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Analysis of the friction and wear mechanisms of multilayered plasma-sprayed ceramic coatings
201
Wear, 160 (1993) 201-212
Analysis of the friction and wear mechanisms plasma-sprayed ceramic coatings
of multilayered
Jen Fin Lin and Tzuen Ren Li Department
ofMechanicai Engine&q,
National Cheng Kung University, Tainun (Taiwan)
(Received December 17, 1991; revised and accepted July 16, 1992)
Abstract Wear tests were conducted on a rotor-vane-disk adaptor where three rotating vanes were pressed against a disk. Vanes were coated by WC and used as the upper specimen while the disk was coated by Cr,O, and used as the lower specimen. A buffer layer of various thicknesses and contents was placed between the top coating and the bulk steel of the disk to alleviate the effects of the large difference in thermal properties of the two materials. The experimental results reveal that correct placement of a buffer layer can indeed improve the wear resistance. Factors such as the temperature to which the specimen was heated before testing, the proportion by weight of each individual constituent in the buffer layer, and the thickness of each coating layer, were also important for the volume of wear of the lower specimen. As the specimens were heated to higher temperatures, the wear volume decreased with increasing proportions of Cr,O, in the buffer layer. Elevating the preheating temperature of the specimens can diminish the wear volume but increases the friction coefficient. The steady-state wear rate is not much influenced by the constituents of the buffer layer and the coating thickness. Brittle fracture, abrasion, adhesion and oxidation were found to be the primary wear mechanisms in the tests.
1. Introduction Several engineering applications which require high wear resistance often employ ceramics as coating materials. Wear resistance is not only a fundamental material property but also depends on the entire system which includes the ambient atmosphere, material thermal properties, specimen temperature and operation conditions. In the literature [l-25], ceramic-ceramic couples or ceramic-metal couples have been employed in most experiments aimed at understanding wear mechanisms in ceramic materials. Since ceramics are nominally brittle, wear is expected to occur by brittle fracture, abrasion, adhesion and oxidation. However, ample evidence suggests that at high temperature plastic deformation becomes an important wear mechanism. This plasticity is attributed to the inhibition of fracture by the hydrostatic pressure associated with the contact conditions. The wear resistance of materials can be effectively improved by (1) planting a buffer layer between the top coating and the bulk material, (2) altering the proportion by weight of constituents in the buffer layer, and (3) changing the starting temperature of specimens before testing. The pa~icipation of the buffer layer can alleviate stresses arising owing to the notable dif-
0043-1648/93,‘$6.00
ference in thermal properties between the top coating and bulk materials, and can thus improve wear resistance. For the purpose of examining the effect of changing the coating thickness and the proportion of each individual constituent of the buffer layer on wear volume and wear mechanisms, experiments were conducted by designating specimen preheating temperatures before testing. However, surfaces contacting simultaneously at both high load and high specimen temperature were excluded from the present investigation because these conditions induce rapid cracking of the surface. The coating layers were deposited by plasma spraying onto AISI steels. The coatings of the rotor vane (upper specimen) consisted of two materials: NI-Cr alloy as the bottom layer and WC as the top layer. The experimental conditions were classified in three groups. To investigate the effect of coating thickness on wear resistance, the disk (bottom specimen) coatings were placed in six sets marked with Arabic numerals, as shown in Table 1. The specimens in set 1 had singlelayer coatings; CrzO, was the sole coating material. Sets 2 and 3 had double-layer coatings, with Ni-Cr alloy as the bottom layer and Cr,O, ceramic material as the top layer. The specimens of sets 4-6 had threelayer coatings, with Ni-Cr as the bottom layer, mixtures of different proportions of Ni-Cr and Cr,O, as the
0 1993 - Elsevier Sequoia. All rights reserved
202 TABLE
J. E: Lin, T. H. Li I Plasma-sprayed ceramic coatings 1. Arrangement
of rotor vanes and disks in pairs and operation
Experimental conditions
Group Load (N) Temperature (“C) Roughness R, (pm) Sliding velocity (m SK’) Lubrication
Coating thickness
Top specimen, WC
I 89 22 0.4 1 Dry contacts
conditions
11 33.4 150 0.4 1 Dry contacts
III 33.4 250 0.4 1 Dry contacts
Bottom layer (100% Ni-Cr Top layer (100% WC)
alloy)
0.05 0.35
(mm) Bottom Cr203
specimen,
Set 1 0.30 0.45 0.60
One-layer coating (100% Cr,O,) Twq;layer coatmg
Bottom
layer (100% Ni-Cr
alloy)
Top layer (100% Cr20,)
Three-layer coating
Bottom layer (100% Ni-Cr
alloy)
Buffer layer (75% Ni-Cr+2.5% CrzOs) (50% Ni-Cr + 50% CrzOJ (25% Ni-Cr + 75% Cr203) Top layer (100% Cr,O,)
buffer layer, and 100% Cr,O, as the top layer. It should be noted that the specimens of sets 5 and 6 actually have the same constituents, but they were intentionally arranged to examine the effect of preheating temperature on the wear behavior. These specimens were provided with three coating thicknesses: 0.3 mm, 0.45 mm and 0.6 mm. The variations of the constituents in the buffer layer are listed in Table 1. All experiments were carried out under dry contacts on a Falex multi-specimen test machine. The wear characteristics including wear volume, friction coefficient, temperature rise at the substrate and wear mechanisms were obtained at a fixed rotating speed. The wear mechanisms were assessed by investigating surface topography, surface roughness after wear, the production of oxides, and the variation in friction coefficient.
2. Friction tester and operating conditions The friction and wear tests were conducted on a Falex multi-specimen testing machine using a rotorvane-disk adaptor. The rotor-vanes rotated at a constant
Set 2 0.05
Set 3 0.15
0.25 0.40 0.55
0.15 0.30 0.45
Set 4 0.05
Set 5 0.10
Set 6 0.10
0.05
0.10
0.10
0.20 0.35 0.50
0.10 0.25 0.40
0.10 0.25 0.40
velocity of 1 m s-’ while the disk was stationary. A schematic diagram of two specimens in contact is shown in Fig. 1. The cross-section of each vane is rectangular and has a contact area of 25.5 mm’; the circular disk has a diameter of 25.15 mm. Possible operating conditions included two levels of applied load (33.4 N and 89 N) and three levels of preheating temperatures (22 “C, 150 “C, and 250 “C). The mean contact pressures on the friction surface corresponding to the loads of 33.4 N and 89 N were 4.37 x 105 N rnp2 and 1.16~ lo6 N me2 respectively. To avoid the probable occurrence of brittle fracture, the experimental conditions were arranged according to two rules: (1) high load and low preheating temperature, or (2) low load and high preheating temperature. The disk specimens were heated by an electrical heater which was assembled around the disk supporter. The substrate temperature was measured by a thermocouple mounted 0.2 mm beneath the substrate surface. The experiments were conducted under dry contacts, and constant sliding velocity. Each datum shown in the curves was obtained by averaging three experimental readings.
203
J. F. Lin, T. R. Li I Plasma-sprayed ceramic coatings
upper
ower
shaft
shaft
(a)
/r//
A
/I
/
--., t \
,’
0 -t
\ \
’\ ‘1 ’\
q
\\
‘.
/
x._‘_-
‘I._
\
\ \ --+
I’ I/ <_g
\\
\\
\
,,----I--‘-,
4
I’ I’I /I
,A/’
//
_-d
(b) Fig. 1. Schematic diagram of (a) the adaptor and (b) the upper and lower specimens: 1 Vickers holder, 2 retaining ring, 3 specimen ring-vane, 4 spring plunger, 5 ball, 6 vane specimen.
