within the solidified No. 3 and No. 4 composite alloys were coarser than those within the No. 1 and No. 2 composite alloys. The values in Table Ⅲ are the peak temperatures for the first and second endothermic vertices during the heating. The melting temperature was significantly 9
reduced and the melting range of the Ni–Cr composite alloy was expanded with the Cu–Ce alloy addition. Table Ⅲ Liquidus (Tst) and solidus (Tir) temperatures of the Cu–Ce alloy in the composite alloys during
the heating.
Sample
Tst (℃)
Tir (℃)
Melting range (℃)
No. 1 Ni–Cr composite alloy
978.7
933.7
45
No. 2 Ni–Cr composite alloy
975.7
879.7
96
No. 3 Ni–Cr composite alloy
972.7
876.7
96
No. 4 Ni–Cr composite alloy
972.7
861.7
111
3.2. Microstructures of the alloy segments Figs. 3(a), 4(a), 5(a), and 6(a) show back-scattered-electron images of the characteristic microstructures of the four solidified composite alloys. The phases in the composite alloys were observed. Their compositions were evaluated by EDS, as shown in Figs. 3(b)–(e), 4(b), 4(c), 5(b), and 5(c). The brighter regions consisted of mainly α-Ni phases with approximately 3.4 wt.% of chromium. The brightest regions consisted of mainly NiCu-rich phases with different weight percentages of copper, chromium, and silicon. The darker regions comprised α'-Ni phases with approximately 6.6 and 4.7 wt.% of chromium and silicon, respectively. The black regions comprised Cr and CrB phases. Other phases such as Ni3Si were also detected. These results are in agreement with the XRD results in Fig. 13. Owing to the smaller atomic number of boron, precise detection of the weight composition of the Ni3B phase by EDS was not possible.
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Fig. 3 (a) Microstructure of No. 1 Ni–Cr composite alloy. EDS analyses of the (b) α-Ni, (c) α'-Ni, (d) Ni3Si, and (e) NiCu-rich regions.
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Fig. 4 (a) Microstructure of No. 2 Ni–Cr composite alloy. EDS analyses of the (b) α'-Ni and (c) NiCu-rich
regions.
Fig. 5 (a) Microstructure of No. 3 Ni–Cr composite alloy. EDS analyses of the (b) CrB and (c) NiCu-rich
regions.
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Fig. 6 Microstructure of No. 4 Ni–Cr composite alloy.
As shown in Figs. 3(a), 4(a), 5(a), and 6, the α-Ni and α'-Ni phase contents decreased with the Cu–Ce alloy addition, while the segregation of Ni3Si phases increased. Furthermore, the NiCu-rich regions and Ni3Si phases became coarser. This is consistent with the degrees of undercooling in Table Ⅱ. Notably, the NiCu-rich regions were surrounded by α-Ni, α'-Ni, and Ni3Si phases. In the NiCu-rich regions, the weight percentages of Ni were considerably higher than those of Cu. According to the analysis results in Sections 3.1 and 3.2, likely initially Cu–Ce alloy particles melted and mainly Ni atoms from the Ni–Cr alloy diffused into the Cu–Ce liquid, thereby inducing the formation of NiCu-rich regions. According to the Arrhenius equation, D = D0exp(−Q/RT), where D is the diffusion coefficient, Q is the activation energy, R is the gas constant, T is the absolute temperature, and D0 is the diffusion constant, the diffusion coefficient increases with the temperature. 13
Initially, at approximately 860 °C, Cu–Ce alloy particles in the composite brazing alloy started to melt. At this temperature, the diffusion coefficient of Ni in Cu is approximately 2.9 × 10−15 m2s−1, while the diffusion coefficient of Cu in Ni is 0.6 × 10−15 m2s−1. Thus, the diffusion of Ni in Cu is faster than that of Cu in Ni [27, 28], as the diffusion in the liquid is 104–106 times faster than that in the solid [29]. With the increase in temperature, a liquid Cu–Ce is formed and the diffusion of Ni in the liquid Cu–Ce is considerably faster than that of Cu in the solid Ni. Thus, a large number of Ni atoms from the Ni–Cr alloy particles diffused into the Cu–Ce liquid, which was the main diffusion direction. This led to the formation of NiCu-rich regions. When the temperature was increased to approximately 950 ℃, Ni–Cr alloy particles started to melt and Ni and Cu mutually diffused into each other, which accelerated the formation of NiCu-rich regions. 3.3. Performance of the brazed diamond segments Figure 7 shows the Vickers microhardness of the five alloy segments. The mean microhardness values of No. 1 and No. 2 composite alloys were decreased by approximately 1.7 and 4.1%, while those of No. 3 and No. 4 composite alloys were decreased by approximately 6.9 and 12.6%, respectively, compared to that of the Ni– Cr alloy. The hardness of the alloy segment decreased with the Cu–Ce alloy addition, as the contents of α-Ni and α'-Ni phases having hardness higher than that of the NiCu-rich regions decreased (Section 3.2). Simultaneously, the NiCu-rich regions and
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Ni3Si phases became coarser (Sections 3.1 and 3.2), and thus their strengths decreased.
Fig. 7 Comparison of the microhardness of the different alloys.
