Accelerated wear testing of PTFE composite bearing materials

Accelerated wear testing of PTFE
composite bearing materials
J.K. Lancaster*
Hertzian line contact stresses between a rotating metal cylinder and a reciprocating flat strip of the bearing material are used in an accelerated wear test
described in this paper. The test readily demonstrates differences in the wear
behaviour between different PTF E-based d r y bearing materials. For one
widely-used woven PTFE fibre composite, the relationships between wear
and stress, temperature, counterface roughness and fluid c o n t a m i n a t i o n were
shown to be similar to those f o u n d in long-term bearing tests
Current military and civil aircraft contain from several
hundred to several thousand dry bearings, respectively.
Many of these are in non-critical areas, such as hinges or
latches etc, but a significant proportion are crucial to the
performance and reliability of the aircraft itself. Until the
present decade, one particular type of dry bearing material
has predominated in aircraft applications: a thin layer of
interwoven PTFE and glass fibres impregnated with synthetic resin and adhesively bonded on to a more rigid
backing. The load-carrying capacity of such bearings is
relatively high, up to 350 MPa, but thermal considerations
limit operation at high stresses to low sliding speeds
( < 0.01 m/s). As aircraft components become increasingly
sophisticated, a number of requirements have arisen, or can
be foreseen, which necessitate improvements in performance
from this type of bearing, eg greater load-carrying capacity,
greater stiffness, longer life (lower wear), greater tolerance
to temperature extremes, and insensitivity to contamination
by solid particles or fluids. To meet these requirements,
a 'second generation' of liner materials is now emerging
based either on variants of the woven PTFE fibre/reinforcing fibre concept or on other forms of construction. Some
examples of the various products currently available commercially are given in Table 1. The combinations of
materials and methods of construction which have been
examined so far, however, represent only a small fraction
of a whole range of possibilities, some of which are shown
in Table 2. There would consequently appear to be considerable scope remaining for the further development of dry
bearing materials.
Evaluation of the wear life of thin-layer dry bearings is
both costly and time-consuming. The aircraft manufacturers
require information primarily for design purposes, but to
ensure the relevance of wear data to particular applications
it has always been deemed necessary to undertake bearing
tests in conditions closely simulating those of the intended
application. Because the stresses involved are usually high,
large and expensive apparatus is required. Further, because
sliding speeds are always low, and wear rates (of the most
successful materials) are also low, long test periods are
needed to establish a defined wear rate, or a total wear life.
The situation is further exacerbated by the fact that the
operating conditions imposed on aircraft dry bearings cover
*Procurement Executive, Ministry o f Defence, MaterialsDepartment, Royal Aircraft Establishment, Farnborough, Hampshire, UK
0301-679X/79/02065-11 $02.00 © 1979 IPC Business Press
Table 1 Some types of thin-layer, dry-bearing materials
available commercially
Interwoven PTFE fibres/other fibres + resins
Glass - Fiberslip, F a b r o i d ,
D a c r o n - Fiberglide,
Duralon
N o m e x - Faftex,
Fibriloid
Copper- P / d a n e
Airflon
Resins + PTFE additions + fabric reinforcements
Nomex
Dacron
-
Uniflon
Kahrlon,
Fraslip
Filled/reinforced PTFE
Carbon-graphite,
nickel, ceramic,
bronze
Bronze, stainless steel
meshes
Saimap,
Rulon
- Unimesh,
Saimap,
Metalloplast
Porous bronze + polymer
PTFE + Pb
Acetal
Poly(phenylene sulphide} -
Glacier D U
Glacier D X
Glarnat 5 8
a wide spectrum: some of the most important are listed in
Table 3. Unfortunately, it has not so far proved possible to
extrapolate data obtained in one particular set of sliding
conditions to predict performance with an acceptable
degree of certainty in other conditions. A similar range of
problems faces bearing manufacturers. Information is needed
on their own, and competitors', products: it is also necessary
to monitor performance for quality control purposes. In
addition, experimental materials emerging from development programmes usually require rapid evaluation to
provide feed-back to guide further development.
All the above considerations clearly point to the need for
some form of accelerated wear-test procedure for thin-layer,
dry bearing materials. It would be unreasonable to expect
such a test to provide design data directly applicable to
service conditions, but at the very least it could focus
attention on the most promising materials and so reduce
the volume of simulation testing ultimately needed. In
addition, an accelerated test should, in principle, be able to
TRIBOLOGY international April 1979
65
Table 2 Possible material variants in dry bearings
LINER
Resin-adhesive
Weave Structure
Phenolic
'FriedeI-Crafts'
Vinyl-phenolic
Epoxies
Silicones
Polyimides
Fibre size and spacing
Single/multifilament fibres
Mixed multifjlaments
Proportion of PTFE to other fibres
Type of
PS
HT Thermoplastics PES
PPS
weave
Fibre-filler type
Plain
Satin
Twill
Leno
Gauze
Filament winding
Polyimide
Polyimide-amide
Nomex
Dacron
Fibres Glass
Kevlar
Carbon/graphite
Metals (Cu, B, steel)
PTFE flock or particles
Lamellar solid libricants (Graphite, MoS2)
COUNTERFACE
Metals
Coatings
Steels - heat treatments
Aluminium alloys
Titanium alloys
Plasma spraying
CVD
carbides
Diffusion treatments - boriding, nitriding
Electrochemical - Cr, Co-Cr2 C3
Platings
Ion - AI on Ti, Cu
Bonded solid lubricants
Table 3 Operational parameters in aerospace dry
bearing applications
Load
Type of
Motion
T e m p e r a - Environment
ture
Undirectional
Fluctuating
Reversing
Axial/radial
ratio
(sphericals)
Continuous
Steady
Oscillatory
Cycling
Vibration
Misalignment
Humidity
Contamination fluids, solids
Apparatus
Bearing size and shape," Speed," Counterface roughness;
Heat dissipation parameters
Motor for
reciprocefion
Drive m o t o r
~
~
~
bearing
Test
strip
Melal counterface
ring
Fig I R e c i p r o c a t i n g line c o n t a c t apparatus
66
establish qualitative trends relating wear to the controlling
variables, such as load, speed, temperature, counterface
roughness, environment, contamination etc. This paper
describes the development of an apparatus for accelerated
wear testing of thin, dry bearing, liner materials and
presents preliminary results obtained on several types of
products in different conditions of sliding.