3. Vane and disk specimens The substrates for the vane and disk specimens were AISI M2 and AISI 52100 steels respectively. The constituents of these two steels are shown in Table 2. Ceramic powders were deposited by plasma-spraying onto the friction surface of the specimen. The spraying parameters required to manufacture the ceramic coat-
ings were obtained from the plasma spraying handbook for the Plasma-Technik System (Switzerland). The metal substrate of the vane was first coated with Ni-Cr alloy as the bottom layer and then pure WC as the top material. The disk was coated with Ni-Cr alloy as the bottom layer; in sets 4-6 there was then one layer of mixed ceramic material as buffer layer; then 100% Cr,O, was deposited as the top coating. All vane
204 TABLE
J. F. Lin, 2. The
constituent
Specimen substrate
elements
of two steel
T. R. Li / Plasma-sprayed substrates
of vane
and disk
C
CT
Si
Mn
P
S
W
MO
V
= 0.80 - 0.90 =0.95-1.10 = 0.20 - 0.25
= 3.80 - 4.50 =1.30-1.60 = 17.0- 22.0
= 0.15 - 0.35 = 0.15 - 0.35 = 0.5 - 2.0
= 0.25 - 0.45 < 0.5 = 0.5 - 2.5
GO.025 Q0.025
GO.01 $0.025
= 6.00 * 7.00
= 4.80 - 5.80
= 1 .&Jo - 2.30
specimens had a coating thickness of 0.4 mm while disks had three coating thicknesses varying from 0.3 mm to 0.6 mm. The experiments were conducted using three vanes and a disk connecting in pairs in accordance with the arrangements shown in Table 1. Table 3 presents the properties of WC and Cr,O,. It can be seen that the hardnesses of the two coating materials are fairly close. All specimens before friction testing were well ground to a roughness R, of 0.42f0.04 pm. The specimens were cleaned for 20 min ultrasonically with pure alcohol.
4. Wear volume and friction coefficient The amount of wear of a disk (lower specimen) after testing was obtained in terms of the wear profiles of the frictional track. In this analysis, the wear volume is the product of the length of the annular wear trace and the mean cross-sectional area of the wear profile. The mean area of the wear profile was the average of six wear areas measured at a uniform angle span of 60” along the friction track. The annular wear track was calculated to have the plane area 221.5 mm*. It should be mentioned that the wear volumes in the present experiments were obtained by disassembling the tribo-contact after each test. Each datum was the average value of three readings taken for the same operating conditions. The friction coefficient during frictional contact was calculated from the normal load P and the induced frictional torque T 3. Properties
Coating material
coatings
Constituent elements (%)
ASIS M2 steel AISI 52100 steel Ni-Cr alloy powder
TABLE
ceramic
of hvo ceramic
EDS analysis
where r is the mean contact radius of the rotating vane; it was found to have the value 16.03 mm.
5. Results and discussion 5.1. Wear behavior
The wear resistance of coating materials can be improved by selecting properly the coating thickness and the constituents of the buffer layer. The experiments were conducted at three preheating temperatures by the sliding friction of rotor vanes rotating against a disk. The wear behavior was evaluated in each test based on factors including the wear volume, friction coefficient, and the substrate temperature of the lower specimen. Figures 2(a) and 2(b) show the wear volume of the lower specimen and the substrate temperature with sliding time. The tests were conducted using a load of 89 N, a sliding velocity of 1 m s-‘, a starting temperature of 22 “C, and a coating thickness of 0.3 mm. Both the wear volume and the substrate temperature increased with increasing sliding time. Of five coating types, the largest wear volume resulted from a single-layer coating of 100% Cr,O, material; the double-layer coating without buffer layer showed the next greatest wear volume. Furthermore, the coating mode producing the maximum wear volume normally presented the lowest substrate temperature. When the specimens were at ambient temperature (22 “C) before testing, the three-layer
materials
Density (%)
f = T/(Pr)
(g cm-‘)
Tungsten carbide, WC
WC 88.397 Ca 11.603
= 12.5 - 13.2
Chromium oxide, Crz03
CT203 95.262 TiOZ 2.742 SiOz 1.995
~5.2-5.6
Vickers hardness (N mm-*)
=looo-1150 =lloo-1200
Thermal conductivity (Kcal mh-’
7.8 26.6
Crystalloid ‘C-l)
Melting point (“C)
Hexagonal Hexagonal
2850 2100
Thermal expansion coefficient (x10-6) 6.2 (=22-800 10.5 (=22-1100°C)
“C)
J. F. Lin. T. R. Li I ~~a-s~ra~d 0.70 speknm thm bd:g9N sliding v&x&y
: 0.34 nml
: I ntfs rtr-awd allw.:22oc
0.60 n0.50 j go.40
$ o.30 0.30 nnn CCr,O,, (whole)
2 3 k 0.20 z B
0.05 mm M-0) 0.25 min tCr,O,,
tEmk.) (Top)
0.05 mm W-0) (Lt0a.l 6 : 0.05 mm (75% N&C?+ 25% C1~0,l fBuff.)
0.10
A : 0.05 mm 60% Ni-Cr + 50% Cr,O,) (Buff.) 0 : 0.05 mm (25% Ni-Cr + 75% Cr&I
0.20 mm (CrzOo,) (Tog)
0.00 (a)
sliding wlwity pnmrd
temp.
205
ceramic coatings
Figures 3(a) and 3(b) show the wear volume and substrate temperature for 0.6 mm thick coatings. To examine the influence of coating thickness on wear resistance, the operation conditions shown in Fig. 2 remained unchanged, except that the coating thickness was increased. Basically, the curves for the coating thickness of 0.6 mm are similar to those for the thickness of 0.3 mm; the order of the magnitudes of wear volume and substrate temperature remain unchanged. However, the wear volumes for the thicker coatings were increased substantially. Thus, the wear rate for the thicker coating is higher, and the corresponding temperatures measured at the substrate were relatively lower, compared with the results shown in Fig. 2(a). Interpretation of the lower substrate temperature is difficult because the thermal resistance of heat conduction between the top ceramic material and the steel substrate is closely related to factors including the
: 1 m/S :22OC
1.20
specimen thickness : 0.60 mm kad:89N sliding vciwity
: 1m/s pnhcated temp, :22*c
:
* 0.30 nun fCrzOo,) (whole)
0.05 mm WI-Cr) &km) 0 : 0.05 mm (7.5% Ni-Cr + 25% CrzO,) (Buff,) A : 0.05 mm 60% NiX:r + 50% Cr20,) (Buff,)
l
: 0.60 mm (Cr,O,)(whale
0 : 0.0s mm (25% Ni-Cr + 75% CrzO,) @uff.) 0.20 mm b&O,) 100 @I
O
200
Sliding
(Tap) 300
Time
400
(Min.)
500 A : 0.05 mm (50% Ni-Cr + Xx% r&O,) (Buff,) 0 : 0.05 mm (25% Ni-Cr + ?SW f&0$
Fig. 2. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.3 mm, load 89 N, sliding velocity 1 m s-l, temperature of specimen before testing 22 “C.
(a)
O-O
spccima thichlas
coating with a buffer layer of 75% Ni-Cr + 25% Cr,O, finally produced the minimum wear volume and the maximum substrate temperature at large sliding distances. The reason why the single-layer coatings produced the largest wear volume may be partly attributed to the relatively weak bonding force between the ceramic coating and the bulk steel. Obviously, the presence of the buffer layer alleviated the significant difference in material properties between the ceramic material and the metal substrate. The temperature of the disk at the measured point of the substrate was related to the heat generated due to frictional contacts and the rate of heat flow transferred from the frictional surface to the bulk material (metal substrate). The heat flow rate is closely related to the variations in thermal resistance of all coating layers at various working temperatures.
(Buff.)
0.00
1 -
lmd:89N sliding vclacity
-
plchuncdemp.:22~C
: 0.60 mm
: I m/s
150G 0
A : 0.03 mn 60% Ni-cr + Xv% CrzO$ (Buff.1 0 : 0.05 tntn (25% Ni-Cr + 75% Crz03) @UN.)
100 03)
’
200
Sliding
300
Time
(Min.)
400
Fig. 3. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.6 mm, load 89 N, sliding velocity 1 m s-*, temperature of specimen before testing 22 “C.