The wear losses of the four multi-layer brazed diamond segments are presented in Fig. 8. The wear loss of the multi-layer brazed diamond segment initially gradually increased, and then rapidly increased with the Cu–Ce alloy addition. The wear losses of No. 2 and No. 3 multi-layer brazed diamond segments were increased by 10 and 55%, respectively, while that of No. 4 segment was increased by 380% compared to that of the segment with the Ni–Cr alloy. The wear resistance of the multi-layer brazed diamond segment decreased with the Cu–Ce alloy addition. No. 2 multi-layer brazed diamond segment based on No. 2 composite alloy containing the 5% Cu–Ce alloy exhibited a higher wear resistance than those of No. 3, and No. 4 segments. Notably, the mean microhardness of No. 4 composite alloy was decreased by approximately 12.6%, whereas the wear resistance of No. 4 multi-layer brazed diamond segment was decreased by approximately 79.2% compared to those of the segment with the Ni–Cr alloy, as analysed in Section 3.4. 15
The wear ratios of the diamond segments are also presented in Fig. 8, which initially increased, and then rapidly decreased with the Cu–Ce alloy addition. The wear ratio of No. 2 multi-layer brazed diamond segment was increased by 13.6%, while those of No. 3 and No. 4 multi-layer brazed diamond segments were decreased by 10.7 and 68.2%, respectively, compared to that of the segment with the Ni–Cr alloy. A higher wear ratio of the diamond segment implies a higher machining performance. No. 2 multi-layer brazed diamond segment based on No. 2 composite alloy containing the 5% Cu–Ce alloy exhibited the highest machining performance among those of the segments. Notably, the wear loss of No. 2 multi-layer brazed diamond segment was increased by 10%, but its wear ratio was increased by 13.6%, which can be attributed to the reduction in thermal damage of the diamonds and increase in their performance.
Fig. 8 Wear losses of the four multi-layer brazed diamond segments.
3.4. Morphologies of the brazed diamond segments Figures 9–12 show SEM images of the four multi-layer brazed diamond segments after the wear test. As shown in Fig. 9, large corrosion pits were observed 16
on the diamonds’ surfaces. In addition, integral fractures occurred in a few diamonds. Partial edges of the diamonds with high exposures were rarely observed after the wear test, which are forms of thermal damage of the diamonds. The Ni–Cr alloy led to a large unfavourable thermal damage to the diamond grits during the high-temperature process, which had negative influences on the processing efficiency and wear resistance of the brazed diamond segment with the Ni–Cr alloy. As shown in Figs. 10–12, few large corrosion pits on the diamonds’ surfaces were observed. The edges of the diamonds with high exposures were still observed. The Cu–Ce alloy reduced the thermal damage to the diamonds during the high-temperature process. However, some irregular holes with lengths and widths of approximately 50–400 µm are observed in Figs. 10–12. The number of irregular holes had the highest value for the sample in Fig. 12, while that of the sample in Fig. 11 was higher than that of the sample in Fig. 10. The area of the irregular holes in Fig. 12 had also the largest value. The change tendency of the irregular holes is consistent with that of the NiCu-rich phases in Section 3.2. The appearance of the irregular holes can be attributed to the NiCu-rich phases falling off from the solidified Ni–Cr composite alloy owing to the low bond strengths between the NiCu-rich phases and other phases and low strengths of the NiCu-rich phases. Furthermore, as shown in Fig. 10, micro- and macrofractures occurred in a large number of diamonds. Mainly macrofractures are observed in Figs. 11 and 12, as the rapid wear of the Ni–Cr composite alloy led to a large grinding force acting on each 17
diamond grain, which tended to cause macrofractures. No. 2 multi-layer brazed diamond segment based on the Ni–Cr composite alloy containing the 5% Cu–Ce alloy exhibited the best surface morphology after the wear test. This indicates that No. 2 multi-layer brazed diamond segment exhibited the highest machining performance. Therefore, the large number of coarse NiCu-rich phases falling off from the solidified Ni–Cr composite alloy were the main origin of the rapid decrease in wear resistance of No. 4 brazed diamond segment, as shown in Fig. 8. Based on the above analytical results on the wear morphology and experimentally determined microhardness, wear loss, and wear ratio in Section 3.3, the optimal amount of Cu–Ce alloy addition was 5% in this study.
Fig. 9 Morphology of the brazed diamond segment based on the Ni–Cr alloy after the wear test.
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Fig. 10 Morphology of No. 2 brazed diamond segment after the wear test.
Fig. 11 Morphology of No. 3 brazed diamond segment after the wear test.
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Fig. 12 Morphology of No. 4 brazed diamond segment after the wear test.
3.5. Mechanism of the reduction in thermal damage
Figures 13(a) and (b) show XRD patterns of the alloy segments based on No. 2 and No. 3 composite alloys, respectively. As shown in Fig. 13, the dominant phases were Ce2Ni7, CeNi2, Ni3Si, Ni4B3, CrB, and NiCr solid solution. The comparison of Figs. 13(a) and (b) suggests that the CeNi and Cu5Si phases appear with the Cu–Ce alloy addition. These results are mainly consistent with the microstructure evolutions of the four solidified Ni–Cr composite alloys in Section 3.2.
