Interface phenomena in processing of ceramic reinforced amorphous metal composites

,O”RNAL

OF

NON-CRYSTAGIINESO~ ELSEVIER

Journal of Non-Crystalline Solids 205-207 (1996) 742-747

Interface phenomena in processing of ceramic reinforced amorphous metal composites G. Kaptay a,* , P. Bkzy

a, F. Szigeti b, A. Lovas ‘, Z. G&i

a, L. Boly6n a

a University of Miskolc, 3515 Miskolc Egyetemonros, Hungn~y b Agriculiural Uniuersiv of Gijdijll~?, Gb’diilB, Hungary ’ KFKI, SziIn’rdtes@zikai KI, Hungary

Abstract The role of interface adhesion on the wear resistance of ceramic reinforcedamorphous metalmatrix composites (AMMC) was studied both theoretically and experimentally. Theoretically, the wear resistance is expected to improve with wettability. Calculationsshowthat the Fe,,Ni,,Si,,B6 melt shouldwet WC particles well, but wet Sic particles poorly. Therefore

betterwearresistance is expectedfor AMMC usingWC ratherthan Sic particles.Evidencefor betterwetting, higherwear resistance and better abrasive ability of the Fe,,Ni,,Si,,B,-WC presented.

1. Introduction

Amorphous metallic matrixes have been produced by the melt spinning method for over 15 years. A number of papers have recently been published in the field of ceramic reinforced amorphous metal matrix composites(AMMC) as well (for a review, see Ref. [l]). AMMCs can have certain improved properties compared to those of ‘simple’ metal matrix composites (MMCS). One of the properties of interest is the wear resistanceof the compositematerial [2]. The amorphousmetallic ribbons are produced by the melt spinning method. Ceramic particles can be blown into the melt puddle formed on the roller of the app’aratus.The wear resistanceof the composite

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system compared to the Fe,,Ni,,Si,,B,-SiC

system is

will be higher with a higher surface concentration of the ceramic particles, and with higher wear resistance of the particles themselves. However, this is not always the case.If the particles delaminateunder the influence of friction forces, the composite material will actually have a smaller wear resistancethan the matrix itself. Therefore it is essentialto choose ceramic-metallic couples with as high a total adhesion between the ceramic particles and the solidified matrix as possible. In the present paper, first, the total adhesionbetween the matrix and the particles will be addressed theoretically. Then, for two metal/ceramic couples (Fe,,Ni,,Si,,B,/SiC and Fe,,Ni,,Si,,B,/WC), real numbers will be evaluated and used for calculations in order to find the better candidate (Sic or WC) for producing AMMC based on Fe,,Ni,,Si,,B, alloys. Finally, results of wear resistanceand abrasiontests will be presented for the two AMMCs studied.

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G. Kapray

2. Theoretical

et al./Joorrmal

of Non-Crystalline

considerations

2.1. The spec@ the particles

adhesion between the matrix and

The adhesion energy between the solid ceramic and solid metal can be calculated based on the adhesion energy between the ceramic and the liquid metal in the following form [3]: Wad(sA

- sB) =

(v,~/V,~)2’3W,d,,*-,~)~

(1)

where W denotes adhesion energies between solid A-liquid B and solid A-solid B, and V the molar volume of the liquid and solid metal B.Based on Eq. (l), one can see that higher adhesion energy between the ceramic and the melt will lead also to higher adhesion energies between the same ceramic and the solid, resulting from the solidification of the given liquid (we suppose there are no major structural changes during solidification, which is definitely the case for amorphous metal composites). In other words, the more the liquid metal wets the ceramics, the higher will be the specific adhesion between the ceramics and the metallic matrix in the composites. 2.2. The contact interface area between the matrix and the particles Fig. 1 shows a spherical particle partially immersed into the melt, but not yet reached by the solidification front. Parameter n in Fig. 1 denotes the depth of immersion of the particle into the melt, while parameter d is the closest distance between the

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particle and the solidification front. During processing of AMMCs, n: and d are time dependent variables. First, the particle moves through the gas phase approaching the gas/liquid interface, then crosses it, going deeper and deeper into the melt (x increases). At the same time the entire system (liquid metal with a partially immersed particle in it) moves with the cold roller, and therefore the solidification front~approaches the particle (d decreases). #en d becomes zero, i.e., the solidification front touches the particle, the particle will be engulfed by the solid metal at a particular final depth of immersion X. The total surface area between the particle and the matrix can be increased if the final depth of immersion of the particle is as great as possible. In the ideal case x equals the diameter of the particle (total immersion and maximum interface area). It is, however, usually not possible, due to the extraordinarily short period of time (less than 1 ms) during which the particle is in contact with the melt before the solidification front engulfs it. As particles used for AMMCs are small (less than 100 pm in size), interfacial forces play a major role in determining the final depth of immersion of the particle. The resulting interfacial force is the sum of the liquid/gas and the solid/liquid interfacial forces: F = F,, + F,, ,

