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Study on impact fusion at particle interfaces and its effect on coating microstructure in cold spraying
Applied Surface Science 254 (2007) 517–526 www.elsevier.com/locate/apsusc
Study on impact fusion at particle interfaces and its effect on coating microstructure in cold spraying Wen-Ya Li a,*, Chao Zhang a,b, Xueping Guo a, Chang-Jiu Li b, Hanlin Liao a, C. Coddet a a b
LERMPS, Universite´ de Technologie de Belfort-Montbe´liard, site de Se´venans, 90010 Belfort cedex, France State Key Laboratory for Mechanical Behavior of Materials, School of Materials Science and Engineering, Xi’an Jiaotong University, Xi’an 710049, PR China Received 12 October 2006; received in revised form 24 March 2007; accepted 16 June 2007 Available online 24 June 2007
Abstract This paper deals with the impact melting phenomenon at the interfaces between the deposited particles in cold-sprayed coatings and its effect on coating microstructure and particle bonding mechanism. Al-12Si, Al2319, Ti, Ti-6Al-4V, Ni and NiCoCrAlTaY powders were selected as feedstocks, which have various thermal and mechanical properties. The analytical results showed that most of the used materials possibly experienced the local melting at the contact interfaces of particles under certain impact conditions. Low melting point, relatively high gas temperature and chemical reaction with the atmosphere are the main factors contributing to the impact fusion during cold spraying. The results also indicated that the local melting would benefit the formation of a metallurgical bonding between the deposited particles and enhance the coating cohesion. # 2007 Elsevier B.V. All rights reserved. Keywords: Cold spraying; Particles; Aluminium alloy; Titanium alloy; Nickel alloy; Impact fusion
1. Introduction Cold spraying, as an emerging coating process in recent years, has attracted worldwide interest for its high deposition efficiency and volume production of many metallic coatings [1–16] and composites [16]. In some cases, even cermets [16,17], nanostructured coatings [17,18] and functional ceramic layer [19] could be deposited by cold spraying. In this process, spray particles (typically 5–50 mm) are accelerated to a high velocity by a high-speed gas jet generated through a convergent-divergent Laval type nozzle. A coating is usually formed through the intensive plastic deformation of particles impacting on a substrate at a temperature well below the melting point of the spray material. Therefore, cold spraying was endowed with some unique characteristics as compared to the conventional thermal spraying processes, such as almost no oxidation and phase transformation occur.
* Corresponding author. Tel.: +33 3 84583160; fax: +33 3 84583286. E-mail address: [email protected] (W.-Y. Li). 0169-4332/$ – see front matter # 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.apsusc.2007.06.026
Although many investigators have examined cold spraying process by both experiments [3–18,23–25] and numerical simulations [1–3,19–22,24] on the processing parameters, microstructure and corresponding properties of the deposits, there are still some underlying problems to be clarified concerning the bonding mechanism. The most prevailing hypothesis is that plastic deformation may disrupt thin surface films, such as oxides, and provide intimate conformal contact under high local pressure, thus permitting bonding to occur [24]. This kind of bonding is something like the so-called physical bonding by van der Waals force. And thus, this bonding process was considered to be comparable to that in processes such as explosive welding or shock wave powder compaction [1–3,20–23]. This bonding hypothesis can explain the observed critical velocity necessary to achieve a successful deposition and is consistent with the fact that a wide range of ductile materials, such as metals and alloys, have been deposited by cold spraying and the spray particles experience extensive deformation to form lens-like shapes. Non-ductile materials, such as ceramics, can only be deposited if they are co-cold-sprayed with a ductile (matrix) material or sprayed on a
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ductile substrate to form a thin layer [1–3,20–22,24]. Assadi et al. [1,2] and Grujicic et al. [20,21] argued that the adiabatic shear instability at the particle/substrate or particle/deposited material interfaces upon impacting contribute to the interfacial bonding. Grujicic et al. [21] further suggested that this interfacial instability may give rise to material mixing at the interface and provide mechanical interlocking between the two materials. However, many experimental results on coating adhesive strength (usually 15–60 MPa depending on materials and spray parameters [1,2,4,8,11,14,15]) suggested that the mechanical interlocking would be the main bonding mechanism in cold spraying, like that in the conventional thermal spraying processes [26]. It is known that, in explosive welding or shock wave powder compaction, the local melting of the interfaces provides a reliable metallurgical bonding. However, there are few reports indicating the local melting of coldsprayed particles except in some simulation results where the possible melting was estimated [1–3,20–22,24]. Zhang et al. indicated out the impact-induced melting of Sn substrate with Al particles generated by using helium as driving gas under a pressure of 2.5 MPa [25]. The previous study also showed that the impact fusion occurred in cold spraying of Sn and Zn [4,5]. The recent study indicated that the local melting between the deposited Zn particles could benefit the formation of metallurgical bonding and thus increased both the adhesion and cohesion of cold-sprayed Zn coatings [14]. Therefore, the study of impact fusion during cold spraying is of certain significance in understanding the nature of process. In this study, the impact-melting phenomenon at the interfaces of high-speed metallic particles in cold spraying and its effect on coating microstructure was investigated by both analytical description and experimental validation. 