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Research Papers: Applications

A Numerical Investigation Into Cold Spray Bonding Processes

[+] Author and Article Information
Baran Yildirim, Teiichi Ando, Andrew Gouldstone

Department of Mechanical Engineering,
Northeastern University,
Boston, MA 02115

Hirotaka Fukanuma

Plasma Giken Co., Ltd.,
4-1 Imaichi, Yoriimachi, Osato-gun,
Saitama 369-1214, Japan

Sinan Müftü

Department of Mechanical Engineering,
Northeastern University,
Boston, MA 02115
e-mail: s.muftu@neu.edu

1Corresponding author.

Contributed by the Tribology Division of ASME for publication in the JOURNAL OF TRIBOLOGY. Manuscript received March 10, 2014; final manuscript received July 11, 2014; published online October 3, 2014. Assoc. Editor: George K. Nikas.

J. Tribol 137(1), 011102 (Oct 03, 2014) (13 pages) Paper No: TRIB-14-1053; doi: 10.1115/1.4028471 History: Received March 10, 2014; Revised July 11, 2014

Specific mechanisms underlying the critical velocity in cold gas particle spray applications are still being discussed, mainly due to limited access to in situ experimental observation and the complexity of modeling the particle impact process. In this work, particle bonding in the cold spray (CS) process was investigated by the finite element (FE) method. An effective interfacial cohesive strength parameter was defined in the particle–substrate contact regions. Impact of four different metals was simulated, using a range of impact velocities and varying the effective cohesive strength values. Deformation patterns of the particle and the substrate were characterized. It was shown that the use of interfacial cohesive strength leads to a critical particle impact velocity that demarcates a boundary between rebounding and bonding type responses of the system. Such critical bonding velocities were predicted for different interfacial cohesive strength values, suggesting that the bonding strength in particle–substrate interfaces could span a range that depends on the surface conditions of the particle and the substrate. It was also predicted that the quality of the particle bonding could be increased if the impact velocity exceeds the critical velocity. A method to predict a lower bound for the interfacial bonding energy was also presented. It was shown that the interfacial bonding energy for the different materials considered would have to be at least on the order of 10–60 J/m2 for cohesion to take place. The general methodology presented in this work can be extended to investigate various materials and impact conditions.

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References

Figures

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Fig. 1

Mesh structure and the schematic view of the FE model. Average mesh size around the impact region is dP/25.

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Fig. 2

Adiabatic stress–strain curves for each material at ɛ· = 107 based on the JC model (model parameters are provided in Table 2)

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Fig. 3

Deformation and plastic strain distributions for (a) copper, (b) aluminum, (c) 316 L steel, (d) titanium, (e) Cu-on-Al, and (f) Al-on-Cu at t = 200 ns for different impact velocities. Rebound/bonding behavior of the particle for σc = 200 MPa is shown. Note that simulations with different σc values show that it has no or little effect on the particle and substrate deformation patterns. Particle bonding is observed at (a-2), (a-3), (b-4), (c-4), (d-4), (f-3), and (f-4). Particle entrapment due to mechanical interlocking can be seen in (e-2) and (e-3).

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Fig. 4

Rebound kinetic energy of the particle (KER) for different effective cohesive strength values. Critical velocity is the velocity at which KER of the particle goes to zero. Notice the scale change of vertical axis in (a) and (f).

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Fig. 5

Calculated critical velocities (solid lines) for a dP = 25 μm particle for each particle–substrate material system along with the experimental results from literature (hollow circles). In (a), calculated critical velocities for different particles sizes are marked with solid triangles. Experimental data are taken from Refs. [5,6,23,28,50].

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Fig. 6

Time evolution of the contact area at different impact velocities (a) and effective cohesive strength values (b). (c) Area in contact (shown in dark gray) at different times for impact with VP = 750 m/s and σc = 400 MPa. Bonded area is the region that stays in contact during the entire duration of rebound phase. Particle and substrate materials are titanium.

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Fig. 7

(a) The normalized contact area (A¯c = Ac/2πdp2) and (b) the lower bound of the interfacial bonding energy, predicted by using the results of σc = 0, as a function of the damage parameter (ρVp2/2Y). The ratio of the kinetic energy of rebound to contact area (KER/Ac) is used as the lower bound of the interfacial bonding energy. The density and the dynamic yield stress values (ρ,Y) used in this analysis are (8940 kg/m3, 401 MPa) for copper, (4500 kg/m3, 959 MPa) for titanium, (2710 kg/m3, 391 MPa) for aluminum, and (8000 kg/m3, 829 MPa) for steel.

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Fig. 8

The effects of COF on (a) compression ratio, (b) normalized crater depth, and (c) rebound velocity

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Fig. 9

The effects of COF on the total interfacial contact force due to contact pressure and frictional stresses for the cases COF = 0.1 and COF = 0.9, at 400 m/s impact

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Fig. 10

Compression ratio (CR), normalized crater depth (CD), rebound velocity and maximum temperature for different mesh sizes

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Fig. 11

Deformed mesh structure of the particle and the substrate (a), and rebound velocity as a function of particle impact velocity (b) for failure strains εf = 1, εf = 2, and no failure condition (εf = ∞). Particle and substrate are both copper.

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