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The study of scuffing and pitting failure of cam-tappet rubbing pair
Wear, 140 (1990)
135-147
136
The study of scuf%ingand pitting failure of cam-tappet rubbing pair Liu Jiajun, Lu Zhiqiaug and Cheng Yinqian Mechanical Engineering Department, Tsinghua University
(China.
(Received September 7, 1989; revised November 28, 1989; accepted February 6, 1990)
Abstract The mechanism of scufbng is generally considered to be adhesive wear. However, in this research it has been found that the main mechanism is fatigue wear in the form of delamination and that plastic deformation plays a very important role in the scuffing formation, which shows an obvious strain fatigue nature. The analyses of wear debris also verify this conclusion. Pitting is opposite to scuEing and displays a stress fatigue nature. All cracks are formed without detectable plastic deformation. The statistic analysis showed that the majority of pitting cracks initiate on or near the surface. In order to comirm the viewpoint that scffig and pitting failure belong to strain and stress fatigue separately, simulating wear tests were conducted using materials with different structures and hardnesses. On the basis of failme analyses and results of wear tests the mechanism of scufEingand pitting is discussed in detail.
1. Introduction
The cam-tappet, which works in the most severe conditions, is one of the three main rubbing pairs in an internal-combustion engine. Medium plain carbon steel (containing 3.46-3.65wt.%C, 2.1-2.2wt.%S, O&OShvt.%Mn, 0.2-0.4wt.%Cr and 0.3-0.5wt.WMo) has been commonly used as both cam and tappet materials [ 11. The main failure modes of the cam and tappet are scufhng and pitting, and much has been published about the mechanism of these in recent years, but investigations of wear mechanisms directed towards the cam and tappet are still lacking. The nature of scuffing and pitting formation on these actual machine parts remains unclear. The aim of this research is to clarify the mechanism of scuBng and pitting through systematic wear failure analyses and wear tests, and, on the basis of these studies, to define the principles and directions in which to improve the wear resistance of materials. The conclusions obtained also contribute to the better understanding of the mechanisms of scufbng and pitting.
0043-1648/90/$3.50
(8 Elsevier Sequoia/Printed in The Netherlands
2. Failure analysis of scuffing The wear tracks of a large quantity of failed cams and tappets with scuffing had been observed frequently using stereo-microscopy, optic metallographic microscopy and scanning electron microscopy @EM). The observations indicate clearly that various features of macroscopic scuffing have almost similar microscopic morphologies, which may be described as follows. (1) Significant plastic flow occurs on the worn surface. (2) The sctied surface shows the damage feature in the form of delamination in almost all obsemed fields (Fig. 1). (3) Fatigue striation ch~acteristics can be seen in some places where the delaminated layers have just flaked off (Fig. 2). These microscopic features show that the mechanism of scuffing formed on cams and tappets is certainly not adhesive wear as considered generally, but strain fatigue wear [ 21. In order to confirm this, the wear debris was also analysed in detail. Almost all the collected wear debris was of a thin plate-like nature. The dimensions of the platelets are usually larger than the platelet thickness by one to two orders of magnitude (Fig. 3). Using SEM it can be observed that the sIiding surface retains some smoothness, and the back side of some debris shows fatigue fracture. ~he~ore, in the ferrograms a considerable quantity of spherical wear particles was found, which may be considered as evidence of a fatigue process 13, 41 (Fig. 4). In addition, the plastic deformation of the sublayer under the wear track is the most important characteristic of scufhng failure. In the highly deformed region the flow lines are almost parallel to the surface. The cracks can easily initiate in places where microdefects (microvoids, non-metallic inclusions, second-phase particles, cellular structure of dislocations, etc.) exist and
Fig. 1. SEM micrograph showing delaminationon the wear surface of a steel cam. (Magnification, 40X.) Fig. 2. SEM micrograph showing fatigue striation on the wear surface of a steel cam. (Ma~cation, 2400 X .)
Fig. 3. SEM micrograph
showing the smooth surface of wear debris. (Magnification,
Fig. 4. SEM micrograph
showing
Fig. 5. SEM micrograph (Magnification, 1000 X .) Fig. 6. SEM micrograph
(Magnification,2000 X
showing
showing
spherical
particles
130
X .)
among the wear debris. (Magnification,
nucleation
propagation
of a crack in the
highly deformed sublayer.
.)
with the development of plastic deformation. Figures 5 and 6 show the nucleation and propagation of a crack in the highly deformed sublayer of steel cam. In order to express quantitatively the degree of plastic deformation the method of Dautzenberg and Zaat [5] had been adopted, by which the effective deformation 8 at a certain distance from the surface can be calculated. The change in effective deformation along the depth below surface of cam and tappet is shown in Fig. 7. Several points may then be summarized and inferred from this analysis. (1) Much greater effective deformation near the surface (within 1-2 pm) occurs on the tappet than on the cam owing to the higher contact probability. This situation shows that the tappets work under more severe conditions and naturally have a shorter service life. propagate
138
Dept.h
Fig. ‘7. Change tappet.
