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Influence of the heat treatment on the abrasive wear resistance of electroless Ni-P
ELSEVIER
Surface and Coatings Technology
91-95
(1997) 513-548
Influence of the heat treatment on the abrasive wear resistance of electroless Ni-P
Abstract Electroless Ni-P coatings on plain carbon steel werr prepared by using an acid bath based on NiClz as source of nickel cations. -___ Hear-_--__-_ treatments were performed at 260’ and 300°C for 25 h and 1 h, respectively. The abrasive resistances of the coatings which suffered different heat treatments were determined by using the modified ball cratering method, recently proposed by Staia et al. The abrasive media employed was a mixture of 100 ml ethylene glycol and I g of 3 pm diamond paste. A RB3S.UGlOOP3 ball-bearing was used as tribological pair. The individual wear constants for the substrate and coatings were determined from the recorded plots of the crater depth vs. sliding distances with the aid of theory of the imposed shape wear scar reported in the literature by Kassman et al. Electron microscopy was used to study the morphology of the worn surfaces. As a result, it was found that the heat treatment decreases the abrasive wear resistance of the coatings. A comparison between the methods proposed by Rutherford and Hutchings and ours for assessing the wear constants is carried out, particularly in relation to the relevance of including in the computation the value of the friction force. 0 1997 Elsevier Science S.A. Keytrvrds:
Heat treatment; Electroless Ni-P coatings; Wear resistance
1. Introduction In the last two decades a variety of surface engineering processes have been developed with the aim of enhancing the performance of materials mainly from the wear, fatigue, corrosion and biocompatibility point of view. Parallel with the development of these coatings technologies, there has been an increase in the number of methods employed to determine the main properties and characteristics of the produced coatings and, at the same time, an increase of accelerated (laboratory tests) and in situ techniques able to evaluate their tribological performance and/or corrosion behavior. Without doubt, the accelerated laboratory tests used to assess the tribological behavior of the coatings, in most of the cases, are far away from describing the tribological performance of the system coating/substrate in service conditions. Nevertheless, they have permitted the screening of the quality of these surface engineered products, thus being able to accomplish the assessment of their performance related to the intrinsic and extrinsic
* Corresponding author. Tel.: +58 582 7529017; fax: ~58 582 7529017; e-mail: [email protected]
0257~8972/97/$17.00 Q 1997 Elsevier Science S.A. Ail rights reserved PZZ SO257-8972(97)00463-5
parameters characteristic to each test. Moreover, combined with the information regarding fundamental properties, they were able to contribute to optimization of the coating processing. Previous work conducted by us [I] has allowed us to develop a novel test by modifying the cratering ball technique, which is a standard method employed in measuring the coatings thickness, in order to quantify the individual abrasion wear coefficient of coatings and substrates with the aid of the theory of the imposed shape wear scar proposed for the first time by Kassman et al. [2]. The materials under study were silicon wafers, to test the data reproducibility of the method, and Ni-P coatings as-deposited and heattreated. Wear test5 were conducted by using as abrasive medium a mixture of 3 pm diamond paste and ethylene glycol. Kassman et al.‘s [2] theory has proved to be a simple and reproducible method for checking the mechanical quality of a thin coating by using the dimple grinder normally employed for electron microscopygple preparation. By applying it they have achieved a major improvement in relation to all the existing laboratory methods for determining wear resistance [3], which until then were unable to assess the separate influence of the substrate and the coating on the overall tribological performance of the system.
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Some other improvements, accordingly explained [ 1, 51, were obtained when using the ball cratering method instead of the dimple grinder. These are regarding two main features: firstly, the fact that the modified ball cratering equipment permits the delivery of the a fresh abrasive mixture during the experiments, impeding its degradation; and secondly, the possibility of collecting the results for different values of sliding distances without errors which can arise from the removal of the sample. At the same time, Rutherford and Hutchings [4], have proposed the use of the ball cratering method to determine the individual abrasion wear coefficients for a variety of thin hard films and different bulk materials by using the theory of the imposed wear scar shape proposed by Kassman et al. [2]. There are many differences between their method and that proposed by us. These differences can be divided into two groups: those regarding the theoretical basis (mainly calculation of the forces) and those regarding the test method per se as sliding velocity, nature of the abrasive mixture and its delivery, and the way of data collecting, among others. In the calculation of the forces in this test, Rutherford and Hutchings [4] have pointed out the influence of the friction force between the sphere and the sample in determining the normal load for a water-based slurry. Also, they have assumed that the friction force is equal to zero for a more viscous liquid (glycerol) when it was used as carrier of the abrasive particles, due to formation of a fully lubricating hydrodynamic film. Therefore, the present study has been conducted with the aim of determining whether the frictional force between the sample and the ball could be neglected for our experimental abrasive mixture and, if not, to determine the extent of the variation of the independent wear coefficients for electroless Ni-P coatings on plain carbon steel, taking into account this influence. At the same time, it will be very interesting to discuss if this modification, not taken into consideration by the authors in previous publications, is indeed necessary in performing this test as a screening method in assessing the three body abrasion wear resistance parameters of metallurgical coatings.
