Dry bearings: a survey of materials and factors affecting their

Transcription

Dry bearings: a survey of
materials and factors affecting
their performance
J. K. Lancaster*
Following a general discussion of performance criteria and testing of dry bearings, the various
materials currently available commercially are classified into four general groups - polymers,
carbons-graphites, solid film lubricants and composites, and ceramics-cermets. The material
properties relevant to bearings are discussed in detail for each group, and the special features
required in design are noted. Finally, an attempt is made to develop a selection procedure for
dry bearings, based on identifying the major requirements in a given application.
Applications ofunlubricated
'dry' bearings have been
expanding rapidly in recent years, and there are three main
areas in which their use is indicated:
a where fluids are ineffective, as at low or high temperatures, or in reactive environments,
b where fluids cannot be tolerated because of the possibility of contamination of the product or the
environment,
c where fluids are undesirable because of lack of opportunity for. or the impossibility of, maintenance.
Cost, although always an important consideration, is not
usually the decisive reason for choosing a dry bearing. Some
of the most successful dry bearing fornmlations are significantly more expensive than their mass produced metallic
counterparts intended for lubricated service.
Most of the discussion in this survey is concerned with
sliding bearings, and for these there are two primary property requirements. Firstly. the materials must be able to
support an applied load in the environment concerned
without significant distortion, deformation or loss in
strength. Secondly, both the coefficient o f friction and the
rate of wear must be acceptably low and preferably also
insensitive to minor changes in the conditions of sliding
likely to be encountered, eg temperature, humidity,
contamination, etc. Four groups of materials satisfy all,
or most of these requirements. The largest group consists
of materials based on synthetic polymers to which are
added various fillers or reinforcements intended to enhance
particular properties. The second group, in order of general
usage, comprises carbons and graphites, together with additives, and these are particularly important for applications
at temperatures above those which can be tolerated by
most of the common polymer-based products. Thirdly,
for applications in which a low coefficient of friction is the
main requirement, solid film lubricants, based on ptfe or
lamellar solids such as graphite or MoS2, can be used in
* Materials Department, Procurement Executive, Ministry of
Defenee, Royal Aircraft Establishment, Farnborough, Hampshire,
England
conjunction with suitable metallic or non-metallic substrates. Finally, for high temperature applications (above
about 400°C) where the magnitude of the wear rate is
usually the major consideration, a variety of hard metals,
ceramics and cermets is available, either in bulk form, or
as coatings on a metallic substrate.
The properties of a variety of materials falling within
these four groups are discussed in detail in this survey and
their advantages and limitations for dry sliding bearings
are compared and contrasted. Some of these materials
can also be used as constituents of rolling element bearings
intended for operation in the absence of fluid lubricants
and this aspect is also examined. First, however, it is
pertinent to discuss some of the more general features
associated with the operation of dry bearings and, in
particular, the problem of wear.
Performance criteria
Strength
There is no single strength parameter which uniquely
defines the load-carrying capacity of a dry bearing. During
sliding, both tensile and compressive stresses are present
within the contact area, and shear stresses within the subsurface layers. A widely-used limit is one-third of the
maximum compressive stress, but this is relevant only to
those materials which exhibit similar tensile and compressive properties. For carbon/graphites and ceramics, the
ultimate tensile strengths may be very much lower than
the compressive strength. A further complication which
can arise with polymer-based materials is visco-elastic
behaviour: the dependence of mechanical properties on
time. The stress-strain relationships at any particular rate
of loading are non-linear and even under static loading
creep or permanent deformation may occur. Ptfe and
many of its composites are particularly prone to the
latter. Although carbons and graphites do not exhibit
visco-elasticity, the stress-strain relationships of these
materials also tend to be non-linear. It is therefore impossible, in general, to quote unique values of elastic moduli
for many of the materials of interest in dry bearings.
TRIBOLOGY December 1973
219
Fortunately, except in a few specialized applications such
as air-frame bearings, the stiffness o f a dry bearing
assembly is seldom a critical parameter. F o r design
purposes it is usually adequate to use tensile or flexurai
moduli obtained b y conventional test methods.
Wear
13y far tb.e greatest uncertainty associated with the operation o f dry bearings is wear. Eoth simple theory and
numerous experiments have shown that once the surface
conditions during sliding attain a steady-state the volume
of wear. v. is proportional to the distance of sliding d.
Further, if changes in the applied load, g', do not cause
significant changes in any other variable, and particularly
in the surface temperature, then the volume of wear per
unit sliding distance is directly proportional to the load.
Thus,
,, = k W d
(i)
The constant k is usual!y called the 'specific wear rate'
and is basically a material property. However, because the
two assumptions made during the derivation of Equation
1 are not always valid, the value o f k may depend on the
conditions o f sliding and on the precise geometrical
sliding arrangement used to determine wear. It is this
last complication which m a y give rise to uncertainties
when attempting to extrapolate the results of accelerated
laboratory wear tests to the practical case.
The specific wear rate, k = v/Wd, has the dimensions
of (stress) -1 , but it is more meaningful to use units whose
physical significance is more readily apparent - mm3/N m.
However, in the operation o f a journal bearing, it is not
usually the volumetric loss of material which is the important factor but the radial wear h, which leads to increased
Table 1 Conversion of wear rates and PV factors into Sm
units from British, American and non-SI metric units
in 3
Specific wear rate
10 - 1 2
10 -13
mm 3
= 1.2 X 10 . 8 - ft - lbf
Nm
in3 = 1.4 X 10 - 8 ram3
in - lbf
Nm
cm 3 ~ 1 0 - 8 m m 3
10-12 _ _
cm - kgf
Nm
Wear factor
10-10
10 - 8
PV factor
in3 _ rain
mm 3
-~ 1.9X 1 0 - 8 - ft - l b f - h
Nm
cm 3 - rain
mm 3
~ 1.7X 10 - 8 m - kgf- h
Nm
lbf
ft
MN m
1000 i ~ X min-- -- 0.035 ~ m2 X -s
kgf
m
MN
m
t00-X--~0.17
Xcm 2 rain
m-2
s
220
Stdtic l o a d - c a r r v i n c
D_
CD
O
d
Log V
Fig 1
Relationships between P and V for dry bearings
clearance. To a first approxhmation, h = via where A !s
the projected area of contact, and hence h = kPd where P
is the nominal applied pressure. If the assumption is now
made that changes in speed. V. do not cause changes in
other variabIes, and in particular the surface temperature.
then d = Vt and
h =kevt
(2)
The constant k = h/PVt is numerically the same as the
specific wear rate defined above, but when used in the
context o f radial wear is often referred to in US literature
in 3 mm
as a w e a r factor' and expressed in units
ff I b f - h
For convenience, some conversion factors between the
various units, Imperial. Metric and SI. are given in Table i.
PV factors
Expressing the wear rate of a bearing in terms o f radiai
wear per unit time, Equation 2 leads directly to the con°
cept o f a PV factor as a performance criterion for bearings.
PV factors are widely quoted in the literature and m a y
take two forms:
a the 'limiting PV' above which wear increases rapidly
either as a consequence of thermN effects or of stre>
sos approaching the elastic limit,
b the P V f a c t o r for continuous operation at some arbitrarily specified wear rate, eg 25 # m / 1 0 0 h
In neither case is the PV factor a unique criterion of performance because the assumptions made in the derivation
of Equations 1 and 2 (changes in P and V do not introduce
changes m any other variable) are only usually valid ave:
a very restricted range of P and V. In Fig i, curve (a) shows
a schematic relationship between P and V fo~ constant, wear
rate and it m a y be noted that there is an inverse propor
tionality, ie a c o n s t a n t P V factor, only over the central
range o f P and V. At low speeds, the m a x i m u m pressure
which can be used is limited b y the strength of the material
and as this pressure is approached the specific wear ra-Ee
no longer remains independent of load but begins to
increase. The surface temperature m a y also increase
and #lese two factors manifest themselves in Lhe P - V
curve for constant wear rate as an asymptotic approach
to a limiting P as V decreases. At high speeds, the generation of frictional heat again raises the temperature of the
surface layers and this also tends to increase the specific
wear rate. Consequently, there is an asymptotic approach
to a limiting V as P decreases. Similar considerations apply
to the 'limiting P V' relationships. The general shape of
the curve is the same, see Fig 1 curve (b) but is displayed
to higher values of P and V. The extent of this displacement depends on the particular material involved and the
way in which its specific wear rate depends upon temperature and stress.