below
in effective
surface
deformation
along the depth
below the surface
of the cam and
(2) Scufllng takes place in the sublayer with a very large amount of plastic deformation; this determines the strain fatigue nature of this kind of faihire. (3) Scuffing which appeared in the early stage can also be explained by the large amount of plastic deformation on the asperities when the new surfaces come into contact. If the surface smoothness cannot be improved soon afterwards, the development of plastic deformation will continue under the worn surface and certainly lead to severe scuffig. 3. Fahre
analysis of pitting
It is well known that the mechanism of pitting is the fatigue process, including the initiation and propagation of cracks; so it is important to study the characteristics of pitting cracks on the cross-section of sublayer in order to obtain a better understanding of the regularity of pitting formation. Many metallographic specimens displaying pitting cracks have been observed systematically. On the basis of these observations their general features can be obtained. (I) The majority of cracks initiate on the very surface or from the bottom of micropits, which are easily formed on the surface of chilled cast iron (Fig. 8), and propagate with a certain inclination downwards. (2) A smaller percentage of cracks initiate at a certain depth of sublayer and propagate parallel to the surface (Fig. 9). This situation is related closely to the smaller amount of plastic deformation in the sublayer. If an equal quantity of defects exist in the material, cracks will not be induced at the defect because there is insufficient cumulative deformation for scufllng to occur. So the pitting shows a typical stress fatigue nature. The reason that plastic deformation does not occur is related to many factors, such as the load, sliding speed, lubrication and surface smoothness. If other factors are kept unchanged, the smoothness will influence the
Fig. 8. SEM micrograph showing a pitting crack initiated from the bottom of a micropit. (Magnification, 1000 X .) Fig. 9. SEM micrograph showing a pitting crack propagated parallel to the surface. (Magnification, 1000 x .)
b. Fo&ation cementite
a. Original structure
c.
Appearance of Crack at cm. h
e.
Formation micropitting
Fig. 10. Schematic
of
d.
of relief
Propagation of crack
Induction secondary
Of crack
diagram of the pitting formation
mechanism
on the chilled cast iron surface.
plastic deformation significantly. When the running-in period is over, the smoothness increases and the friction coefficient p decreases. In consequence, the plastic deformation will be hardly developed. The fact that the cracks initiate on the surface is obviously a result of the tangential force, which not only increases the maximum shear stress but also shifts the position of the ~maximum shear stress to the surface [ 61. Whether the pitting cracks will occur or not depends on the strength of material, the criterion for which is 7max > a,. The pearlitic matrix of chilled cast iron has a lower strength; consequently, the cracks can initiate more easily in the pea&tic matrix of the surface, which results in the micropitting. The mechanism of the pitting formation induced by micropits could arise because owing to the appearance of micropits at the pearlite matrix the cementite will take on a relief or large asperity nature, and the hertzian
140
stress distribution will be modulated according to the model of Berthe and Michau [ 71, i.e. each cementite asperity can generate a micro-hertzian stress field. Then the size of the micropits can increase and secondary cracks may be induced from weaker places such as the bottom of the micropits. This process can be shown in the scheme of Fig. 10. 4. Wear
tests on the cam and tappet
4.1. Experimental d&ads 4.1. I. Apparatus and specimens
Experiments were conducted on the sliding-wear-testing machine commonly used for evaluating the wear resistances of cam and tappet. The load is applied by the pressure of a spiral spring; the rotative speed of cam can be adjusted through a silicon-controlled rectifier; both parameters can be set up so as to be continuous. The cam specimens were manufactured from the same material and were subjected to the same heat treatment as real cams (the diameter of the cam base circle was 33.7 mm, the cam width was 16.5 mm, the cam lift was 7.2 mm and the radius of nose curvature was 5.2 mm). The Rockwell C hardness of the nose was about 55-60 HRC. The tappet specimens were cut off directly from the head and then heat treated to different hardnesses. This rubbing pair was lubricated with No. 10 engine oil (the viscosity was 7-13 cSt at 372 K without additives). 4.1.2. Test conditions The wear tests were conducted under the following conditions. (I) In tests for measuring the weight loss of tappet specimens worn by scuffing, the tappet specimen must be kept fixed without any rotation. A load of 1100 N and a rotative speed of 800 rev min-’ were used in one group of experiments. In another group the load was changed from 200 to 3000 N and the rotative speed remained the same. In aII these experiments, tappet specimens with various hardnesses were sliding against the same steel cams. (II) In tests for determining the pitting life of tappet specimens, pitting was formed when the tappet specimen was rotating naturaIIy around its own axis during the wear test. The contact stress was changed from 650 to 1000 N mmP2 and a rotative speed of 800 rev min-’ was used. Two hardness levels of chilled cast iron tappets were compared in this experiment. 4.1.3. Wear
rneasur~ts
The wear rate of tappet specimens was determined by measuring their weight loss on a balance with a sensitivity of 0.1 mgf after a certain time of sliding until the total sliding distance reaches about 12.5 X lo4 m. Before measurement, the specimens were ultrasonically cleaned in acetone. Weight measurements were repeated until at least three consecutive weighings agreed to within 0.5 mgf.
141
The pitting life was determined by the number of cam cycles to when the first pitting on the tappet surface just occurred.