(1997) 543-548
2.2. Mnterinls The abrasive tests were performed on plain carbon AISI 1020 steel uncoated and coated with electroless Ni-P. All samples were coated in an acid electroless deposition bath by employing the procedure described elsewhere [6]. All coated samples are practically identical, with an average phosphorus content of 6.67% determined by using glow discharge optical spectroscopy (GDEOS) and an average coating thickness of 9.3 Frn determined by using the ball cratering method. Table I presents some of the characteristic features of the sample under study. Values of Knoop microhardness determined by using a load of 0.98 N for 10 s and of roughness determined by using a Mitutoyo Surftest 301 profilometer are also included. 2.3. Main test parameters As was stated previously [l], there are three main parameters that should to be taken into consideration: load, test velocity and abrasive medium. In the following treatment only changes in load calculations will be addressed, keeping the same sliding velocity (1.11 m s-‘) and the same abrasive mixture. 2.4. Load cdcrhtion When friction forces are not considered, the load exerted on the sample, N,, has been reported [I] to be equal to:
Air bed
--WI
Abrasive + Lubricant
2. Experimental details 2.1. Test
equipnerzr
A general view of the instrument and its attachments is presented in Fig. 1. The abrasive mixture of 100 ml of ethylene glycol and 1 g of 3 pm diamond paste, magnetically stirred, was delivered by using a glass syringe in such a manner as to ensure rolling of the ball in contact with the shaft without slip, and at the same time to provide the mixture continuously at a constant flow value of 2.5 ml/h.
Fig. 1. The test equipment and its attachments.
2.5. Determinatiorz of the cl~namic friction
Table 1 Sample characteristics and heat treatments Sample
Heat treatment
Roughness, KnoopIoo microR, k-d) hardness (kg mm-‘)
Ni-P, as-deposited 0.40 260°C for 25 h, air-cooled 0.41 Ni-P Ni-P 400°C for 1 h, air-cooled 0.41
Nt =
=~n~cosw case + 7 sinwcos8
497 795 916
-.-
(1)
where 0 = the angle between the direction of the normal force N, and the weight of the ball, IV; w = the angle between the weight of the ball, IV, and the distance, K, between the two points of contact on the axis with the ball; 0 = the angle between the reaction force on the axis, N2, and the x direction. This load depends on the size of the ball-bearing, the angle between the sample and the horizontal plane, the friction force between the ball-bearing and the sample in the presence of the slurry, and the distance from the sample to the rotational axis. When friction forces are taken into account, some additional geometrical parameters have to be determined together with the value of the dynamic friction coefficient for the tribological system under study. These measurements are related to the angle between the direction of the friction forcefi and the x axis, which is equal to the angle between the direction of the normal force N1 and the weight of the ball (0) and the angle between the friction forcefz and the x axis (w). As the ball achieves a constant velocity, in a few seconds from the start, it can be assumed that the net torque acting on the ball is equal to zero, which implies that fi and f2 are approximately equal. The normal force exerted on the sample (Ni ), obtained by combining the summation of forces in the z and x directions represented in Fig. 2, is expressed as:
coeficient
The pin-on-disk method was employed for evaluating the dynamic friction coefficient. The mandrel-holding rod was substituted by a container which allows the positioning of the coated sample and the addition of the abrasive mixture. The maximum sliding velocity of 0.1 m s-’ was used in order to avoid the abrasive mixture ejection. The experiment was performed in air at 22°C and 35 F 5% humidity. The pin, a ball-bearing of 6 mm diameter, was loaded with a normal load of 1 N. The contact radius was of 4 mm and five tests were carried out.
2.6. Collection of the results fi-orn the nbrnsive w’ear test In a tested sample there are three sets of five craters each. A wear test is made of three single tests and a single test is made of five different runs, one for each sliding distance. After each single run, the sample is rotated counter-clockwise and another identically but longer single run is performed. When the test is finished, the diameters of the craters are measured in-an image analyzer. Subsequently, the worn-off volume (V, and the crater depth (h) are easily calculated from the measured crater diameters (D) for a determined sliding-distance (S) and the ball radius (R).
3. Results and discussion 3. I. Deteeation
~~OJ the dynamic friction
coeficient
The average value of the dynamic friction coefficient, p, was found to be approximately 0.26 and no changes of it with time were observed. With this value of ,LLI,the normal load N, calculated by using eqn (2) gives a value of 0.12 N, which is nearly 30% smaller than the value of the normal load evaluated when the friction force between the ball and
N1 = (WsinQcosa)/ ( cosQ[sinO~i (2cosw - cos@] + sinwcosa[cos0 +p, (sin0 + 2sino)j) (2) Since the S and K distances are measured on the equipment, the radius I’ is known, the angle 8 is set, the angles Q, 0 and w can be readily found. From trigonometric considerations and the measured values for K = 16.25 mm and S =23.92 mm, the angles Q = 25.48”, u = 80.56” and w = 3.46”have been determined. Knowing that 0 = 80”, r = 19.05 mm and W = 0.225 Kgf, a value of the normal load of 0.17 N was obtained on this specific setup, when using eqn (1). If eqn (2) is employed to calculate the normal load N1 the dynamic friction coefficient, p, has to be determined.