(2)

where F denotes the forces acting on the particle (1, v, s are for liquid metal, vapor, solid metal). Interfacial forces are treated as negative when they attract the particle into the melt, and treated as positive when they push the particle out of the melt. The more negative the resulting interfacial force F, the higher will be the final depth of immersion of the particle and the total adhesion between the particle and the matrix. According to Ref. [4], the interfacial force at the liquid/vapor interface is the following function of the depth of immersion x: F,, = -2~~5,,[R(cos

Fig. 1. A spherical particle partially immersed into the melt x > 2 R - the particle is totally immersed in the liquid metal.

742-747

Of 1) -x],

(3)

where R denotes the radius of a particle, u the surface tension of the melt, and .!J the contact angle between the particle and liquid metal. An analysis of Eq. (3) shows that at any x, the particle will be attracted more into the melt (or

i

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pushed out less), if the contact angle is less, i.e., when the liquid wets the particle better. Hence, better wetting will provide deeper final depth of immersion, and thus stronger total adhesion in the composite. It is also shown in Ref. [4] that the role of the gravity force driven by a density difference between the particle and the melt is unimportant and can be totally disregarded for any metal/ceramic couples, if the size of the particle is less than 100 p,rn (which is the case for all AMMCs). Generally, the interface force arising from the approaching solid/liquid interface repulses (‘pushes’) the particle. This is one of the key problems in producing ‘normal’ MMCs (see Ref. [S] and references therein). The interfacial force at the solid/liquid interface can be written as follows [6]: F,, = 2R7ia(2aP,

- qpl - q,)

a i a+d

2 1

cos P,

(4) where cr denotes the ratio of heat conductivities of the particle and the melt, a the radius of atoms of the liquid metal, d the smallest distance between particle and the solidification front, m the interface energies (p - particle, s - solid metal, 1 - liquid metal), and /3 the angle between direction of rotation of the roller and solid-liquid interface at the closest point to the particle. Taking into account the Young equation, and also Eq. (l), the expression containing surface energy terms of Eq. (4) can be written approximately as 2a,, - CT@- a,, = a,, - vs,, = ap, - a,, - cT1,cos 8. (5) From Eq. (5) it follows that better wetting between the melt and the particle will provide a lower repulsive force from the solid/liquid interface. Hence improved wetting will lead to deeper final immersion of the particle into the melt. 2.3. Total adhesion between the particles matrix in AMMCs

and the

The total adhesion between a particle and the matrix can be defined as the product of specific adhesion (adhesion energy per unit interface area) and the total interface area. According to Eqs. (l)-

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(5), improved wetting between the liquid metal and the ceramic particles will provide (i) stronger specific adhesion between the particles and the matrix (Eq. (I)); (ii) stronger attractive force acting on the particles by the melt/vapor interface, i.e., deeper final depth of immersion, thus improved total interface area between the matrix and the particles (Eq. (3)); (iii) weaker repulsive force acting on the particles by the solidification front, i.e., deeper final depth of immersion, thus improved total interface area between the matrix and the particles (Eqs. (4) and (5)); (iv) improved real contact interface area between the melt (later the matrix) and the particles of nonideal shape compared to the geometrical interface area between the melt (later the matrix) and the ideally shaped particles. All the above consequences of the improved wetting between the melt and the particles will lead to a higher total adhesion energy between the particles and the matrix of AMMCs. Therefore, if all other parameters are kept constant, ceramic-liquid metal couples with improved wetting will provide AMMCs with higher total adhesion between the particles and the matrix, and thus provide higher wear resistance.

3. Results 3.1. Calculations for Sic and WC particles All calculations were performed for 1300 K. First, the properties of the Fe,,Ni,,Si,,B, melt were estimated. Activities of the components of the melt were found based on the Gibbs energies of formation of intermetallic compounds between the components (data taken from Ref. [7]) in the framework of the ideal associated model [8]. The following activity data have been found: Fe (0.35), Ni (0.25), Si (0.0002>, B (0.00005). The surface tension of the melt was calculated based on Gibbs equation, using the above activity data and the surface tension data of the pure liquid components and available binary systems [9]. For a melt with a low level of nonmetallic contamination (such as 0, S, N and P), the surface tension was found to be 1.6 J/m*. In Table 1, equilibrium contact angle data are given between the pure phases of interest as mea-

G. Kaptay Table 1 Equilibrium Ceramic

contact angles between

et al. / Journal

of Non-Crystalline

pure phases

Melt

T (“Cl

Time (min)

Atmosphere

WC WC WC

Si Fe Ni

1.500 1550 1450

15 15 15

Ar Ar, vacuum Ar, vacuum

Sic Sic Sic

Si Fe Ni

1370 1550 1450

20 15 15

vacuum He He

B (deg) 0 0 0 43 80 48

Refs.