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Fig. 13 XRD patterns of the alloy segments based on (a) No. 2 and (b) No. 3 composite alloys,
respectively.
Thus, the formation of Ce2Ni7 and CeNi2 phases promoted the diffusion of Ni atoms into the Cu–Ce liquid and led to the formation of NiCu-rich regions containing a high weight content of Ni. Therefore, the weight content of Cu decreased to below approximately 25 wt.% (Figs. 3, 4, and 5). According to the experimental results in Sections 3.1, 3.2, and 3.4, the mechanisms of the reduction in thermal damage to the brazed diamonds by the Cu–Ce alloy are illustrated in Figs. 14 and 15.
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Fig. 14 Mechanism of the thermal damage to diamonds by the Ni–Cr alloy.
Fig. 15 Mechanism of the reduction in thermal damage to diamonds by the Cu–Ce addition.
In the high-temperature brazing of diamonds with the Ni–Cr alloy, a large number of Ni atoms can move toward the diamonds and dissolve carbon elements from the diamonds, thereby promoting chemical reactions between the dissolved carbon atoms and Cr elements [17, 30]. The high brazing temperature can promote the above dissolution and reaction process. Therefore, the erosion of diamonds was significant, as shown in Fig. 14. In the brazing of diamonds with the Ni–Cr composite alloy, initially the Cu–Ce alloy particles melted (Fig. 2) and mainly Ni atoms diffused into the Cu–Ce alloy liquid, thereby leading to the formation of NiCu-rich regions. Simultaneously, the formation of Ce2Ni7 and CeNi2 phases (Fig. 13) could promote the diffusion of Ni into the Cu–Ce liquid. The melting temperature of the Ni–Cr composite alloy was significantly reduced by the addition of the Cu–Ce alloy (Section 3.1). The number of Ni atoms moving toward the diamonds could be significantly reduced, and thus the erosion of diamonds was not considerable, as shown in Fig. 15. The formation of 22
NiCu-rich regions could reduce the differences in material properties between the bonding alloys and substrates. Thus, microcracks near the diamond grits were rarely observed. Therefore, the thermal damage to the diamonds was significantly decreased.
4. Conclusions (1) The undercooling degrees of No. 1, No. 2, No. 3, and No. 4 composite alloys were 12, 27, 45, and 51 ℃, respectively. The undercooling degrees of the Ni–Cr alloy in the composite alloys increased, thereby leading to coarse NiCu-rich regions and Ni3Si phases. (2) The solidified Ni–Cr composite alloy contained α-Ni, α'-Ni, Cr, CrB, and Ni3Si phases and NiCu-rich regions. Mainly Ni atoms from the Ni–Cr alloy diffused into the Cu–Ce liquid and thus induced the formation of NiCu-rich phases. (3) The brazed diamond segment based on the composite alloy with the 5-wt.% Cu–Ce alloy addition exhibited the highest wear resistance and machining performance and the best surface morphology after the wear test. The excessive Cu– Ce alloy addition led to a rapid decrease in wear resistance of the brazed diamond segment owing to the large number of coarse NiCu-rich phases falling off from the composite alloy during the wear test. (4) The mechanism of the reduction in thermal damage to the brazed diamonds by the Cu–Ce alloy addition was elucidated. Initially the Cu–Ce particles melted and mainly Ni atoms diffused into the Cu–Ce liquid, thereby leading to the formation of NiCu-rich regions and Ce2Ni7 and CeNi2 phases, which in turn promoted the diffusion. 23
The melting temperature of the Ni–Cr composite alloy was significantly reduced by the addition of the Cu–Ce alloy. Acknowledgment The authors gratefully acknowledge the financial support of this research by China Postdoctoral Science Foundation (Grant No. 2019M653589), National Natural Science Foundation of China (Grant No. 51720105016), National Postdoctoral Program for Innovative Talents (Grant No. BX20180250). The authors also gratefully acknowledge Instrumental Analysis Center of Xi'an Jiaotong University for their technical support in SEM experiments. References [1] C. Artini, M.L. Muolo, A. Passerone, Diamond-metal interfaces in cutting tools: a review, J. Mater. Sci. 47 (2012) 3252–3264. [2] W.F. Ding, J.H. Xu, Z.Z. Chen, C.Y. Yang, C.J. Song, Y.C. Fu, Fabrication and performance of porous metal-bonded CBN grinding wheels using alumina bubble particles as pore-forming agents, Int J Adv Manuf Technol 2013; 67:1309-15. [3] Q.P. Wu, Z.Y. Ouyang, Y. Wang, H. Yang, K. Song, Precision grinding of engineering ceramic based on the electrolytic dressing of a multi-layer brazed diamond wheel, Diam. Relat. Mater.100 (2019) 107552. [4] Q.P. Wu, Z. Luo, Z.H. Deng, Y. Wang, Z.Q. Fu, Fabrication of diamond whiskers by powder injection molding, Int. J. Refract. Met. Hard Mater.74 (2018) 114-122. 24
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Declaration of interest statement The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
S0272-8842(20)30414-4
DOI:
https://doi.org/10.1016/j.ceramint.2020.02.092
Reference:
CERI 24315
To appear in:
Ceramics International
Received Date: 11 December 2019 Revised Date:
4 February 2020
Accepted Date: 11 February 2020
Please cite this article as: D. Duan, C. Li, J. Ding, Y. Liu, L. Sun, Q. Lin, Z. Jiang, Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys, Ceramics International (2020), doi: https://doi.org/10.1016/j.ceramint.2020.02.092. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2020 Published by Elsevier Ltd.