Fig. 2. Schematic diagram for load calculation.
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M. H. Stain et al. / Srtrjke
rind Coarings Techdog):
the sample is not considered and therefore, a variation in both individual wear coefficients of coating and substrate is expected.
IIhr =-r
( > [ 1 +h-t
Kc
ifh>r
14)
Ks
where h represents the depth of the crater. From the values experimentally determined of S, h and t the constants Kc and K, can be readily determined by means of the leastsquare method. Thus: rb !gl 11: I&=----L ‘5 Si i=l
(6)
where N represents the number of experimental data. When dealing with only one homogeneous material as steel, eqn (5) can be applied for the case corresponding to h 2 t. For the heat-treated nickel electroless samples at 260” and 4OO”C, the equation corresponding to h > t was used. In all cases, as indicated previously [l], a high degree of correlation was obtained between the experimental and theoretical points and, as an example, Fig. 3 presents the variation of the crater depth vs. sliding distance for the Ni-P coating heat-treated at 400°C for 1 h, irrespective of the value of the normal load. The calculated values of the wear constants for every material under study with and without considering the friction force N1 are presented in Table 2. As can be noticed, there is an increase of 30% in the values of the wear constants for each material under study when the normal load is calculated by taking into account the friction force. This increase, without doubt, cannot be considered insignificant. However, as was mentioned above, the proposed experiment is a good accelerated laboratory test which ensures reliability, as demonstrated by the experimental results. This method can be used for optimizing the deposition conditions, and to quantify the relative wear constant of different coatings and substrates produced by the same process under different conditions. Nevertheless, this test is unable to simulate the real operat-
h-i
2t-
i= 1
KS=
(3)
SVPT.
(1997) 543-548
5 (hi-t)3
As reported by Kassman et al. [2] by using the special case of a spherical cap shaped wear scar produced on a coated material with a coating thickness, t, the approximate sliding distance, SapPr.,is given by:
whereas:
94-95
(5)
and
M-P Heat treated (400°C x 1h)
300
400 Sliding
500
600
700
800
900
Distance Im]
Fig. 3. Crater depth vs. sliding distance for the Ni-P coating heat-treated at 3OOT for 1 h. Velocity I. 1I m s-’ for both normal load values of0.17 N and 0.12 N.
517 Table 2 Calculated values for the wear constant of the samples under study for different values of the normal load A’, depending whether (with f.f.) or not (no f.f.) the friction force was considered in its calculation Sample Ni-P as-deposited Ni-P heat-treated (260°C x 26 h) Ni-P heat-treated (400°C x 1 h)
No f.f. With f.f. No f.f. With f.f. No f.f. With f.f.
Load (N)
K, (m’&I)
0.17 0.12 0.17 0.12 0.17 0.12
3.57 5.06 8.00 11.39 1.00 1.42
ing conditions and cannot ensure that the coating will have the same behavior. From the above discussion it is possible to conclude three important facts: (a) the inclusion of the friction force in the computation of the wear constants give rise to a significant increase in the values; (b) however, in order to take into account such a force it is required to evaluate two additional geometrical parameters as well as to determine the dynamic friction coefficient, which renders the method much more complicated that the simpler and easier method earlier proposed by us; (c) since for the screening of the wear properties of different materials only relative values of the wear constants are required, the inclusion of the friction force in the calculation is irrelevant. However, it must be specified wether or not such a force was taken into account into the calculation if a comparison between results of different laboratories is to be made. Thus, this specification will become an intrinsic parameter of the test in the same manner as is, for example, the abrasive mixture employed or the diamond tip radius used in the scratch test to determine adhesion [6]. The inclusion of the friction force implies additional experiments in order to evaluate the dynamic friction coefficient for the system under study, which makes the use of this simple technique dependent on other equipments with their inherent experimental errors. The relative values of the wear constants obtained for different conditions of electroless Ni-P indicate that there is not a direct relationship between the coating hardness and abrasion wear resistance, since other factors are injluencing this behavior, underlying the fact that hardness could not be considered as a key guide to the coating abrasive wear. As can be observed, the abrasive wear constant, Kc, for the Ni-P coating heat-treated at 400°C (a microhardness of 916 HKrm) is higher than the wear constant obtained when the heat treatment was conducted at 260°C (microhardness of 795 HKr,), results which contradict the explanation presented by Duncan [7] which attributed the improvement of abrasive resistance of Ni-P coatings to the Ni3P particle coarsening, as the heat treatment temperatures increase. Also, it has to be mentioned that due to the difference between the thermal coefficients of the coating and the steel substrate, the coating heat-treated at 400°C presented cracking at the surface which could negatively influence its abrasive resistance behavior.