Because of the general non-tinearity of P - V curves,
design information for dry bearing materials is best expressed by presenting the complete P - V relationship. However, data for many commercially available materials are
insufficient for this purpose and the only information often
available is a PV factor at one or two speeds. In the latter
circumstances, extrapolation of the data to other sliding
conditions can be somewhat uncertain.
Temperature
As already mentioned, the major complicating factor
responsible for the deviations of the P V relationship
from linearity, particularly at high speeds, is the temperature rise due to frictional heating. In view of the
importance of temperature, it is worth devoting some
attention at this stage to ways in which its magnitude can
be estimated. The temperature developed at the localized
asperity contacts involves two components:
Ta, the mean temperature of the interface between
the bearing and its journal resulting from the general
dissipation of frictional heat around the system. The
value of T a depends on the geometrical construction
and materials of the whole bearing assembly.
Tf, the 'flash' temperature rise at the localized asperity
contacts, which is largely independent of geometry.
Thus, T = Ta + Tf. Neglecting heat losses, the mean surface temperature Ta is proportional to the total energy
dissipated.
T~ = To + CuIt'V
(3)
materials whose thermal conductivities are very much less
than those of a mating steel journal, eg polymers,
Tf = 1 X 10 2 gp~2w1/2 V
for low apeeds, (4)
and
Tf = 5.7 X 10 -5 p.p~4w1/4V1/2
forhigh speeds.(5)
Pm is the flow pressure of the bearing material and is
approximately equal to the indentation hardness. W is the
absolute load applied. The two separate cases, low and high
speeds, arise because at low speeds the frictional heat is
shared between the two sliding components and time is
available to reach thermal equilibrium. At high speeds, the
thermal conditions are transient, and the majority of the
heat generated is conducted into the moving shaft. For
bearing materials whose thermal properties are similar to
those of a steel journal, eg carbons-graphites, tile values of
Tfat low speeds are about half those given by Equation 4,
t at high speeds they remain the same as those given
by Equation 5. As mentioned earlier when discussing the
mean surface temperature rise, the thermal conductivity
of tile journal is a much more important factor at high
speeds than that of the bearing material itself. Finally,
it should be noted that the temperatures derived are the
maximum possible, assuming that the whole of the applied
load is supported by a single contact. In practice, this
will not be so and there will be a distribution of lower
asperity temperatures depending on their number, size
and proportion of load supported by each.
It was also assumed in the derivation of Equations 4
and 5 that the hardness of the bearing material did not
vary with temperature. ~Nilst this is generally true for
carbon/graphites and ceramics and cermets, the hardness
of polymers varies appreciably with temperature. To a
first approximation, the variation is exponential
r m = Po e aT
where Po is the room temperature hardness, and a is a
material constant of the order of 0.005. The effect of
hardness variations will be most pronounced at higll
speeds and for this situation
Tf = T/oe 0.0038Tf
where 1o is the temperature of the environment, # is the
coefficient of friction and C is a constant characterizing
the thermal properties of the materials and the particular
geometrical configuration. Values of C usually lie within
the range 0.1 1°C s/N m, and a typical value for
25 m m × 25 mm polymer-based bearings sliding against
a steel journal is about 0.5°C s/N m. In general, an
increase in the thermal conductivity of the journal material
causes a much greater reduction in the value of C than a
corresponding increase in the conductivity of the bearing
itself.
For any geometrical bearing arrangement, the mean
surface temperature, Ta, is easily measurable with a surface
thermocouple. The flash temperatures at the asperity
contacts, however, cannot be measured in this way and
it is only possible to make theoretical estimates. A convenient method is to apply Jaeger's 1 analysis for two semiinfinite bodies in relative motion making contact over a
small area of square cross-section. Simplification of the
relevant fommlae has been discussed elsewhere 2 and
it is sufficient here merely to quote the end results. For
where Tfo is the temperature derived from Equation 5,
assuming no change in hardness with temperature. In
summary, Fig 2 gives curves for T¢'/I~ against W 1/2 V for
materials of different hardnesses. The full lines are for
polymers whose hardness varies with temperature and
whose thermal conductivities are negligible in comparison
with the steel counter-face. The dotted lines are for
carbons-graphites whose hardness is constant and whose
thermal conductivities are comparable with those of the
steel journal, and the hatched lines are for ceramics and
cermets, again with constant hardness, but with negligible
thermal conductivities compared to steel. It may be noted
that despite the higher thermal conductivities of carbonsgraphites, the values of Tf at high W 1/2 V exceed those for
polymers of similar hardness because thermal softening
does not occur.
Comparison of the expressions for the mean temperature rise, Equation 3, and the flash temperature rise,
Equations 4 and 5, leads to the general conclusion that
at high speeds and low loads, the temperature limit of
221
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C ....
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cerrnets
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[0 -3
I / ~ z l
[0 .2
PI
l
IO -I
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/5o0
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...~
. " /
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5~
interest, and this can be measured mos~ conveniently by
monitoring the relative movement of the bearing 'Mth
respect to a fixed datum. The housing or base-plate supporting the rolling-element bearings of the shaft o h e n
fulfils this function. For comparative purposes, it is aIso
desirable :o compute the specific wear ra~e e ~ J l e matefiai.
With thrust-beariv~gs there is no oroblem, but with jour~,~a]
bearings the depth of wear does r.,ot become directly pro°
portionat ~o the volume until the width of the wear scar
becomes equal to the bear/ng diameter~ Until this stage
is reached, the volume of wear, v. is given by:
I II
1
10 2
W½V(N½m/s}
Fig2 Flash temperatures during sliding on steel. Fulllines,
polymers; dotted lines, carbons-graphites; chain lines,
ceramics-cermets. FigUres adiacent to each line are hardnesses, VPN
a bearing is mainiy the result of the flash temperature,
whereas at low speeds and high loads the mean temperature rise may become more significant. Clearly, therefore,
the temperature conditions attained during the operation
of a particu]ar size of bearing are dependent upon the
individual values o f P and V and not direct1?} on the PV
product. It is also apparent that increasing the size of a
bearing for constant P and V will again affect the temperature conditions because of the changes in absolute load
involved in maintaining constant pressure.
Testing
The most satisfactory way in which to assess the performance o f any dry bearing material is undoubtedly a practical
trial in the intended application. However this is often
impractical for reasons o f time, cost or ]ack o f opportunity,
arid in such circumstances it is necessary to resort to
laboratory testing. A variety o f simplified bearing test
rigs has been developed, and two examples are shown in
Fig 3. Ideally, all test rigs should be able to m o n i t o r
continuously the frictional torque, wear and temperature,
and should also be capable of operating over a wide range
of loads, speeds and ambient temperatures. As yet there
has been no agreement on a standard size for dry bearing
tests and various sizes between 12.5 and 25 m m bore are
used b y manufacturers of different products. P.ecause
of the complicating effects o f temperature already described,
it is not always valid to compare the reported performance of bearings of different types and sizes.
Determination of the limiting P V curve for a partic'alar
material is a relatively simple matter. A constant velocity
is chosen and the ioad on the bearing is then increased
ir~ stages, sufficient time being allowed at each stage for
the friction and temperature to stabilize. Fig 4 shows the
results of a typical experiment of this type. At high loads
it is no longer possible for the friction or temperature to
stab~ize, and the load corresponding to the last region of
stab~ity is taken as the limiting value. Repetition o f the
whole procedure at different speeds enables the limiting
P V curve to be built up.
The determination of P - V relationships corresponding
to a specified wear rate is much more time-consuming as
it involves measurements o f the rate of wear over a wide
range o f loads and speeds. As already mentioned, the
depth or radial wear rate of a dry bearing is of particular
222
TRJBOLOGY December 1973
where D is the shaft diameter, l is the bearing length, h is
the maximum de~th of wear and e is the il~itial clearance.