2. Analytical description It is well-known that in high-speed processes, such as highspeed machining, ballistic impact and penetration, the obvious heat generation at the contact interfaces is usually observed due to plastic deformation, especially under the adiabatic shearing condition [27,28]. It was considered that the adiabatic shearing would also occur when cold-sprayed particles impact on a substrate or the previously deposited coating surface under the velocities ranging from 300 to 1000 m/s depending on the spray conditions [1–3,20–23]. However, because of the very short duration of particle impact and the tiny particle size, it is very difficult to observe the whole deformation process of particles. Only the deformed particles can be observed by microscopy. Therefore, numerical methods were employed by some investigators to study the particle deformation process [1–3,20–23]. In this study, a simple analytical description was presented to investigate the particle impact melting. Firstly, taking into account the short-impacting process, typically tens of nanoseconds, it was assumed that the kinetic energy of a particle was converted to local thermal energy input, which originated from the adiabatic shearing [1]. The heat conduction within the particle was neglected as a first approximation. This assumption has been proved to be
reasonable by the simple calculation of thermal diffusion [1,2,20]. Moreover, variations of material properties such as density (r) and heat capacity (Cp) with plastic strain, stress or temperature are ignored. The temperature rise is based on the empirical assumption that part of the plastic work under adiabatic conditions is dissipated as heat, which is defined as b. Therefore, the temperature rise, DT, can be simplified to [28]: Z ep b DT ¼ sdep (1) rC p 0 where s is the flow stress and ep is the plastic strain. The kinetic energy (Ek) of a particle of a diameter of dp and a velocity of vp is as follows: 1 Ek ¼ mp v2p 2
(2)
where mp is the particle mass. It was assumed that only the thin contact zone, which experiences intensive deformation, was heated. Therefore, based on the energy conservation law, the temperature rise under the adiabatic shearing condition can be deduced as: Z T aEk ¼ f v mp C p dT ¼ f v mp C p DT (3) Tr
DT ¼
av2p aEk ¼ f v mp Cp 2 f v C p
(4)
where, a is the kinetic energy conversion factor to generate the temperature rise. f v is the volume fraction that the heated zone takes in the particle including the counter zone in the substrate. If the melting just occurs at a certain velocity ðvpm Þ, the energy balance will be as follows taking into consideration latent heat of fusion (HL): 1 amp v2pm ¼ f v mp ðCp ðT m T r Þ þ H L Þ 2
(5)
where Tm is the melting point of particle and Tr is the reference temperature, here it is taken as 20 8C. Therefore, the velocity for interface melting can be calculated as: rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 f v ðCp ðT m T r Þ þ H L Þ (6) vpm ¼ a According to Eq. (6), the velocity for interface melting of different materials can be estimated. Table 1 gives the properties of several typical materials [29–31] used in cold spraying. Fig. 1 shows the effects of ratio of kinetic energy conversion into heat (a) and volume fraction of melt ( f v) on the minimum velocity for impact fusion of different materials. It is clear that the less a and the higher f v are, the higher vpm is for all the materials. It is found that for the materials of low melting point, such as Sn and Zn, the interface melting should be easily satisfied with the relatively low particle velocities regardless of b and f v. This has been proved in the previous studies [4,5]. For Ti and Ti-6Al-4V, the calculated impact fusion velocity is relatively high owing to their high melting point and latent heat of fusion. However, the experimental results with these two
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Table 1 Properties of the selected materials [29–31] Material
Sn Zn Cu Al2319 Ni Al NiCoCrAl Al12Si Mg Ti-6Al-4V Ti
Density (kg/m3)
Heat capacity (J/(kg K))
Thermal conductivity (W/(m K))
Melting point (8C)
Latent heat of fusion (J/kg)
Thermal diffusivity (m2/s)
Tensile strength (MPa) Yield
Ultimate
7170 7131 8880 2840 8899 2688 8180 2660 1740 4420 4506
228 389 386 880 447 905 461 963 1020 537 522
66.6 121 398 120 90.5 237 11.5 121 156 7.6 21.9
232 420 1083 638 1455 660 1427 532 649 1660 1680
60600 102000 205000 382800 290000 395000 334000 496600 372000 418200 440000
4.07E-05 4.36E-05 1.16E-04 4.80E-05 2.28E-05 9.74E-05 3.05E-06 4.72E-05 8.79E-05 3.20E-06 9.31E-06
25 30 90 80 150 44 917 70 90 880 140
52 55 210 240 317 80 1207 155 115 950 220
materials showed that the impact fusion could also be easily fulfilled due to their reactions with oxygen during spraying [12], which will be shown in detail in the following section. As for Mg, although no experimental results could be given in this study, it is believed that its reactivity can also promote its impact fusion according to the previous primary experiment. As for Cu, it is seen from Fig. 1 that the calculated impact fusion velocity is also low. However, according to the experiment results, up to now, no evidence of interface melting was found for Cu particles within a large velocity range [1,2,6,24]. It was considered that the relatively high thermal diffusivity of Cu, as shown in Table 1, might contribute to this fact. But it needs further studies to clarify this point. As for other materials, they may also experience the interface melting under certain conditions in cold spraying. The detailed results and discussion will be presented in the following section. On the other hand, only some of the material properties, including heat capacity, latent heat of fusion and melting point, were considered in Eq. (6) for the simple calculation. Actually, the other material properties, such as density, thermal conductivity and strength, also exert influences on the complex
Fig. 1. Effects of ratio of kinetic energy conversion into heat (a) and volume fraction of melt ( fv) on the minimum velocity for impact fusion of different materials.