x x x x x x
K,~$& 10-11 10-l” IO“’ 10-14 lo“:’ 10-13
~
-1.85 2.62 5.17 7.74
x x x x
R_’
IO-” lo-” 10-I” 10-I”
0.97 0.97 0.95 4.95 0.91 0.91
As was indicated by Rutherford and Hutchings [A-], the action of the abrasive particles within the contact will depend on the ratio between the film thickness of a lubricant between the sample and the ball and the abrasive mean particle diameter, d. They proposed the following equation in order to evaluate the lubricant film thickness, h,. This equation represents the solution to Reynolds equation in two dimensions for a sphere of radius R, sliding against a plane in the presence of a viscous fluid and is expressed as:
where IVY is the normal load, ~1is the sliding speed and r represents the fluid viscosity. In the present investigation for values of V = 1.11 ms-.!, and R = 0.01905 m, qEthY]e,,e slycOl= 4.3 x IO-’ N m-?-S N, = 0.12, the film thickness has been determined to be about 0.3 1 pm, that is to say smaller than the mean abrasive particle diameter. In these conditions, the ratio hJd is equal to 0.1, indicating that the abrasive particle has a major contribution to the wear process. This can be corroborated by analyzing the scanning electron micrographs of the worn surfaces of the Ni-P coating, heat-treated at 4OO”C, which are presented in Fig. 4a and b. Well-defined wear scars of nearly the same width as the me-anva1u.e of the abrasi-ve particle diameter are-observed, indicat%lg- a pure abrasion wear mechanism, which results in the material removal due to the cutting performed by the abrasive particles. Similar morphologies of the worn surfaces were obtained for all the samples under study. It was considered [8] that when the tribosystem is composed of a two triboelements and lubricant,it~qossible to obtain an indication of the dominating lubrication or wear mechanism by monitoring suitable test parameters such as friction, lubricant film thickness, temperature, wear and surface roughness. In these conditions, three main lubrication regimes have been identified [9] as a function of the variation of the friction coefficient, ,LL,and the film thickness-toroughness ratio, A. In the present work a value of film thickness-to-roughness ratio, X, equal to 0.77 has been determined, which together with a friction coefficient of 0.26 calculated previously possibly could place our system in a boundary lubrication mode, were the tribological behavior is governed by solid-solid friction and wear processes.
548
MH. Slain er al. / Surfiire cd
Coatings Technology 94-95
11997) 543-548
without ensuring the reproducibility of the real operating conditions. A simple indication of the fact that the friction force was not taken into account when reporting the wear constants, as an intrinsic parameter, will be enough to allow comparisons to be made with the work performed in other laboratories. Also, it has to be pointed out the simplicity of the test proposed by us in reference to: {a) use of a mnch simpler expression to calculate the load which involves parameters that can be easily determined, (b) easy measurement of the craters diameters for different sliding distances without the misfit produced when moving the sample or the complicated setup to measure crater in situ, and (c) practical and lowcost setup for the slurry delivery attachment.
Acknowledgements The authors wish to acknowledge the financial support of National Council for Scientific and Technological Research of Venezuela (CONICIT) through the project Sl-2612, to the Council for Scientific and Humanistic Development (CDCH) of the Central University of Venezuela through the project O&17-2789/95, and to the Organization of the American States (OAS) through the Multinational Project on Materials. The financial support of the Postgraduate Studies Council of Central University of Venezuela is also acknowledged.
References Fig. 3. (a and bj Scanning electron micrographs of the worn surface of the heat-treated Ni-P coating [at 300°C ~for 1 h) at different magnifications.
4. Conclusions The analysis carried out in the present work has shown that our initially proposed method is able to quantify the relative abrasive wear resistance of electroless Ni-P coatings which have undergone different post-deposition treatment conditions. Although a significant increase in the values of the individual abrasive wear constants of the coatings were obtained when the friction force was taken into account, in our opinion, this approach will make this simple test troublesome. The inclusion of the friction force implies additional experiments in order to evaluate the dynamic friction coefficient for the system under study, which makes this test costly and dependent on other equipment
[1] M.H. Staia, C. Enriquez, ES. Puchi, B. Lewis and M. Jeandin, in T.S. Sudarshan, M. Jeandin and M. Khor (eds.), S&zce Modificntion Technologies X, Inst. of Materials, UK, 1997, p. 130. [2] A. Kassman. L. Erickson, M. Olsson, P. Hedenqvist, S. Jacobson and S. Hogmark, Su$ Cont. Technoi., 50 (1991) 75. [3] B. Bhushan, in W.B.Harding and G.A. DiBari teds.), O~rrvie~~ of Comings Mareriais, Su&ce Trearmenrs md Screening Techniques for Tribological Appiicntions, Part 2: Screening Techniques, Testing of Mefallic and inorganic Coatings, STP-947, 1987, p. 310. [3] K.L Rutherford and I.M. Hutchings, &lj-: Coot. Technol., 79 (1996), 231. [5] M.H. Staia. E. Castillo, E.S. Puchi, D. B. Lewis and H.E. Hintermann, Surj Cont. Technol., 86-87 (1996) 598. [6] P.A. Steinmann and H.E. Hintermann, J. Var. Sci. Techno/., A7 (3) ( 1989) 2257. [7] R. Duncan, Meral finishing, March (1990) 11. [8] K. Holmberg and A. Matthews, in D. Dowson (ed.), Coating Tribology, Tribology Series, 28, Elsevier, Amsterdam, 1994, p. 307. 191 H. Czichos, Basic Tribological Parameters, ASM Handbook vol. 18, ASM International, 1992, 474.