TRIBOLOGY international April 1979
One of the difficulties associated with the choice of any
accelerated wear test is uncertainty about whether the
dominant parameter influencing wear is the absolute load
or the nominal stress. Simple theoretical considerations 1
lead to the conclusion that the rate of wear (per unit
distance of sliding) should be directly proportional to the
load and independent of the apparent area of contact;
experiments with relatively small apparent areas of contact
have generally confirmed this prediction 2 . With large
contact areas, however, complications begin to intrude due
to difficulties associated with the escape of wear debris
from beneath the contact zone. It is then found that wear
rates can depend to some extent on both the size and shape
of the contact area a's. A further possibility also exists that,
due to material strength limitations, wear rates may begin
to increase when the nominal stress exceeds some critical
value. For metals, Burwell and Strang 6 have suggested that
this critical stress is of the order of H/3 where His the
indentation hardness, but corresponding criteria for polymers and composites have never been established.
For the present purposes, small-scale, wear-testing equipment was devised in which the absolute loads between the
sliding surfaces were, of necessity, relatively low, but which
induced high contact stresses via Hertzian line contacts. The
apparatus (Fig 1) was based on a concept originally described
by Michalon e t al 7 . A strip of the bearing material of
Lood,
~o
(.9
F
D
4
N
•
150
x
o
o
300
450
600
\
o
x
E
[3
3
8
m
LU
2
--
I
--
0
I
I
I
10
20
30
I
40
I
50
60
B a l l diameter, m m
Fig 2 Elastic modulus v. ball diameter for PTFE-fibre composite, K, 0.28mm thick on steel
interest, 6.35 mm wide x 38 mm long is loaded in fine
contact against the curved surfaceof a rotating ring, 25 mm
in diameter. The bearing strip is mounted in a self-aligning
arm and arranged to reciprocate slowly ( ~ 3 cycles/min)
over a stroke length of 12.5 mm on a rolling element slide
attached to an electrically driven cam. The particular merit
of this arrangement is that, apart from at the end of each
stroke, fine-contact conditions are always maintained
despite wear o f the bearing material and the metal ring.
Friction was measured from the output o f a torque transducer interposed between the rotating ring and a variable
speed driving motor, and displayed continuously on a chart
recorder. Wear was determined, in terms o f depth, by
traversing the strip beneath a vertically mounted displacement transducer and again displaying the output on a chart
recorder. In general two measurements of wear depth were
made, approximately 1.5 mm from each side of the bearing
strip.
and the recovered depth, h, after removal of the major
load but with a minor load remaining. Assuming that there
is no change in the radius of the contact circle, a, following
elastic recovery9 it is then possible to determine a and R
from measured values of d and h using the sagittal relationship. Fig 2 shows values of the elastic modulus of one
particular composite dry bearing liner material, 0.275 mm
thick, on a steel backing, obtained with five different ball
sizes and four loads. It can be seen that the modulus is
essentially independent of load and also of ball diameter
from 12.5 mm upwards. The increased modulus obtained
with smaller ball diameters is presumably attributable to
an increasing contribution from the steel backing as the
ball diameter decreases and hence the depth o f penetration
increases. The elastic modufi of a number of other commercially-available, dry-bearing finer materials were also
measured, and values are given in Table 4, assuming
Poisson's ratio = 0.35. There are few published values of
the moduli of these types of materials available for
comparison, but estimates by Rowland and Wyles 1° from
static compression tests on small pads give 4.2 GPa for
material M and 4 GPa for material K, in reasonable agreement with the present results.
From the measured values of the elastic moduli, the maximum compressive line-contact stress relevant to the wear
experiments can be computed directly from elasticity
theory; ac = 0.066 ~/PE MPa, for a 25 mm dia steel ring,
a pad width o f 6.35 ram, and a total load of P newtons.
Table 4 Elastic moduli of different types of polymerbased dry bearing liners
Code
A
Material Description
Elastic
Modulus
GPa
PTFE flock in synthetic resin
reinforced with 'Terylene' fabric
1.6
B
Interwoven PTFE/'Nomex' fibres
+ high temperature resin
1.65
C
As above, B, but lower temperature resin
1.7
D
PTFE flock in synthetic resin
reinforced with 'Nomex' fabric
2.2
E
Interwoven PTFE/glass fibres +
polyimide resin
2.2
As D, but with coarser fabric
weave
2.45
G
Interwoven PTFE/'Nomex' fibres
and resin/PTFE overlay
2.5
H
Granulated PTFE in vinylphenolic resin + 'Dacron' fabric
2.65
J
Filled PTFE-impregnated bronze
mesh
3.2
0.75 (1 - v2)LrR
E =
a 3 (R - r)
K
Interwoven PTFE/glass fibres +
phenolic resin - surface all PTFE
3.65
where L is the applied load, r is the radius of the ball
indenter, a is the radius of the contact circle, R is the
recovered radius of the indentation after removal o f the
load and v is Poisson's ratio. The Rockwell hardness tester
measures the depth o f penetration, d, under a major load
L
As K, but with some glass fibre
exposed at surface
4.3
Porous bronze impregnated
with PTFE/Pb
5.0
Elastic Modulus Measurements
To calculate the contact stress between the metal ring and
the bearing strip from elasticity theory, it is necessary to
determine the elastic modulus o f the bearing material. This
is a somewhat imprecise quantity for polymers and
composites since it depends markedly on both the magnitude of the imposed strain and its rate of application. The
simplest way to derive a modulus relevant to the wear
experiments is to measure the elastic recovery o f ball
indentations, and a convenient apparatus for this purpose
is a Rockwell hardness tester.