Values of the fn (h/e) are shown in Fig 5. Measurements
of the depth of wear are sometimes complicated by a
deformation componen'; which may be significant for the
more elastic materials in the early stages o f sliding when linecontact conditions prevail. Deformation wil! not, however.
normaity enter into the determination o f the steadyosta~e
wear rate.
Because wear-testing of bearings can be very rimeconsuming, a great deal of the available data on the wear
of dry bearing materials has been obtained from accelerated
laboratory tests with simpler geometrical configurations.
In general, these take the form of a smal] rider c f the
bearing material loaded against a larger, moving, metal
surface. The essential feature of this type of test is the
non-conforming geometry, and a number of different
T a N e 2 Retationshipsfordetermimngwearvo!umesin
accelerated wear tests of the pin-ring and pM-disc type
Plane-ended
cylinder
on ring
v~
[ d2
6R
-
a : cylinder radius
R = ring radius
d = scar width
~r
Crossed
cylinders
3a2 "\~
d3 1
v ~
/~-
64R J R
.[
d~
i
d2 [
1+--16 ~ b k 7 -
a = small cylinder radius
R : large cylinder radius
d = long dimension of elliptical scar
7i"
Cone on
V~
d 3 cot 0
24
disc
0 = semi-apex angle of cone
d = scar diameter
7/
Sphere on
disc
V~
d4
+ - -
64R
R = s~here radius
d = scar diameter
d2)
12 R ~
o]
Fig3
a Journal bearing test r i g - DowCorningLFW5
b Thrust bearing test rig - Dow Coming LFW6
load, speed, temperature, environment, counter-face inetal
and counter-face roughness. The major differences in test
conditions arise as a consequence of the non-conforming
geometry which is deliberately used to increase the sensitivity of the wear measurements. With non-conforming
geometry, potentially abrasive wear debris can readily
escape from the contact zone without adverse effects,
whereas in a bearing the debris may contribute significantly
to increased wear. Further, the contact stress for a given
load is often greater in non-conforming geometry than in
a bearing, and may thus lead to material breakdown in the
former but not in the latter. These effects will be discussed
in more detail in the various sections devoted to particular
groups of materials. It is worth noting here, however that
despite these differences there can nevertheless be a very
specimen configurations has been used. A selection is
shown in Fig 6. The end of the rider may be shaped to
enable the volume of wear to be computed from the dimensions of the wear scar developed, and fornmlae for some of
the more common arrangements are given in Table 2.
It should be mentioned that determinations of the wear
volume from the dinnensions o1 wear scars is accurate only
if the elastic modulus of the material being worn is sufficiently high to ensure that elastic recovery is negligible
following removal of the load. The limiting modulus is
about 0.2 GN/m 2. and below this value it becomes necessary for wear measurements to be made from weight losses.
Accelerated wear tests are often considered to be an
acceptable alternative to full-scale bearing testing in the
early stages o f materials selection or a development programme. However, there is still a considerable amount of
controversy about the value of accelerated tests in predicting the wear behaviour o f bearings in practical conditions.
In an accelerated test there is little difficulty involved in
simulating the practical operating parameters of absolute
200
%
Pressure
150
,oo~~,
~
1-
50
,
\
Failure
Temperature
Friction
___L
I
1
2
3
4
5
Time (h)
Fig4 Typical variation of friction and temperature during
the determination of 'limiting PV' for a polymer-based
bearing (data from Willis;)
223
is too great for accurate predictions of bearing wear to be
made from the accelerated tests, the correlatior_ is sad
ficiently high to enable the most promising materials to
be isolated.
!-O
O-8~
Z~
t-
(I- ¼hI
O.Oi
O.2![
o
i
2
;
4
h
Fig 5 Values of fn(h/e) for calculating the volume of wear
of a journal bearing from depth measurements:
V = Olh fn(h/e}
reasonable correlation between accelerated wear measurements and those on bearings. Fig 7 shows a comparison
obtained for a series of polymers containing carbon fibres
as reinforcement. Bearing wear measurements were obtained
over a period of 500 h in the conditions given. The bearings were then removed from their housings and wear
rates determined on their outer surfaces during sliding
against mild steel in a crossed-cylinders geometry, see
Fig 6a, for a period of about 12 h. Whilst the 'scatter
Polymer-based materiNs
Of the very large number of polymers available commercially, oniy a stuN] proportion is in signigcant use for dr-f
bearing applications. These are listed in Tab]e 3.
Nate~4als may conveniently be divided into XhermoptasticF,
which soften appreciably or melt at a characteristic cemperamre, and 'thermosets" wl-~ich cross-l~k under the
influence of heat and do not subsequently melt. There
are two main effects of tempera-rare on mechanical properties. Firstly, there is the gener~ reduction in strengtb
and stiffness which results mainly from. a weakening of
the interactions between the polymer chains. The temperarare tLrr,.it from this cause is conveniently specified in
terms of a 'heat-distortion temperature' at which tl~e
deflection of a loaded cantilever beam reaches an arbit rarily defined value (ASTM-D-648). Secondiy~ s~rength
changes also occur with it±creasing temperature as a
result of chemical reactions within the polymer itself,
or between the polymer and the e n w r o n m e ~ , usuatly
oxygen° [t is thus possible to identify both a °therm±
stability' and an 'oxidatwe stabilityL However, the quota-
W
w
b
n
W
~L~v
v
i
\
,
\
\
\
\
I
@
~2
W
CO2v'
w
]
@V2
i
n
I
C
[
E:ZE3
R
d
Fig 6 Accelerated wear test arrangements for polymer-based materials: a 'Pin and ring', b 'Pin and disc', c 3-body abrasive
wear, d 2-body abrasive wear (Taber abraser)
224
IO-5
tion of precise values of temperature corresponding to
these limits is generally precluded because the extent of
chemical reaction is time, as well as temperature, dependent. This may be illustrated by Fig 8 which shows the
temperatures for different times of exposure at which
various groups of polymers retain 50% of their room
temperature strength. It is important to note that there
appears to be no polymer currently available commercially
which ret~:ins 5(F/c of its room temperature strength for
more than ~O 000 h at temperatures above 300°C.
To put the thermal and mechanical properties of polymers as a general class of materials into perspective, it is
useful to quote order-of-magnitude comparisons of some
properties with respect to mild steel. The tensile strengths
of polymers are lower by a factor of about 10, their elastic
moduli and thermal conductivities are lower by a factor
of about 100, and their thermal expansion coefficients
are greater by a factor of about 10. These deficiencies
can be partially offset by suitable choice of fillers or
reinforcing fibres, and a list of the more widely used types
is also given in Table 3. Before describing the properties
of individual groups of polymers and polymer-based composites, a few general comments on friction and wear may
be helpful.
o
_-g
oZ
o
iO-6 -
t-co
o
°
o
t_
o
13~
iO-7-
o
o
o
o
o
o
o
o
o
o
o
/
I
iO-8
iO-7
I
iCr•
I
io-S
iO-4
~WGor rate in pin and ring tests(mnna/Nrn)
Correlation between pin and ring and journal bearing
tests for carbon fibre reinforced polymers sliding on m i l d
steel. Each p o i n t represents a different material
Bearing tests: load = 8.9 N (P = 0.05 M N / m m 2 ) , V =
0.65 m/s
counterface - cast iron, 0.3 # m cla
Pin and ring tests: load = 11.7 N (P variable), V = 0.54 m/s
counterface - mild steel, 0.18/Jm cla
Fig 7
Friction
It is often assumed that, in comparison to metals, tile
coefficients of friction of polymers are relatively low, but
this is not generally correct. Because most polymers
exhibit visco-elastic behaviour, the magnitude of the
coefficient of friction involves an appreciable component
arising from deformation and elastic hysteresis loss, and
this component varies markedly with the conditions of
sliding, and particularly with speed. Thus, although the
coefficient of friction of ptfe is commonly quoted as
0.05- O. 1, these levels are obtained only at very heavy
loads and low speeds of sliding, or when tile ptfe is present
as a thin film on a harder substrate. At high speeds or
light loads, ptfe sliding on metals or against itself may
exhibit friction coefficients as high as 0.3. Similar, though
less marked, effects occur with other polymers. The
addition of fillers to polymers also affects the coefficient
of friction. If the filler is a solid lubricant, particles may
transfer to the counter-face, establish a lubricating film
and reduce friction..alternatively with hard, rigid fillers
Table 3
o
such as glass, mica or asbestos, the load may be preferentially supported by the filler particles and the coefficient
of friction is then largely characteristic of the Idler/counterface interactions. For all these reasons, values of the coefficient of friction quoted in the literature, or, in fact,
even in this section, should be regarded as order-ofmagnitude values only: there is no such thing as a unique
coefficient of friction for a given material.