DOI DOI [lOI 1111 I121 u21

sured by different authors. From this table, one can conclude that WC is perfectly wetted by the three main components of the melt, hence the Fe,,Ni,,Si,,Bb melt will probably also wet it perfectly (contact angle will be zero degrees). On the other hand, the same three components of the melt wet Sic, but not perfectly (contact angle between 0” and 90”). Using our activity data given above and the corrected Gibbs equation for the solid/liquid interface, the equilibrium contact angle between the Fe,,Ni,,Si,,B6 melt and Sic was calculated to be approximately 65”. In Table 1, the measured equilibrium contact angles are shown. However, during production of AMMCs the phases have a very limited time of contact (less than 1 ms). Therefore dynamic contact angles must be estimated. The subject is reviewed in [13]. It is shown that 100-1000 s are needed for an ordinary liquid metal-ceramic system to reach the equilibrium contact angle. However, that is mainly due to kinetic limitations of the non-wetting-wetting transition. In our case, we need the data for dynamic adhesion energy rather than the dynamic (macroscopic) contact angle. It can be shown (see Ref. [14]) that the (equilibrium) wetting of Sic in our case is mainly due to chemical reactions between the phases. However, perfect (equilibrium) wetting of WC cannot be explained by chemical reactions. For explaining perfect wetting of WC the electron-electron interaction between the conductive WC and the conductive liquid metal should be taken into account. This kind of interaction does not exist in the SiCliquid metal system, simply because Sic is not a conductor. In our opinion the electron-electron interaction (for WC) can be fully realized during the ms time available; however, chemical reactions (for

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745

Sic) cannot. Therefore, using the Young-Dupre equation we finally can conclude that (i) for WC the dynamic adhesion energy corresponds to the equilibrium contact angle (0”) i.e., the dynamic adhesion energy is higher than (1 + cos0) X 1.6 J/m2, i.e., > 3.2 J/m2; (ii) for SI’C the dynamic adhesion energy is less than the adhesion energy corresponding to the equilibrium contact angle (65”), i.e., the dynamic adhesion energy is less than (1 + cos 65) X 1.6 J/m’, i.e., < 2.3 J/m’. In Section 2 of this paper, it has been shown that if the wetting (adhesion energy) of small ceramic particles by the melt is improved, an AMMC with improved wear resistance can be obtained. Therefore, based on the results of the above calculations, one can conclude that the Fe,,Ni,,Si,,B,-WC amorphous metal matrix composite material will probably have a higher wear resistance than the Fe,,Ni,,Si,,B,-SiC system. 3.2. Experimental

euidence

In order to test the above theoretical results, amorphous ribbons with ceramic particles were prepared by a modified melt spinning method. The apparatus used was similar to the one given in fig. 12 of Ref. Lll, but our apparatus had no field concentrator and had a different blower situated on the other side of the crucible (compared to fig. 12 of Ref. [l]). Sic and WC particles with a size of 30-100 pm were blown into the melt puddle formed on the roller of the apparatus. The thickness of the Fe,,Ni,,Si,,B, amorphous ribbon was about 60 Km. The ribbons were analyzed by SEM for the distribution of particles on the surface. The cross-section and the length section of the composite were studied with a microscope in order to evaluate the incorporation of single particles. The following quantitative results (being in full agreement with the theoretical results described above) were obtained. (i) The majority of WC particles are deeply incorporated in the matrix, some of them are fully situated in the bulk of the matrix. The observable ‘contact angle’ is small, in some cases close to zero degrees. The contact interface is almost equal to the surface area of the particles. (ii) All the Sic particles are only partly incorporated into the matrix, actually more than half of their

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‘body’ is outside of the matrix. The observable ‘contact angle’ is large. The contact interface is much smaller than the surface area of the particles. The wear resistance of the composites was evaluated using the ‘pin on disk’ method described in [15]. The test was performed using abrasive cloth P600 with SIC particles of 16-49 km. The nominal load was 0.5 N/mm2 and slip velocity range was 0.08-0.16 m/s. The weight loss of the composite as function of wear path was measured. The wear resistance of the composite is defined as the inverse of the weight loss of the composite per unit wear path (in m/mg). Average data for the matrix material and the two AMMCs studied are shown in column 3 of Table 2. The abrasive ability of the composite materials was evaluated using the following apparatus. The composite material to be studied was welded to the outer surface of a steel cylinder with a diameter of 135 mm. The cylinder was fit on a turning-machine and was turned with a constant peripheral speed of 5.1 m/s. A piece of dry oak (acting as a sample to measure abrasive ability of the composites) was pressed with a constant pressure of 0.1 N/mm’ to the surface of the cylinder (Le., to the AMMCs) with the total contact surface of curved rectangular shape (4.5 x 25 mm2>. The weight loss of the oak sample as a function of the wear path was measured. The abrasive ability of the composite material was defined as the weight loss of the oak sample per unit abrasion path (in mg/m). In column 4 of Table 2 the average abrasive abilities of the two AMMCs and the commercial P120 Sic abrasive cloth (with a 106-125 pm particle size of Sic on it> are given for the first 30000 m of wear path.