Surface and Coatings Technology
91-95
(1997) 513-548
Influence of the heat treatment on the abrasive wear resistance of electroless Ni-P
Abstract Electroless Ni-P coatings on plain carbon steel werr prepared by using an acid bath based on NiClz as source of nickel cations. -___ Hear-_--__-_ treatments were performed at 260’ and 300°C for 25 h and 1 h, respectively. The abrasive resistances of the coatings which suffered different heat treatments were determined by using the modified ball cratering method, recently proposed by Staia et al. The abrasive media employed was a mixture of 100 ml ethylene glycol and I g of 3 pm diamond paste. A RB3S.UGlOOP3 ball-bearing was used as tribological pair. The individual wear constants for the substrate and coatings were determined from the recorded plots of the crater depth vs. sliding distances with the aid of theory of the imposed shape wear scar reported in the literature by Kassman et al. Electron microscopy was used to study the morphology of the worn surfaces. As a result, it was found that the heat treatment decreases the abrasive wear resistance of the coatings. A comparison between the methods proposed by Rutherford and Hutchings and ours for assessing the wear constants is carried out, particularly in relation to the relevance of including in the computation the value of the friction force. 0 1997 Elsevier Science S.A. Keytrvrds:
Heat treatment; Electroless Ni-P coatings; Wear resistance
1. Introduction In the last two decades a variety of surface engineering processes have been developed with the aim of enhancing the performance of materials mainly from the wear, fatigue, corrosion and biocompatibility point of view. Parallel with the development of these coatings technologies, there has been an increase in the number of methods employed to determine the main properties and characteristics of the produced coatings and, at the same time, an increase of accelerated (laboratory tests) and in situ techniques able to evaluate their tribological performance and/or corrosion behavior. Without doubt, the accelerated laboratory tests used to assess the tribological behavior of the coatings, in most of the cases, are far away from describing the tribological performance of the system coating/substrate in service conditions. Nevertheless, they have permitted the screening of the quality of these surface engineered products, thus being able to accomplish the assessment of their performance related to the intrinsic and extrinsic
* Corresponding author. Tel.: +58 582 7529017; fax: ~58 582 7529017; e-mail: [email protected]
0257~8972/97/$17.00 Q 1997 Elsevier Science S.A. Ail rights reserved PZZ SO257-8972(97)00463-5
parameters characteristic to each test. Moreover, combined with the information regarding fundamental properties, they were able to contribute to optimization of the coating processing. Previous work conducted by us [I] has allowed us to develop a novel test by modifying the cratering ball technique, which is a standard method employed in measuring the coatings thickness, in order to quantify the individual abrasion wear coefficient of coatings and substrates with the aid of the theory of the imposed shape wear scar proposed for the first time by Kassman et al. [2]. The materials under study were silicon wafers, to test the data reproducibility of the method, and Ni-P coatings as-deposited and heattreated. Wear test5 were conducted by using as abrasive medium a mixture of 3 pm diamond paste and ethylene glycol. Kassman et al.‘s [2] theory has proved to be a simple and reproducible method for checking the mechanical quality of a thin coating by using the dimple grinder normally employed for electron microscopygple preparation. By applying it they have achieved a major improvement in relation to all the existing laboratory methods for determining wear resistance [3], which until then were unable to assess the separate influence of the substrate and the coating on the overall tribological performance of the system.
544
hl.H, Shin et a/. / Surjcrrce and Coatings Terhnoiog)~ 94-95
Some other improvements, accordingly explained [ 1, 51, were obtained when using the ball cratering method instead of the dimple grinder. These are regarding two main features: firstly, the fact that the modified ball cratering equipment permits the delivery of the a fresh abrasive mixture during the experiments, impeding its degradation; and secondly, the possibility of collecting the results for different values of sliding distances without errors which can arise from the removal of the sample. At the same time, Rutherford and Hutchings [4], have proposed the use of the ball cratering method to determine the individual abrasion wear coefficients for a variety of thin hard films and different bulk materials by using the theory of the imposed wear scar shape proposed by Kassman et al. [2]. There are many differences between their method and that proposed by us. These differences can be divided into two groups: those regarding the theoretical basis (mainly calculation of the forces) and those regarding the test method per se as sliding velocity, nature of the abrasive mixture and its delivery, and the way of data collecting, among others. In the calculation of the forces in this test, Rutherford and Hutchings [4] have pointed out the influence of the friction force between the sphere and the sample in determining the normal load for a water-based slurry. Also, they have assumed that the friction force is equal to zero for a more viscous liquid (glycerol) when it was used as carrier of the abrasive particles, due to formation of a fully lubricating hydrodynamic film. Therefore, the present study has been conducted with the aim of determining whether the frictional force between the sample and the ball could be neglected for our experimental abrasive mixture and, if not, to determine the extent of the variation of the independent wear coefficients for electroless Ni-P coatings on plain carbon steel, taking into account this influence. At the same time, it will be very interesting to discuss if this modification, not taken into consideration by the authors in previous publications, is indeed necessary in performing this test as a screening method in assessing the three body abrasion wear resistance parameters of metallurgical coatings.