F
From elasticity theory s ,
M
TRIBOLOGY international April 1979
67
Foilure
T
'
Plateou '
C
'Knee '
Time (distance of sliding,cycles of oscillation )
Fig 3 Idealized wear depth-time relationship for thin-layer,
dry-bearing liners
300
E
25O
C
K
E
200
~
J
~" 150
I00
A
H
5O
P
0
?;
;
5
I0
TT
15
;
i
20
25
Time, h
I
30
I
35
Results
It is appropriate at this stage to discuss, briefly, the various
ways in which the wear properties of dry bearings can be
defined. The quantity of practical interest to the user is
the depth of wear, since this determines the degree of
'back-lash' introduced into a bearing assembly. Ideally, the
detailed relationship between depth of wear and time (or
distance) of sliding should be determined for each bearing
material, size and particular combination of sliding
conditions: this situation is seldom realised in practice.
Many publications merely quote the total duration of
sliding before failure, or before reaching an arbitrarily
defined depth of wear. Specification requirements, eg
MIL-B-81820C n also invoke a maximum tolerable depth
of wear within a given time of operation of a bearing under
defined conditions of sliding. Data of this type are o f only
very limited value when attempting to compare the wear
behaviour of different bearing materials and sizes in different conditions.
Wear depth-time relationships for most types of t;tin-layer
dry bearings are usually of the form idealized in Fig 3. Wear
is initially rapid, AB, and rises to a 'knee' beyond which
there is a 'plateau' region, BC, of constant, or almost constant, wear rate. After reaching a critical depth of wear, the
approach to failure, CD, is usually comparatively rapid.
TRIBOLOGY international April 1979
Typical wear depth-time relationships obtained with the
reciprocating line contact (RLC)apparatus for a number
of commercially available, dry bearing liners are shown in
Fig 4 la. In all cases, the counterface material was AISI
440C stainless steel, hardened to 700 VPN and randomly
abraded to give a surface roughness of 0.05-0.08 prn Ra.
Although the absolute load was constant for all materials,
the stress-levels vary slightly because of the differing
elastic moduli. The general shapes o f the curves resemble
those typical of bearing tests on these types of materials
and the results clearly demonstrate that there are marked
differences between the various products. It is not,
unfortunately, possible to compare these results directly
with those from bearing tests because data are not available
in the literature on such a wide range of materials in
constant conditions of sliding. In order to attempt an
assessment of the relevance of the wear results in the
RLC test to bearings, further experiments were therefore
made concentrating on two particular materials for which
some bearing test data are readily available; a porous bronze
layer impregnated with PTFE and Pb (M) and an interwoven PTFE fibre/glass fibre cloth impregnated with
phenolic resin (K).
I
40
Fig 4 Wear depth-time relationships for 10 dry-bearing
materials in reciprocating line contact 12. Load = 450N,
Speed = O.13m/s, counterface 440C. (Details of materials
in Table 4)
68
From relationships of this type, a 'specific wear rate' can
be defined for the plateau region, BC, in terms of the
volume of wear per unit distance of sliding per unit load.
The magnitude of this specific wear rate, together with
information about the volume of initial wear, if significant,
then provides sufficient data to permit ready cross-comparisons between the wear of different materials in different
sliding conditions.
Stress
Wear depth-time relationships were obtained for the
porous bronze-PTFE-Pb composite at a range o f loads,
and hence stresses, and the values of the specific wear rates
over the plateau region are shown in Fig 5, curve 1, plotted
against the maximum compressive Hertzian line contact
stress. Data for comparison purposes, obtained from a
variety of sources, is shown in curves 2 - 7 . In transposing
this published wear data into its present form, it has been
necessary to make a number of assumptions for those
situations where depth-time relationships were unavailable;
details are also given in the legend to the figure. For journal
bearing tests, the volumetric loss, V, was computed from
the depth of wear, d, using the relationship,
V = 1.57 Dwd fn(~)
where D is the bearing diameter, w is the width, and e is the
initial clearance. The latter is seldom reported, and was
assumed to be 25 #m; the numerical value o f fn(d/e) is
therefore taken as 0.913 . Data is only included in the comparison for those tests in which the mean surface temperatures are either known, or can reasonably be assumed, to lie
within the range 20°C to about 50°C.
Although Fig 5 shows that there is clearly no exact agreement between the specific wear rates in different sliding
conditions at any one stress level, the results taken as a
whole appear to define a general pattern. The specific wear
rates begin to increase rapidly when the stress exceeds
some critical value. No significant wear pattern is evident,
however, if the wear rates are plotted against load rather
than stress. To see whether this increase in wear rate with
stress was more general, a similar series o f RLC experiments,
together with a comparison with published data, was made
for an isotropic material, polyacetal. These results are given
in Fig 6 and show no significant variation of wear rate with
stress up to the maximum values attainable. It must therefore be concluded that the wear pattern exhibited by the
porous bronze composite is a feature of the material itself.
i0-4
E
z
x
10-5
2E
'_/// t
10-6
x
×
o
10-7
lO-e
Q
I
I0
I
I0.
Stress, MPo
I
I00
I000
Curve Source
Conditions
Assumptions
1 Present work
Reciprocating pad on rotating
ring; unidirectional motion,
0.65m/s; counterface AID 71B
tool steel, 800 VPN, 0.15 ~tm
Ra; loads 70-400N. Wear depth
v. time measured.