J. F. Lin,
206
T R. Li I Plasma-sprayed
complex geometry of frictional contacts, the variations in thermal conductivity of the coating layers, and the thickness of each coating layer. Nevertheless, increasing the coating thickness, simply from the viewpoint of heat conduction, can increase the thermal resistance and thus decrease the heat flow rate in the axial direction. The increase in coating thickness tends to lower the substrate temperature. The substantial increase in wear volume for the coating 0.6 mm thick is probably also related to the relatively weak bonding force on the frictional surface due to the too large thickness of the coating layer. Figure 4(a) shows the wear volume of a 0.3 mm thick Cr,O, coating operating at a load of 33.4 N and the higher preheating temperature of 150 “C. Comparing the present results with those shown in Fig. 2(a), the curves of wear volume for the higher preheating temperature (150 “C) were flatter; the wear volumes cor0.60
o.55
_ T
-
2 0, 0.25
specimen lhickncss: 0.30 mm lOad : 33.4 N slidmg velocity
: I m/s preheated temp. : ,50°c
/
k= 0.20
CLlOmm N-W (Bo!t.)(whole) 0 0.10mm (75% Ni-Cr + 25% CrzO,) (Buff.) A : 0.10 mln (50% NbCr
+ 50% CI*O,)
(BurrI
0:
+ 75% Cr,O,)
(Buff
E 150
1
0IOmm Kr*O,) (Top)
i
E
0.10 tnm (25% Ni-Cr
0.10
mm (Ni-Cr) (Bolt.) 0 : 0.10 mm (7% NI-Cr + 25% Cr,O,) (Buff,)
E
A : 0.10 mm (50% Ni-Cr + M% Cr,O,) (Buff) 0: 0.10mm (25% Ni-Cr + 75% 13~0,) (Buff.) 0.10 mm K&O,) (Top)
,,OtiI
@I
’
100
Sliding
200
Time
300
400
(Min.)
Fig. 4. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.3 mm, load 33.4 N, slidingvelocity 1 m s-l, temperature of specimen before testing 150 “C.
ceramic
coatings
responding to large sliding distances were relatively smaller. Of three different constituent modes, the buffer layer with 50% Ni-Cr+50% Cr,O, produced the least wear volume. The wear rates for all three modes are quite small over the whole range of sliding distances. Figure 4(b) shows the variations in substrate temperature with sliding distance. The substrate temperatures showed a mild increase, and reached a stable state in a rather short time. The specimens producing the smaller wear volumes usually showed the higher substrate temperatures. A comparison between the temperatures in Figs. 4(b) and 2(b) illustrates that the steady substrate temperatures resulting from the higher preheating temperature (150 “C) were lower. The reason why the frictional surface preheated to the higher temperature (150 “C) generated the smaller wear volume was thought to be due to the temperature rise on the contact surface causing thermal softening, and increasing the capability of wear resistance. This behavior resulted in a lower wear rate compared with wear by brittle fracture at low temperatures. The experimental results shown in Figs. 5(a) and Fig. 5(b) show the effect of the increased preheating temperature on wear volume. The curves in Fig. 5(a) show the wear volumes obtained at the preheating temperature of 150 “C. The wear volume corresponding to the buffer contents of 50% Ni-Cr+50% Cr,O, was the smallest. The sharper variations in wear volume at the beginning of sliding motion were mild at larger sliding distances; the wear rates are almost the same for the three constituent modes. Figure 5(b) shows the variations in wear volume of Cr,O, coatings where the specimens were heated to 250 “C before testing. The experimental data reveal that the buffer layer of 25% Ni-Cr +75% Cr,O, produced the least wear volume. After a steep increase in wear volume within short sliding distances, the wear volume became steady, and the wear rate was insignificant. Differences between the experimental data corresponding to the two preheating temperatures show that the optimum constitutions of Ni-Cr alloy and Cr,O, material in the buffer layer which produce the minimum wear volume, are closely related to the preheating temperature of the specimen. Elevating the preheating temperature resulted in a reduction in wear volume; and the specimen with the higher content of Cr,O, in the buffer layer produced the smaller wear volume. The two figures show that the effect on wear volume of changing the Cr,O, content is smaller than the range of scatter of tribological data. The curves in Fig. 6 illustrate the variations in wear rate under various testing conditions and for three coating thicknesses. The difference in steady wear rate for each individual curve is nearly negligible, except for the wear rates at the beginning of sliding motion. This means that the
.T. F. Lin, T. R. Li I P~s~a-sp~~~
ceramic coatings
207
steady-state wear rate is hardly influenced by the constituents and the thickness of the buffer layer, and the preheating temperature of the specimen.
rlidingveksiry : t mfs prdl~temp.:ISO’C 0.10 mm (Ni-Cr> (Bau.) 4 (75% NiiG + 25% Cr,O,f (Buff.)
: &,(Imm
i) : 0.10 mm (50% Ni-Cr + SD% Cr.& 0
: 0.10 mm (25% Ni-G
(Buff.)
+ 75% Ciz03) awf.~
Fig. 5. Variation in wear volume with sliding distance: coating thickness 0.3 mm, load 33.4 N, sliding velocity 1 m s-l, temperature of specimen before testing (a) 150 “C and (b) 250 “C.
35.00
_i
y 0 4
30.00
-
*
25.00
-
20.00
-
F =E
sliding
m/s
N temperature:150*C thickness: :0.30 mm, Inark * :0.45 mm in set 5 :o.co mmI
load:33.4 preheated
N temperature:250’C thickness:
specimen
? 15.00
2 0
velocity:1
load:33.4 preheated specimen 0-o A--3 cl-_0
load:89 N Preheated temperature:22?! specimen thickness: *-+ :0.30 mm mark * x-x :0.45 m* in set 4 *-* :0.60 mmI
10.00
2 z
~
-5.00
0.0
5.0
Sliding
10.0
15.0
Distance
I”“““‘I’i’W 20.0 25.0
* 1000
(m)
Fig. 6. Variation in wear rate with sliding distance.
rr 5‘O.G
5.2. Friction The variation of friction coefficient with sliding time for three coating thicknesses is shown in Figs. 7(a)-7(c). The thickness of each layer of the three-layer coating was designated according to the specifications of set 4. The wear experiments were conducted starting with three different preheating temperatures. The coefficient of friction of the specimen with a preheating temperature of 22 “C shows a substantial increase with increasing sliding distance; however, it is expected to obtain a steady value after a longer sliding time. The curves shown in Figs. 7(b) and 7(c) illustrate that the time required to reach a stable friction coefficient is shorter when the specimens are preheated to higher temperatures. The effect of coating thickness on friction coefficient was also dependent on the starting temperature of the specimen. When the tests began at a low preheating temperature (22 “C), no immediate connection between the friction coefficient and coating thickness could be identified. The time required to reach the steady frictional coefficient is relatively longer, but when the specimens were heated to higher temperatures (150 “C and 250 “C), a larger friction coefficient resulted from thicker coatings and the time required to reach steady state was shorter. As the curves in Figs. 7(b) and 7(c) show, there is a drop in friction during the entire sliding procedure, which is probably related to the formation of a softer oxide (WO,) at higher temperatures. However, the friction drop was absent when the low preheating temperature (22 “C) was used. In the region of oxidetilm formation, the friction is decreased. The significant increase in friction is thought to be related to the runin failure where ceramic contacts through the film. Ceramic contacts began at the tips of asperities, friction began to increase, and failure was very sudden. Friction at this stage was also important in determining the wear behavior. High interface temperatures due to frictional heat may be responsible for enhancing softening; softening in ceramics increases the coefficient of friction. The curves in Fig. 8 show the variations of friction coefficient with surface roughness R,. With increasing sliding distance, the worn surface shows continuous reduction in surface roughness. The greatest reduction resulted from the test with the highest preheating temperature (250 “C). The data corresponding to the two higher preheating temperatures (150 “C and 250 “C) reveal that the friction coefficients at large sliding distances were nearly independent of the variations in surface roughness. However, the friction coefficient for
J. F. L&z, 7: R, Li f Plasma-strayed
208
i
c -0.80
A:
coatings
0.4s mm
O:O.Mfmm : 1m/s
I 1 :
g
ceramic
laad:89N sliding velocity preheated amp.:22"C
3 .: 2 0.60 1
0.00
?,,,~,~,~,,.~,,,,,~,,~~~.,,~~‘Tm,Iml,iiII,,,I,/ 0 100 200 300 400 500 Time
Sliding
(a)
(Min.)
0.00
0.20
I/(/I/1II,,I,J I<,,, l,tI/j,I
Coefficient
of
Friction
Fig. 8. Variation of friction coefficient mean surface roughness R,.
coefficients were generated to higher temperatures.
IV O.OO 10
specimen lhi&“css 0 : 0.30 mm A
: 0.45 mm 0:0.60mm load :33.4 N sliding ~elcaty: I *s
Sliding
(b)
Time
(min.)
e=f==d 0.00
tmnp. : 230 “C
,)I))II,‘I,II1I1(//I,/O)I)/II 100
(c)
:
O
200
Sliding
300
Time
400
(Min.)