Wear
A great deal of the information on the wear properties of
polymers and polymer-based composites has been derived
from accelerated wear tests of the type already shown in
Fig 6. The main objective of these tests has been to determine how wear depends upon the conditions of sliding,
such as load, speed, temperature, etc, and it is reasonable
to assume that the trends observed will be relevant to
bearings. As discussed earlier, however, the absolute magnitudes of the wear rates obtained from accelerated tests
Main plastics and fillers of interest for bearings
Polymers
Fillers and reinforcements
Thermosetting
resins
To improve
mechanical
properties
To
reduce
friction
To improve
thermal
properties
Polyethylene, high molecular
phenolics
asbestos
graphite
bronze
weight
Acetal-homo-, and co-polymer
polyesters
glass
MoS2
silver
epoxies
ptfe
(particles or fibre)
carbon/graphite.
Thermoplastics
Polyamides (Nylon 6,
6.6 and 11)
silicones
carbon
textile fibres
Ptfe
polyimides
mica
Polyphenylene oxide
metals and oxides
Polycarbonate
T R I B O L O G Y December 1973
225
~A
O
A
B
E
C
F-
Oq
I
I0
[O2
tO 3
IO 4
IC
Time {h]
Fig 8
50% of
Area A
Area B
Area C
Area D
Temperature-time relationships for the retention of
mechanical properties.
Polyamide-imide, pNyimide
Silicone, polyphenyiene, polybenzimidazo!e
Epoxy, polyester, phenol-formaldehyde, ptfe
Melamine-formaldehyde, fep, polyphenyJene oxide,
polysulfone (data from Ref 4)
may not always be sufficiently reliable to be used for the
prediction o f wear in bearings.
In the early stages of sliding of a polymer composite
over a freshly prepared metal counter-face, the wear rate
is relatively high. However, during subsequent traversals
over the same track on the counter-face, the metat surface
can_ be modified by transfer o f polymer or Nler, abrasion
by the filler, or abrasion by contaminants from the surrounding environment. In most cases these modifications lead
to a smoother counter-face surface and the wear rate then
gradually decreases to a constant value. If the initiN
counter-face roughness is increased, the magnitude of the
initial wear rate increases markedly, but that of the final
wear rate is less affected, because of counter-face modification by transfer or abrasion. These effects are illustrated
schematically in Fig 9. In service conditions, the wear of
bearings is sometimes estimated from single measurements
at the end of a given period of operation, and the dotted
lines in Fig 9 show how wear rates derived from such single
measurements may be greatly in error. At least two wear
determinations at different times are required to define
the approximate shape of the:_wear volume-time relationship.
To a very rough approximation, the initial wear rate of
a polymer is proportional to the centre-line-average surface
roughness of the counter-face raised to a power of 2 - 3 .
The wear occurring in this regime is abrasive in type and
similar to that caused by hard abrasive particles. The
metal asperities penetrate the softer polymer and produce
wear debris by shear or micro-cutting, or by low cycle
fatigue. For abrasive wear, there is an inverse proportionality between the wear rate of po!ymers and the product
so, the breaking strength times the elongation to break.
This result therefore accords with the general observation
that the abrasive wear resistance of rubbers is appreciably
greater than that o f the more rigid polymers. V,h e n fillers
or reinforcing fibres are added to a polymer, the strength
increases, typically by a factor of 1.5-4, but the elongation to break may decrease by a factor of up to t 00.
Following reinforcement of a polymer therefore, the
product s e is frequently reduced, and this, in turn, leads
to a reduction in its resistance to abrasive wear.
226
As mentioned earlier during repeated traversals of a
counter-face beneath a polymer coro_posite~ the meta~ sur~
face may be modified by transfer or by abrasion. When
steady state conditions have been reached, the wear process
then becomes prLmarily one of fatigue on a locNized
scale, the number of cycles to failure depending ".upon the
1ocaiized s~resses occumng, wMch. in turn depend on the
t o p o g a p h y generated on the counteroface. Transfer from
polymers with high elongations, such as pt.',~e, acetals or
nylons, leads to smoother counter-face surfaces and Lence
to lower rates o f wear. However~ transfer from brittle
polymers such as polyesters, some epoxies or polystyrene.
is often in the form o f irregular lumps: the effective surface
roughness of the counter-face may then increase during
sliding, together with the rate of weal Solid iubricast
fitlers, such as ptfe, MoS2, etc. which are intended ~.o re4-ace
friction, also contribute to the formation of relative!}
smooth transfer films and so assist in reducing wear. tn
general, ptfe ~s more effective in this respect than graphite.
which in turn is more effective than MoS2. ~n some poiyo
mer composites, notabiy those filled witb~ carboa/gra?hite
particles or fibres, transfer film formation, an4~hence the
magnitude of the wear ra~e, is very susceptible ;o fluid
contamination. As an extreme examp!e, FN i0 shows the
wear volume-th'ne relationship for a carbon fibre-frilled
epoxy resin sliding on stainless steeI. In the early stages
of s!iding a transfer film deve!ops on the counter-face
and the wear rate gradually decreases to a 1~iting value.
When water is added to the system, however, the transfer
film is removed almost immediately and the orlginai recta1
surface is exposed, leading to. an increase in wear rate
by a factor of 500. Composites containing ptfe show
//
f/
f
(3
i
~6
E
20
>
. . . - -- .-
j
///
/
@1
Initial
wear
//./"
ng go%i2e
roughness
i-
1
Steady-state wear
T~m¢
Fig 9 Typicamwear volume-time re~ationsNps for po!yme~
based bearings on journais of increasing roughness
Factor increase in wear rate if journal roughness is doubled
~nitiai
Steady°
state
No transfer.film
No fil!er or non-abrasive filler
X4-X6
X4.-X5
Transfer fimm formation
No filler or non-abrasive filler
X4--X6
X2
Xt-Xt.6
X I - X ~ o4
Moderately abrasive f[Her
(reiative to shaft)
tive figures for the abrasiveness of different fillers, based
on an arbitrary test in which a bronze ball is oscillated
over a filled polymer surface. When the filler is very
abrasive, such as glass, the counter-face may be seriously
damaged leading to an increased roughness, and in turn,
an increased wear rate of tile composite. It is wellestablished that glass-filled polymers should not be used
as bearings against a relatively soft shaft material such as
aluminium alloy or bronze. Mildly abrasive fillers such
as graphite, or high modulus carbon fibres merely remove
the peaks of the cotinter-face asperities, reduce the localized
stresses, and in turn the wear rate of the composite. The
effect of fillers in modifying the counter-face surface
may influence the wear properties of a composite far
more than the improved mechanical properties of the
filled material.
0"4
gEO.3
E
"6 o . 2
E
-v/
Water
added
0.1
I
0
02
I
I
04
0-6
A
/~.
1
V
2.8
3.0
i0 5 rev
Fig 10 Wear volume-time relationship for an epoxy-carbon
fibre composite on stainless steel, showing the effect of
removing the transfer film by water additions
similar effects, though to a less marked extent as ptfe
can transfer to metals even in the presence of water. This
susceptibility of transfer film formation to fluid contamination may well account for the poor performance of
many dry-bearing materials in wet, or alternate wet and
dry conditions.