Microstructure and performance of brazed diamond segments with NiCr–x(CuCe) composite alloys Duanzhi Duan, Changsheng Li*, Jianjun Ding, Yangpeng Liu, Lin Sun, Qijing Lin, Zhuangde Jiang State Key Laboratory for Manufacturing Systems Engineering, School of Mechanical Engineering, Xi’an Jiaotong University, Xi’an 710049, China
Abstract Novel multi-layer brazed diamond segments were fabricated using NiCr–x(CuCe) composite alloys. Differential scanning calorimetry curves of the composite alloys were measured and analysed. The microstructures of the alloy segments and surface topographies of the brazed diamond segments were characterised. Performance tests of the alloy segments and brazed diamond segments were performed. The undercooling degree of the Ni–Cr alloy in the composite alloy increased with the Cu–Ce alloy addition, which led to coarse NiCu-rich regions and Ni3Si phases. A brazed diamond segment with a 5% Cu–Ce alloy addition exhibited the highest wear resistance and machining performance and the best surface morphology after a wear test. An excessive Cu–Ce alloy addition led to a rapid decrease in wear resistance of the brazed diamond segment owing to the large number of coarse NiCu-rich phases falling off from the composite alloy. The mechanism of the reduction in thermal damage to diamonds by the Cu–Ce alloy is elucidated. Initially the Cu–Ce particles melted and mainly Ni atoms diffused into the Cu–Ce liquid, thereby leading to the formation of NiCu-rich regions and Ce2Ni7 and CeNi2
— — — — *Corresponding author. E-mail address: [email protected] (C.S. Li). 1
phases, which in turn promoted the diffusion. The melting temperature of the Ni–Cr composite alloy was significantly reduced by the addition of the Cu–Ce alloy. Keywords: Ni–Cr alloy; microstructure; differential scanning calorimetry; Cu–Ce alloy; performance
1. Introduction Artificial synthetic diamond grains have been extensively used for the fabrication of grinding tools to process hard, fragile, and other hard-to-machine materials (e.g., jade, marble, engineering ceramics, glasses, and metal-matrix composites) [1]. Most of the existing diamond grit tools are fabricated by electroplating and hot-press sintering. The defects of the diamond tools (e.g., low grain protrusion and low bond strength between grains and metal matrix) make them more susceptible to a grain premature pull-out. Single-layer brazed diamond tools have been suggested as alternatives to overcome the above defects. In brazed diamond tools, the interfacial bonding strength between diamond grits and the bonding alloy is improved by active brazing with “work” element (e.g., Ti and Cr) addition into the brazing alloy owing to the formation of a chemical bond at the interface. However, these brazed diamond tools have only one layer of diamonds and short lifetimes [2]. Multi-layer brazed diamond tools are considered promising because of the high interfacial bonding strengths and long lifetimes. However, no extensive studies on multi-layer brazed diamond tools have been reported [3, 4].
2
The most commonly used active brazed alloys are based on the Ni–Cr and Cu– Sn–Ti alloys [5–14]. The Ni–Cr alloy has become the prevalent material in the development of single-layer brazed diamond grit tools owing to its high hardness and good abrasion resistance [15, 16]. However, possible thermal damage attributed to the Ni–Cr alloy, including brazing residual stresses near the interfaces between grains and the bonding alloy and corrosion pits on grain surfaces, have been reported [16–18]. The thermal damage limits the potentially superior machining performance of the brazed diamond tools [19, 20]. In this study, to decrease or even avoid thermal damage, multi-layer brazed diamond segments were fabricated using the Ni–Cr alloy and four NiCr–x(CuCe) composite alloys (x = 2–20 wt.%). Differential scanning calorimetry (DSC) curves of the Ni–Cr composite alloys were measured upon heating and cooling and analysed. The microstructures of the alloy segments and surface topographies of the brazed diamond segments were studied. In addition, a wear test was performed on the brazed diamond segments. The mechanism of the reduction in thermal damage of the brazed diamonds was analysed.
2. Materials and methods A self-made NiCrBSi brazing alloy powder containing 10 wt.% of chromium, Cu–20-wt.%-Ce alloy powder (Changsha Tianjiu Metal Materials Co. Ltd., China), artificial synthetic diamond grits (35/40 US mesh), and cylindrical 45-steel substrate with a diameter of 16 mm was used. The size of the Cu–Ce powder alloy was smaller 3
than 74 µm. The compositions of the different brazing alloys used in the experiment are listed in Table Ⅰ. Table Ⅰ Compositions of five brazing alloys (wt.%).