The friction coefficient p between the steel cam and chilled cast iron tappet was measured on a standard Timken testing machine. ,Uwas found to be about 0.12-0.15. 4.2. Results The results of wear tests showing the effect of different tappet hardnesses on their wear resistance are shown in Figs. 11 and 12. It can be seen that the wear rates are quite different in the early unsteady state and in the subsequent steady state. The wear rate R0 = AW*/~~ in the unsteady state is reduced gradually on increase in the sliding distance; this corresponds to the running-m period. The wear rate in the steady state is lowered and has the linear relation AW-AW, L-Lo
R=
AW, and R represent qu~titatively the trend of the wear resistance of materials; so they can be considered as two characteristic values of sliding wear under a certain condition. Their values obtained in this experiment show that the change in AW, for specimens with various hardnesses is more obvious than the change in R in the steady state. This phenomenon can be explained as owing to the contact of asperities in the early stage. When the unsteady state is over, plastic deformation approaches a saturated value [ 81; meanwhile the real contact area will increase with ~provement in surface smoothness and therefore a decrease in wear rate can be expected. However, in both stages the effects of hardness show similar trends, i.e. the higher the hardness the greater are the values of AW, and R. The values of AWO
o/do+
5
Sliding Fig.
I0
distance
I5 L.XIO~SII
Sliding
distance
LX~O~~
11. Wear behaviour of steel tappets with different hardnesses (1100 N; 800 rev min-I).
Fig. 12. Wear behaviour of chilled cast iron tappets with ditferent hardnesses (1100 N; 800 rev mir-‘).
Martensitic
Pearlitic
Fig. 13. Relation between R,, and the load for a steel tappet (800 rev mir-‘). Fig. 14. Relation between R,, and the load for a chilled cast iron tappet (800 rev mir-I).
KPa c
60. to6
Cycles of cam N
10'
Fig. 15. S-N curve of chilled cast iron tappets with different hardnesses. and R for chilled cast iron are smaller than those for steel, this obviously shows the better wear resistance of the former owing to its composite structure. The wear rate R,,of tappet specimens with different structures vs. load are plotted in Figs. 13 and 14. The results show that the wear rate increases with increase in load. A sudden rise in R. appears when the load reaches a certain level, and severe heating owing to friction was observed at the same time. It can be considered that the sudden change in R. is due to the intensification of plastic deformation and the higher temperature rise, which may cause an increase in the crack propagation rate of materials. The data on wear rate also become more scattered because of the random temperature rises.
143
It shows the opposite feature to that of scuffing resistance, i.e. the pitting resistance of chilled cast iron with a martensitic matrix, (a Vickers hardness of 610 HV) is higher than that of a pearlitic matrix (a Vickers hardness of 320 HV). This result agrees quite well with the research on chilled cast iron tappets with different matrixes [9] (Fig. 15) It proves that the initiation of cracks is the dom~ant factor for determining the fatigue life of pitting formation; so, the higher the strength of material is, the greater the pitting resistance will be.
5. Discussion about the scufFing and pitting mechanism Archard’s adhesion theory is commonly accepted nowadays to explain the mechanism of scuEng, i.e. the sliding wear volume of materials is inversely proportional to their hardness [lo], but in this experiment completely opposite results were obtained. The results of wear tests and previous wear failure analyses state clearly that the scuffig mechanism of a cam-tappet pair does not follow Archard’s model and is strain fatigue wear. Therefore it can be considered that the dominant factor determining the wear rate is not the nucleation but the propagation stage of cracks, and the crack propagation rate of materials will become one of the main parameters influencing the scuffing resistance. The crack propagation rates of steels at various states of heat treatment are different; for instance, the experimental results of Barson [ 1 l] are as follows: for martensitic steel,
and, for pea&tic steel,
Of course, the propagation of cracks in the wear process is very complicated; it cannot be described simply using the crack models of type I or II, but there must be common tendencies. In order to evaluate qualitatively the crack propagation rate of a tappet material, a simple model (Fig. 16) has been suggested. Suppose that the specimen is worn off by plate-like debris of equal thickness h, the weight loss will be AW=nbBhm The number of asperities necessary for wearing off one layer will be bBhm V=RA
144
h
Fig. 16. The simple model of tappet wear.