It is not generally appreciated that most of the fillers
and reinforcing fibres added to polymers are abrasive
towards metals. This applies even to lamellar solid lubricant fillers, such as graphite and MoS2, where the abrasiveness is partly intrinsic, resulting from their crystal structure,
and partly a consequence of the inevitable impurity content in these materials. Mthough the abrasiveness of
lamellar solid fillers is very slight, it is nevertheless sufficient
to influence the wear process. Table 4 shows some comparaTable 4
fillers
Relative abrasiveness of plastics with various
Composite
Abrasiveness
wear rate of
bronze ball
(10 7 mm3/N m)
Ptfe
+ 30%
+ 25%
+ 25%
+ 30%
+ 25%
+ 25%
+ 40%
+ 33%
<0.1
620
100
80
31
17
2
0.8
0.5
glass fibre
asbestos fibre
carbon fibre (high strength)
mica
coke
carbon fibre (high modulus)
bronze
graphite
Phenol-formaldehyde
Mineral filled
Asbestos paper filled
Wood filled
Paper filled
Cotton cloth reinforced
730
75
17
2.6
1.8
Ptfe and p tfe-based composites
The major disadvantages of unfilled ptfe as a bearing
material are its poor mechanical properties and a marked
tendency to cold-flow under load. Its use in the unfilled
form is therefore restricted to a few applications at very
light loads, or as thin films bonded to a harder substrate.
The mechanical and wear properties of ptfe improve
dramatically with the addition of fillers, and a very wide
variety of compositions is available; one manufacturer
lists 49. The properties of a selection of the more popular
are given in Table 5. A particularly widely used composite
is one containing bronze, graphite and lead, and one
reason for the success of tiffs material appears to be the
facility with which it can form a transfer film on a steel
cotinter-face. Unfortunately, fillers which are most effective in reducing deformation under load are not necessarily
file most effective in reducing wear. The deformation under
load for mica-filled ptfe, for example is only about 1%
and thus very much lower than that of most other filled
ptfe formulations. However, the specific wear rate of
this composite, ~ 7 0 X 10 7 mm3/Nm is much higher
than that of the glass, bronze or graphite-filled materials,
see Table 5. The ultimate selection of a particular filled
composite for a specific application is therefore almost
always a compromise.
Most filled ptfe compositions are available either as
moulding powders or as bar-stock from which bearings
can be machined. Ptfe cannot be injection-moulded and
the usual production technique is compression moulding
followed by free-sintering in a vacuum or inert gas
atmosphere. Material containing small proportions of
fillers can be extruded. An alternative, but more expensive, method for composite preparation is hot compression
moulding followed by slow cooling under pressure. There
is some evidence to suggest that materials prepared in
this way exhibit superior mechanical and wear properties
to those produced by free-sintering. Even when filled or
reinforced, however, tile mechanical properties of ptfe
composites still remain relatively poor in comparison to
those of other filled polymers, see Tables 8 and 9. This
situation can be improved by using ptfe fornmlations in
thin layers attached to a harder backing. The most common of these materials, and one of the most successful,
is produced by incorporating a ptfe-Pb mixture into a
layer of porous bronze sintered on to a steel backing.
The maximum load-carrying-capacity of this type of
bearing is of the order of 150 MN/m 2 and the bronze
content greatly increases the thermal conductivity. This
bearing material has been extensively described and tested,
227
Table 6 Changesin iife of porous bronze-ptfe-Pb bushes
produced by Ntering sliding conditions, Relative to mlid
steet shafts of 0.4 #m cta roughness, and continuous operation at 20°C
(Data from Pratt s)
Bronze gridceramic filled #
Factor
c~.ange
Sliding conditions
i
:o.,_
5i
OO ~ ~
[~"
Y_
/
, S
~
1
Shaft roughness
~ Porous
bF6nze-ptfePb bearings
A
Shaft material
,# decreased to 0.2 #zm
increased to 0.6 g m
X2
t
X½
I stainless steel
I
chromium plate
anodized aluminium
×2
Temperature increased to
f 120°C
1 200°C
Intermittent operation at high PVs
Time
Fig 11 Typical variations of depth of wear with time for
a Porous bronze - ptfe - Pb bearings
b Bronze grid - ceramic filled ptfe bearings
and exhibits a characteristic variation of wear with time
as shown in Fig 11. The initial period of rapid wear is
conffmed to an overlay of ptfe approximately 12 #m thick,
and bronze begins to appear at the sliding surface as soon
as this overtay is removed. There is then a period of low
wear at a constant rate until, after removal of a further
depth of about 50/~m, the bronze content of the surface
increases to an unacceptably high level and the wear rate
begins to increase rapidly. Because of this characteristic
pattern of behaviour these bearings possess a well-defined
!ife, and the relationship between life and P V factor is
shown in Fig 12. The approximate way in which the life
X~A
x ~/_,~
×2
is affected by various changes in the sliding conditions is
given in Table 6.
Bearings with essentiaily the same ingredients, but of
different construction are also available. Ceramic-fil!ed
ptfe tape can be pressure-rolled into a bronze grid to a
depth of about three-quarters of the thickness of the grid.
and the layer is then cemented on to a metal backing.
This bearing exhibits a similar characteristic wear-time
relationship to the porous bronze-ptfe-Pb type, but with
greater 'bedding-m wear' and a somewhat higher steady.state wear rate, see Fig 11, cu.we B. To compensate in
part for the latter, however, We wear rate remains constant
until a depth of about 200 #m has been removed. It has
been known for some time that ptfe in fibrous form exhibits
much greater strength than in bulk, possibty because of a
greater degree of chain orientation. Severn bearing con-
Table 5 Properties of ptfe and filled ptfe composites
F ler,
Properties
~
Specific gravity
Tensile strength
MN/m 2
Elongation
%
Ftexural modulus
GN/m 2
Deformation under %
load at 25°C and
14 MN/m 2
Expansion coef10-5/°C
ficient 2 5 100°C
Thermal conducW/m°C
tivity
Specific wear rate
10 -7 mm3/N m
Friction coefficient
on steel at
0.01 m/s
[ 0.05 m/s
Limiting y~0.5 m/s MN/m z X m/s
P V at
t 5 m/s
228
glass
12½% weight
glass
12½% weight 15% weight
MoS 2
graphite
20% weight
carbon
5% weight
graphite
55~ weight
bronze
5% weight
MoS 2
2.19
t7.5
300
1.1
11
2.3
13
230
1.1
4
2.12
9.5
130
1.4
8.1
2.1
11.6
70
1.2
2°9
3.9
i3.0
90
15
4.6
12.1
11
12.5
8.4
10.1
None
2.2
9
400
0.6
Continuous
flow
17
0.43
0.5 t
0.45
0.44
0.72
@.25
1.4
0.09
1.2
0.09
6.8
0.12
L2
0.12
1 o0
0.13
0.33
0.39
0.5
0.5
0.35
0.53
0.44
0.04
0.62
0.62
0.6
0.95
0.53
0.42
0.44
0.44
0.06
0,09
4000
0.!
view of the earlier remarks about the abrasiveness of fillers
on the counter-face, it is important to note that the provision of a smooth counter-face is much more important
when the fillers are non-abrasive. This is illustrated in
Fig 13 which compares the wear-time relationships of
ptfe/carbon fibre (high modulus) with those of ptfe/glass
fibre on mild steel counter-faces of different roughnesses.
In the case of the ptfe/glass, the composite is sufficiently
abrasive to smooth tile steel surface and the steady state
wear rate is virtually independent of counter-face roughness. Increasing the counter-face hardness, however, will
tend to offset this smoothing effect, and careful attention
should always be paid to minimizing the roughness of very
hard counter-faces for all types of composites.
In summary, Fig 14 gives the P V relationship corresponding to a wear rate of 25 #m/100 h for many of
the ptfe-based formulations discussed.
5OOC
2000
"Z
..J
I000
o
500
o
200
IOC
5C
20
OI
l
0.2
I
I
0.5
I
PV(MN/m2x
I
l
2
5
m/s}
Fig 12 Life - PV relationship for porous bronze-ptfe-Pb
bushes on mild steel shafts. 15.8 mm bore, 19 mm long.