E = 5.0 GPa
Journal bearings, 15.9mm id X 19
mm wide; unidirectional motion,
speeds 0.62, 1.24, 2.48 m/s; mild
steel shafts. Life measured.
P calculated
assuming constant V of
1.24 m/s Initial overlay of
25/~m removed
very rapidly.
Effective
depth of bearing layer =
38/~m
2 Pratt t4
3 Anderson I s
The influence of stress on the mean coefficient of friction
during the 'plateau' regime o f wear is shown in Fig 9. For
any one set of sliding conditions the values decrease with
increasing stress, but the absolute magnitudes of friction
E
z
g"
E
E
8
10-6
As above, 3, but with grooves
machined in surface to permit
escape of debris
5 Rowland and
Wyles 1°
Pad on reciprocating track;
pad size 36ram X 7.2ram,
track FV520 CR steel, 0.050.1 ttm Ra; speed = 0.087m/s;
stroke length = 0.244 m; temperature controlled, 20°C. Cycles
to failure measured.
Effective depth
of bearing
layer = 38~,m
Pad on reciprocating track; pad
size 31 X 31 ram, track nitrided
H50 steel, ~ 1000 VPN, 0.05-0.1
/~m Ra. Speed = 0.025 m/s, stroke
length = 0.15m, Total distance
of sliding to failure determined.
As 5.
7 Rowland and Journal bearings,'127 m m i d X
Wyles 1°
45.6 mm wide; oscillatory
motion, -+25°, 3 cycles/rain,
shaft FV520 CR steel, 0,050.1 ~m Ra; temperature controlled 20°C. Cycles to failure
measu red.
The RLC experiments also show that as the stress, or load,
increases, the initial amount o f wear, prior to reaching the
wear 'plateau', also increases (Fig 8(a)). This initial depth
of wear is often rather ill-defined and is not always quoted
explicitly in reports of practical bearing tests. However,
the same general trend is implicit in results given by
Rowland and Wyles 10 and shown in Fig 8(b); the number
o f cycles of oscillation to reach an initial depth of wear
of 0.127 mm decreases with increasing stress. Similar results
have also been reported by Barrett 27 .
10-5
Thrust washer, 51ram od X
38mm id against mild steel, 0.080.12#m Ra. Unidirectional
motion, spee~l: 0.01 m/s. Wear
measured by weight loss &
depth.
4 Anderson is
6 Cheesman 16
The effects of stress on the specific wear rate of the PTFE
fibre/glass fibre-phenolic resin composite, K, are shown in
Fig 7, curve 1 for the present RLC tests, and in curves 2 - 1 1
for other conditions of sliding. Again, the temperatures
induced during sliding in all cases are known, or believed,
to have been within the range 2 0 - 5 0 ° C . In contrast to the
results with the porous bronze, there is now no clear pattern
of wear behaviour; some experiments, including these in
RLC conditions, show wear rates increasing with stress,
whereas others show either decreasing or approximately
constant wear rates. Replotting the wear rates against
absolute load, rather than stress, again fails to reveal any
general trend.
IS
IT
e4
-P
b
5
t9
03
10-7
I
I
O. I
I0
___L_
. . . . . . . .
tO
IO0
Stress, MPo
As 5.
Fig 5 Variation o f wear rate with stress for porous bronze
impregnated with PTFE/Pb, M
Fig 6 Variation o f wear rate with stress for polyacetal
1 Present work. Reciprocating line contact against 18%W
tool steel
2 Lancaster 17. ~ o s s e d cylinders, against 18%W tool steel
3 Du Pont ~
4 Theberge a9 Thrust washers against 1040 steel
5 Clerico & Rosetto 2°. Amsler machine against steel - line
contact
6 Tanaka & Ychiyama 21. lh'n on steel disc
7 Shen & Dumbleton 22. Thrust washers against stainless
steel - oscillatory
8 ESDU a3. Journal and thrust bearings - manufacturer's
data
TRIBOLOGY international April 1979
69
appear to depend greatly on the sliding conditions. However,
a more generally consistent pattern of behaviour now
emerges if the coefficients of friction are plotted against
the absolute load. From this, it is clear that there is an
overall trend towards decreasing friction with increasing
load, similar to that typical of pure PTFE 2s and several
other polymers 29 .
Counterface
Roughness
It is generally accepted that the wear life of thin-layer,
PTFE-based, dry bearings increases as the counterface
surface roughness decreases. For high performance aerospace bearing applications, counterface roughnesses are
normally in the range 0,05-0.10/lrn Ra. To obtain a more
quantitative appreciation of the importance of counterface
roughness, RLC experiments were made with the PTFE
fibre/glass fibre-phenofic resin composite, K, at a range of
stresses on counterfaces of four different roughnesses.
Fig 10 shows the variation of the total life to failure with
stress, and the curves clearly demonstrate that a significant
life on the roughest surfaces is only attainable at very low
stresses. Conversion o f these results to provide the
relationship between specific wear rate and counterface
roughness at one particular stress is illustrated in Fig 11,
curve 1. The only published data available for comparison
on a generally similar type o f bearing material are shown
in ctrrves 2 and 3.
Fig 7 Variation o f wear rate with stress for PTFE Fibre composite, K
I0
abc
10-41
4 Cheesman 16
i0-5 E
z
E
E 10-6 -o"
10-7
i
Pad on reciprocating track; pad
size 31 X 31 ram; track nitrided
H50 steel, ~ 10000 VPN, 0.050.1#m Ra; speed = 0.025 m/s,
stroke length 0.15m. Wear
depth v. time measured.