where h can be measured with a profilometer, h can be obtained from the wear debris analysis and other values are all measurable. So the crack propagation rate duldv can be calculated approximately from A/v. The wear rate of material corresponds well to the crack propagation rate. Therefore it may be considered that all the factors affecting the crack propagation rate will be the main parameters of wear rate, e.g. hardness, structure and temperature. As for the pitting resistance, we can express the total pitting life Nf as the sum of the non-cracking life No and the crack propagation life N,: N;=N*+N, They can be converted to the passing numbers of asperities for convenient calculation: 2777-N N’=A Np=-
D da/d v
If A= 200 pm, r= 20 mm, N= 5.6 X lo6 cycles, the average diameter of pitting is about 2 mm and da/dv= 2 x 10W3 (obtained from the equation A./Y), we have N =40xrrX5*6X f 200x10-3 N,,=
106__3 5X IO9 *
2000 =106 2x 10-3
NP F=o.3
x 10-3
f
This states clearly that the crack propagation life is only less than 0.1% of the total pitting life. Therefore the key to increasing the pitting resistance
145
lies in raising the resistance to the initiation of cracks, i.e. in improving the strength of materials. Considering extensively the regularities of scuffing and pitting formation, it is appropriate to describe the wear process by Morrow’s equation for the fatigue life [ 121: %t = %e+ Gip=$
(UVf)b+
q’(2NJ
where b and c are the exponents of strength and plasticity. When E,,~=O, the fatigue life is determined by the first term, corresponding to the situation of stress fatigue, which may be expressed by the S-N curve. When the plastic deformation becomes larger, the fatigue life will be determined by the second term, corresponding to strain fatigue, which must be described by the Manson-CofIin curve. When both kinds of deformation (elastic and plastic) exist simultaneously, the total amplitude of strain is composed of two terms as shown in Fig. 17. In the running-in period the surface is relatively rough; the wear mechanism is precisely the strain fatigue, which behaves as scuffing, corresponding to the left-hand side of the ‘ntersection in Fig. 17. Along with the development of wear process, the smoothness of surface improves, and the wear mechanism becomes stress fatigue, which behaves as pitting, corresponding to the righthand side of the intersection. On the basis of the above analyses, we can imagine that on the same cam-tappet pair a conversion of wear mode from scuffing to pitting may occur. The scufhng phenomenon appears generally in the early stage; once this stage has passed safely, scufhng will hardly be encountered hereafter in normal working conditions. However, if the early scuffig cannot be made to cease effectively, a vicious circle will happen and lead finally to severe wear failure.
LogZNf
Fig. 17. Schematic diagram for Morrow’s equation.
146
6. Conclusions
(1) The dominant scuffing mechanism of a cam-tappet pair is not adhesive wear but is fatigue wear in the form of delamination, which is due to strain fatigue. The adhesion between metals may be avoided to a great extent by the existence of a boundary lubrication film. (2) The controlling factor of strain fatigue life is the crack propagation rate. Therefore, the higher the hardness of a material, the quicker the crack propagation rate and the greater the tendency to scuffing will be. (3) The mechanism of pitting is stress fatigue. The main factor determining the stress fatigue life is the strength of the material. Therefore, the stronger the material is, the higher the pitting resistance will be. The majority of pitting cracks initiate on or near the surface. (4) The wear mode may be converted from scuG.ng to pitting with a change in smoothness from a lower to a higher level. (5) The application of general fatigue theory to the wear process is an important development in the wear theory, but this application is still in the preliminary stage. Many problems need to be studied extensively.
Acknowledgements
This work was carried out in the metallographic laboratory of the Mechanical Engineering Department. The authors would like to thank the technicians for their help, and No. 2 Automobile Factory for offering the failed cams and tappets.
References 1 The chilled cast iron tappet of gas valve, I@rm. Sci. Techml. Rep. I, 1978 (No. 2 Automobile Factory). 2 R. G. Campany and R. W. Wilson, The metalhugy of scoring and scuf3ng failure, Proc. 9th Leeds-Lgon Sywzp. cm Tribologg, 1982, p. 201. 3 D. Scott, Debris examination-a prognostic approach to failure prevention, Wear, 34 (1975) 15. 4 E. Rabinowicz, The formation of spherical wear particles, Wear 42 (1977) 149. 5 J. H. Dautzenberg and J. H. Zaat, Quantitative determination of deformation by sliding wear, Wear, 23 (1973) 9. 6 J. 0. Smith and C. K. Liu, Stresses due to tangential and normal loads on an elastic solid with application to some contact stress problems, Iprans. ASIDE, 75 (1953) 157. 7 D. Berthe and B. Michau, Effect of roughness ration and Hertz pressure on micropits and spalls in concentrated contacts, Proc. 4thLeed.-L@onSymp. on l%%bology, 1977, Mechanical Engineering Publications, London, 1989, p. 233. 8 D. A. Rigney and J. P. Hirth, Plastic deformation and sliding friction of metals, Wear, 53 (1979) 345. 9 Cheng Jiaxiang, Ao Bmgqiu, Zhang Hongying and Peng Guangqi, The measures of chilled cast iron tappet, Tech. Rep. 1982 (No. 2 Automobile Factory). 10 T. A. Storske, Adhesive wear of lubricated contacts, Ipribol. ht., 12 (4) (1979) 169.
147 11 J. M. Barson, J. Eng. Ind., (4) (1971) 1190. 12 S. Kocanda, Fatiguefailure of metals, The Netherlands 1978.
Appendix
Sijthoff and Noordhoff International,
A: Nomenclature
da/dN crack propagation rate crack length width of wear track length of wear track average diameter of pitting modulus of elasticity thickness of platelike wear debris total sliding distance sliding distance of unsteady state specific gravity number of worn layers number of cycles of pitting life total pitting life crack propagating life non-cracking life radius of the cam specimen in the simple model wear rate of the steady state wear rate of the unsteady state total weight loss of the tappet specimen weight loss of the tappet specimen in the unsteady state effective deformation total strain amplitude elastic strain amplitude plastic strain amplitude plasticity coefficient of fatigue average spacing of asperities friction coefficient number of asperities strength coefficient of fatigue shear strength of material maximum shear stress
135-147
136
The study of scuf%ingand pitting failure of cam-tappet rubbing pair Liu Jiajun, Lu Zhiqiaug and Cheng Yinqian Mechanical Engineering Department, Tsinghua University
(China.