Speeds 0.62, 1.25 and 2.5 m/s (750, 1500 and 300 rpm)
(from Pratt s )
structions utilize ptfe fibre. In one, fibre/phenolic resin
mixtures are attached to a woven polyamide backing,
and in turn to metal strip. Another version involves ptfe
fibre interwoven with glass fibre in such a way that the
rubbing surface is largely ptfe and the opposite surface is
mainly glass. The fabric is then cemented with a phenolic
resin to a metal substrate. This particular type of bearing
is capable of supporting extremely high loads (up to about
350 MN/m 2) at low speeds and is particularly suitable for
applications involving oscillating motion. The wear-time
curves are of similar shape to those in Fig 11 but a unique
feature is that at light or moderate loads several distinct
'plateaus' may occur in sequence, separated by short
periods of more rapid wear. Performance tends to deteriorate rapidly at high speeds~ partly because of the low
thermal conductivity.
For less arduous applications, a variety of filled ptfe
liners and tapes is marketed. Some ceramic-filled tapes
are provided with a self-adhesive backing, and graphitefilled materials are available as liners already cemented to
steel or reinforced phenolic resin backings. Separate
liners. 0.4 0.8 mm thick have the advantage that they can
be allowed to 'float', which facilitates heat-dissipation
because some heat is then liberated directly at the housingliner interface.
The counter-face roughness generally recommended
for use with ptfe composite bearing materials is within the
range 0.2 0.4/xm cla. Improvements to the shaft roughness usually reduce the wear rate, particularly in the early
stages of sliding as already mentioned. However, there
is some evidence to suggest that there are few advantages
to be gained by bnproving surface finish beyond about
0.05/~m cla. In any case, manufacturing to roughnesses
below this level can become inordinately expensive. In
Other thermoplastics
The particular advantage of thermoplastics over thermosetting resins is that they can generally be injection
moulded: mass-produced bearings are thus potentially
extremely cheap. The most widely used thermoplastic
bearing materials are polyacetals and polyamides (nylons)
and different types of each exist. Polyacetals may be
either a homopolymer of formaldehyde or a copolymer
of formaldehyde and acetaldehyde, but the differences
in properties are slight. At room temperature, the homopolymer is about 10% stronger and stiffer than the cope
lymer, but the position is reversed when the temperature
exceeds about 100°C. Some typical properties of the
homopolymer are given in Table 7. The polyamides are
a complex family: nylons 6 and 11 are self-condensed
amino acids, the number corresponding to the number of
carbon atoms in the parent acid, and nylons 6.6 and
6.10 are reaction products of hexamethylene diamine
(with 6 carbon atoms) and dibasic acids with 6 and 10
carbon atoms respectively. Nylon 6.6 is tile most widely
used of all types, and has the higl, est strength and stiff
ness. Some properties are given in Table 7. Its main
disadvantage is tile relatively high moisture absorption,
3
2
A
B
c
'S
b
I
2
3
4
5
I
2
3 4
S
IO5 revolutions of disc
Fig 13 Wear volume-time relationships on mild steel
a. 0 . 0 4 / l m cla, b. 0 . 1 8 # m cla, c. 0.5/*m cla
A ptfe-carbon fibre (high modulus) - transfer to steel
B ptfe-glass fibre. Abrasion of steel
229
I02,~.
\-~.
Wnvcn nff~/nln~,c, fibre'
I
E
Z
Q.,.
o-ol
04
{
IO
V(rn/s)
Fig 14
P-V retationships for ptfe-based bearings at a depth
wear rate of 25 # m / l O 0 h (mainly from Ref 6)
and consequent low dimensional stability, and nylons
6.!0 and 11 are better in this respect (24 h cnoisture
absorption ~0.4%). The thermal expansion coefficient
of nylon i i, however, (~18 X 10-5°C 1) is about twice
that of the other types. Very large bearings have been
made from cast nylon monomers which are subsequently
polymerized in situ and the properties of this materiai
are very similar to those of nylon 6.6. gearings fon-ned
from sintered nylon powder of high molecular weight
and crystallinity are also available and appear to show
some advantages from the wear resistance standpoint.
By controlling the sintering conditions, the degree of
porosity can be increased and the material impregnated
with oil for use as an alternative to oil-filled porous
TaMe 7
metals or for rolling-element bearing cages. £{.thoag£
nylon 6.6 has a higher melting point and room~.;ernpe<ature:
tensile strength than acetai, the s,rength decreases much
more rapidly with temperature, as shown by compam~g
the values for the heat-distortion temperatures m 7abie Where frictional heating is likely tc be appreciabie aceta
is therefore a preferred choice
Some properties of ? olycarbonate are a.tso gl,reu it:,'
Table 7, and the main claim tc distinction of this po_~y.
mar is its extremely high Lmoact strength Poiycarbonatc
is therefore a potentially val~able material for abrasive
or erosive wear situations, although its bearing properues
in clean environments are nor remarkab]e The outstand.ing temperature stabiIity of me polyimides has a!ready
been noted in Fig 8. These polymers are a~ain availao!e
in several different types, some of which are p:-oeessed
as thermoplastics by hot-comoression moulding and some
as thermosetting resins. Table 7 includes data for fi~e sore>
pression moulded type. [n gener£ the rr'_,ecaanicai ~roparties are similar at room ten~ perature to d?ose of nv!ons
and acetats, but the heat distortion ~empera~ure is ,~crv
much higher, and because of ti~is, tt~e !irnitm.g PV ~ac :ors
are also higher. Whilst the specific wea, rare of poiyo
imide at room temperature is not particuiar!y ;.ow. > e r e
is iitt!e or no increase with temperature until abo~t ?CO°C.
It has aiso been reported that ttse wear rate (n n,_t:
. . . .Oh,A1
.
is much lower than in air. by a Nctor of aboux ~O0.
As with ptfe, ti~e mechanical properties of other ;:herrno
plastics are also improved by the addition of fillers o~
reinforcing fibres. Some typical values for glass :fibre
reinforced and solid-lubricant filied cornposr~es are givers
in Table 8. Provided that the filler contents renta~ be}ow
about 20% by volume, there ~s no serious rmsedimem ~s
injection moulding. Some of the improverr~en[s obtained
in the friction and wear properties by ? fie/glass or -~tfe
additions can be very large, eg with polycarbonate ar~d
nylon 6.6. With acetaL however, ptfe/giass causes a
significant increase in tLe specific wear rate and ptfe
alone is a more effective addition, despite the slight impair
ment of the mechanical properties. The wear grepert~es
of polyimide are improved by graphite additions° our
MoS2 is ineffective in air and on!y becomes a useful
lubricant for applications in ssace or ultraoh~gh vacuum~
Additions of si~nail amounts of MoS2 ,.2 5%)i to nylon 6.6
Properties of some unfilled thermoplastics
Potycarbonate
Acetal
(homopolymer)
1.2
65
2.34
225
130
1 o4
] ,i
69
2,8
175
125
79
2o9
260
70
Polymer
Properties
Specific gravity
Tensile strength
Fiexural modulus
Softening/melting point
Heat distortion temperature at
1.8 MN/m 2
[zod notched impact strength
Moisture absorption, 24 h
Specific wear rate
Friction coefficient on steei
MN/m 2
GN/m 2
°C
°C
J/cm
%
10 - 7 mm3/N m
{ 0.05 m/s
Limiting ~ 0.5 m/s
P V at
[ 5 m/s
230
TR1BOLOGY December t973
MN/m 2 X rots
8.7
0.35
480
0.35
0.03
0.01
< 0.01
0.76
0.25
12.5
0.2
0.14
0.12
0.09
Nylon 6.6
0.49
.5
38
0.25
Po!yimide
1,43
~7q
3.2
No melt
>26,0
,% A
~.~;-8
0,32
30
0°42 0.6
4
0:11
0°09
< 0.09
-
Table 8
Properties of thermoplastics with lubricating fillers
Polycarbonate
~f i latieps rn c )d p e ~r e r s
Specific gravity
Tensile strength
Flexural modulus
Heat distortion temperature
at 1.8 MN/m 2
Izod notched impact strength
Moisture absorption, 24 h
Specific wear rate
Friction coefficient on steel
0.0S m/s
Limiting ~ 0.5 m/s
PV at
! 5 m/s
22~,4
ptfe
Acetal
Nylon 6.6
15%
15%
15%
ptfe
3(~:
glass
ptfe
3~4
glass
44%
ptfe
ptfe
3~:~
glass
22%
ptfe
Polyimide
15%
15%
graphite MoS 2
1.33
1.55
1.43
1.4o
1.51
1.50
45
1.3
130
120
8.3
145
1.5
40
2.1
100
1.75
MN/m 2
GN/m 2
°C
107
0.7
160
38
2.1
82
163
~).3
250
45
3.8
>260
41
3.5
>260
J/cm
1.1
0.14
1.1
0.06
5.8
0.2
0.97
I .05
0.46
0.27
0.25
3.2
0.15
>1.4
0.4
0.17
0.38
0.2
38
0.28
0.44
0.42
0.28
0.27
0.55
2.3
0.18
>1.4
0.95
0.28
0.98
0.5
3.1
0.26
0.61
0.7
0.46
%
10 7 mm3/N m
0.15
MN/m 2 X m/s
0.06
are marginal in their effects on friction, but tend to improve
wear resistance by affecting the hardness and crystallinity
developed in the polymer during processing.