impacting process. For example, very low thermal diffusivity and high strength will be helpful to concentrate the thermal energy at the thin interfacial layers. However, a high strength will also increase the difficulty to deposit the particles through plastic deformation, such as for superalloys (MCrAlY) [13,16]. The low density of some light materials, such as Mg, Al and their alloys, will also be unfavorable because of their low kinetic energy upon impact as compared with the heavy particles at the same velocity and diameter. Finally, if the particles rebound after impacting, they will take off part of the kinetic energy. Therefore, the kinetic energy conversion ratio, a, will be very low and the impact fusion will be difficult to occur at the contact interfaces. 3. Experimental procedures The cold spray system was installed in LERMPS (UTBM, France) with a commercial cold spray gun (CGT GmbH, Germany). An optimized nozzle designed by LERMPS was employed in spraying, which had an expansion ratio of about 4.9 and a divergent section length of 170 mm. The highpressure compressed air was used as the accelerating gas and argon was used as powder carrier gas. The standoff distance from the nozzle exit to the substrate surface was 30 mm. Six types of powders were selected as feedstocks, including Al2319 (LERMPS, 5–63 mm), Al-12Si (Sulzer-Metco, Metco52C NS, +45–90 mm), Ti (Northwest Institute for Non-ferrous Metal Research, China, 5–45 mm), Ti-6Al-4V (LERMPS, 5–90 mm), Ni (Medipure 28 99.0%, 5–44 mm) and NiCoCrAlTaY (SulzerMetco, Amdry 997, 5–37 mm). The morphologies of these powders as investigated by scanning electron microscopy (SEM) are shown in Fig. 2. Ti powder was manufactured through a hydride–dehydride (HDH) process and exhibited an angular morphology. The other powders were produced through an atomization process. Mild steel plates were used as substrate and they were grid-blasted using alumina grits prior to spraying. The detailed spray conditions for the different powders are given in Table 2. The coating microstructures were examined by optical microscopy (OM) (Nikon, Japan), SEM (JSM5800LV, JEOL, Japan). For a better observation of the coating microstructure,
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Fig. 2. Powder morphologies according to SEM of the used (a) Al2319, (b) Al-12Si, (c) Ti, (d) Ti-6Al-4V, (e) Ni and (f) NiCoCrAlTaY powders.
some of the polished coatings were etched by a solution of 100 ml H2O + 3 ml HF + 6 ml HNO3. The porosities of deposits were estimated from more than 10 cross-sectional OM micrographs through image analysis with Scion Image software (NIH, USA). The adhesive strength of the coating was characterized using the ASTM standard C633-01. Coatings were produced on the mild steel disks having a 25.4 mm diameter and 10 mm thickness. After deposition, the uncoated surface of sample was sand-blasted, and then both the coating and uncoated surfaces were glued to two prepared samples (with internal screw) using an adhesive (FM1000, MOSAVIA Corporation, USA), having a nominal tensile strength up to 60 MPa. For each coating, three specimens were used. The
Table 2 Cold spray gas conditions for different powders Powder
Air pressure (MPa)
Air temperature (8C)
Adhesive strength (MPa)
Tensile failure position
Al2319 Al12Si Ti Ti-6Al-4V Ni NiCoCrAlTaY
2.8 2.7 2.8 2.8 3.0 3.0
520 560 520 520 585 620
34 4 >50 15 4 10 2 25 3 –
C/S in adhesive C/S C/S C/S –
Note: pressures and temperatures are measured at the upstream of nozzle; C/S: coating/substrate interface.
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cross-sectional fracture surfaces of different coatings obtained by bending the substrates were also examined by SEM. 4. Experimental results and discussion 4.1. General microstructure of the coatings After cold spraying, it was found that, except for the NiCoCrAlTaY powder, all the other powders could be deposited to form a thick coating. According to the previous study and the reported results [13,16], a thick MCrAlY coating could be produced solely by using helium as accelerating gas. In the present study, only a thin NiCoCrAlTaY coating (about 10 mm) was formed by using air as accelerating gas. Fig. 3 shows the OM micrographs of as-sprayed Al2319, Al-12Si, Ti, Ti-6Al-4V and Ni coatings. It can be seen that, except for the dense Al-12Si coating, the other coatings presented many pores, especially Ti and Ti-6Al-4V coatings as shown in Fig. 3(c) and (d). The average porosities of these coatings were estimated to be about 1.9, 5.1, 22.4 and 0.9% for Al2319, Ti, Ti6Al-4V and Ni coatings, respectively. Generally, the extent of deformation of the deposited particles accounts mainly for the porosity of the cold-sprayed coatings. The extent of deformation of a particle is determined by its strength, as well as the density, which will influence the kinetic energy of particle at the same velocity. As reported in literature, it is easy to produce dense coatings of pure Cu, Zn, etc. owing to their good deformability and high density [1,2,4–6,16,24], whereas it is difficult to form a dense Al coating due to its low density [7,8]. For some materials of relatively high strength, such as Ni, stainless steel and MCrAlY [13,16], it is also difficult to form a dense coating except at a relatively high particle velocity, such as using helium as the accelerating gas. As for Ti and Ti-6Al-4V coatings, the porous structure mainly resulted from their reactions with the oxygen in the entrained air or the employed air [12]. If one just looks at the strength of Ti and its reported critical velocity of about 750 m/s [2], a deposition efficiency of less than 5% should be obtained under the spray conditions in this study according to the size distribution of Ti powder and the calculated particle velocities [12]. However, the actual deposition efficiency of Ti powder in this study was experimentally estimated to be about 75%. As for Ti-6Al-4V powder, there would be almost no coating deposition, taking into consideration its high strength according to its size distribution and the calculated particle velocities [12]. But the experimental deposition efficiency of Ti-6Al-4V powder was estimated to be about 60% in this study. During spraying of Ti and Ti-6Al-4V, a much bright flashing jet was clearly observed as shown in Fig. 4. It was considered that the particles rub with the high speed driving gas and/or nozzle inner wall, which could break up the oxide film and generate a relatively high particle surface temperature. When the particles fly through the nozzle, their fresh surfaces will be oxidized owing to the presence of oxygen in the employed air, and thus the flashing jet is generated [12]. In the local contact interfaces of the deposited Ti and Ti-6Al-4V particles, a metallurgical bonding would be formed owing to the relatively high
Fig. 3. OM micrographs of (a) Al2319, (b) Al-12Si, (c) Ti, (d) Ti-6Al-4V and (e) Ni coatings.
Fig. 4. Photo of the flashing jet with Ti particles outside nozzle exit with air preheating at 520 8C. Jet length from nozzle exit was about 40 cm.