Brazing alloy
Ni–Cr alloy powder
Cu–Ce alloy powder
Ni–Cr alloy
100
0
No. 1 Ni–Cr composite alloy
98
2
No. 2 Ni–Cr composite alloy
95
5
No. 3 Ni–Cr composite alloy
90
10
No. 4 Ni–Cr composite alloy
80
20
Diamond grits were ultrasonically irrigated in acetone. The brazing alloy or mixture of the brazing alloy and diamond grits was filled into a graphite mould cavity. The graphite moulds were placed into a furnace and were subjected to a brazing heat treatment at a peak temperature of 1050 ℃ with a hold up of 10 min. Alloy segments and diamond segments were obtained upon furnace cooling, which are denoted according to the brazing alloy used in the experiment. For example, the alloy segments and diamond segments based on the Ni–Cr alloy are denoted as “conventional”. The alloy segments and diamond segments based on No. 1 composite alloy are denoted as No. 1. Figure 1 shows the multi-layer brazed diamond segments obtained through the above brazing process.
4
Fig. 1 Multi-layer brazed diamond segments.
The metallographic process included mechanical grinding, disc polishing, and chemical etching by a solution of HCl, H2SO4, and CuSO4. The microstructures of the alloy segments after the metallographic process and morphologies of the diamond segments were observed using tungsten-filament scanning electron microscopy (SEM, SU3500, Hitachi, Japan) with energy-dispersive spectrometry (EDS). The phase compositions of the composite alloy segments were analysed by X‐ ray diffraction (XRD, D8 ADVANCE, Bruker, Germany). The microhardness of the alloy segments after fine polishing was measured using an HXD-1000TMC/LCD Vickers microhardness tester by following the American Society for Testing and Materials (ASTM) E384-11ɛ1 standard. A force of 500 gf (5 N) was loaded with a dwell time of 10 s. Each presented microhardness is the mean value of nine indentations. The wear resistance of the multi-layer brazed diamond segments were evaluated based on their wear losses. A lower wear loss implies a higher wear resistance. The wear losses of the diamond segments were measured using an MPX-2A friction and 5
wear tester and same machining parameter setup. The load was 100 N, the rotation speed was 800 r/min, and the test time was 30 min. A resin-bonded silicon carbide wheel with a diameter of 200 mm and thickness of 20 mm was used as the counterpart material. The machining performance of the segments was evaluated using their wear ratios (ratio of the weight loss of removed counterpart material to the weight loss of brazed diamond segment wear). The wear losses (∆m, g) of the segments and counterpart material were determined by weighing them on an analytical balance with a measurement accuracy of ±0.001 g before and after the wear test. The wear test for each brazed diamond segment was repeated three times to obtain valid and statistically significant values.
3. Experimental results and discussion 3.1. Melting and solidification of the Ni–Cr composite alloy
To identify their elementary reactions during the heating and cooling, the Ni–Cr composite alloys were studied by DSC (STA 449F3, Netzsch, Germany) under argon atmosphere. The temperature measurement range was 20 to 1050 ℃ with an accuracy of ±0.1 ℃. The heating and cooling rates were 10 ℃/min. Figure 2 shows the DSC curves of the four Ni–Cr composite alloys and Ni–Cr alloy upon heating and cooling. The vertices in the DSC curves correspond to phase transitions during the heating and cooling. The valleys in the heating curves indicate strong melting transitions, whereas the peaks in the cooling curves indicate intense
6
solidification transitions. It is difficult to precisely obtain the onset temperatures for the vertices, and thus the peak temperatures were used.
(a) No. 1 Ni–Cr composite alloy
(b) No. 2 Ni–Cr composite alloy
(c) No. 3 Ni–Cr composite alloy
(d) No. 4 Ni–Cr composite alloy
(e) Ni–Cr alloy
Fig. 2 DSC results during the heating and cooling of the Ni–Cr composite alloys.
7
As shown in Fig. 2, several valleys are observed during the heating, attributed to the melting heat flows originating from the Cu–Ce alloy and Ni–Cr alloy powders. The Ce–Cu phase diagram [21] shows that the liquidus temperature of the Cu– 20-wt.%-Ce alloy used in our study is approximately 860 ℃ (1133.15 K). Thus, when the Ni–Cr composite alloy powders were heated, initially the Cu–Ce alloy powders melted, and then the Ni–Cr alloy particles melted. As shown in Figs. 2(a)–(d), initially the Cu–Ce alloy powders melted and generated the first endothermic vertex. The Ni– Cr alloy particles then melted and generated the second and third vertices. The endothermal valley decreased with the Cu–Ce addition, mainly because when the temperature was up to 860 °C, initially the Cu–Ce alloy melted and liquefied to spread along the Ni–Cr alloy, thereby improving the wettability of the Ni–Cr alloy and decreasing the melting point of the Ni–Cr composite alloy. Furthermore, one endothermal valley is observed in Fig. 2(e) with the highest heat-flow compared to those in Figs. 2(a)–(d). One endothermal valley is observed in Fig. 2(a) with a considerably higher heat-flow than those in Figs. 2(b)–(d), which implies that this valley corresponds to the melting of the Ni–Cr alloy with the highest weight fraction (Table Ⅰ). During the cooling, initially a Ni-rich phase froze and generated the first exothermal vertex. With the decrease in temperature, the Cu–Ni solid solution started to freeze and generated the second vertex, which explains the two vertices in the DSC
8
curves in Figs. 2(a)–(d). The values in Table Ⅱ are the peak temperatures for the vertices attributed to the Ni–Cr alloy during the heating and cooling. Table Ⅱ Comparison of the peak temperatures (°C) and undercooling values (∆T) in the DSC curves for
the Ni–Cr alloy in the composite alloys.