Fig. 7. Coefficient of friction as a function three different coating thicknesses.
500
of sliding time for
the preheating temperature of 22 “C showed a continuous increase within the presently designated sliding distances. At the same sliding distance, higher friction
(f)
with sliding distance and
by the surfaces preheated
In tangential motion, the sliding or rubbing of one surface over the other caused surface fracture on the contact surfaces. Table 4 shows EDS analyses of worn surfaces after dry rubbing tests. The elemental analyses reveal that traces of the two coating materials were detected on the opposite surfaces. Using EDS maps, traces of chromium were found on the coating surface of the WC specimen. Similarly, traces of tungsten were found on the surface of the coating Cr,O,. Wear by adhesion had indeed occurred between the two rubbing surfaces. The severity of adhesion was enhanced when the specimen was preheated to higher temperatures. X-ray diffractometry was used to analyze the wear debris. This is an immediate method for accessing the crystal structure of oxides. Figures 9(a)-9(c) show microregion X-ray diffraction maps for WC coatings. There are many diffraction peaks in each figure which were given by more than one substance. With the aid of computer searching in accordance with the location and intensity of the diffraction peaks, the chemical structure of the oxides on the WC coating was derived. According to the ~nsti~ent element analysis, oxidation of WC occurred on the surface forming WO, (melting point of WO, 1180 “C, melting point of WC 2850 “C). Since WO, is rather soft, the sliding couple would show abrasive wear. The height of the XRD peaks in the area of the WC surface shows that the amount of WO, was increased when the preheating temperature of the specimen was raised. Figures lO(a )-10(c) show images of Cr,O, debris on the frictional surface of the WC specimen. The bright
J. F. Lin, T. R. Li f ~~srna-~p~~d TABLE
eernmk coatirzgs
209
4. EDS analyses of the warn surfaces of WC and Cr2&
Element
Load
Sliding distance
(N)
(m)
22 150 250 22 150 250 22 150 250 22 150 250 22 150 250
89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4
28800
22 150 250 22 150 250 22 150 250 22
89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4
28800
Preheating temperature (“C)
si Ti
W
0
Sliding velocity (m s-l)
28800
28800
28800
28800
Proportion (wt.%)
31.506 30.194 31.073 0.890 1A32 0.614 1.146 1.653 1.015 3.241 4.269 4.887 63.217 62.202 62.411
WC sJP&?m
Cr
Si
150 250 22 150 250
particles distributed over the surface were identified as chromium-containing particles, which implies that these particles are CrzO,. The amount of these bright particles increased notably when the specimen was preheated to higher temperatures. The microregion Xray diffraction traces of the Cr,O, coating were also measured, and the atomic concentration of chromium was high. There are many diffraction peaks which perhaps represent two substances including Cr,O, and Cr,C,; however, the amount of C&C, was too small to be detected. Figures ll(a )-11(c) show the surface structure of the Cr,O, disk after dry friction testing, In Fig. 11(a), brittle fatigue spallings are the primary wear mechanism; hardly any evidence of plastic deformation was observed at such a low temperature. When the preheating temperature of the specimen was elevated to 150 “C, numerous pits were distrluted over the surface in place ofsome fatigue spaflings. Part of the surface was smeared
28800
28800
28800
28800
2.513 2.706 4.487 0.000 O.OofI 0.000 0.222 0.096 0.427 85.143 83.960 83.531 12.122 13.239 11.555
flat owing to abrasive wear. When the specimen was heated further to 250 “C (Fig. 11(c)), a different surface from that of Figs. 11(a) and 11(b) is seen. A few smallscale but evenly distributed fractures are seen on the surface in place of the severe brittle fractures found at low temperature. This kind of fault implies that the wear resistance was improved and the wear depth was decreased. The do~nant wear mechanism at this temperature is still brittle fracture.
6. Summary The experiments dealt with the evaluation of wear resistance in multilayered ceramic coatings, A buffer layer with various proportions of Cr,O, was planted between the top and bottom layers. The buffer layer indeed showed the capability of diminishing the wear volume, The optimum proportion of Cr203 content in
210
200Q
1600
800
(a)
32
2*
r4
56
68
1.5K
! 20
/
40
68
wo,
WC EBB
i
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AJJJ% WC
WC 680
d-
(cl
a
~ ’ 20
wo3
AL
-1
32
44
56
68
88
9. X-ray maps of the worn surfaces of the WC coating and debris after friction testing under various operation conditions: coating thickness 0.3 mm, sliding distance 28 800 m; {a) load 89 N, preheating temperature 22 “C, (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C. Fig.
J. F. Lin, T. R. Li I Plasma-sprayed ceramic coatings
(4
(4
(b)
(cl
(cl
Fig. 10. Images of the Cr,Os debris on the WC surface due to adhesive wear: coating thickness 0.3 mm, sliding distance 28 800 m; (a) load 89 N, preheating temperature 22 “C; (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C.
Fig. 11. Micrographs of the worn surface of the WC coating after wearing tests: coating thickness 0.3 mm, sliding distance 28 800 m; (a) load 89 N, preheating temperature 22 “C; (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C.
212
J. F. Lin, T. R. Li / Plasma-sprayed ceramic coatings
the buffer layer and the preheating temperature were examined, and reasonable interpretations were proposed to illustrate the wear behavior. The interaction between wear volume and the measured temperature at the substrate presented uniform behavior even under various operating conditions and for different coating thicknesses. The primary wear mechanisms occurring on the contact surfaces were identified using EDS and X-ray diffraction analyses.
7. Conclusions
The wear volume of a multi-layered ceramic coating as a function of sliding distance is controlled by factors including the preheating temperature of the specimen before testing, the constituents in the buffer layer, the coating materials of the vane and disk, the thickness of each coating layer, and the applied load. Of these five factors, the preheating temperature of the specimen, the proportion of Cr,O, in the buffer layer, and the coating thickness were the main subjects of the present study. Because ceramics have high strength and brittleness, the testing conditions were designed to prevent brittle fracture during the experiments. The major wear mechanisms occurring at the presently designated specimen temperatures include (1) brittle fracture (spalling and pitting), (2) abrasion (scratching and polishing), (3) adhesion, and (4) ‘oxidation. From the above studies some conclusions can be drawn. (1) If the test starts from higher preheating temperatures, the time required to reach the steady friction coefficient is relatively shorter, the stable friction coefficient is relatively higher, and the wear volume is smaller. However, the effect of Cr,O, content on the change in wear volume is usually smaller than the range of scatter of tribological data. (2) The participation of the buffer layer can indeed reduce the wear volume. Similar to raising the preheating temperature of the specimen, increasing the proportion by weight of Cr,O, in the buffer layer reduces the wear volume. (3) The wear volume is relatively smaller when the specimen is preheated to higher temperatures. The specimens producing smaller wear volume usually have higher substrate temperatures. (4) The steady-state wear rate is not much influenced by the constituents, the preheating temperature, and the thickness of the buffer layer.