5 Rowland and Pad on reciprocating track; pad
Wyles 1°
size 36 mm X 7.2 rnm; track
FV520 CR steel 0.05-0.1 ~m
Ra. Speed = 0.087 m/s, stroke
length 0.244 m; temperature
controlled, 20°C. Cycles to
wear 0.127 ram, 0.254 mm and
0.305 mm determined
Wear rate assumed constant
between 0.127
mm and 0.254
mm
6 Ampep =4
Spherical bearings, 19.05 mm
bore; counterface AISI 440C,
0.05 #m Ra; oscillatory motion,
+-25°, 10-13.5 cycles/min
Volume calc.
as for journal
bearings assuming a mean
dia f o r the
ball.
7 Wade 2s
Journal bearings, 38 mm dia, 9
mm wide; counterface AISI
440C; continuous rotation,
0.02 m/s. Wear depth v. time
measu red.
8 Wade 2s
As above, 7, but oscillatory
motion +25°; max speed, 0.013
m/s. Wear depth v. time
measu red.
9 Anderson ~s
Thrust washer, 51 mm od X 38
mm id. a) on mild steel, 0.08-0.12
#m Ra; unidirectional motion,
speed = 0.01 m/s; b) as (a), but
oscillatory motion, 10 cycles/
rain, -+25°; c) as (b) but on AISI
440C. Wear measured by depth and
weight loss,
10 Craig 26
Journal bearings (in glass-phenolic
housings); a) 38 mm X 6.35 ram;
b) 25 mm X 9.5 ram. c) 9.5 mm
X 12.7 ram; counterfaces stainless steel and Cr plate, 0.13-0.18
#m Ra; oscillatory motion,
+-25°, 10 cycles/rain. Life
measured to 0.127 mm wear.
11 ESDU 23
Range for filled PTFE formulations generally. Bearing tests,
journal and thrust, under 'light
duty' conditions of light loads
and low speeds for which temperature increases are small. Manufactu rer's data.
G0
/
I
I
I
lO-e~i
9b~,l
I
i
I
~9o
I
10-9
OI
I
Io
Stress, MPo
IO
I
IOO
_
I000
Curve Source
Conditions
Assumptions
1 Present work
Reciprocating pad on rotating
ring: unidirectional motion,
0.13 m/s; counterface AISI
440C, ~ 700 VPN, Wear
depth v. time measured.
E = 3.6 GPa
2 Lancaster ~7
3 pins, 5ram dia on disc; unidirectional motion, 0.09 m/s;
counterface Cr-plate, ~ 850
VPN,,0.05 #m Ra; loads 4008000N. Wear depth v. time
measured.
3 Rowland and
Wyles 1°
Journal bearings, 127 m m i d
× 45.6 mm wide on FV520 CR
steel shafts, 0.05-0.1#m Ra;
oscillatory motion, +-25°, 3
cycles/rain. Temperature controlled at 20°C. Cycles determined for 0,13 mm wear and
to failure, 0,30 mm wear.
70
TRIBOLOGY
international April 1979
Wear rate assumed constant
between 0.13
mm and failure,
0.30 ram.
0.05 mm assumed to be removed very
rapidly followed by constant
wear rate to
0.127 ram.
Temperature
or no effect. Craig 32 has concluded that the total life of
spherical bearings with woven PTFE liners during oscillatory motion is reduced in sea-water to between 1/3 and ¼
of that obtained in dry conditions, and Barrett 27 has also
reported deleterious effects due to salt-spray contamination.
The variation of specific wear rate with temperature
obtained from RLC experiments with the PTFE fibre
based composite, K, is shown in Fig 12, curve 1, and data
for comparison from other tests on similar materials in
curves 2 - 7 . Apart from a possible shallow minimum in
curve 2, all the results show a general increase in wear rate
with temperature. The absolute values of specific wear rate
at any one temperature, however, depend markedly on the
particular conditions of sliding imposed. The RLC experiments further showed that the initial depth of wear during
the early stages of sliding also increased slightly with
increased temperature, and a similar trend is implicit in
data from joumal bearing tests reported by Rowland and
Wyles x° .
Discussion
The first point to note from the results presented above is
that the wear rates of both the porous bronze and the
PTFE fibre composites in RLC conditions are generally
greater, at any one stress, than those found in most bearing
tests. Only with polyacetal is the wear rate largely independent of stress and other conditions of sliding. The main
reason for the high wear rates of the composites probably
results from the fact that in line-contact conditions there
is an appreciable penetration of the steel cylinder into the
composite surface layer. The depth of penetration can be
calculated from a modification of an analysis by Crook 3s
which leads to the relationship:
0.636 (1 v2)I~a
E
1/2 3
Fluid Contamination
The effects of water and a mineral oil-based hydraulic fluid,
DTD 585, on the wear depth-time relationships for the
PTFE fibre composite, K, in RLC experiments are shown
in Fig 13. Both fluids clearly reduce the total wear life,
either when introduced following a period of dry sliding,
curves B and C, or when present from the onset of sliding,
curves D and E. Data for comparison from bearing tests on
a similar finer material are somewhat inconclusive. Fig 14(a)
shows similar deleterious effects of both fluids for spherical
bearings in conditions of reversing load, but under steady
loads Fig 14(b) shows that water does not reduce life, and
Fig 14(d) shows that DTD 585 reduces wear in the early
stages of sliding but has little effect on the steady-state
wear rate. The results of pin and disc tests 34 on PTFE
+ 25% chopped glass fibre, in Fig I4(c), again show that
water increases the wear rate, but that DTD 585 has little
a=
wE
~.½+ln[2"Sw3/2(PD(1-u2)) ]~
-
where P is the total load applied, w is the width of the
bearing strip, D is the ring diameter, E is the elastic modulus
of the bearing material, and v is Poisson's ratio. For w =
6.35 mm,D = 25.4 mm, and v = 0.35 the depth of penetration into the PTFE fibre composite, K, ranges from 0.011
to 0.063 mm over a load range from 100 to 750 N. Similar
values are obtained from a more recent analysis by Brewe
and Hamrock 36 . It follows that in RLC conditions, elastic
deformation will affect a much greater depth of the surface
layer than in bearings where the contact area is distributed.