Sample
Heating Tst (℃)
Cooling Tir (℃)
Undercooling (Tst – Tir) (℃)
No. 1 Ni–Cr composite alloy
978.7
966.7
12
No. 2 Ni–Cr composite alloy
975.7
948.7
27
No. 3 Ni–Cr composite alloy
972.7
927.7
45
No. 4 Ni–Cr composite alloy
972.7
921.7
51
As shown in Table Ⅱ, the exothermal peak upon the cooling attributed to the Ni-rich phase is observed at a lower temperature than that of the endothermic peak upon the heating attributed to the Ni–Cr alloy particles. The degree of undercooling was calculated as the difference between the melting temperature upon the heating and solidification temperature upon the cooling, which is associated with the difficulties in the nucleation of a solid phase in a liquid state [22, 23]. Remarkably, the degrees of undercooling were 12, 27, 45, and 51 ℃ for No. 1, No. 2, No. 3, and No. 4 composite alloys, respectively. Notably, the Cu–Ce alloy addition could lead to an increase in degree of undercooling. A low undercooling can promote the nucleation and enhance the nucleus during liquid brazing alloy solidification and thus lead to fine structures [24–26]. Thus, the Ni-rich phases within the solidified No. 3 and No. 4 composite alloys were coarser than those within the No. 1 and No. 2 composite alloys. The values in Table Ⅲ are the peak temperatures for the first and second endothermic vertices during the heating. The melting temperature was significantly 9
reduced and the melting range of the Ni–Cr composite alloy was expanded with the Cu–Ce alloy addition. Table Ⅲ Liquidus (Tst) and solidus (Tir) temperatures of the Cu–Ce alloy in the composite alloys during
the heating.
Sample
Tst (℃)
Tir (℃)
Melting range (℃)
No. 1 Ni–Cr composite alloy
978.7
933.7
45
No. 2 Ni–Cr composite alloy
975.7
879.7
96
No. 3 Ni–Cr composite alloy
972.7
876.7
96
No. 4 Ni–Cr composite alloy
972.7
861.7
111
3.2. Microstructures of the alloy segments Figs. 3(a), 4(a), 5(a), and 6(a) show back-scattered-electron images of the characteristic microstructures of the four solidified composite alloys. The phases in the composite alloys were observed. Their compositions were evaluated by EDS, as shown in Figs. 3(b)–(e), 4(b), 4(c), 5(b), and 5(c). The brighter regions consisted of mainly α-Ni phases with approximately 3.4 wt.% of chromium. The brightest regions consisted of mainly NiCu-rich phases with different weight percentages of copper, chromium, and silicon. The darker regions comprised α'-Ni phases with approximately 6.6 and 4.7 wt.% of chromium and silicon, respectively. The black regions comprised Cr and CrB phases. Other phases such as Ni3Si were also detected. These results are in agreement with the XRD results in Fig. 13. Owing to the smaller atomic number of boron, precise detection of the weight composition of the Ni3B phase by EDS was not possible.
10
Fig. 3 (a) Microstructure of No. 1 Ni–Cr composite alloy. EDS analyses of the (b) α-Ni, (c) α'-Ni, (d) Ni3Si, and (e) NiCu-rich regions.
11
Fig. 4 (a) Microstructure of No. 2 Ni–Cr composite alloy. EDS analyses of the (b) α'-Ni and (c) NiCu-rich
regions.
Fig. 5 (a) Microstructure of No. 3 Ni–Cr composite alloy. EDS analyses of the (b) CrB and (c) NiCu-rich
regions.
12
Fig. 6 Microstructure of No. 4 Ni–Cr composite alloy.