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interfacial temperature resulting from both the chemical reaction and the adiabatic impacting process. The detailed results will be shown in the following section. 4.2. Interface microstructures of the coatings Fig. 5 shows the SEM micrographs of the etched cross-section and of the fractured surface of an Al2319 coating. It is observed that the Al2319 particles in this coating have experienced intensive plastic deformation. But there are still some pores between the particles. It was found that a metallurgical bonding could be formed at some local interfaces in Al2319 coating as marked by arrows in Fig. 5(b). The adhesive strength of Al2319 coating from the pull-off test (Table 2) was about 34 MPa, which is relatively high, taking into account its strength (Table 1). The tensile fracture occurred at the coating/substrate interface with many Al2319 particles remaining on the substrate. It was considered that the metallurgical bonding could be associated with the impact fusion during deposition. From the fractured surface, by bending, the local melting could be observed with a ductile fracture as marked by arrow in Fig. 5(d). The first spray trials with Al-12Si powder showed that a very thin coating was produced with a gas temperature of about 400 8C. However, it was found that the deposited particles have intensively flattened like those of a thermally sprayed coating and the limited surface melting was observed on the flattened particles due to the relatively high gas temperature [15]. When the gas temperature was increased to 560 8C (close to Al-12Si
melting point of 577 8C), these phenomena were more obvious [15]. The adhesive strength of Al-12Si coating was not less than 50 MPa (Table 2), which is relatively high taking into account its strength (Table 1). Fig. 6 shows the as-sprayed and fractured surface morphologies of Al-12Si coatings deposited at a gas temperature of 560 8C. The evidence of melting on a crater caused by the rebounded particle could be observed as marked by arrows in Fig. 6(b). That is similar to results observed for cold-sprayed Zn coatings [4,5]. From the fractured surface by bending, it is seen that the failure mainly occurred across the flattened particles with a typical ductile fracture. Many dimples were presented at the fractured zones. This fact may suggest a fine bonding between the deposited particles. Fig. 7 shows the SEM micrographs of the etched crosssection and of the fractured surface of a Ti coating. It is found that the deposited Ti particles did not deform so much taking into account the angular morphology of the feedstock. It is also observed that at some interfaces between the deposited particles, metallurgical bonding could be formed as marked by arrows in Fig. 7(b). The possible melting zones could also be observed from the fractured surface as indicated by arrows in Fig. 7(d). Although the adhesive strength of Ti coating was about 15 MPa, this metallurgical bonding may enhance the coating cohesion taking into its relatively high porosity. Consequently, Ti particles can be deposited with high deposition efficiency, even without an extensive deformation. Such phenomena could be observed more clearly for Ti-6Al-4V coatings. Fig. 8 shows the SEM micrographs of the etched
Fig. 5. SEM micrographs of cross-section of Al2319 coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrow in (d) indicates the ductile fracture appearing as small dimples.
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Fig. 6. SEM surface morphologies of as-sprayed Al-12Si coating (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local melting. (d) indicates the ductile fracture appearing as small dimples.
Fig. 7. SEM micrographs of cross-section of Ti coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrows in (d) indicate the ductile fracture appearing as small dimples.
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Fig. 8. SEM micrographs of cross-section of Ti-6Al-4V coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrows in (d) indicate the ductile fracture appearing as small dimples.
Fig. 9. SEM surface morphologies of Ni coating (a,b) and fractured surface morphologies (c,d).
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Fig. 10. SEM surface morphologies of NiCoCrAlTaY coating. The arrow in (b) may indicate the interface melting.
cross-section and of the fractured surface of a Ti-6Al-4V coating. It is clear that the deposited Ti-6Al-4V particles experienced little deformation. At the local contact interfaces, the metallurgical bonging could be formed as marked by the arrows in Fig. 8(b). The melting zones can also be observed from the fractured surface as indicated by arrows in Fig. 8(d). Therefore, the particles can be adhered together with a little contact area and without obvious plastic deformation thanks to their active surfaces. However, the strength of this porous Ti6Al-4V coating was about 10 MPa (Table 2), which is relatively low and similar to that of not more than 15 MPa as reported by Blose using helium as an accelerating gas [11]. Therefore, as structural coatings, it is necessary to improve their mechanical properties. It has been proved that post-spray heat treatment would be an effective method [11]. Fig. 9 shows the as-sprayed and fractured surface morphologies of a Ni coating. Not like other coatings, although the analytical description showed the possibility of melting for Ni particles, it was difficult to observe the local melting zone under the spray conditions in this study. Similar to that observed for cold-sprayed Cu coatings [1,6,16], the interfaces between the flattened particles appeared as a little smooth surfaces as shown in Fig. 9(c) because no ductile fracture was observed. The adhesive strength of Ni coating was relatively low, which was about 25 MPa as shown in Table 2. As for the NiCoCrAlTaY coating, although it seemed the local melting might occurred in some zones as marked by arrow in Fig. 10(b), further studies are necessary to clarify the possible localized melting in the Ni and Ni based alloy coatings deposited using helium as accelerating gas. 5. Concluding remarks Based on analytical descriptions and experimental results on the impact fusion of particles during cold spraying, the impact fusion and its effect on coating microstructure were investigated in the present study. It is found from analytical results that most metallic spray materials possibly experience the local melting at the contact interfaces between the deposited particles in cold-sprayed coatings, which is dependent on the adopted spray conditions. According to analytical descriptions, a high kinetic energy conversion ratio and low volume fraction of
intensive deformation will yield the most possible impact fusion. The experimental results showed that low melting point, relatively high gas temperature and chemical reaction with the atmosphere are the main factors contributing to the impact fusion for different spray materials under certain impact conditions in cold spraying. The poor thermal conductivity also helps the occurrence of local melting. The experimental results also showed that the local melting would benefit the formation of a metallurgical bonding between the deposited particles and enhance the coating cohesion.
Acknowledgements This work was finanically supported by the Franche-Comte Regional Council of France. The authors would like to thank Lucas Dembinski of LERMPS for the supply of Al2319 and Ti6Al4V powders.