Table 2 The adhesion energy, of the AMMCs

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The following conclusions can be drawn from Table 2. (i> The wear resistance of the Sic based AMMC is 10% greater than the wear resistance of the matrix material. This is obviously due to the fact that the Sic particles fell out of the matrix. Otherwise our measurement would have provided infinitely large wear resistance, as the abrasive material was also Sic. The wear resistance of the WC based AMMC was 25% higher than the wear resistance of the matrix material. Consequently, WC improved the wear resistance of the matrix more than Sic, due to higher adhesion of WC particles to the matrix material. (ii) The abrasive ability of AMMC with WC is four times higher than the abrasive ability of the AMMC with Sic, and is 12 times higher then the abrasive ability of the commercial P120 abrasive cloth. This result is in full agreement both with improved adhesion of WC to the matrix and with the higher wear resistance of AMMC with WC.

4. Conclusions It has been shown that AMMCs with higher wear resistance can be produced using metal-ceramic couples with improved wettability. Calculations show that the Fe,,Ni,,Si,,Bb melt. wets WC particles perfectly, while it only poorly wets Sic particles. Therefore AMMCs with better wear resistance can be produced using WC particles rather than Sic particles. Experiments show evidence for better wetting, higher wear resistance and better abrasive ability in the Fe,,Ni,,Si,,B6-WC system compared with the Fe,ONi,OSi,,B,-SiC system.

Acknowledgements the wear resistance,

and the abrasive

Material

Adhesion energy (J/m’)

Wear resistance h/mg)

Abrasive ability

Commercial P120 Matrix of the AMMC AMMC with SIC AMMC with WC

< 2.3 > 3.2

2.9 3.2 3.6

0.15 0.43 1.8

ability

This work has been supported by the Hungarian Academy of Sciences (OTKA T016825).

(mg/m)

References [l]

E.R. Murray and D.G. Ast, in: Rapidly Solidified Alloys, H.H. Liebennann (Marcel Dekker, New York) p. 119.

ed.

G. Kaptay

et aE./JownaE

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[21 F. Szigeti and H. Kreye, Schweis. Schneid. 44 (1992) 13. [3] G. Kaptay, ‘Method for estimating solid-solid interface energies in metal-ceramic systems’, in: Key Engineering Materials (TramTech, Aedermannsdorf, 1996) to be published. [4] G. Kaptay, ‘Interfacial phenomena during melt processing of metal matrix composites, part I’, in: Key Engineering Materials (TransTech, Aedermannsdorf, 1996) to be published. [5] D.M. Stefanescu, A.A. Moitra, AS. Kacar and B.K. Dhindaw, Metall. Trans. 21A (1990) 231. [6] G. Kaptay, ‘Interfacial phenomena during melt processing of metal matrix composites, part II’, in: Key Engineering Materials (TransTech, Aedermannsdorf, 1996) to be published. [7] L. Barin, Thermochemical Data of Pure Substances, Vol. I-2,2nd Ed. (VCH, 1993). [8] K. Wasai and K. Mukai, J. Jpn. Inst. Met. 46 (1‘982) 266.

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[9] V.I. Nizhenko and L.I. Floka. Metallurgiia (1981) (in Russian). [IO] G.V. Samsonov, A.D. Panasyuk and G.K. Kozina, Poroshk. Metall. 11 (1968) 42. [ll] T. Whalen and A. Anderson, J. Am. Ceram. Sot. 58 (1975) 396. 1121 G.V. Samsonov and A.D. Panasyuk et al., DAN USSR A5 (1969) 468. [13] S.I. Pope], Adgeziia Raspl. Paika Mater. 1 (1976) 3. [14] G. Kaptay, Mater. Sci. Forum 77 (1991) 315. [15] P. Barczy, F. Szigeti and J. Szucs, in: Increased Wear Resistance and Performance Through PM Techniques and Surface Modifications, ed. T. Pieczonka and J. Frydrych, Network Proc. held in Krakow, Poland, 3-7 Apr. (1995) p. 124.