(Received September 7, 1989; revised November 28, 1989; accepted February 6, 1990)
Abstract The mechanism of scufbng is generally considered to be adhesive wear. However, in this research it has been found that the main mechanism is fatigue wear in the form of delamination and that plastic deformation plays a very important role in the scuffing formation, which shows an obvious strain fatigue nature. The analyses of wear debris also verify this conclusion. Pitting is opposite to scuEing and displays a stress fatigue nature. All cracks are formed without detectable plastic deformation. The statistic analysis showed that the majority of pitting cracks initiate on or near the surface. In order to comirm the viewpoint that scffig and pitting failure belong to strain and stress fatigue separately, simulating wear tests were conducted using materials with different structures and hardnesses. On the basis of failme analyses and results of wear tests the mechanism of scufEingand pitting is discussed in detail.
1. Introduction
The cam-tappet, which works in the most severe conditions, is one of the three main rubbing pairs in an internal-combustion engine. Medium plain carbon steel (containing 3.46-3.65wt.%C, 2.1-2.2wt.%S, O&OShvt.%Mn, 0.2-0.4wt.%Cr and 0.3-0.5wt.WMo) has been commonly used as both cam and tappet materials [ 11. The main failure modes of the cam and tappet are scufhng and pitting, and much has been published about the mechanism of these in recent years, but investigations of wear mechanisms directed towards the cam and tappet are still lacking. The nature of scuffing and pitting formation on these actual machine parts remains unclear. The aim of this research is to clarify the mechanism of scuBng and pitting through systematic wear failure analyses and wear tests, and, on the basis of these studies, to define the principles and directions in which to improve the wear resistance of materials. The conclusions obtained also contribute to the better understanding of the mechanisms of scufbng and pitting.
0043-1648/90/$3.50
(8 Elsevier Sequoia/Printed in The Netherlands
2. Failure analysis of scuffing The wear tracks of a large quantity of failed cams and tappets with scuffing had been observed frequently using stereo-microscopy, optic metallographic microscopy and scanning electron microscopy @EM). The observations indicate clearly that various features of macroscopic scuffing have almost similar microscopic morphologies, which may be described as follows. (1) Significant plastic flow occurs on the worn surface. (2) The sctied surface shows the damage feature in the form of delamination in almost all obsemed fields (Fig. 1). (3) Fatigue striation ch~acteristics can be seen in some places where the delaminated layers have just flaked off (Fig. 2). These microscopic features show that the mechanism of scuffing formed on cams and tappets is certainly not adhesive wear as considered generally, but strain fatigue wear [ 21. In order to confirm this, the wear debris was also analysed in detail. Almost all the collected wear debris was of a thin plate-like nature. The dimensions of the platelets are usually larger than the platelet thickness by one to two orders of magnitude (Fig. 3). Using SEM it can be observed that the sIiding surface retains some smoothness, and the back side of some debris shows fatigue fracture. ~he~ore, in the ferrograms a considerable quantity of spherical wear particles was found, which may be considered as evidence of a fatigue process 13, 41 (Fig. 4). In addition, the plastic deformation of the sublayer under the wear track is the most important characteristic of scufhng failure. In the highly deformed region the flow lines are almost parallel to the surface. The cracks can easily initiate in places where microdefects (microvoids, non-metallic inclusions, second-phase particles, cellular structure of dislocations, etc.) exist and
Fig. 1. SEM micrograph showing delaminationon the wear surface of a steel cam. (Magnification, 40X.) Fig. 2. SEM micrograph showing fatigue striation on the wear surface of a steel cam. (Ma~cation, 2400 X .)
Fig. 3. SEM micrograph
showing the smooth surface of wear debris. (Magnification,
Fig. 4. SEM micrograph
showing
Fig. 5. SEM micrograph (Magnification, 1000 X .) Fig. 6. SEM micrograph
(Magnification,2000 X
showing
showing
spherical
particles
130
X .)
among the wear debris. (Magnification,
nucleation
propagation
of a crack in the
highly deformed sublayer.
.)
with the development of plastic deformation. Figures 5 and 6 show the nucleation and propagation of a crack in the highly deformed sublayer of steel cam. In order to express quantitatively the degree of plastic deformation the method of Dautzenberg and Zaat [5] had been adopted, by which the effective deformation 8 at a certain distance from the surface can be calculated. The change in effective deformation along the depth below surface of cam and tappet is shown in Fig. 7. Several points may then be summarized and inferred from this analysis. (1) Much greater effective deformation near the surface (within 1-2 pm) occurs on the tappet than on the cam owing to the higher contact probability. This situation shows that the tappets work under more severe conditions and naturally have a shorter service life. propagate
138
Dept.h
Fig. ‘7. Change tappet.