Several thermoplastics, notably nylon 6.6 and acetals,
can benefit greatly from marginal lubrication with fhfids,
and in these conditions their performance often exceeds
that of conventional metallic bearing materials. To optimize the load-carrying-capacity, several types of bearings
are available comprising thin layers of nylon or acetal
bonded to a steel backing. A more sophisticated construction is a porous bronze layer on a steel backing, impregnated
with acetal, and leaving a layer of polymer about 100/Ira
thick over the surface of the bronze. Regularly spaced
recesses on the surface are provided to retain small amounts
of lubricant. This type of bearing can exhibit a life of
above 1000 h at PV factors up to about 2 MN/m 2 X m/s.
A summary of the available P V relationships for the
various thermoplastics, unfilled and filled, is given in
Fig 15.
0.32
0.32
5
50
0.1 0.3 0.1 0.3
6
5
bearing is almost inevitable, unit costs tend to be rather
high.
A characteristic feature of reinforced thermosetting
laminates is their anisotropy in mechanical properties.
Both tensile and compressive strengths may vary by
factors of up to 5 depending on the direction of testing
relative to that of the laminations. Because of th is, the
data given in Table 9 only provides order-of-magnitude
'°°I
Thern?osetting rosins
Although the raw material costs of the more common
thermosetting resins, such as phenol-fornlaldehyde, are
less than those of most thermoplastics, fabrication costs
tend to offset this advantage. The precursors are usually
fluid and the addition of hardeners and a catalyst induces
cross-linking of the molecular chains leading to an irreversible transformation to the solid state. The cheaper phenolformaldehyde and polyester resins are somewhat brittle
and are almost always used in fibre-reinforced form.
Epoxies and silicones, on the other hand, can be produced
with a whole spectrum of properties by the addition of
suitable flexibilizing agents. ~fhe simplest method of
producing reinforced thermosetting bearing materials
is by impregnation of fibrous mat or cloth cellulose,
cotton, asbestos, glass, carbon etc by the liquid resin,
or a solution thereof, pressing into an appropriate shape
and curing at elevated temperature. Alternatively, tubes
may be produced by filament winding techniques.
Neither method is amenable to rapid, large scale production, and since some machining to produce a finished
E
g
EL
nrnide
"aphit¢
0'
\
\
0"0
\
0.01
O4
I
IO
V(m/s)
Fig 15 P-Vrelationships for thermoplastic and filled
thermoplastic bearings at a depth wear rate of 25/lm/1 O0 h
(partly from Ref 6)
231
TaNe 9 Properties of reinforced thermosetting resins
MelaminePhenolformaldehyde formaldehyde Silicone
+ cellulose
+ cellulose
+ asbestos
Properties
Specific gravity
Tensile strength
Flexural modulus
Izod notched impact
strength
Water absorption, 24 h
Maximum allowable
temperature
MN/m 2
GN/m 2
J/cm
1.4
104
6.9
0.68
%
°C
1.5
130
Specific wear rate 10 - 7 mm3/N m
i .35
76
8.3
i .!
1.5
i50
1.75
62
I4
I .d
0.8
250
Epoxy
+ cellulose
1.2
90
3.5
1.1
1.5
175
Polyester
+celluiose
1.25
69
4.8
t .6
t .5
140
With solid lubricants added, values normally within the range 3 - 5 0
Polyimide
+ carbo~
fibre
1.6
450
40
0~5
2
300
0.5
Vaiues may vary appreciably depending on form of reinforcement (cloth, chopped fibre mat, unidirectional ~bres); direction of testing relative
to iaminations; degree of resin cure.
properties. The figures are nevertheless sufficient to show
that the strengths and stiffnesses of some composites
can greatly exceed those of filled or reinforced thermoplastics. In bearings, advantage can only be taken o f these
high strengths if some means are made available either
for dissipating the frictional heat effectively, or minimizing
its generation by reducing the coefficient of friction.
A major area o f application of reinforced thermosets is
thus for water4ubricated bearings in marine engineering,
rolling mills, etc. For operation as dry bearings, the addition of solid lubricants such as ptfe, graphite or MoS2 is
virtually essential, but even with these additions the specific wear rates seldom fall as low as those obtainable with
,°2t
:~
~
"~
I-
graphite, MoS2,pt
f¢
\\
×
\
the better filled ptfe composites. Because reinforced
thermosets do no~ soften appreciably on heating, they
continue to be useable up to the temperatures ar which
thermal or oxidative degradation of the resin or reinforcement begin to be significant. The rates of wear then tend
to increase, but not so rapidly as for thermoptastgcs near
their softening points, t~einforced thermosets will continue to operate for short periods at temperatures wetl
in excess of those quoted in Table 9 because therma! or
oxidative degradation are bo-th time and temperature
dependent.
Epoxy resins are considerably more expensive than
phenolics, and as shown in Table 9 the mechanical properties of reinforced composites of both types are broadly
shmilar. However, the fact that some epoxy formulations
can be made less brittle than phenoiics has led ro their
use in a filled rather than reinforced form, which partly
offsets the cost. Epoxies containing graphite or MoS2 are
available as solid bars for machining into bearings, or as
a two-component fluid for casting. Filled epoxies
can also be sprayed as thin layers on to a metat backing.
Bearkr~gs based on reinforced p olyhqqides are st~ll a com~
parative rarity, partly due to the expense of the po!ymer
and partly to the difficulties in fabrication. The introduction of carbon fibre as reinforcement further increases
cost, but this material is neverthe!ess of interest for
specialized aircraft applications because o f the combination
of very high strength and a specific wear rate as tow. or
lower, than those obtained with ptfe composites, see
Table 5.
There is little information avai!able on the F - V re!atmno
ships o f many of the reinforced therm_osetting resin composites described, but a generalized grouping is given in
Fig 16, based, as in the earlier results, on an arbitra~
wear rate of 25 gin/100 h.
Design points
o.oJ
0 •~01
0 ' II
~,,"/ I
-
:
V{m/s}
Fig 36 P-V relationships for filled and reinforced therrnosetting resin bearings at a depth wear rate of 25/~mtl00 h
(partly from Ref 6)
232
~0
It is convenient here to summarize some of t2c~especific
design factors which are invo!ved in using polymer-based
dry bearings.
Counter-face roughness
As already discussed, this is one of the most [mportar-t
parameters affecting the magnitude of the wear, particulariy
in the early stages of sliding. Values of 0.2 0.4/am cla,
with an upper limit of 0.7/am cla, are usually recommended,
but reductions down to 0.05/am cla will almost always
reduce wear unless abrasive fillers, such as glass, are present
in the composite. Abrasive fillers necessitate harder counterfaces. It is important to note that the magnitude of the
cla roughness is not a unique guide to the effect of the
counter-face on wear of the bearing. A ground surface of
0.2/am cla will generally cause less wear than an abraded
surface of the same value, and in turn an abraded surface
is preferable to one of the same roughness produced by
grit-blasting. The reasons for these differences lie in the
detailed nature of the topographies produced by each
finishing treatment.
Aspect ratk)
In theory, the performance of a dry hearing should be
independent of the aspect ratio length/internal diameterbut in practice there are two complicating features. With
large aspect ratios, distortions or misalignment may cause
stress concentrations and excessive localized heating. In
such cases a short period of 'running-in' at reduced loads,
or initial marginal lubrication, may be helpful. Short
aspect ratios introduce problems associated with the location of bearings in their housing. In general, the optimum
ratio appears to be of the order of unity.