In consequence, wear resulting from localised stress concentrations in RLC conditions may be supplemented by a
E
=L 200
,..*
-6
150
QI4
E I00
"6
,,/I I
123
o
20
40
60
80
MPa
~ 800
~ 600
400
2
~
IOO
200
o
x
300
MPo
Fig 8 Effect of stress on initial
wear of PTFE fibre composite,
K
(a) Reciprocating line contact
(b) Rowland and Wyles x°
002
t
0
zK)O
500
10 2
103
[3 £3
x.
"~.
0.02 }[
I0
10
Stress, MPo
I
I
o
~ oo6
~- 0 0 4
010
0.08t006}004 [-
o
b
I00
i e°rings
200
._~
028
- - g 026
"6 0.24
"5 022
020
o 18
016
014-1
0121
I
a
E
i'E,,.~-- I000
"6
g ol2
o.10
"6 0.08
030
~ 50
I02
•
I03
[....
I04
~x 6
[
!05
__
106
I07
Load, N
Fig 9 Variation of friction coefficient with stress and load for PTFE fibre
composite K (Sources of data as for Fig 7)
TRIBOLOGY
international A p r i l 1979
71
50h~
2o
data in Fig 7 is not, however, available for comparison. It is
evident that a detailed examination of the basic wear
mechanisms of PTFE-fibre composites, and of the relationships between surface film formation and the conditions of
sliding, is long overdue.
[]
i
16
14
IO
×
8
6
4
i
065
2
,
20
40
60
80
Stress, MPo
I00
120
Fig 10 Effect of counterface roughness on life-stress relationship for PTFE fibre composite, K, in reciprocating line
contact (counterface tool steel, AID 71B, 800 VPN}
contribution arising from bulk deformation. For an isotropic material, such as acetal, this additional contribution
appears to be negligible, but for composites a very different
situation arises. Repeated, bulk deformation during reciprocation will encourage disruption at the interface between
the various constituents in the composite leading to weakening of the surface layer and thus to an enhanced wear rate.
As the load (stress) increases, the depth of penetration also
increases, and so, in turn, will the bulk deformation contribution to wear. It is apparent that at any one level of
stress the sliding conditions imposed by the RLC test are
appreciably more severe than those characteristic of
bearings.
All the wear results for the porous bronze composite, M,
in Fig 5 clearly point to the conclusion that the performance of this material is stress-limited. This is, presumably,
a consequence of its partially sintered structure. The
dependence of wear rate on stress in RLC conditions is
very similar to that characteristic of bearing conditions
at high stresses. The situation with the PTFE-fibre
composite, K, however, is much less well-defined. From the
results in Fig 7 it is only possible to conclude that the
dependence of wear rate on stress in RLC conditions is at
least not wholly dissimilar from that obtained in some
bearing tests. One characteristic feature of the wear
behaviour of most PTFE-based composites is the formation
of transfer films on the opposing metal counterface and/or
films of consolidated wear debris over the surface of the
composite itself. It is generally accepted that wear tends
to decrease as these surface films become more uniform and
coherent. A preliminary optical examination of surfaces
worn during RLC conditions has shown little evidence of
film formation on either the PTFE-fibre composite, K, or
the metal counterface, and this is at least consistent with
the relatively high wear rates obtained. Information on the
extent of debris film formation relating to the other wear
72
T R I B O L O G Y international April 1979
Despite the uncertain effects of stress on the wear rate of
the PTFE-fibre composite, K, the RLC test appears to
provide trends relating the wear rate to counterface roughness and temperature which are broadly similar to those
found in bearing tests. The results in Figs 10 and 11 all
confirm the value, in current dry bearing practice, of
providing counterface roughnesses within the range
0.05-0.1 #m Ra. Moreover, they also suggest that further
possible reductions in wear, which might result from using
counterfaces smoother than 0.05 p_m Ra, could well be
insufficient in magnitude to justify the extra cost involved
in preparing such surfaces. Increases in wear rate with
increasing temperature are a general feature of most types
of PTFE composites 31 and the curves given in Fig 12
demonstrate that the PTFE fibre composite, K, is no
exception. It has been suggested that an increase in temperature affects wear by inhibiting the formation of debris films
on the surface of either the counterface, the composite, or
both al . An additional possibility also arises in RLC conditions since, as the elastic modulus decreases with increasing
temperature, there will be greater penetration of the
counterface into the composite leading to a greater degree
of surface disruption.
10-4 l
o/
2
E
z
E
E
o 10-5
/.
(3
U)
/
x
x
Y
iO-6
/
I
I
O.OI
I
I
I
I
I
I
I
O.I
Surfoce roughness,/z rn Ro
I
I
I
I0
Fig 11 Variation of wear rate with counterface roughness
for PTFE fibre composite, K. (1) Reciprocating line contact, O.13m/s, 76 MPa {2) Craig26. Journal bearings, 25 mm
x 9.5 mm;oscillatory motion +-25 °, 10 cycles/min," IOMPa.
(3) Rowland and Wyles 1°. Pad on track, O.087 m/s, 186 MPa
iO-5
--
1
/
E
Z
t~
E
E
d id e
iO-~
1(3"
I
40
I
80
I
120
I
160
I
200
I
240
I
280
Temperature, °C
Conditions
Assumptions
1 Present work
Reciprocating pad on rotating
ring; unidirectional motion, 0.13
m/s; counterface AISI 440(3,
~ 700 VPN, 0.05-0.1/~m Ra, 80
MPa. Wear depth v. time
measured.
E, and hence
stress assumed
independent
of temperature
2 Rowland and
Wyles 1°
Journal bearings, 127mm id X
45.6 mm on FV520 CR steel
shafts, 0.05-0.1jura Ra. Oscillatory motion, +25°, 3 cycles/
rain; 310 MPa Cycles determined for 0.127 mm and 0.305
mm wear.