below
in effective
surface
deformation
along the depth
below the surface
of the cam and
(2) Scufllng takes place in the sublayer with a very large amount of plastic deformation; this determines the strain fatigue nature of this kind of faihire. (3) Scuffing which appeared in the early stage can also be explained by the large amount of plastic deformation on the asperities when the new surfaces come into contact. If the surface smoothness cannot be improved soon afterwards, the development of plastic deformation will continue under the worn surface and certainly lead to severe scuffig. 3. Fahre
analysis of pitting
It is well known that the mechanism of pitting is the fatigue process, including the initiation and propagation of cracks; so it is important to study the characteristics of pitting cracks on the cross-section of sublayer in order to obtain a better understanding of the regularity of pitting formation. Many metallographic specimens displaying pitting cracks have been observed systematically. On the basis of these observations their general features can be obtained. (I) The majority of cracks initiate on the very surface or from the bottom of micropits, which are easily formed on the surface of chilled cast iron (Fig. 8), and propagate with a certain inclination downwards. (2) A smaller percentage of cracks initiate at a certain depth of sublayer and propagate parallel to the surface (Fig. 9). This situation is related closely to the smaller amount of plastic deformation in the sublayer. If an equal quantity of defects exist in the material, cracks will not be induced at the defect because there is insufficient cumulative deformation for scufllng to occur. So the pitting shows a typical stress fatigue nature. The reason that plastic deformation does not occur is related to many factors, such as the load, sliding speed, lubrication and surface smoothness. If other factors are kept unchanged, the smoothness will influence the
Fig. 8. SEM micrograph showing a pitting crack initiated from the bottom of a micropit. (Magnification, 1000 X .) Fig. 9. SEM micrograph showing a pitting crack propagated parallel to the surface. (Magnification, 1000 x .)
b. Fo&ation cementite
a. Original structure
c.
Appearance of Crack at cm. h
e.
Formation micropitting
Fig. 10. Schematic
of
d.
of relief
Propagation of crack
Induction secondary
Of crack
diagram of the pitting formation
mechanism
on the chilled cast iron surface.
plastic deformation significantly. When the running-in period is over, the smoothness increases and the friction coefficient p decreases. In consequence, the plastic deformation will be hardly developed. The fact that the cracks initiate on the surface is obviously a result of the tangential force, which not only increases the maximum shear stress but also shifts the position of the ~maximum shear stress to the surface [ 61. Whether the pitting cracks will occur or not depends on the strength of material, the criterion for which is 7max > a,. The pearlitic matrix of chilled cast iron has a lower strength; consequently, the cracks can initiate more easily in the pea&tic matrix of the surface, which results in the micropitting. The mechanism of the pitting formation induced by micropits could arise because owing to the appearance of micropits at the pearlite matrix the cementite will take on a relief or large asperity nature, and the hertzian
140
stress distribution will be modulated according to the model of Berthe and Michau [ 71, i.e. each cementite asperity can generate a micro-hertzian stress field. Then the size of the micropits can increase and secondary cracks may be induced from weaker places such as the bottom of the micropits. This process can be shown in the scheme of Fig. 10. 4. Wear
tests on the cam and tappet
4.1. Experimental d&ads 4.1. I. Apparatus and specimens
Experiments were conducted on the sliding-wear-testing machine commonly used for evaluating the wear resistances of cam and tappet. The load is applied by the pressure of a spiral spring; the rotative speed of cam can be adjusted through a silicon-controlled rectifier; both parameters can be set up so as to be continuous. The cam specimens were manufactured from the same material and were subjected to the same heat treatment as real cams (the diameter of the cam base circle was 33.7 mm, the cam width was 16.5 mm, the cam lift was 7.2 mm and the radius of nose curvature was 5.2 mm). The Rockwell C hardness of the nose was about 55-60 HRC. The tappet specimens were cut off directly from the head and then heat treated to different hardnesses. This rubbing pair was lubricated with No. 10 engine oil (the viscosity was 7-13 cSt at 372 K without additives). 4.1.2. Test conditions The wear tests were conducted under the following conditions. (I) In tests for measuring the weight loss of tappet specimens worn by scuffing, the tappet specimen must be kept fixed without any rotation. A load of 1100 N and a rotative speed of 800 rev min-’ were used in one group of experiments. In another group the load was changed from 200 to 3000 N and the rotative speed remained the same. In aII these experiments, tappet specimens with various hardnesses were sliding against the same steel cams. (II) In tests for determining the pitting life of tappet specimens, pitting was formed when the tappet specimen was rotating naturaIIy around its own axis during the wear test. The contact stress was changed from 650 to 1000 N mmP2 and a rotative speed of 800 rev min-’ was used. Two hardness levels of chilled cast iron tappets were compared in this experiment. 4.1.3. Wear
rneasur~ts
The wear rate of tappet specimens was determined by measuring their weight loss on a balance with a sensitivity of 0.1 mgf after a certain time of sliding until the total sliding distance reaches about 12.5 X lo4 m. Before measurement, the specimens were ultrasonically cleaned in acetone. Weight measurements were repeated until at least three consecutive weighings agreed to within 0.5 mgf.
141
The pitting life was determined by the number of cam cycles to when the first pitting on the tappet surface just occurred.