As shown in Figs. 3(a), 4(a), 5(a), and 6, the α-Ni and α'-Ni phase contents decreased with the Cu–Ce alloy addition, while the segregation of Ni3Si phases increased. Furthermore, the NiCu-rich regions and Ni3Si phases became coarser. This is consistent with the degrees of undercooling in Table Ⅱ. Notably, the NiCu-rich regions were surrounded by α-Ni, α'-Ni, and Ni3Si phases. In the NiCu-rich regions, the weight percentages of Ni were considerably higher than those of Cu. According to the analysis results in Sections 3.1 and 3.2, likely initially Cu–Ce alloy particles melted and mainly Ni atoms from the Ni–Cr alloy diffused into the Cu–Ce liquid, thereby inducing the formation of NiCu-rich regions. According to the Arrhenius equation, D = D0exp(−Q/RT), where D is the diffusion coefficient, Q is the activation energy, R is the gas constant, T is the absolute temperature, and D0 is the diffusion constant, the diffusion coefficient increases with the temperature. 13
Initially, at approximately 860 °C, Cu–Ce alloy particles in the composite brazing alloy started to melt. At this temperature, the diffusion coefficient of Ni in Cu is approximately 2.9 × 10−15 m2s−1, while the diffusion coefficient of Cu in Ni is 0.6 × 10−15 m2s−1. Thus, the diffusion of Ni in Cu is faster than that of Cu in Ni [27, 28], as the diffusion in the liquid is 104–106 times faster than that in the solid [29]. With the increase in temperature, a liquid Cu–Ce is formed and the diffusion of Ni in the liquid Cu–Ce is considerably faster than that of Cu in the solid Ni. Thus, a large number of Ni atoms from the Ni–Cr alloy particles diffused into the Cu–Ce liquid, which was the main diffusion direction. This led to the formation of NiCu-rich regions. When the temperature was increased to approximately 950 ℃, Ni–Cr alloy particles started to melt and Ni and Cu mutually diffused into each other, which accelerated the formation of NiCu-rich regions. 3.3. Performance of the brazed diamond segments Figure 7 shows the Vickers microhardness of the five alloy segments. The mean microhardness values of No. 1 and No. 2 composite alloys were decreased by approximately 1.7 and 4.1%, while those of No. 3 and No. 4 composite alloys were decreased by approximately 6.9 and 12.6%, respectively, compared to that of the Ni– Cr alloy. The hardness of the alloy segment decreased with the Cu–Ce alloy addition, as the contents of α-Ni and α'-Ni phases having hardness higher than that of the NiCu-rich regions decreased (Section 3.2). Simultaneously, the NiCu-rich regions and
14
Ni3Si phases became coarser (Sections 3.1 and 3.2), and thus their strengths decreased.
Fig. 7 Comparison of the microhardness of the different alloys.
The wear losses of the four multi-layer brazed diamond segments are presented in Fig. 8. The wear loss of the multi-layer brazed diamond segment initially gradually increased, and then rapidly increased with the Cu–Ce alloy addition. The wear losses of No. 2 and No. 3 multi-layer brazed diamond segments were increased by 10 and 55%, respectively, while that of No. 4 segment was increased by 380% compared to that of the segment with the Ni–Cr alloy. The wear resistance of the multi-layer brazed diamond segment decreased with the Cu–Ce alloy addition. No. 2 multi-layer brazed diamond segment based on No. 2 composite alloy containing the 5% Cu–Ce alloy exhibited a higher wear resistance than those of No. 3, and No. 4 segments. Notably, the mean microhardness of No. 4 composite alloy was decreased by approximately 12.6%, whereas the wear resistance of No. 4 multi-layer brazed diamond segment was decreased by approximately 79.2% compared to those of the segment with the Ni–Cr alloy, as analysed in Section 3.4. 15
The wear ratios of the diamond segments are also presented in Fig. 8, which initially increased, and then rapidly decreased with the Cu–Ce alloy addition. The wear ratio of No. 2 multi-layer brazed diamond segment was increased by 13.6%, while those of No. 3 and No. 4 multi-layer brazed diamond segments were decreased by 10.7 and 68.2%, respectively, compared to that of the segment with the Ni–Cr alloy. A higher wear ratio of the diamond segment implies a higher machining performance. No. 2 multi-layer brazed diamond segment based on No. 2 composite alloy containing the 5% Cu–Ce alloy exhibited the highest machining performance among those of the segments. Notably, the wear loss of No. 2 multi-layer brazed diamond segment was increased by 10%, but its wear ratio was increased by 13.6%, which can be attributed to the reduction in thermal damage of the diamonds and increase in their performance.
Fig. 8 Wear losses of the four multi-layer brazed diamond segments.
3.4. Morphologies of the brazed diamond segments Figures 9–12 show SEM images of the four multi-layer brazed diamond segments after the wear test. As shown in Fig. 9, large corrosion pits were observed 16
on the diamonds’ surfaces. In addition, integral fractures occurred in a few diamonds. Partial edges of the diamonds with high exposures were rarely observed after the wear test, which are forms of thermal damage of the diamonds. The Ni–Cr alloy led to a large unfavourable thermal damage to the diamond grits during the high-temperature process, which had negative influences on the processing efficiency and wear resistance of the brazed diamond segment with the Ni–Cr alloy. As shown in Figs. 10–12, few large corrosion pits on the diamonds’ surfaces were observed. The edges of the diamonds with high exposures were still observed. The Cu–Ce alloy reduced the thermal damage to the diamonds during the high-temperature process. However, some irregular holes with lengths and widths of approximately 50–400 µm are observed in Figs. 10–12. The number of irregular holes had the highest value for the sample in Fig. 12, while that of the sample in Fig. 11 was higher than that of the sample in Fig. 10. The area of the irregular holes in Fig. 12 had also the largest value. The change tendency of the irregular holes is consistent with that of the NiCu-rich phases in Section 3.2. The appearance of the irregular holes can be attributed to the NiCu-rich phases falling off from the solidified Ni–Cr composite alloy owing to the low bond strengths between the NiCu-rich phases and other phases and low strengths of the NiCu-rich phases. Furthermore, as shown in Fig. 10, micro- and macrofractures occurred in a large number of diamonds. Mainly macrofractures are observed in Figs. 11 and 12, as the rapid wear of the Ni–Cr composite alloy led to a large grinding force acting on each 17
diamond grain, which tended to cause macrofractures. No. 2 multi-layer brazed diamond segment based on the Ni–Cr composite alloy containing the 5% Cu–Ce alloy exhibited the best surface morphology after the wear test. This indicates that No. 2 multi-layer brazed diamond segment exhibited the highest machining performance. Therefore, the large number of coarse NiCu-rich phases falling off from the solidified Ni–Cr composite alloy were the main origin of the rapid decrease in wear resistance of No. 4 brazed diamond segment, as shown in Fig. 8. Based on the above analytical results on the wear morphology and experimentally determined microhardness, wear loss, and wear ratio in Section 3.3, the optimal amount of Cu–Ce alloy addition was 5% in this study.