Wear rate assumed constant
between 0.127
and 0.305ram.
3 Rowland and
Wyles 1°
Pad on track, 31 X 31 mm pads;
track F V 5 2 0 C R steel, 0.050.1 jura Ra; speed 0.087 m/s;
stroke length 0.244 m: 186 MPa.
Cycles determined for 0.13,
0.254, and 0.305 mm wear.
As above, 2
Curve Source
4 Williams a°
5 Wade ~s
6 Ampep 24
7 Evansal
Journal bearings, 15.9mm X
15.9mm. Oscillatory motion,
-+32°, 10 cycles/min; Counterface Cr plate; 70 MPa. Depth
v. time measured.
Sphericals, 19 mm bore; oscillatory motion, +-25° , 10 cycles/
rain. Counterface AISI 440C;
182 MPa. Total wear measured
after 10 s cycles
The effects of fluid contaminants on the wear of the PTFE
fibre composite, K, in both RLC conditions (Fig 13) and in
reversing load tests on spherical beatings (Fig 14(a)) are
clearly deleterious to performance. In bearing tests with
unidirectional, steady loads, however, the results are more
ambiguous. Whether or not the apparent reduction in wear
with the mineral oil based hydraulic fluid in the early steps
of sliding in RLC Conditions (Fig 13, curve E) is significant
remains to be confirmed, but if so, it is consistent with the
results for spherical bearings (Fig 14) where the duration
of Sliding is also relatively short, and for the pin and disc
results (Fig 14). It has been suggested on the basis of pin
and disc tests that the deleterious effects of water on the
wear of PTFE composites result mainly from the inhibition
of transfer film formation 37. A contributory factor may
also be the tendency for the water to penetrate the interface between the reinforcing fibres and the matrix and
weaken the surface layer, eg with PTFE-glass fibre composites. During sliding in the presence of excess fluid,
hydrostatic stresses could then develop and facilitate
disruption of the surface layer, leading to increased wear.
This last process could well be particularly significant in
RLC conditions and in the presence of reversing loads.
One difficulty associated with the interpretation of literature data on bearing tests in fluid-contaminated conditions
is that it is not always very clear how much fluid is actually
present during wear. The specification requirements for
compatibility between bearing materials and fluids, eg
MIL-B-81820C 11 , merely call for the material to be soaked
in the appropriate fluid for 24 h and subsequently tested
under nominally 'dry' conditions. This procedure is likely
to give very different results than when testing in the
presence of a continuous supply of fluid.
Of the three main requirements, outlined in the introduction, which an accelerated wear test should satisfy quality control, preliminary screening of new materials,
and establishing trends between wear and its controlling
parameters - attention has been concentrated in the
present work mainly on the latter. As far as quality control
0.25
C
0.20
:
o
/,?/
xt
I
A Dry
E
E
Assume mean
die for ball,
and initial
running-in wear
of 0.05 ram.
Assume mean
Sphericals, 31.75mm bore; oscildia for ball.
latory motion, -+25° , 10-13.5
cycles/rain. Counterface AISI 440C;
175MPa. Wear depth v. time
meesureo
Range of filled PTFE composites
in pin-disc expt; counterface
18/8 stainless steel ($130), 0.15
p.m Ra; 0.7 m/s, load = 12N
(stresses ~ 1-5 MPa). Wear volume
v. time measured
Fig 12 Variation o f wear rate with temperature for PTFE
fibre composite, K
OI5
t
x ..~.__~ ~-~
t
0.10
L/~x/f/
0.05
II/~
,"
~/eter o
L~,,"
O
/
W°tero /
DTD 585
I
50
I
IOO
I
I
150
200
Time. rain
I
250
I
:300
Fig 13 Effect o f fluid contamination on depth wear-time
relationships for PTFE fibre composite, K, in reciprocating
line contact. A dry, B and C water and DTD 585 added
after initial dry sliding, D and E water and DTD 585 present continuously. Load = 450N, speed = 0.13 m/s
TRIBOLOGY
i n t e r n a t i o n a l A p r i l 1979
73
I
-x. . . . .
weeks, and sometimes months, required for tests on
actual bearings.
x
Results obtained using this accelerated test procedure
have shown that:
1. There are marked differences in the wear behaviour of
different types of dry bearing materials.
2. For two types of PTFE-based composites, and polyacetal, the relationships between wear rate and stress
are broadly the same as those obtained in some bearing
tests.
3. For one particular PTFE fibre/glass fibre composite,
the effects of counterface roughness, temperature, and
fluid contamination on wear are generally similar to
those encountered in bearings.
L
Water
!
L
I
I
I03
a
I
104
I
I
I
I
I
105
I
Cycles to wear 0 2 0 mm
Dry
Water
I
I
I05
CycLes to failure
102
b
......
Dry
I
I
Copyright © Controller HMSO, London, 1979
--K
DTD585
Water
I
I
I II
10-7
I
I
10-6
C
I
II
10-5
I
I
I
II
10-4
Weor rote mm 3 / Nm
1
0.10
005
/
/
/ i*
1
1 1
1. Atchatd J.F. Contact of rubbing surfaces. J.Appl. Phys. 24,
981-988 (1953)
2. Atchatd J.F. and Hirst W. The wear of metals under unlubricared conditions. Prec. Roy. Soc. (Lend) A236, 397-410
(1956)
3. Lancaster J.K. Geometrical effects on the wear of polymers and
carbons. Trans ASME, J.Lub.Tech, 97F, 187-1.94 (1975)
4. Weintraub M.H., Anderson A.E. and Gaeler R.L. Wear of phenolic
resin-asbestos friction materials. Advances in polymer friction and
Dry
E
E
References
wear, Polym. Sci. and Tech. 5B, 623-647, (ed. L-H Lee) Plenum
Press (1975)