The friction coefficient p between the steel cam and chilled cast iron tappet was measured on a standard Timken testing machine. ,Uwas found to be about 0.12-0.15. 4.2. Results The results of wear tests showing the effect of different tappet hardnesses on their wear resistance are shown in Figs. 11 and 12. It can be seen that the wear rates are quite different in the early unsteady state and in the subsequent steady state. The wear rate R0 = AW*/~~ in the unsteady state is reduced gradually on increase in the sliding distance; this corresponds to the running-m period. The wear rate in the steady state is lowered and has the linear relation AW-AW, L-Lo
R=
AW, and R represent qu~titatively the trend of the wear resistance of materials; so they can be considered as two characteristic values of sliding wear under a certain condition. Their values obtained in this experiment show that the change in AW, for specimens with various hardnesses is more obvious than the change in R in the steady state. This phenomenon can be explained as owing to the contact of asperities in the early stage. When the unsteady state is over, plastic deformation approaches a saturated value [ 81; meanwhile the real contact area will increase with ~provement in surface smoothness and therefore a decrease in wear rate can be expected. However, in both stages the effects of hardness show similar trends, i.e. the higher the hardness the greater are the values of AW, and R. The values of AWO
o/do+
5
Sliding Fig.
I0
distance
I5 L.XIO~SII
Sliding
distance
LX~O~~
11. Wear behaviour of steel tappets with different hardnesses (1100 N; 800 rev min-I).
Fig. 12. Wear behaviour of chilled cast iron tappets with ditferent hardnesses (1100 N; 800 rev mir-‘).
Martensitic
Pearlitic
Fig. 13. Relation between R,, and the load for a steel tappet (800 rev mir-‘). Fig. 14. Relation between R,, and the load for a chilled cast iron tappet (800 rev mir-I).
KPa c
60. to6
Cycles of cam N
10'
Fig. 15. S-N curve of chilled cast iron tappets with different hardnesses. and R for chilled cast iron are smaller than those for steel, this obviously shows the better wear resistance of the former owing to its composite structure. The wear rate R,,of tappet specimens with different structures vs. load are plotted in Figs. 13 and 14. The results show that the wear rate increases with increase in load. A sudden rise in R. appears when the load reaches a certain level, and severe heating owing to friction was observed at the same time. It can be considered that the sudden change in R. is due to the intensification of plastic deformation and the higher temperature rise, which may cause an increase in the crack propagation rate of materials. The data on wear rate also become more scattered because of the random temperature rises.
143
It shows the opposite feature to that of scuffing resistance, i.e. the pitting resistance of chilled cast iron with a martensitic matrix, (a Vickers hardness of 610 HV) is higher than that of a pearlitic matrix (a Vickers hardness of 320 HV). This result agrees quite well with the research on chilled cast iron tappets with different matrixes [9] (Fig. 15) It proves that the initiation of cracks is the dom~ant factor for determining the fatigue life of pitting formation; so, the higher the strength of material is, the greater the pitting resistance will be.
5. Discussion about the scufFing and pitting mechanism Archard’s adhesion theory is commonly accepted nowadays to explain the mechanism of scuEng, i.e. the sliding wear volume of materials is inversely proportional to their hardness [lo], but in this experiment completely opposite results were obtained. The results of wear tests and previous wear failure analyses state clearly that the scuffig mechanism of a cam-tappet pair does not follow Archard’s model and is strain fatigue wear. Therefore it can be considered that the dominant factor determining the wear rate is not the nucleation but the propagation stage of cracks, and the crack propagation rate of materials will become one of the main parameters influencing the scuffing resistance. The crack propagation rates of steels at various states of heat treatment are different; for instance, the experimental results of Barson [ 1 l] are as follows: for martensitic steel,
and, for pea&tic steel,
Of course, the propagation of cracks in the wear process is very complicated; it cannot be described simply using the crack models of type I or II, but there must be common tendencies. In order to evaluate qualitatively the crack propagation rate of a tappet material, a simple model (Fig. 16) has been suggested. Suppose that the specimen is worn off by plate-like debris of equal thickness h, the weight loss will be AW=nbBhm The number of asperities necessary for wearing off one layer will be bBhm V=RA
144
h
Fig. 16. The simple model of tappet wear.