Fig. 9 Morphology of the brazed diamond segment based on the Ni–Cr alloy after the wear test.
18
Fig. 10 Morphology of No. 2 brazed diamond segment after the wear test.
Fig. 11 Morphology of No. 3 brazed diamond segment after the wear test.
19
Fig. 12 Morphology of No. 4 brazed diamond segment after the wear test.
3.5. Mechanism of the reduction in thermal damage
Figures 13(a) and (b) show XRD patterns of the alloy segments based on No. 2 and No. 3 composite alloys, respectively. As shown in Fig. 13, the dominant phases were Ce2Ni7, CeNi2, Ni3Si, Ni4B3, CrB, and NiCr solid solution. The comparison of Figs. 13(a) and (b) suggests that the CeNi and Cu5Si phases appear with the Cu–Ce alloy addition. These results are mainly consistent with the microstructure evolutions of the four solidified Ni–Cr composite alloys in Section 3.2.
20
Fig. 13 XRD patterns of the alloy segments based on (a) No. 2 and (b) No. 3 composite alloys,
respectively.
Thus, the formation of Ce2Ni7 and CeNi2 phases promoted the diffusion of Ni atoms into the Cu–Ce liquid and led to the formation of NiCu-rich regions containing a high weight content of Ni. Therefore, the weight content of Cu decreased to below approximately 25 wt.% (Figs. 3, 4, and 5). According to the experimental results in Sections 3.1, 3.2, and 3.4, the mechanisms of the reduction in thermal damage to the brazed diamonds by the Cu–Ce alloy are illustrated in Figs. 14 and 15.
21
Fig. 14 Mechanism of the thermal damage to diamonds by the Ni–Cr alloy.
Fig. 15 Mechanism of the reduction in thermal damage to diamonds by the Cu–Ce addition.
In the high-temperature brazing of diamonds with the Ni–Cr alloy, a large number of Ni atoms can move toward the diamonds and dissolve carbon elements from the diamonds, thereby promoting chemical reactions between the dissolved carbon atoms and Cr elements [17, 30]. The high brazing temperature can promote the above dissolution and reaction process. Therefore, the erosion of diamonds was significant, as shown in Fig. 14. In the brazing of diamonds with the Ni–Cr composite alloy, initially the Cu–Ce alloy particles melted (Fig. 2) and mainly Ni atoms diffused into the Cu–Ce alloy liquid, thereby leading to the formation of NiCu-rich regions. Simultaneously, the formation of Ce2Ni7 and CeNi2 phases (Fig. 13) could promote the diffusion of Ni into the Cu–Ce liquid. The melting temperature of the Ni–Cr composite alloy was significantly reduced by the addition of the Cu–Ce alloy (Section 3.1). The number of Ni atoms moving toward the diamonds could be significantly reduced, and thus the erosion of diamonds was not considerable, as shown in Fig. 15. The formation of 22
NiCu-rich regions could reduce the differences in material properties between the bonding alloys and substrates. Thus, microcracks near the diamond grits were rarely observed. Therefore, the thermal damage to the diamonds was significantly decreased.
4. Conclusions (1) The undercooling degrees of No. 1, No. 2, No. 3, and No. 4 composite alloys were 12, 27, 45, and 51 ℃, respectively. The undercooling degrees of the Ni–Cr alloy in the composite alloys increased, thereby leading to coarse NiCu-rich regions and Ni3Si phases. (2) The solidified Ni–Cr composite alloy contained α-Ni, α'-Ni, Cr, CrB, and Ni3Si phases and NiCu-rich regions. Mainly Ni atoms from the Ni–Cr alloy diffused into the Cu–Ce liquid and thus induced the formation of NiCu-rich phases. (3) The brazed diamond segment based on the composite alloy with the 5-wt.% Cu–Ce alloy addition exhibited the highest wear resistance and machining performance and the best surface morphology after the wear test. The excessive Cu– Ce alloy addition led to a rapid decrease in wear resistance of the brazed diamond segment owing to the large number of coarse NiCu-rich phases falling off from the composite alloy during the wear test. (4) The mechanism of the reduction in thermal damage to the brazed diamonds by the Cu–Ce alloy addition was elucidated. Initially the Cu–Ce particles melted and mainly Ni atoms diffused into the Cu–Ce liquid, thereby leading to the formation of NiCu-rich regions and Ce2Ni7 and CeNi2 phases, which in turn promoted the diffusion. 23
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Declaration of interest statement The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.