References 1 T. Sugita and H. Yasunaga, Friction and wear of MgO single crystals on sintered Al24 in vacuum, Wear, 24 (1976) 223-234. 2 J. A. Kirk and W. D. Syniuta, Scanning electron microscopy and microprobe investigation of high speed sliding wear of aluminum oxide, Wear, 27 (1974) 367-381. 3 R. Komanduri and M. C. Shaw, Wear of silicon carbide in high speed diding, Wear, 36 (1976) 363-371. 4 U. Dworak, H. Olapinski and W. Stannek, Wear characteristics of ceramic materials, Ber. Dtsch. Keram. Ges., 54 (1977) 416-419. 5 M. I. Mendelson, Theoretical evaluation of wear in piasmasprayed TiO against gray cast, Wear, 50 (1978) 71-84. 6 P. K. Mehrotra, in K. C. Ludema (ed.), Mechanirms of wear in ceramic material, Wear of Materials, ASME, New York, 1983, pp. 194-201. 7 C. S. Yust and F. J. Carignan, Observations on the sliding wear of ceramics, ASLE Trans., 28 (1985) 245-252. 8 K. Kato, Tribology and ceramics, Wear, 136 (1990) 117-133. 9 D. H. Buckley and M. Kazuhisa, Friction and wear of ceramics, Wear, 100 (1984) 333-353. 10 J. Breznak, E. Berval and N. H. Macmillian, Sliding friction and wear of structural ceramics, part 1: room temperature behavior, J. Mater. Sci., 20 (198.5) 465742680. 11 D. C. Cranmer, Ceramic tribology - needs and opportunities, STLE, TriboZ. Trans., 31 (1988) 164-173. 12 A. V. Levy and N. Jee, Unlubricated sliding wear of ceramic materials, Wear, 121 (1988) 363-380. 13 J. F. Braza, Mechanical failure mechanisms in ceramic sliding and rolling contacts, STLE, Ttibol. Trans., X12 (1990) 1-9. 14 C. Ding, Plasma-sprayed wear-resistant ceramic and cermet coating materials, Thin Solid Films, 118 (1984) 485-493. 15 Y. Wang, The analysis of the chemical structure and properties of ceramic surface films in friction using scanning electron microscopy, auger electron spectroscopy and microregion Xray diffraction, Wear, 128 (1988) 277-290. 16 A. Skopp, M. Woydt and K. H. Habig, Unlubricated sliding friction and wear of various S&N, pairs between 22 “C and 1000 “C, Ttibol. ht., 23 (1990) 189-199. ceramic ma17 B. Longson, Lubrication of high temperature terials, Ttibol. Znt., 16 (1983) 221-225. 18 M. Woydt and K. H. Habig, High temperature tribology of ceramics, Tribal. ht., 22 (1989) 75-87. 19 Y. Wang, The analysis of the friction and wear mechanisms of plasma-sprayed ceramic coating at 450 “C, Wear, 128 (1988) 265-276. 20 M. D. Thouless, Spalling of flat plates by thermal shock, J. Am. Cerum. Sot., 68 (1985) Clll-112. 21 S. Rangaswamy, H. Herman and S. Safai, Thermal expansion study of plasma-sprayed oxide coating, Z&in Solid Films, 73 (1980) 43-52. 22 P. K. Mehrotra, in K. C. Ludema (ed.), Mechanisms of Wear in Ceramic Materials, Wear of Materials, ASME, New York, 1983, pp. 194-201. 23 0.0. Ajayi and K. C. Ludema, Surface damage of st~ctural ceramic: implications for wear modeling, Wear, 124 (1988) 237-257. 24 G. R. Miller and L. M. Keer, Interaction between a rigid indenter and a near-surface void or inclusion, J. Appt. Me&.. 50 (1983j 615620. of plasma25 Y. Wang, The friction and wear performance sprayed ceramic coatings at high temperature, Wear, 129 (1989) 223-234.
Wear, 160 (1993) 201-212
Analysis of the friction and wear mechanisms plasma-sprayed ceramic coatings
of multilayered
Jen Fin Lin and Tzuen Ren Li Department
ofMechanicai Engine&q,
National Cheng Kung University, Tainun (Taiwan)
(Received December 17, 1991; revised and accepted July 16, 1992)
Abstract Wear tests were conducted on a rotor-vane-disk adaptor where three rotating vanes were pressed against a disk. Vanes were coated by WC and used as the upper specimen while the disk was coated by Cr,O, and used as the lower specimen. A buffer layer of various thicknesses and contents was placed between the top coating and the bulk steel of the disk to alleviate the effects of the large difference in thermal properties of the two materials. The experimental results reveal that correct placement of a buffer layer can indeed improve the wear resistance. Factors such as the temperature to which the specimen was heated before testing, the proportion by weight of each individual constituent in the buffer layer, and the thickness of each coating layer, were also important for the volume of wear of the lower specimen. As the specimens were heated to higher temperatures, the wear volume decreased with increasing proportions of Cr,O, in the buffer layer. Elevating the preheating temperature of the specimens can diminish the wear volume but increases the friction coefficient. The steady-state wear rate is not much influenced by the constituents of the buffer layer and the coating thickness. Brittle fracture, abrasion, adhesion and oxidation were found to be the primary wear mechanisms in the tests.
1. Introduction Several engineering applications which require high wear resistance often employ ceramics as coating materials. Wear resistance is not only a fundamental material property but also depends on the entire system which includes the ambient atmosphere, material thermal properties, specimen temperature and operation conditions. In the literature [l-25], ceramic-ceramic couples or ceramic-metal couples have been employed in most experiments aimed at understanding wear mechanisms in ceramic materials. Since ceramics are nominally brittle, wear is expected to occur by brittle fracture, abrasion, adhesion and oxidation. However, ample evidence suggests that at high temperature plastic deformation becomes an important wear mechanism. This plasticity is attributed to the inhibition of fracture by the hydrostatic pressure associated with the contact conditions. The wear resistance of materials can be effectively improved by (1) planting a buffer layer between the top coating and the bulk material, (2) altering the proportion by weight of constituents in the buffer layer, and (3) changing the starting temperature of specimens before testing. The pa~icipation of the buffer layer can alleviate stresses arising owing to the notable dif-
0043-1648/93,‘$6.00
ference in thermal properties between the top coating and bulk materials, and can thus improve wear resistance. For the purpose of examining the effect of changing the coating thickness and the proportion of each individual constituent of the buffer layer on wear volume and wear mechanisms, experiments were conducted by designating specimen preheating temperatures before testing. However, surfaces contacting simultaneously at both high load and high specimen temperature were excluded from the present investigation because these conditions induce rapid cracking of the surface. The coating layers were deposited by plasma spraying onto AISI steels. The coatings of the rotor vane (upper specimen) consisted of two materials: NI-Cr alloy as the bottom layer and WC as the top layer. The experimental conditions were classified in three groups. To investigate the effect of coating thickness on wear resistance, the disk (bottom specimen) coatings were placed in six sets marked with Arabic numerals, as shown in Table 1. The specimens in set 1 had singlelayer coatings; CrzO, was the sole coating material. Sets 2 and 3 had double-layer coatings, with Ni-Cr alloy as the bottom layer and Cr,O, ceramic material as the top layer. The specimens of sets 4-6 had threelayer coatings, with Ni-Cr as the bottom layer, mixtures of different proportions of Ni-Cr and Cr,O, as the
0 1993 - Elsevier Sequoia. All rights reserved
202 TABLE
J. E: Lin, T. H. Li I Plasma-sprayed ceramic coatings 1. Arrangement
of rotor vanes and disks in pairs and operation
Experimental conditions
Group Load (N) Temperature (“C) Roughness R, (pm) Sliding velocity (m SK’) Lubrication
Coating thickness
Top specimen, WC
I 89 22 0.4 1 Dry contacts
conditions
11 33.4 150 0.4 1 Dry contacts
III 33.4 250 0.4 1 Dry contacts
Bottom layer (100% Ni-Cr Top layer (100% WC)
alloy)
0.05 0.35
(mm) Bottom Cr203
specimen,
Set 1 0.30 0.45 0.60
One-layer coating (100% Cr,O,) Twq;layer coatmg
Bottom
layer (100% Ni-Cr
alloy)
Top layer (100% Cr20,)
Three-layer coating
Bottom layer (100% Ni-Cr
alloy)
Buffer layer (75% Ni-Cr+2.5% CrzOs) (50% Ni-Cr + 50% CrzOJ (25% Ni-Cr + 75% Cr203) Top layer (100% Cr,O,)
buffer layer, and 100% Cr,O, as the top layer. It should be noted that the specimens of sets 5 and 6 actually have the same constituents, but they were intentionally arranged to examine the effect of preheating temperature on the wear behavior. These specimens were provided with three coating thicknesses: 0.3 mm, 0.45 mm and 0.6 mm. The variations of the constituents in the buffer layer are listed in Table 1. All experiments were carried out under dry contacts on a Falex multi-specimen test machine. The wear characteristics including wear volume, friction coefficient, temperature rise at the substrate and wear mechanisms were obtained at a fixed rotating speed. The wear mechanisms were assessed by investigating surface topography, surface roughness after wear, the production of oxides, and the variation in friction coefficient.