(1997) 543-548
2.2. Mnterinls The abrasive tests were performed on plain carbon AISI 1020 steel uncoated and coated with electroless Ni-P. All samples were coated in an acid electroless deposition bath by employing the procedure described elsewhere [6]. All coated samples are practically identical, with an average phosphorus content of 6.67% determined by using glow discharge optical spectroscopy (GDEOS) and an average coating thickness of 9.3 Frn determined by using the ball cratering method. Table I presents some of the characteristic features of the sample under study. Values of Knoop microhardness determined by using a load of 0.98 N for 10 s and of roughness determined by using a Mitutoyo Surftest 301 profilometer are also included. 2.3. Main test parameters As was stated previously [l], there are three main parameters that should to be taken into consideration: load, test velocity and abrasive medium. In the following treatment only changes in load calculations will be addressed, keeping the same sliding velocity (1.11 m s-‘) and the same abrasive mixture. 2.4. Load cdcrhtion When friction forces are not considered, the load exerted on the sample, N,, has been reported [I] to be equal to:
Air bed
--WI
Abrasive + Lubricant
2. Experimental details 2.1. Test
equipnerzr
A general view of the instrument and its attachments is presented in Fig. 1. The abrasive mixture of 100 ml of ethylene glycol and 1 g of 3 pm diamond paste, magnetically stirred, was delivered by using a glass syringe in such a manner as to ensure rolling of the ball in contact with the shaft without slip, and at the same time to provide the mixture continuously at a constant flow value of 2.5 ml/h.
Fig. 1. The test equipment and its attachments.
2.5. Determinatiorz of the cl~namic friction
Table 1 Sample characteristics and heat treatments Sample
Heat treatment
Roughness, KnoopIoo microR, k-d) hardness (kg mm-‘)
Ni-P, as-deposited 0.40 260°C for 25 h, air-cooled 0.41 Ni-P Ni-P 400°C for 1 h, air-cooled 0.41
Nt =
=~n~cosw case + 7 sinwcos8
497 795 916
-.-
(1)
where 0 = the angle between the direction of the normal force N, and the weight of the ball, IV; w = the angle between the weight of the ball, IV, and the distance, K, between the two points of contact on the axis with the ball; 0 = the angle between the reaction force on the axis, N2, and the x direction. This load depends on the size of the ball-bearing, the angle between the sample and the horizontal plane, the friction force between the ball-bearing and the sample in the presence of the slurry, and the distance from the sample to the rotational axis. When friction forces are taken into account, some additional geometrical parameters have to be determined together with the value of the dynamic friction coefficient for the tribological system under study. These measurements are related to the angle between the direction of the friction forcefi and the x axis, which is equal to the angle between the direction of the normal force N1 and the weight of the ball (0) and the angle between the friction forcefz and the x axis (w). As the ball achieves a constant velocity, in a few seconds from the start, it can be assumed that the net torque acting on the ball is equal to zero, which implies that fi and f2 are approximately equal. The normal force exerted on the sample (Ni ), obtained by combining the summation of forces in the z and x directions represented in Fig. 2, is expressed as:
coeficient
The pin-on-disk method was employed for evaluating the dynamic friction coefficient. The mandrel-holding rod was substituted by a container which allows the positioning of the coated sample and the addition of the abrasive mixture. The maximum sliding velocity of 0.1 m s-’ was used in order to avoid the abrasive mixture ejection. The experiment was performed in air at 22°C and 35 F 5% humidity. The pin, a ball-bearing of 6 mm diameter, was loaded with a normal load of 1 N. The contact radius was of 4 mm and five tests were carried out.
2.6. Collection of the results fi-orn the nbrnsive w’ear test In a tested sample there are three sets of five craters each. A wear test is made of three single tests and a single test is made of five different runs, one for each sliding distance. After each single run, the sample is rotated counter-clockwise and another identically but longer single run is performed. When the test is finished, the diameters of the craters are measured in-an image analyzer. Subsequently, the worn-off volume (V, and the crater depth (h) are easily calculated from the measured crater diameters (D) for a determined sliding-distance (S) and the ball radius (R).
3. Results and discussion 3. I. Deteeation
~~OJ the dynamic friction
coeficient
The average value of the dynamic friction coefficient, p, was found to be approximately 0.26 and no changes of it with time were observed. With this value of ,LLI,the normal load N, calculated by using eqn (2) gives a value of 0.12 N, which is nearly 30% smaller than the value of the normal load evaluated when the friction force between the ball and
N1 = (WsinQcosa)/ ( cosQ[sinO~i (2cosw - cos@] + sinwcosa[cos0 +p, (sin0 + 2sino)j) (2) Since the S and K distances are measured on the equipment, the radius I’ is known, the angle 8 is set, the angles Q, 0 and w can be readily found. From trigonometric considerations and the measured values for K = 16.25 mm and S =23.92 mm, the angles Q = 25.48”, u = 80.56” and w = 3.46”have been determined. Knowing that 0 = 80”, r = 19.05 mm and W = 0.225 Kgf, a value of the normal load of 0.17 N was obtained on this specific setup, when using eqn (1). If eqn (2) is employed to calculate the normal load N1 the dynamic friction coefficient, p, has to be determined.