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Study on impact fusion at particle interfaces and its effect on coating microstructure in cold spraying Wen-Ya Li a,*, Chao Zhang a,b, Xueping Guo a, Chang-Jiu Li b, Hanlin Liao a, C. Coddet a a b
LERMPS, Universite´ de Technologie de Belfort-Montbe´liard, site de Se´venans, 90010 Belfort cedex, France State Key Laboratory for Mechanical Behavior of Materials, School of Materials Science and Engineering, Xi’an Jiaotong University, Xi’an 710049, PR China Received 12 October 2006; received in revised form 24 March 2007; accepted 16 June 2007 Available online 24 June 2007
Abstract This paper deals with the impact melting phenomenon at the interfaces between the deposited particles in cold-sprayed coatings and its effect on coating microstructure and particle bonding mechanism. Al-12Si, Al2319, Ti, Ti-6Al-4V, Ni and NiCoCrAlTaY powders were selected as feedstocks, which have various thermal and mechanical properties. The analytical results showed that most of the used materials possibly experienced the local melting at the contact interfaces of particles under certain impact conditions. Low melting point, relatively high gas temperature and chemical reaction with the atmosphere are the main factors contributing to the impact fusion during cold spraying. The results also indicated that the local melting would benefit the formation of a metallurgical bonding between the deposited particles and enhance the coating cohesion. # 2007 Elsevier B.V. All rights reserved. Keywords: Cold spraying; Particles; Aluminium alloy; Titanium alloy; Nickel alloy; Impact fusion
1. Introduction Cold spraying, as an emerging coating process in recent years, has attracted worldwide interest for its high deposition efficiency and volume production of many metallic coatings [1–16] and composites [16]. In some cases, even cermets [16,17], nanostructured coatings [17,18] and functional ceramic layer [19] could be deposited by cold spraying. In this process, spray particles (typically 5–50 mm) are accelerated to a high velocity by a high-speed gas jet generated through a convergent-divergent Laval type nozzle. A coating is usually formed through the intensive plastic deformation of particles impacting on a substrate at a temperature well below the melting point of the spray material. Therefore, cold spraying was endowed with some unique characteristics as compared to the conventional thermal spraying processes, such as almost no oxidation and phase transformation occur.
* Corresponding author. Tel.: +33 3 84583160; fax: +33 3 84583286. E-mail address: [email protected] (W.-Y. Li). 0169-4332/$ – see front matter # 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.apsusc.2007.06.026
Although many investigators have examined cold spraying process by both experiments [3–18,23–25] and numerical simulations [1–3,19–22,24] on the processing parameters, microstructure and corresponding properties of the deposits, there are still some underlying problems to be clarified concerning the bonding mechanism. The most prevailing hypothesis is that plastic deformation may disrupt thin surface films, such as oxides, and provide intimate conformal contact under high local pressure, thus permitting bonding to occur [24]. This kind of bonding is something like the so-called physical bonding by van der Waals force. And thus, this bonding process was considered to be comparable to that in processes such as explosive welding or shock wave powder compaction [1–3,20–23]. This bonding hypothesis can explain the observed critical velocity necessary to achieve a successful deposition and is consistent with the fact that a wide range of ductile materials, such as metals and alloys, have been deposited by cold spraying and the spray particles experience extensive deformation to form lens-like shapes. Non-ductile materials, such as ceramics, can only be deposited if they are co-cold-sprayed with a ductile (matrix) material or sprayed on a
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ductile substrate to form a thin layer [1–3,20–22,24]. Assadi et al. [1,2] and Grujicic et al. [20,21] argued that the adiabatic shear instability at the particle/substrate or particle/deposited material interfaces upon impacting contribute to the interfacial bonding. Grujicic et al. [21] further suggested that this interfacial instability may give rise to material mixing at the interface and provide mechanical interlocking between the two materials. However, many experimental results on coating adhesive strength (usually 15–60 MPa depending on materials and spray parameters [1,2,4,8,11,14,15]) suggested that the mechanical interlocking would be the main bonding mechanism in cold spraying, like that in the conventional thermal spraying processes [26]. It is known that, in explosive welding or shock wave powder compaction, the local melting of the interfaces provides a reliable metallurgical bonding. However, there are few reports indicating the local melting of coldsprayed particles except in some simulation results where the possible melting was estimated [1–3,20–22,24]. Zhang et al. indicated out the impact-induced melting of Sn substrate with Al particles generated by using helium as driving gas under a pressure of 2.5 MPa [25]. The previous study also showed that the impact fusion occurred in cold spraying of Sn and Zn [4,5]. The recent study indicated that the local melting between the deposited Zn particles could benefit the formation of metallurgical bonding and thus increased both the adhesion and cohesion of cold-sprayed Zn coatings [14]. Therefore, the study of impact fusion during cold spraying is of certain significance in understanding the nature of process. In this study, the impact-melting phenomenon at the interfaces of high-speed metallic particles in cold spraying and its effect on coating microstructure was investigated by both analytical description and experimental validation. 