2. Friction tester and operating conditions The friction and wear tests were conducted on a Falex multi-specimen testing machine using a rotorvane-disk adaptor. The rotor-vanes rotated at a constant
Set 2 0.05
Set 3 0.15
0.25 0.40 0.55
0.15 0.30 0.45
Set 4 0.05
Set 5 0.10
Set 6 0.10
0.05
0.10
0.10
0.20 0.35 0.50
0.10 0.25 0.40
0.10 0.25 0.40
velocity of 1 m s-’ while the disk was stationary. A schematic diagram of two specimens in contact is shown in Fig. 1. The cross-section of each vane is rectangular and has a contact area of 25.5 mm’; the circular disk has a diameter of 25.15 mm. Possible operating conditions included two levels of applied load (33.4 N and 89 N) and three levels of preheating temperatures (22 “C, 150 “C, and 250 “C). The mean contact pressures on the friction surface corresponding to the loads of 33.4 N and 89 N were 4.37 x 105 N rnp2 and 1.16~ lo6 N me2 respectively. To avoid the probable occurrence of brittle fracture, the experimental conditions were arranged according to two rules: (1) high load and low preheating temperature, or (2) low load and high preheating temperature. The disk specimens were heated by an electrical heater which was assembled around the disk supporter. The substrate temperature was measured by a thermocouple mounted 0.2 mm beneath the substrate surface. The experiments were conducted under dry contacts, and constant sliding velocity. Each datum shown in the curves was obtained by averaging three experimental readings.
203
J. F. Lin, T. R. Li I Plasma-sprayed ceramic coatings
upper
ower
shaft
shaft
(a)
/r//
A
/I
/
--., t \
,’
0 -t
\ \
’\ ‘1 ’\
q
\\
‘.
/
x._‘_-
‘I._
\
\ \ --+
I’ I/ <_g
\\
\\
\
,,----I--‘-,
4
I’ I’I /I
,A/’
//
_-d
(b) Fig. 1. Schematic diagram of (a) the adaptor and (b) the upper and lower specimens: 1 Vickers holder, 2 retaining ring, 3 specimen ring-vane, 4 spring plunger, 5 ball, 6 vane specimen.
3. Vane and disk specimens The substrates for the vane and disk specimens were AISI M2 and AISI 52100 steels respectively. The constituents of these two steels are shown in Table 2. Ceramic powders were deposited by plasma-spraying onto the friction surface of the specimen. The spraying parameters required to manufacture the ceramic coat-
ings were obtained from the plasma spraying handbook for the Plasma-Technik System (Switzerland). The metal substrate of the vane was first coated with Ni-Cr alloy as the bottom layer and then pure WC as the top material. The disk was coated with Ni-Cr alloy as the bottom layer; in sets 4-6 there was then one layer of mixed ceramic material as buffer layer; then 100% Cr,O, was deposited as the top coating. All vane
204 TABLE
J. F. Lin, 2. The
constituent
Specimen substrate
elements
of two steel
T. R. Li / Plasma-sprayed substrates
of vane
and disk
C
CT
Si
Mn
P
S
W
MO
V
= 0.80 - 0.90 =0.95-1.10 = 0.20 - 0.25
= 3.80 - 4.50 =1.30-1.60 = 17.0- 22.0
= 0.15 - 0.35 = 0.15 - 0.35 = 0.5 - 2.0
= 0.25 - 0.45 < 0.5 = 0.5 - 2.5
GO.025 Q0.025
GO.01 $0.025
= 6.00 * 7.00
= 4.80 - 5.80
= 1 .&Jo - 2.30
specimens had a coating thickness of 0.4 mm while disks had three coating thicknesses varying from 0.3 mm to 0.6 mm. The experiments were conducted using three vanes and a disk connecting in pairs in accordance with the arrangements shown in Table 1. Table 3 presents the properties of WC and Cr,O,. It can be seen that the hardnesses of the two coating materials are fairly close. All specimens before friction testing were well ground to a roughness R, of 0.42f0.04 pm. The specimens were cleaned for 20 min ultrasonically with pure alcohol.
4. Wear volume and friction coefficient The amount of wear of a disk (lower specimen) after testing was obtained in terms of the wear profiles of the frictional track. In this analysis, the wear volume is the product of the length of the annular wear trace and the mean cross-sectional area of the wear profile. The mean area of the wear profile was the average of six wear areas measured at a uniform angle span of 60” along the friction track. The annular wear track was calculated to have the plane area 221.5 mm*. It should be mentioned that the wear volumes in the present experiments were obtained by disassembling the tribo-contact after each test. Each datum was the average value of three readings taken for the same operating conditions. The friction coefficient during frictional contact was calculated from the normal load P and the induced frictional torque T 3. Properties
Coating material
coatings
Constituent elements (%)
ASIS M2 steel AISI 52100 steel Ni-Cr alloy powder
TABLE
ceramic
of hvo ceramic
EDS analysis
where r is the mean contact radius of the rotating vane; it was found to have the value 16.03 mm.
5. Results and discussion 5.1. Wear behavior
The wear resistance of coating materials can be improved by selecting properly the coating thickness and the constituents of the buffer layer. The experiments were conducted at three preheating temperatures by the sliding friction of rotor vanes rotating against a disk. The wear behavior was evaluated in each test based on factors including the wear volume, friction coefficient, and the substrate temperature of the lower specimen. Figures 2(a) and 2(b) show the wear volume of the lower specimen and the substrate temperature with sliding time. The tests were conducted using a load of 89 N, a sliding velocity of 1 m s-‘, a starting temperature of 22 “C, and a coating thickness of 0.3 mm. Both the wear volume and the substrate temperature increased with increasing sliding time. Of five coating types, the largest wear volume resulted from a single-layer coating of 100% Cr,O, material; the double-layer coating without buffer layer showed the next greatest wear volume. Furthermore, the coating mode producing the maximum wear volume normally presented the lowest substrate temperature. When the specimens were at ambient temperature (22 “C) before testing, the three-layer
materials
Density (%)
f = T/(Pr)
(g cm-‘)
Tungsten carbide, WC
WC 88.397 Ca 11.603
= 12.5 - 13.2
Chromium oxide, Crz03
CT203 95.262 TiOZ 2.742 SiOz 1.995
~5.2-5.6
Vickers hardness (N mm-*)
=looo-1150 =lloo-1200
Thermal conductivity (Kcal mh-’
7.8 26.6
Crystalloid ‘C-l)
Melting point (“C)
Hexagonal Hexagonal
2850 2100
Thermal expansion coefficient (x10-6) 6.2 (=22-800 10.5 (=22-1100°C)
“C)
J. F. Lin. T. R. Li I ~~a-s~ra~d 0.70 speknm thm bd:g9N sliding v&x&y
: 0.34 nml
: I ntfs rtr-awd allw.:22oc
0.60 n0.50 j go.40
$ o.30 0.30 nnn CCr,O,, (whole)
2 3 k 0.20 z B
0.05 mm M-0) 0.25 min tCr,O,,
tEmk.) (Top)
0.05 mm W-0) (Lt0a.l 6 : 0.05 mm (75% N&C?+ 25% C1~0,l fBuff.)
0.10
A : 0.05 mm 60% Ni-Cr + 50% Cr,O,) (Buff.) 0 : 0.05 mm (25% Ni-Cr + 75% Cr&I
0.20 mm (CrzOo,) (Tog)
0.00 (a)
sliding wlwity pnmrd
temp.
205
ceramic coatings
Figures 3(a) and 3(b) show the wear volume and substrate temperature for 0.6 mm thick coatings. To examine the influence of coating thickness on wear resistance, the operation conditions shown in Fig. 2 remained unchanged, except that the coating thickness was increased. Basically, the curves for the coating thickness of 0.6 mm are similar to those for the thickness of 0.3 mm; the order of the magnitudes of wear volume and substrate temperature remain unchanged. However, the wear volumes for the thicker coatings were increased substantially. Thus, the wear rate for the thicker coating is higher, and the corresponding temperatures measured at the substrate were relatively lower, compared with the results shown in Fig. 2(a). Interpretation of the lower substrate temperature is difficult because the thermal resistance of heat conduction between the top ceramic material and the steel substrate is closely related to factors including the
: 1 m/S :22OC
1.20
specimen thickness : 0.60 mm kad:89N sliding vciwity
: 1m/s pnhcated temp, :22*c
:
* 0.30 nun fCrzOo,) (whole)
0.05 mm WI-Cr) &km) 0 : 0.05 mm (7.5% Ni-Cr + 25% CrzO,) (Buff,) A : 0.05 mm 60% NiX:r + 50% Cr20,) (Buff,)
l
: 0.60 mm (Cr,O,)(whale
0 : 0.0s mm (25% Ni-Cr + 75% CrzO,) @uff.) 0.20 mm b&O,) 100 @I
O
200
Sliding
(Tap) 300
Time
400
(Min.)