where h can be measured with a profilometer, h can be obtained from the wear debris analysis and other values are all measurable. So the crack propagation rate duldv can be calculated approximately from A/v. The wear rate of material corresponds well to the crack propagation rate. Therefore it may be considered that all the factors affecting the crack propagation rate will be the main parameters of wear rate, e.g. hardness, structure and temperature. As for the pitting resistance, we can express the total pitting life Nf as the sum of the non-cracking life No and the crack propagation life N,: N;=N*+N, They can be converted to the passing numbers of asperities for convenient calculation: 2777-N N’=A Np=-
D da/d v
If A= 200 pm, r= 20 mm, N= 5.6 X lo6 cycles, the average diameter of pitting is about 2 mm and da/dv= 2 x 10W3 (obtained from the equation A./Y), we have N =40xrrX5*6X f 200x10-3 N,,=
106__3 5X IO9 *
2000 =106 2x 10-3
NP F=o.3
x 10-3
f
This states clearly that the crack propagation life is only less than 0.1% of the total pitting life. Therefore the key to increasing the pitting resistance
145
lies in raising the resistance to the initiation of cracks, i.e. in improving the strength of materials. Considering extensively the regularities of scuffing and pitting formation, it is appropriate to describe the wear process by Morrow’s equation for the fatigue life [ 121: %t = %e+ Gip=$
(UVf)b+
q’(2NJ
where b and c are the exponents of strength and plasticity. When E,,~=O, the fatigue life is determined by the first term, corresponding to the situation of stress fatigue, which may be expressed by the S-N curve. When the plastic deformation becomes larger, the fatigue life will be determined by the second term, corresponding to strain fatigue, which must be described by the Manson-CofIin curve. When both kinds of deformation (elastic and plastic) exist simultaneously, the total amplitude of strain is composed of two terms as shown in Fig. 17. In the running-in period the surface is relatively rough; the wear mechanism is precisely the strain fatigue, which behaves as scuffing, corresponding to the left-hand side of the ‘ntersection in Fig. 17. Along with the development of wear process, the smoothness of surface improves, and the wear mechanism becomes stress fatigue, which behaves as pitting, corresponding to the righthand side of the intersection. On the basis of the above analyses, we can imagine that on the same cam-tappet pair a conversion of wear mode from scuffing to pitting may occur. The scufhng phenomenon appears generally in the early stage; once this stage has passed safely, scufhng will hardly be encountered hereafter in normal working conditions. However, if the early scuffig cannot be made to cease effectively, a vicious circle will happen and lead finally to severe wear failure.
LogZNf
Fig. 17. Schematic diagram for Morrow’s equation.
146
6. Conclusions
(1) The dominant scuffing mechanism of a cam-tappet pair is not adhesive wear but is fatigue wear in the form of delamination, which is due to strain fatigue. The adhesion between metals may be avoided to a great extent by the existence of a boundary lubrication film. (2) The controlling factor of strain fatigue life is the crack propagation rate. Therefore, the higher the hardness of a material, the quicker the crack propagation rate and the greater the tendency to scuffing will be. (3) The mechanism of pitting is stress fatigue. The main factor determining the stress fatigue life is the strength of the material. Therefore, the stronger the material is, the higher the pitting resistance will be. The majority of pitting cracks initiate on or near the surface. (4) The wear mode may be converted from scuG.ng to pitting with a change in smoothness from a lower to a higher level. (5) The application of general fatigue theory to the wear process is an important development in the wear theory, but this application is still in the preliminary stage. Many problems need to be studied extensively.
Acknowledgements
This work was carried out in the metallographic laboratory of the Mechanical Engineering Department. The authors would like to thank the technicians for their help, and No. 2 Automobile Factory for offering the failed cams and tappets.
References 1 The chilled cast iron tappet of gas valve, I@rm. Sci. Techml. Rep. I, 1978 (No. 2 Automobile Factory). 2 R. G. Campany and R. W. Wilson, The metalhugy of scoring and scuf3ng failure, Proc. 9th Leeds-Lgon Sywzp. cm Tribologg, 1982, p. 201. 3 D. Scott, Debris examination-a prognostic approach to failure prevention, Wear, 34 (1975) 15. 4 E. Rabinowicz, The formation of spherical wear particles, Wear 42 (1977) 149. 5 J. H. Dautzenberg and J. H. Zaat, Quantitative determination of deformation by sliding wear, Wear, 23 (1973) 9. 6 J. 0. Smith and C. K. Liu, Stresses due to tangential and normal loads on an elastic solid with application to some contact stress problems, Iprans. ASIDE, 75 (1953) 157. 7 D. Berthe and B. Michau, Effect of roughness ration and Hertz pressure on micropits and spalls in concentrated contacts, Proc. 4thLeed.-L@onSymp. on l%%bology, 1977, Mechanical Engineering Publications, London, 1989, p. 233. 8 D. A. Rigney and J. P. Hirth, Plastic deformation and sliding friction of metals, Wear, 53 (1979) 345. 9 Cheng Jiaxiang, Ao Bmgqiu, Zhang Hongying and Peng Guangqi, The measures of chilled cast iron tappet, Tech. Rep. 1982 (No. 2 Automobile Factory). 10 T. A. Storske, Adhesive wear of lubricated contacts, Ipribol. ht., 12 (4) (1979) 169.
147 11 J. M. Barson, J. Eng. Ind., (4) (1971) 1190. 12 S. Kocanda, Fatiguefailure of metals, The Netherlands 1978.
Appendix
Sijthoff and Noordhoff International,
A: Nomenclature
da/dN crack propagation rate crack length width of wear track length of wear track average diameter of pitting modulus of elasticity thickness of platelike wear debris total sliding distance sliding distance of unsteady state specific gravity number of worn layers number of cycles of pitting life total pitting life crack propagating life non-cracking life radius of the cam specimen in the simple model wear rate of the steady state wear rate of the unsteady state total weight loss of the tappet specimen weight loss of the tappet specimen in the unsteady state effective deformation total strain amplitude elastic strain amplitude plastic strain amplitude plasticity coefficient of fatigue average spacing of asperities friction coefficient number of asperities strength coefficient of fatigue shear strength of material maximum shear stress