2. Analytical description It is well-known that in high-speed processes, such as highspeed machining, ballistic impact and penetration, the obvious heat generation at the contact interfaces is usually observed due to plastic deformation, especially under the adiabatic shearing condition [27,28]. It was considered that the adiabatic shearing would also occur when cold-sprayed particles impact on a substrate or the previously deposited coating surface under the velocities ranging from 300 to 1000 m/s depending on the spray conditions [1–3,20–23]. However, because of the very short duration of particle impact and the tiny particle size, it is very difficult to observe the whole deformation process of particles. Only the deformed particles can be observed by microscopy. Therefore, numerical methods were employed by some investigators to study the particle deformation process [1–3,20–23]. In this study, a simple analytical description was presented to investigate the particle impact melting. Firstly, taking into account the short-impacting process, typically tens of nanoseconds, it was assumed that the kinetic energy of a particle was converted to local thermal energy input, which originated from the adiabatic shearing [1]. The heat conduction within the particle was neglected as a first approximation. This assumption has been proved to be
reasonable by the simple calculation of thermal diffusion [1,2,20]. Moreover, variations of material properties such as density (r) and heat capacity (Cp) with plastic strain, stress or temperature are ignored. The temperature rise is based on the empirical assumption that part of the plastic work under adiabatic conditions is dissipated as heat, which is defined as b. Therefore, the temperature rise, DT, can be simplified to [28]: Z ep b DT ¼ sdep (1) rC p 0 where s is the flow stress and ep is the plastic strain. The kinetic energy (Ek) of a particle of a diameter of dp and a velocity of vp is as follows: 1 Ek ¼ mp v2p 2
(2)
where mp is the particle mass. It was assumed that only the thin contact zone, which experiences intensive deformation, was heated. Therefore, based on the energy conservation law, the temperature rise under the adiabatic shearing condition can be deduced as: Z T aEk ¼ f v mp C p dT ¼ f v mp C p DT (3) Tr
DT ¼
av2p aEk ¼ f v mp Cp 2 f v C p
(4)
where, a is the kinetic energy conversion factor to generate the temperature rise. f v is the volume fraction that the heated zone takes in the particle including the counter zone in the substrate. If the melting just occurs at a certain velocity ðvpm Þ, the energy balance will be as follows taking into consideration latent heat of fusion (HL): 1 amp v2pm ¼ f v mp ðCp ðT m T r Þ þ H L Þ 2
(5)
where Tm is the melting point of particle and Tr is the reference temperature, here it is taken as 20 8C. Therefore, the velocity for interface melting can be calculated as: rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 f v ðCp ðT m T r Þ þ H L Þ (6) vpm ¼ a According to Eq. (6), the velocity for interface melting of different materials can be estimated. Table 1 gives the properties of several typical materials [29–31] used in cold spraying. Fig. 1 shows the effects of ratio of kinetic energy conversion into heat (a) and volume fraction of melt ( f v) on the minimum velocity for impact fusion of different materials. It is clear that the less a and the higher f v are, the higher vpm is for all the materials. It is found that for the materials of low melting point, such as Sn and Zn, the interface melting should be easily satisfied with the relatively low particle velocities regardless of b and f v. This has been proved in the previous studies [4,5]. For Ti and Ti-6Al-4V, the calculated impact fusion velocity is relatively high owing to their high melting point and latent heat of fusion. However, the experimental results with these two
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Table 1 Properties of the selected materials [29–31] Material
Sn Zn Cu Al2319 Ni Al NiCoCrAl Al12Si Mg Ti-6Al-4V Ti
Density (kg/m3)
Heat capacity (J/(kg K))
Thermal conductivity (W/(m K))
Melting point (8C)
Latent heat of fusion (J/kg)
Thermal diffusivity (m2/s)
Tensile strength (MPa) Yield
Ultimate
7170 7131 8880 2840 8899 2688 8180 2660 1740 4420 4506
228 389 386 880 447 905 461 963 1020 537 522
66.6 121 398 120 90.5 237 11.5 121 156 7.6 21.9
232 420 1083 638 1455 660 1427 532 649 1660 1680
60600 102000 205000 382800 290000 395000 334000 496600 372000 418200 440000
4.07E-05 4.36E-05 1.16E-04 4.80E-05 2.28E-05 9.74E-05 3.05E-06 4.72E-05 8.79E-05 3.20E-06 9.31E-06
25 30 90 80 150 44 917 70 90 880 140
52 55 210 240 317 80 1207 155 115 950 220
materials showed that the impact fusion could also be easily fulfilled due to their reactions with oxygen during spraying [12], which will be shown in detail in the following section. As for Mg, although no experimental results could be given in this study, it is believed that its reactivity can also promote its impact fusion according to the previous primary experiment. As for Cu, it is seen from Fig. 1 that the calculated impact fusion velocity is also low. However, according to the experiment results, up to now, no evidence of interface melting was found for Cu particles within a large velocity range [1,2,6,24]. It was considered that the relatively high thermal diffusivity of Cu, as shown in Table 1, might contribute to this fact. But it needs further studies to clarify this point. As for other materials, they may also experience the interface melting under certain conditions in cold spraying. The detailed results and discussion will be presented in the following section. On the other hand, only some of the material properties, including heat capacity, latent heat of fusion and melting point, were considered in Eq. (6) for the simple calculation. Actually, the other material properties, such as density, thermal conductivity and strength, also exert influences on the complex
Fig. 1. Effects of ratio of kinetic energy conversion into heat (a) and volume fraction of melt ( fv) on the minimum velocity for impact fusion of different materials.