500 A : 0.05 mm (50% Ni-Cr + Xx% r&O,) (Buff,) 0 : 0.05 mm (25% Ni-Cr + ?SW f&0$
Fig. 2. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.3 mm, load 89 N, sliding velocity 1 m s-l, temperature of specimen before testing 22 “C.
(a)
O-O
spccima thichlas
coating with a buffer layer of 75% Ni-Cr + 25% Cr,O, finally produced the minimum wear volume and the maximum substrate temperature at large sliding distances. The reason why the single-layer coatings produced the largest wear volume may be partly attributed to the relatively weak bonding force between the ceramic coating and the bulk steel. Obviously, the presence of the buffer layer alleviated the significant difference in material properties between the ceramic material and the metal substrate. The temperature of the disk at the measured point of the substrate was related to the heat generated due to frictional contacts and the rate of heat flow transferred from the frictional surface to the bulk material (metal substrate). The heat flow rate is closely related to the variations in thermal resistance of all coating layers at various working temperatures.
(Buff.)
0.00
1 -
lmd:89N sliding vclacity
-
plchuncdemp.:22~C
: 0.60 mm
: I m/s
150G 0
A : 0.03 mn 60% Ni-cr + Xv% CrzO$ (Buff.1 0 : 0.05 tntn (25% Ni-Cr + 75% Crz03) @UN.)
100 03)
’
200
Sliding
300
Time
(Min.)
400
Fig. 3. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.6 mm, load 89 N, sliding velocity 1 m s-*, temperature of specimen before testing 22 “C.
J. F. Lin,
206
T R. Li I Plasma-sprayed
complex geometry of frictional contacts, the variations in thermal conductivity of the coating layers, and the thickness of each coating layer. Nevertheless, increasing the coating thickness, simply from the viewpoint of heat conduction, can increase the thermal resistance and thus decrease the heat flow rate in the axial direction. The increase in coating thickness tends to lower the substrate temperature. The substantial increase in wear volume for the coating 0.6 mm thick is probably also related to the relatively weak bonding force on the frictional surface due to the too large thickness of the coating layer. Figure 4(a) shows the wear volume of a 0.3 mm thick Cr,O, coating operating at a load of 33.4 N and the higher preheating temperature of 150 “C. Comparing the present results with those shown in Fig. 2(a), the curves of wear volume for the higher preheating temperature (150 “C) were flatter; the wear volumes cor0.60
o.55
_ T
-
2 0, 0.25
specimen lhickncss: 0.30 mm lOad : 33.4 N slidmg velocity
: I m/s preheated temp. : ,50°c
/
k= 0.20
CLlOmm N-W (Bo!t.)(whole) 0 0.10mm (75% Ni-Cr + 25% CrzO,) (Buff.) A : 0.10 mln (50% NbCr
+ 50% CI*O,)
(BurrI
0:
+ 75% Cr,O,)
(Buff
E 150
1
0IOmm Kr*O,) (Top)
i
E
0.10 tnm (25% Ni-Cr
0.10
mm (Ni-Cr) (Bolt.) 0 : 0.10 mm (7% NI-Cr + 25% Cr,O,) (Buff,)
E
A : 0.10 mm (50% Ni-Cr + M% Cr,O,) (Buff) 0: 0.10mm (25% Ni-Cr + 75% 13~0,) (Buff.) 0.10 mm K&O,) (Top)
,,OtiI
@I
’
100
Sliding
200
Time
300
400
(Min.)
Fig. 4. (a) The wear volume and (b) the temperature at the substrate as a function of sliding distance: coating thickness 0.3 mm, load 33.4 N, slidingvelocity 1 m s-l, temperature of specimen before testing 150 “C.
ceramic
coatings
responding to large sliding distances were relatively smaller. Of three different constituent modes, the buffer layer with 50% Ni-Cr+50% Cr,O, produced the least wear volume. The wear rates for all three modes are quite small over the whole range of sliding distances. Figure 4(b) shows the variations in substrate temperature with sliding distance. The substrate temperatures showed a mild increase, and reached a stable state in a rather short time. The specimens producing the smaller wear volumes usually showed the higher substrate temperatures. A comparison between the temperatures in Figs. 4(b) and 2(b) illustrates that the steady substrate temperatures resulting from the higher preheating temperature (150 “C) were lower. The reason why the frictional surface preheated to the higher temperature (150 “C) generated the smaller wear volume was thought to be due to the temperature rise on the contact surface causing thermal softening, and increasing the capability of wear resistance. This behavior resulted in a lower wear rate compared with wear by brittle fracture at low temperatures. The experimental results shown in Figs. 5(a) and Fig. 5(b) show the effect of the increased preheating temperature on wear volume. The curves in Fig. 5(a) show the wear volumes obtained at the preheating temperature of 150 “C. The wear volume corresponding to the buffer contents of 50% Ni-Cr+50% Cr,O, was the smallest. The sharper variations in wear volume at the beginning of sliding motion were mild at larger sliding distances; the wear rates are almost the same for the three constituent modes. Figure 5(b) shows the variations in wear volume of Cr,O, coatings where the specimens were heated to 250 “C before testing. The experimental data reveal that the buffer layer of 25% Ni-Cr +75% Cr,O, produced the least wear volume. After a steep increase in wear volume within short sliding distances, the wear volume became steady, and the wear rate was insignificant. Differences between the experimental data corresponding to the two preheating temperatures show that the optimum constitutions of Ni-Cr alloy and Cr,O, material in the buffer layer which produce the minimum wear volume, are closely related to the preheating temperature of the specimen. Elevating the preheating temperature resulted in a reduction in wear volume; and the specimen with the higher content of Cr,O, in the buffer layer produced the smaller wear volume. The two figures show that the effect on wear volume of changing the Cr,O, content is smaller than the range of scatter of tribological data. The curves in Fig. 6 illustrate the variations in wear rate under various testing conditions and for three coating thicknesses. The difference in steady wear rate for each individual curve is nearly negligible, except for the wear rates at the beginning of sliding motion. This means that the
.T. F. Lin, T. R. Li I P~s~a-sp~~~
ceramic coatings
207
steady-state wear rate is hardly influenced by the constituents and the thickness of the buffer layer, and the preheating temperature of the specimen.
rlidingveksiry : t mfs prdl~temp.:ISO’C 0.10 mm (Ni-Cr> (Bau.) 4 (75% NiiG + 25% Cr,O,f (Buff.)
: &,(Imm
i) : 0.10 mm (50% Ni-Cr + SD% Cr.& 0
: 0.10 mm (25% Ni-G
(Buff.)
+ 75% Ciz03) awf.~
Fig. 5. Variation in wear volume with sliding distance: coating thickness 0.3 mm, load 33.4 N, sliding velocity 1 m s-l, temperature of specimen before testing (a) 150 “C and (b) 250 “C.
35.00
_i
y 0 4
30.00
-
*
25.00
-
20.00
-
F =E
sliding
m/s
N temperature:150*C thickness: :0.30 mm, Inark * :0.45 mm in set 5 :o.co mmI
load:33.4 preheated
N temperature:250’C thickness:
specimen
? 15.00
2 0
velocity:1
load:33.4 preheated specimen 0-o A--3 cl-_0
load:89 N Preheated temperature:22?! specimen thickness: *-+ :0.30 mm mark * x-x :0.45 m* in set 4 *-* :0.60 mmI
10.00
2 z
~
-5.00
0.0
5.0
Sliding
10.0
15.0
Distance
I”“““‘I’i’W 20.0 25.0
* 1000
(m)
Fig. 6. Variation in wear rate with sliding distance.