Fig. 2. Schematic diagram for load calculation.
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the sample is not considered and therefore, a variation in both individual wear coefficients of coating and substrate is expected.
IIhr =-r
( > [ 1 +h-t
Kc
ifh>r
14)
Ks
where h represents the depth of the crater. From the values experimentally determined of S, h and t the constants Kc and K, can be readily determined by means of the leastsquare method. Thus: rb !gl 11: I&=----L ‘5 Si i=l
(6)
where N represents the number of experimental data. When dealing with only one homogeneous material as steel, eqn (5) can be applied for the case corresponding to h 2 t. For the heat-treated nickel electroless samples at 260” and 4OO”C, the equation corresponding to h > t was used. In all cases, as indicated previously [l], a high degree of correlation was obtained between the experimental and theoretical points and, as an example, Fig. 3 presents the variation of the crater depth vs. sliding distance for the Ni-P coating heat-treated at 400°C for 1 h, irrespective of the value of the normal load. The calculated values of the wear constants for every material under study with and without considering the friction force N1 are presented in Table 2. As can be noticed, there is an increase of 30% in the values of the wear constants for each material under study when the normal load is calculated by taking into account the friction force. This increase, without doubt, cannot be considered insignificant. However, as was mentioned above, the proposed experiment is a good accelerated laboratory test which ensures reliability, as demonstrated by the experimental results. This method can be used for optimizing the deposition conditions, and to quantify the relative wear constant of different coatings and substrates produced by the same process under different conditions. Nevertheless, this test is unable to simulate the real operat-
h-i
2t-
i= 1
KS=
(3)
SVPT.
(1997) 543-548
5 (hi-t)3
As reported by Kassman et al. [2] by using the special case of a spherical cap shaped wear scar produced on a coated material with a coating thickness, t, the approximate sliding distance, SapPr.,is given by:
whereas:
94-95
(5)
and
M-P Heat treated (400°C x 1h)
300
400 Sliding
500
600
700
800
900
Distance Im]
Fig. 3. Crater depth vs. sliding distance for the Ni-P coating heat-treated at 3OOT for 1 h. Velocity I. 1I m s-’ for both normal load values of0.17 N and 0.12 N.
517 Table 2 Calculated values for the wear constant of the samples under study for different values of the normal load A’, depending whether (with f.f.) or not (no f.f.) the friction force was considered in its calculation Sample Ni-P as-deposited Ni-P heat-treated (260°C x 26 h) Ni-P heat-treated (400°C x 1 h)
No f.f. With f.f. No f.f. With f.f. No f.f. With f.f.
Load (N)
K, (m’&I)
0.17 0.12 0.17 0.12 0.17 0.12
3.57 5.06 8.00 11.39 1.00 1.42
ing conditions and cannot ensure that the coating will have the same behavior. From the above discussion it is possible to conclude three important facts: (a) the inclusion of the friction force in the computation of the wear constants give rise to a significant increase in the values; (b) however, in order to take into account such a force it is required to evaluate two additional geometrical parameters as well as to determine the dynamic friction coefficient, which renders the method much more complicated that the simpler and easier method earlier proposed by us; (c) since for the screening of the wear properties of different materials only relative values of the wear constants are required, the inclusion of the friction force in the calculation is irrelevant. However, it must be specified wether or not such a force was taken into account into the calculation if a comparison between results of different laboratories is to be made. Thus, this specification will become an intrinsic parameter of the test in the same manner as is, for example, the abrasive mixture employed or the diamond tip radius used in the scratch test to determine adhesion [6]. The inclusion of the friction force implies additional experiments in order to evaluate the dynamic friction coefficient for the system under study, which makes the use of this simple technique dependent on other equipments with their inherent experimental errors. The relative values of the wear constants obtained for different conditions of electroless Ni-P indicate that there is not a direct relationship between the coating hardness and abrasion wear resistance, since other factors are injluencing this behavior, underlying the fact that hardness could not be considered as a key guide to the coating abrasive wear. As can be observed, the abrasive wear constant, Kc, for the Ni-P coating heat-treated at 400°C (a microhardness of 916 HKrm) is higher than the wear constant obtained when the heat treatment was conducted at 260°C (microhardness of 795 HKr,), results which contradict the explanation presented by Duncan [7] which attributed the improvement of abrasive resistance of Ni-P coatings to the Ni3P particle coarsening, as the heat treatment temperatures increase. Also, it has to be mentioned that due to the difference between the thermal coefficients of the coating and the steel substrate, the coating heat-treated at 400°C presented cracking at the surface which could negatively influence its abrasive resistance behavior.