DTD 5 8 5
5. Play D. and Godet M. Self-protection of high wear materials.
I
5
O
d
I
10
I
15
I
I
20
25
IO3 cycles
Fig 14 Effects o f fiuid contamination on wear o f PTFE
fibre composite, K
(a) Morton 3a. Spherical bearings, 19. 05 m m shaft dia;
oscillatory motion, +25 °, O.0025 m/s, counterface AIS1
440C, 0.05-0.10 larn Ra; reversing load
(b ) Rowland and Wyles ~°. Pad on reciprocating track," pad
size 36mm x 7.2mm, track FV520 CR steel 0.05-0.10 pm
Ra; speed = O.087 m/s, stroke length 0.244mm; 182 MPa
(c) Evans~ PTFE + 25% vol glass fibre on disc, 18/8
stainless steel 0.15 wn Ra. Load = 12N, speed--- 0.15 m/s
(d) A m p e p 24. Spherical bearings, oscillatory motion +25 °,
10 cycles/min; 182 MPa, unidirectional load
is concerned, preliminary experiments have already
shown 12 that the RLC test readily shows up marked
differences in wear behaviour of interwoven PTFE fibre/
glass fibre composite following slight technological variations in the process of impregnation by synthetic resins.
Whether or not the ability of the RLC test to differentiate
between the wear behaviour of different materials (Fig 4)
is directly relevant to the performance of these materials
in bearings, and if so, in which particular conditions of
sliding, still remains an open question which only further
work can resolve.
ASME/ASLE Lub. Conf., Preprint 77LC-5C-2 (1977)
6. BurwellJ.T. and Strang C.D. On the empirical law of adhesive
wear. J. Appl.Phys. 23, 18-28 (1952)
7. Mi'ehalonD., Gilbert F., Gonin A. and Caubet J.J. Contribution
a une tribometrie industrielle et pedagogique. Mere. Tech. du
CETIM, No 4, 72-95
8. Timoshenko S. and Goodier J.N. Theory of Elasticity. 2nd Ed.
McGraw Hill (1951)
9. Tabor D. The hardness of metals. Oxford (1951)
10. Rowland K.A. and Wyles S.A. Evaluation of dry bearing materials: linear track and large diameter journal bearings. BA C Rep.
PRO 251, (1973)
11. Bearings, plain, self-aligning,self-lubricating, low speed
oscillation. MIL-B-818200 (1974)
12.King R.B. Private communication
13. Lancaster J.K. Dry bearings: a survey of materials and factors
affecting their performance. Tribology 6,219-251 (1973)
14. Pratt G.C. Hastic-based bearings. In-Lubrication and Lubricants,
(ed. E.R. Braithewaite), Elsevier (196 7)
15. Anderson J.C. Private communication
16. Cheeseman K.J. Private communication
17. Lancaster J.K. Unpublished work
18.Delxin; Design handbook. DuPont (196 7)
19. Theberge J.E. Properties of internally-lubricated, glass-fortified
thermoplastics for gears and bearings. Prec. 1st Int. Conf. on
Solid Lub. ASLE SP-3, 166-184 (1971)
Conclusions
20.Clerico M. and Rosette S. Influence of roughness on wear of
thermoplastic on metals pairs; a preliminary analysis.Meccanica,
8,174-180 (1973)
21. Tanaka K. and Uchiyama Y. Friction, wear and surface melting
of crystalline polymers. Advances in polymer friction and wear,
A procedure has been devised for the accelerated wear
testing of thin-layer, dry bearing materials involving the
use of Hertzian line contact stresses. The total life to
failure is determined within hours, in contrast to the
22.Shen C. and Dumbleton J.H. The friction and wear behaviour
of polyoxymethylene in connection with joint replacement.
Wear 38,291-303 (1976)
74
TRIBOLOGY international April 1979
Polym. ScL and Tech. 5B, 499-530 (Ed. L.H. Lee), Plenum
Press (1975)
23.A guide on the design and selection of dry rubbing bearings.
E S D U Data item 76029 (1976)
24.Aerospace Design Manual. Ampep Ltd
25 .Wade D..L Private communication
26.Craig W.D. P T F E bearings for high loads and slow oscillation.
Lub.Eng. 18, 174-181 (1962)
27. Barrett L.D. Teflon fabric bearings in the helicopter rotor
system.ASME Design Eng. Tech. Conf. 1975. Reprint 75-DET.
125 (1975)
28.Allen A.J.G. Plastics as solid lubricants and bearings. Lub. Eng.
1 4 , 2 1 1 - 2 1 5 (1958)
29.Pastor M.W. and Tabor D. The friction and deformation of
polymers. Proc.Roy.Soc. (Lond) A235, 210-224 (1956)
30.Williams F.L High temperature airframe bearings and lubricants. Lub.Eng. 18, 30-38 (1962)
31. Evans D.C. Friction and wear properties of PTFE composites
at elevated temperatures. LMech.E. Tribology. Group Con-
vention, Swansea, Paper 1, 1978
32.Craig W.D. Operation of PTFE bearings in sea water. Lub.Eng.
20, 456-462 (1964)
33.Morton J.G. Private communication
34. Evans D.C. Unpublished work
35. Crook A.W. The elastic deformation of cylinders loaded in
line contact. AEI Report A471, Sept (1955)
36. Brewe D.E. and Hamtock R J . Simplified solution for elliptical
contact deformation between two elastic solids. Trans. ASME
J. Lub.Teeh 99F, 485-487 (1977)
37. Evans D.C. Polymer fluid interactions in relation to wear. Proe.
3rd Leeds-Lyon Symp. on Wear o f non-metallic materials (1978)
p47
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BEARING MATERIAL
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"stick-slip".
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TRIBOLOGY international April 1979
75