rr 5‘O.G
5.2. Friction The variation of friction coefficient with sliding time for three coating thicknesses is shown in Figs. 7(a)-7(c). The thickness of each layer of the three-layer coating was designated according to the specifications of set 4. The wear experiments were conducted starting with three different preheating temperatures. The coefficient of friction of the specimen with a preheating temperature of 22 “C shows a substantial increase with increasing sliding distance; however, it is expected to obtain a steady value after a longer sliding time. The curves shown in Figs. 7(b) and 7(c) illustrate that the time required to reach a stable friction coefficient is shorter when the specimens are preheated to higher temperatures. The effect of coating thickness on friction coefficient was also dependent on the starting temperature of the specimen. When the tests began at a low preheating temperature (22 “C), no immediate connection between the friction coefficient and coating thickness could be identified. The time required to reach the steady frictional coefficient is relatively longer, but when the specimens were heated to higher temperatures (150 “C and 250 “C), a larger friction coefficient resulted from thicker coatings and the time required to reach steady state was shorter. As the curves in Figs. 7(b) and 7(c) show, there is a drop in friction during the entire sliding procedure, which is probably related to the formation of a softer oxide (WO,) at higher temperatures. However, the friction drop was absent when the low preheating temperature (22 “C) was used. In the region of oxidetilm formation, the friction is decreased. The significant increase in friction is thought to be related to the runin failure where ceramic contacts through the film. Ceramic contacts began at the tips of asperities, friction began to increase, and failure was very sudden. Friction at this stage was also important in determining the wear behavior. High interface temperatures due to frictional heat may be responsible for enhancing softening; softening in ceramics increases the coefficient of friction. The curves in Fig. 8 show the variations of friction coefficient with surface roughness R,. With increasing sliding distance, the worn surface shows continuous reduction in surface roughness. The greatest reduction resulted from the test with the highest preheating temperature (250 “C). The data corresponding to the two higher preheating temperatures (150 “C and 250 “C) reveal that the friction coefficients at large sliding distances were nearly independent of the variations in surface roughness. However, the friction coefficient for
J. F. L&z, 7: R, Li f Plasma-strayed
208
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coefficients were generated to higher temperatures.
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300
Time
400
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Fig. 7. Coefficient of friction as a function three different coating thicknesses.
500
of sliding time for
the preheating temperature of 22 “C showed a continuous increase within the presently designated sliding distances. At the same sliding distance, higher friction
(f)
with sliding distance and
by the surfaces preheated
In tangential motion, the sliding or rubbing of one surface over the other caused surface fracture on the contact surfaces. Table 4 shows EDS analyses of worn surfaces after dry rubbing tests. The elemental analyses reveal that traces of the two coating materials were detected on the opposite surfaces. Using EDS maps, traces of chromium were found on the coating surface of the WC specimen. Similarly, traces of tungsten were found on the surface of the coating Cr,O,. Wear by adhesion had indeed occurred between the two rubbing surfaces. The severity of adhesion was enhanced when the specimen was preheated to higher temperatures. X-ray diffractometry was used to analyze the wear debris. This is an immediate method for accessing the crystal structure of oxides. Figures 9(a)-9(c) show microregion X-ray diffraction maps for WC coatings. There are many diffraction peaks in each figure which were given by more than one substance. With the aid of computer searching in accordance with the location and intensity of the diffraction peaks, the chemical structure of the oxides on the WC coating was derived. According to the ~nsti~ent element analysis, oxidation of WC occurred on the surface forming WO, (melting point of WO, 1180 “C, melting point of WC 2850 “C). Since WO, is rather soft, the sliding couple would show abrasive wear. The height of the XRD peaks in the area of the WC surface shows that the amount of WO, was increased when the preheating temperature of the specimen was raised. Figures lO(a )-10(c) show images of Cr,O, debris on the frictional surface of the WC specimen. The bright
J. F. Lin, T. R. Li f ~~srna-~p~~d TABLE
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209
4. EDS analyses of the warn surfaces of WC and Cr2&
Element
Load
Sliding distance
(N)
(m)
22 150 250 22 150 250 22 150 250 22 150 250 22 150 250
89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4
28800
22 150 250 22 150 250 22 150 250 22
89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4 89 33.4 33.4
28800
Preheating temperature (“C)
si Ti
W
0
Sliding velocity (m s-l)
28800
28800
28800
28800
Proportion (wt.%)
31.506 30.194 31.073 0.890 1A32 0.614 1.146 1.653 1.015 3.241 4.269 4.887 63.217 62.202 62.411
WC sJP&?m
Cr
Si
150 250 22 150 250
particles distributed over the surface were identified as chromium-containing particles, which implies that these particles are CrzO,. The amount of these bright particles increased notably when the specimen was preheated to higher temperatures. The microregion Xray diffraction traces of the Cr,O, coating were also measured, and the atomic concentration of chromium was high. There are many diffraction peaks which perhaps represent two substances including Cr,O, and Cr,C,; however, the amount of C&C, was too small to be detected. Figures ll(a )-11(c) show the surface structure of the Cr,O, disk after dry friction testing, In Fig. 11(a), brittle fatigue spallings are the primary wear mechanism; hardly any evidence of plastic deformation was observed at such a low temperature. When the preheating temperature of the specimen was elevated to 150 “C, numerous pits were distrluted over the surface in place ofsome fatigue spaflings. Part of the surface was smeared
28800
28800
28800
28800
2.513 2.706 4.487 0.000 O.OofI 0.000 0.222 0.096 0.427 85.143 83.960 83.531 12.122 13.239 11.555
flat owing to abrasive wear. When the specimen was heated further to 250 “C (Fig. 11(c)), a different surface from that of Figs. 11(a) and 11(b) is seen. A few smallscale but evenly distributed fractures are seen on the surface in place of the severe brittle fractures found at low temperature. This kind of fault implies that the wear resistance was improved and the wear depth was decreased. The do~nant wear mechanism at this temperature is still brittle fracture.
6. Summary The experiments dealt with the evaluation of wear resistance in multilayered ceramic coatings, A buffer layer with various proportions of Cr,O, was planted between the top and bottom layers. The buffer layer indeed showed the capability of diminishing the wear volume, The optimum proportion of Cr203 content in
210
200Q
1600
800
(a)
32
2*
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68
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/
40
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wo,
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WC 680
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wo3
AL
-1
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44
56
68
88
9. X-ray maps of the worn surfaces of the WC coating and debris after friction testing under various operation conditions: coating thickness 0.3 mm, sliding distance 28 800 m; {a) load 89 N, preheating temperature 22 “C, (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C. Fig.
J. F. Lin, T. R. Li I Plasma-sprayed ceramic coatings
(4
(4
(b)
(cl
(cl
Fig. 10. Images of the Cr,Os debris on the WC surface due to adhesive wear: coating thickness 0.3 mm, sliding distance 28 800 m; (a) load 89 N, preheating temperature 22 “C; (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C.
Fig. 11. Micrographs of the worn surface of the WC coating after wearing tests: coating thickness 0.3 mm, sliding distance 28 800 m; (a) load 89 N, preheating temperature 22 “C; (b) load 33.4 N, preheating temperature 150 “C; (c) load 33.4 N, preheating temperature 250 “C.
212
J. F. Lin, T. R. Li / Plasma-sprayed ceramic coatings
the buffer layer and the preheating temperature were examined, and reasonable interpretations were proposed to illustrate the wear behavior. The interaction between wear volume and the measured temperature at the substrate presented uniform behavior even under various operating conditions and for different coating thicknesses. The primary wear mechanisms occurring on the contact surfaces were identified using EDS and X-ray diffraction analyses.
7. Conclusions
The wear volume of a multi-layered ceramic coating as a function of sliding distance is controlled by factors including the preheating temperature of the specimen before testing, the constituents in the buffer layer, the coating materials of the vane and disk, the thickness of each coating layer, and the applied load. Of these five factors, the preheating temperature of the specimen, the proportion of Cr,O, in the buffer layer, and the coating thickness were the main subjects of the present study. Because ceramics have high strength and brittleness, the testing conditions were designed to prevent brittle fracture during the experiments. The major wear mechanisms occurring at the presently designated specimen temperatures include (1) brittle fracture (spalling and pitting), (2) abrasion (scratching and polishing), (3) adhesion, and (4) ‘oxidation. From the above studies some conclusions can be drawn. (1) If the test starts from higher preheating temperatures, the time required to reach the steady friction coefficient is relatively shorter, the stable friction coefficient is relatively higher, and the wear volume is smaller. However, the effect of Cr,O, content on the change in wear volume is usually smaller than the range of scatter of tribological data. (2) The participation of the buffer layer can indeed reduce the wear volume. Similar to raising the preheating temperature of the specimen, increasing the proportion by weight of Cr,O, in the buffer layer reduces the wear volume. (3) The wear volume is relatively smaller when the specimen is preheated to higher temperatures. The specimens producing smaller wear volume usually have higher substrate temperatures. (4) The steady-state wear rate is not much influenced by the constituents, the preheating temperature, and the thickness of the buffer layer.
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