x x x x x x
K,~$& 10-11 10-l” IO“’ 10-14 lo“:’ 10-13
~
-1.85 2.62 5.17 7.74
x x x x
R_’
IO-” lo-” 10-I” 10-I”
0.97 0.97 0.95 4.95 0.91 0.91
As was indicated by Rutherford and Hutchings [A-], the action of the abrasive particles within the contact will depend on the ratio between the film thickness of a lubricant between the sample and the ball and the abrasive mean particle diameter, d. They proposed the following equation in order to evaluate the lubricant film thickness, h,. This equation represents the solution to Reynolds equation in two dimensions for a sphere of radius R, sliding against a plane in the presence of a viscous fluid and is expressed as:
where IVY is the normal load, ~1is the sliding speed and r represents the fluid viscosity. In the present investigation for values of V = 1.11 ms-.!, and R = 0.01905 m, qEthY]e,,e slycOl= 4.3 x IO-’ N m-?-S N, = 0.12, the film thickness has been determined to be about 0.3 1 pm, that is to say smaller than the mean abrasive particle diameter. In these conditions, the ratio hJd is equal to 0.1, indicating that the abrasive particle has a major contribution to the wear process. This can be corroborated by analyzing the scanning electron micrographs of the worn surfaces of the Ni-P coating, heat-treated at 4OO”C, which are presented in Fig. 4a and b. Well-defined wear scars of nearly the same width as the me-anva1u.e of the abrasi-ve particle diameter are-observed, indicat%lg- a pure abrasion wear mechanism, which results in the material removal due to the cutting performed by the abrasive particles. Similar morphologies of the worn surfaces were obtained for all the samples under study. It was considered [8] that when the tribosystem is composed of a two triboelements and lubricant,it~qossible to obtain an indication of the dominating lubrication or wear mechanism by monitoring suitable test parameters such as friction, lubricant film thickness, temperature, wear and surface roughness. In these conditions, three main lubrication regimes have been identified [9] as a function of the variation of the friction coefficient, ,LL,and the film thickness-toroughness ratio, A. In the present work a value of film thickness-to-roughness ratio, X, equal to 0.77 has been determined, which together with a friction coefficient of 0.26 calculated previously possibly could place our system in a boundary lubrication mode, were the tribological behavior is governed by solid-solid friction and wear processes.
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without ensuring the reproducibility of the real operating conditions. A simple indication of the fact that the friction force was not taken into account when reporting the wear constants, as an intrinsic parameter, will be enough to allow comparisons to be made with the work performed in other laboratories. Also, it has to be pointed out the simplicity of the test proposed by us in reference to: {a) use of a mnch simpler expression to calculate the load which involves parameters that can be easily determined, (b) easy measurement of the craters diameters for different sliding distances without the misfit produced when moving the sample or the complicated setup to measure crater in situ, and (c) practical and lowcost setup for the slurry delivery attachment.
Acknowledgements The authors wish to acknowledge the financial support of National Council for Scientific and Technological Research of Venezuela (CONICIT) through the project Sl-2612, to the Council for Scientific and Humanistic Development (CDCH) of the Central University of Venezuela through the project O&17-2789/95, and to the Organization of the American States (OAS) through the Multinational Project on Materials. The financial support of the Postgraduate Studies Council of Central University of Venezuela is also acknowledged.
References Fig. 3. (a and bj Scanning electron micrographs of the worn surface of the heat-treated Ni-P coating [at 300°C ~for 1 h) at different magnifications.
4. Conclusions The analysis carried out in the present work has shown that our initially proposed method is able to quantify the relative abrasive wear resistance of electroless Ni-P coatings which have undergone different post-deposition treatment conditions. Although a significant increase in the values of the individual abrasive wear constants of the coatings were obtained when the friction force was taken into account, in our opinion, this approach will make this simple test troublesome. The inclusion of the friction force implies additional experiments in order to evaluate the dynamic friction coefficient for the system under study, which makes this test costly and dependent on other equipment
[1] M.H. Staia, C. Enriquez, ES. Puchi, B. Lewis and M. Jeandin, in T.S. Sudarshan, M. Jeandin and M. Khor (eds.), S&zce Modificntion Technologies X, Inst. of Materials, UK, 1997, p. 130. [2] A. Kassman. L. Erickson, M. Olsson, P. Hedenqvist, S. Jacobson and S. Hogmark, Su$ Cont. Technoi., 50 (1991) 75. [3] B. Bhushan, in W.B.Harding and G.A. DiBari teds.), O~rrvie~~ of Comings Mareriais, Su&ce Trearmenrs md Screening Techniques for Tribological Appiicntions, Part 2: Screening Techniques, Testing of Mefallic and inorganic Coatings, STP-947, 1987, p. 310. [3] K.L Rutherford and I.M. Hutchings, &lj-: Coot. Technol., 79 (1996), 231. [5] M.H. Staia. E. Castillo, E.S. Puchi, D. B. Lewis and H.E. Hintermann, Surj Cont. Technol., 86-87 (1996) 598. [6] P.A. Steinmann and H.E. Hintermann, J. Var. Sci. Techno/., A7 (3) ( 1989) 2257. [7] R. Duncan, Meral finishing, March (1990) 11. [8] K. Holmberg and A. Matthews, in D. Dowson (ed.), Coating Tribology, Tribology Series, 28, Elsevier, Amsterdam, 1994, p. 307. 191 H. Czichos, Basic Tribological Parameters, ASM Handbook vol. 18, ASM International, 1992, 474.