impacting process. For example, very low thermal diffusivity and high strength will be helpful to concentrate the thermal energy at the thin interfacial layers. However, a high strength will also increase the difficulty to deposit the particles through plastic deformation, such as for superalloys (MCrAlY) [13,16]. The low density of some light materials, such as Mg, Al and their alloys, will also be unfavorable because of their low kinetic energy upon impact as compared with the heavy particles at the same velocity and diameter. Finally, if the particles rebound after impacting, they will take off part of the kinetic energy. Therefore, the kinetic energy conversion ratio, a, will be very low and the impact fusion will be difficult to occur at the contact interfaces. 3. Experimental procedures The cold spray system was installed in LERMPS (UTBM, France) with a commercial cold spray gun (CGT GmbH, Germany). An optimized nozzle designed by LERMPS was employed in spraying, which had an expansion ratio of about 4.9 and a divergent section length of 170 mm. The highpressure compressed air was used as the accelerating gas and argon was used as powder carrier gas. The standoff distance from the nozzle exit to the substrate surface was 30 mm. Six types of powders were selected as feedstocks, including Al2319 (LERMPS, 5–63 mm), Al-12Si (Sulzer-Metco, Metco52C NS, +45–90 mm), Ti (Northwest Institute for Non-ferrous Metal Research, China, 5–45 mm), Ti-6Al-4V (LERMPS, 5–90 mm), Ni (Medipure 28 99.0%, 5–44 mm) and NiCoCrAlTaY (SulzerMetco, Amdry 997, 5–37 mm). The morphologies of these powders as investigated by scanning electron microscopy (SEM) are shown in Fig. 2. Ti powder was manufactured through a hydride–dehydride (HDH) process and exhibited an angular morphology. The other powders were produced through an atomization process. Mild steel plates were used as substrate and they were grid-blasted using alumina grits prior to spraying. The detailed spray conditions for the different powders are given in Table 2. The coating microstructures were examined by optical microscopy (OM) (Nikon, Japan), SEM (JSM5800LV, JEOL, Japan). For a better observation of the coating microstructure,
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Fig. 2. Powder morphologies according to SEM of the used (a) Al2319, (b) Al-12Si, (c) Ti, (d) Ti-6Al-4V, (e) Ni and (f) NiCoCrAlTaY powders.
some of the polished coatings were etched by a solution of 100 ml H2O + 3 ml HF + 6 ml HNO3. The porosities of deposits were estimated from more than 10 cross-sectional OM micrographs through image analysis with Scion Image software (NIH, USA). The adhesive strength of the coating was characterized using the ASTM standard C633-01. Coatings were produced on the mild steel disks having a 25.4 mm diameter and 10 mm thickness. After deposition, the uncoated surface of sample was sand-blasted, and then both the coating and uncoated surfaces were glued to two prepared samples (with internal screw) using an adhesive (FM1000, MOSAVIA Corporation, USA), having a nominal tensile strength up to 60 MPa. For each coating, three specimens were used. The
Table 2 Cold spray gas conditions for different powders Powder
Air pressure (MPa)
Air temperature (8C)
Adhesive strength (MPa)
Tensile failure position
Al2319 Al12Si Ti Ti-6Al-4V Ni NiCoCrAlTaY
2.8 2.7 2.8 2.8 3.0 3.0
520 560 520 520 585 620
34 4 >50 15 4 10 2 25 3 –
C/S in adhesive C/S C/S C/S –
Note: pressures and temperatures are measured at the upstream of nozzle; C/S: coating/substrate interface.
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cross-sectional fracture surfaces of different coatings obtained by bending the substrates were also examined by SEM. 4. Experimental results and discussion 4.1. General microstructure of the coatings After cold spraying, it was found that, except for the NiCoCrAlTaY powder, all the other powders could be deposited to form a thick coating. According to the previous study and the reported results [13,16], a thick MCrAlY coating could be produced solely by using helium as accelerating gas. In the present study, only a thin NiCoCrAlTaY coating (about 10 mm) was formed by using air as accelerating gas. Fig. 3 shows the OM micrographs of as-sprayed Al2319, Al-12Si, Ti, Ti-6Al-4V and Ni coatings. It can be seen that, except for the dense Al-12Si coating, the other coatings presented many pores, especially Ti and Ti-6Al-4V coatings as shown in Fig. 3(c) and (d). The average porosities of these coatings were estimated to be about 1.9, 5.1, 22.4 and 0.9% for Al2319, Ti, Ti6Al-4V and Ni coatings, respectively. Generally, the extent of deformation of the deposited particles accounts mainly for the porosity of the cold-sprayed coatings. The extent of deformation of a particle is determined by its strength, as well as the density, which will influence the kinetic energy of particle at the same velocity. As reported in literature, it is easy to produce dense coatings of pure Cu, Zn, etc. owing to their good deformability and high density [1,2,4–6,16,24], whereas it is difficult to form a dense Al coating due to its low density [7,8]. For some materials of relatively high strength, such as Ni, stainless steel and MCrAlY [13,16], it is also difficult to form a dense coating except at a relatively high particle velocity, such as using helium as the accelerating gas. As for Ti and Ti-6Al-4V coatings, the porous structure mainly resulted from their reactions with the oxygen in the entrained air or the employed air [12]. If one just looks at the strength of Ti and its reported critical velocity of about 750 m/s [2], a deposition efficiency of less than 5% should be obtained under the spray conditions in this study according to the size distribution of Ti powder and the calculated particle velocities [12]. However, the actual deposition efficiency of Ti powder in this study was experimentally estimated to be about 75%. As for Ti-6Al-4V powder, there would be almost no coating deposition, taking into consideration its high strength according to its size distribution and the calculated particle velocities [12]. But the experimental deposition efficiency of Ti-6Al-4V powder was estimated to be about 60% in this study. During spraying of Ti and Ti-6Al-4V, a much bright flashing jet was clearly observed as shown in Fig. 4. It was considered that the particles rub with the high speed driving gas and/or nozzle inner wall, which could break up the oxide film and generate a relatively high particle surface temperature. When the particles fly through the nozzle, their fresh surfaces will be oxidized owing to the presence of oxygen in the employed air, and thus the flashing jet is generated [12]. In the local contact interfaces of the deposited Ti and Ti-6Al-4V particles, a metallurgical bonding would be formed owing to the relatively high
Fig. 3. OM micrographs of (a) Al2319, (b) Al-12Si, (c) Ti, (d) Ti-6Al-4V and (e) Ni coatings.
Fig. 4. Photo of the flashing jet with Ti particles outside nozzle exit with air preheating at 520 8C. Jet length from nozzle exit was about 40 cm.
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interfacial temperature resulting from both the chemical reaction and the adiabatic impacting process. The detailed results will be shown in the following section. 4.2. Interface microstructures of the coatings Fig. 5 shows the SEM micrographs of the etched cross-section and of the fractured surface of an Al2319 coating. It is observed that the Al2319 particles in this coating have experienced intensive plastic deformation. But there are still some pores between the particles. It was found that a metallurgical bonding could be formed at some local interfaces in Al2319 coating as marked by arrows in Fig. 5(b). The adhesive strength of Al2319 coating from the pull-off test (Table 2) was about 34 MPa, which is relatively high, taking into account its strength (Table 1). The tensile fracture occurred at the coating/substrate interface with many Al2319 particles remaining on the substrate. It was considered that the metallurgical bonding could be associated with the impact fusion during deposition. From the fractured surface, by bending, the local melting could be observed with a ductile fracture as marked by arrow in Fig. 5(d). The first spray trials with Al-12Si powder showed that a very thin coating was produced with a gas temperature of about 400 8C. However, it was found that the deposited particles have intensively flattened like those of a thermally sprayed coating and the limited surface melting was observed on the flattened particles due to the relatively high gas temperature [15]. When the gas temperature was increased to 560 8C (close to Al-12Si
melting point of 577 8C), these phenomena were more obvious [15]. The adhesive strength of Al-12Si coating was not less than 50 MPa (Table 2), which is relatively high taking into account its strength (Table 1). Fig. 6 shows the as-sprayed and fractured surface morphologies of Al-12Si coatings deposited at a gas temperature of 560 8C. The evidence of melting on a crater caused by the rebounded particle could be observed as marked by arrows in Fig. 6(b). That is similar to results observed for cold-sprayed Zn coatings [4,5]. From the fractured surface by bending, it is seen that the failure mainly occurred across the flattened particles with a typical ductile fracture. Many dimples were presented at the fractured zones. This fact may suggest a fine bonding between the deposited particles. Fig. 7 shows the SEM micrographs of the etched crosssection and of the fractured surface of a Ti coating. It is found that the deposited Ti particles did not deform so much taking into account the angular morphology of the feedstock. It is also observed that at some interfaces between the deposited particles, metallurgical bonding could be formed as marked by arrows in Fig. 7(b). The possible melting zones could also be observed from the fractured surface as indicated by arrows in Fig. 7(d). Although the adhesive strength of Ti coating was about 15 MPa, this metallurgical bonding may enhance the coating cohesion taking into its relatively high porosity. Consequently, Ti particles can be deposited with high deposition efficiency, even without an extensive deformation. Such phenomena could be observed more clearly for Ti-6Al-4V coatings. Fig. 8 shows the SEM micrographs of the etched
Fig. 5. SEM micrographs of cross-section of Al2319 coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrow in (d) indicates the ductile fracture appearing as small dimples.
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Fig. 6. SEM surface morphologies of as-sprayed Al-12Si coating (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local melting. (d) indicates the ductile fracture appearing as small dimples.
Fig. 7. SEM micrographs of cross-section of Ti coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrows in (d) indicate the ductile fracture appearing as small dimples.
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Fig. 8. SEM micrographs of cross-section of Ti-6Al-4V coating in the etched state (a,b) and fractured surface morphologies (c,d). The arrows in (b) indicate the local bonding between the deposited particles. The arrows in (d) indicate the ductile fracture appearing as small dimples.
Fig. 9. SEM surface morphologies of Ni coating (a,b) and fractured surface morphologies (c,d).
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Fig. 10. SEM surface morphologies of NiCoCrAlTaY coating. The arrow in (b) may indicate the interface melting.
cross-section and of the fractured surface of a Ti-6Al-4V coating. It is clear that the deposited Ti-6Al-4V particles experienced little deformation. At the local contact interfaces, the metallurgical bonging could be formed as marked by the arrows in Fig. 8(b). The melting zones can also be observed from the fractured surface as indicated by arrows in Fig. 8(d). Therefore, the particles can be adhered together with a little contact area and without obvious plastic deformation thanks to their active surfaces. However, the strength of this porous Ti6Al-4V coating was about 10 MPa (Table 2), which is relatively low and similar to that of not more than 15 MPa as reported by Blose using helium as an accelerating gas [11]. Therefore, as structural coatings, it is necessary to improve their mechanical properties. It has been proved that post-spray heat treatment would be an effective method [11]. Fig. 9 shows the as-sprayed and fractured surface morphologies of a Ni coating. Not like other coatings, although the analytical description showed the possibility of melting for Ni particles, it was difficult to observe the local melting zone under the spray conditions in this study. Similar to that observed for cold-sprayed Cu coatings [1,6,16], the interfaces between the flattened particles appeared as a little smooth surfaces as shown in Fig. 9(c) because no ductile fracture was observed. The adhesive strength of Ni coating was relatively low, which was about 25 MPa as shown in Table 2. As for the NiCoCrAlTaY coating, although it seemed the local melting might occurred in some zones as marked by arrow in Fig. 10(b), further studies are necessary to clarify the possible localized melting in the Ni and Ni based alloy coatings deposited using helium as accelerating gas. 5. Concluding remarks Based on analytical descriptions and experimental results on the impact fusion of particles during cold spraying, the impact fusion and its effect on coating microstructure were investigated in the present study. It is found from analytical results that most metallic spray materials possibly experience the local melting at the contact interfaces between the deposited particles in cold-sprayed coatings, which is dependent on the adopted spray conditions. According to analytical descriptions, a high kinetic energy conversion ratio and low volume fraction of
intensive deformation will yield the most possible impact fusion. The experimental results showed that low melting point, relatively high gas temperature and chemical reaction with the atmosphere are the main factors contributing to the impact fusion for different spray materials under certain impact conditions in cold spraying. The poor thermal conductivity also helps the occurrence of local melting. The experimental results also showed that the local melting would benefit the formation of a metallurgical bonding between the deposited particles and enhance the coating cohesion.
Acknowledgements This work was finanically supported by the Franche-Comte Regional Council of France. The authors would like to thank Lucas Dembinski of LERMPS for the supply of Al2319 and Ti6Al4V powders.
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