Wall thickness
This factor is frequently dictated by the overall design
of the component. However, where a choice is at all
possible, a typical value of thickness is of the order of one
tenth of the shaft diameter. This value may usefully be
increased if shock-loadings are anticipated, or decreased
if the pressure approaches the limit at which deformation
under load becomes significant. A reduced wall-thickness
also facilitates the dissipation of frictional heat.
Clearance
Insufficient initial clearance has probably been responsible
for more dry bearing failures than any other single cause.
The clearance required for polymer-based materials is
much greater than that typical of lubricated metallic bearings and for most applications is typically 5/am/mm with a
minimum of 125/am. The necessity for this large value
arises from the combined effects of dimensional instability,
arising from expansion and moisture absorption, and the
development of transfer films of debris on the shaft
surface. With certain proprietary ptfe-based bearing
assemblies, the initial clearance can be greatly reduced by
minimizing the counter-face roughness, and in some cases
reduced to zero by preloading the bearing. In the latter
situation, however, starting-torque may present problems.
Fitting
The clearance may also be affected by the particular method
used for fitting bearings into their housings. The most popular method for polymer-based bearings is press-fitting with
a degree of interference ranging from about 7/am/ram for
unfilled ptfe, through 5/am/mm for nylons and acetals,
to 2-3/am/mm for reinforced thermosets. The closure
in bore resulting from interference fits is very roughly
equal to the interference itself and should be allowed for
when defining the clearance required. Where temperature
fluctuations are likely to be encountered, interference fits
in housings are inadvisable and alternative methods such
as keyways, flanges or adhesives are preferable.
Commercial products
A partial list of proprietary dry bearing materials with trade
names and suppliers is given in Table 10, together with a
brief description of each. The precise compositions of
many of the more complex bearing constructions are not
always divulged by the manufacturers.
Carbons and graphites
Manufactured carbons-graphites comprise a complex series
of materials whose mechanical and tribological properties
can vary very widely. The raw materials take many forms,
as shown in Table 11, and particles of one or more of these
are mixed with binders, pressed into a solid, and finally
heat-treated. A typical composition could be 20% petroleum coke. 6~} of a mixture of other cokes, natural
graphite and lamp black and 20°~ of pitch binder. The
pressure during consolidation affects the density and porosity, the latter being typically within the range 3 l 5%.
Heat treatment up to 1000°C removes volatile constituents,
particularly in the binder, and above 2100°C the amorphous carbon constituents begin to be reordered into a
graphitic structure. The degree of graphitization is a
function of both the temperature and of the structure of
the starting material. Cokes, for example, graphitize
more readily than carbon blacks at a given temperature.
Following heat-treatment any residual porosity may be
impregnated with resins or solid lubricants. Alternatively,
if high proportions of solid lubricants are required, carbon/
graphite powders prepared as above may be mixed with
the lubricant powders and resins and hot-pressed at a
temperature sufficiently high to cure the resin.
It follows from the above oversimplified, and necessarily brief, description that a whole spectrum of materials
can readily be prepared. One of the major difficulties,
in fact, is reproducibility of a desired product. For convenience, the various materials available can be divided
into four main categories as shown in Table 12. The
properties given are typical of each class, and, as usual,
are significant only as orders of magnitude. However,
several characteristic features may be noted:
a the thermal conductivities are of the same order as that
of mild steel, and therefore a factor of about 100
times greater than those of unfilled polymers,
b the coefficients of thermal expansion are only about
¼ ½ those of steel, and thus about 20 times lower
than the values for unfilled polymers,
c the moduli of elasticity are about 10 20 times lower
than that of steel, and roughly of the same order of
magnitude as those of reinforced polymers,
d the tensile strengths are substantially lower than the
compressive strengths, sometimes by a factor of 10.
Carbons-graphites do not exhibit plasticity in their
stress-strain behaviour. The elastic modulus decreases
with increasing strain, and failure occurs by brittle fracture
at strains which are typically 0.5 3%. There are no large
changes in mechanical properties over the temperature
range up to 1000°C in vacuum or inert gas atmospheres.
In air, however, oxidation begins to become significant
at around 350 500°C, depending on the type of carbon
and Table 12 shows that the amorphous materials oxidize
more rapidly at a given temperature than the more graphitic ones. For practical purposes, the temperature limitation of bearings is determined by the rate of oxidation in
relation to the rate of mechanical wear; the oxidation
233
TaNe 10 Trade names and suppliers of some polymer-based bearing materials
Trade name
Material
Manufacturer
Plaslubes
High lubricity FRTP
Nylatron GS, GSM
Nylasint 2G, 6G
Fulton 404
Delrin AF
(Fortified polymers; no
specNc trade names)
Rulon
Fluorosint
CF2
Station
M7
DQ2
DQi
DQ3
Fibreglide
unreinforced thermoplastics with ptfe o r M o S 2
glass reinforced thermoplastics with ptfe or MoS2
nylon + MoS2
nylon + graphite
acetal + ptfe
acetal + ptfe
glass reinforced thermoplastics with ptfe
Fiberfit Inc
Fiberfil inc
',?olypenco Ltd
Polypenco Ltd
LNP Corp
Du Pore
LNP Corp
DNon Corp
Po!ypenco Lid
Crane Packing C o
Morganite Carbon Ltd
Nobrac Carbon
Glacier Metals
Glacier Metals
Glacier Metals
Ampep Ltc
Motynium
ptfe with various fillers
ptfe with inca
ptfe with glass
ptfe with carbon
ptfe with graphite
ptfe with graphite
ptfe with bronze and graphite
ptfe with bronze and lead oxide
interwoven ptfe/cotton cloth with thermosetting resin and
graphite bonded to metal
interwoven ptfe/glass cloth + thermoset bonded to metal
interwoven ptfe/glass cloth + thermoset bonded to metal
filled ptfe on metal backing
porous bronze impregnated with ptfe/Pb on steel backing
porous bronze impregnated with aceta! on a steel backing
filled ptfe on metal backing
ptfeowoven bronze mesh on metal backing
polyimide
polyimide + ] 5% graphite
polyimide + 15% MoS2
polyimide + z~netal and solid lubricant fillers
porous metal impregnated with polyimide and solid lubricants
reinforced po!yimides with sotid lubricants
reinforced polyesters with graphite or MoS 2
thermosetting resins with cotton or cellulose reinforcement and
graphite, MoS2 or ptfe additives
thermosetting resins with cotton or cellulose reinforcement and
graphite, MoS 2 or ptfe additives
thermosetting resins with cotton or celhilose reinforcement and
graphite, MoS2 or ptfe additives
epoxy with MoS2, graphite or ptfe fiilers
epoxy with MoS2, graphite or ptfe fillers
Eccoslip
epoxy with MoS2, graphite or ptfe fillers
Fibreslip
Fabroid
Vandry
DU
DX
Unifton
Unimesh
Vespel SPi
Vespel SP21
Vespel SP3
Feuraton
Feuralloys
Kerimid
Orkot RL, RH, TL
~afnol
Ferrobestos
Raitko
ERB 14J
TaMe 11 Raw materials used in the production of carbonsgraphites
Ampep L;:d
Transport Dynamics !nc
Vandervelt Products
Glacier Metals
Glacier Metals
RoseoForgrove Ltd
Rose-Forgrove Ltd
©u Pont
Du Pont
Du Pont
Berno] inc
Bemoi !nc
Rhone-Pouienc
British Steel Corp
qp,
~ ;r
/[
![~Iflsoi zX ~=
J. W. Roberts Ltd
Raitko
Lid
Nobrac Carbon Lie
Gredimex AG
(Molykote G,K Lid)
Emerson and Cuming
products, being gaseous, do no;: interfere with the wear
process.
F r i c t i o n and wear
Carbons
{petroleum
pitch
h retort
Cokes
Binders
Additives
pitch
white metals
tar
bad bronze
J
~, metallurgical resins
Graphites
natural
artificial
lamp
~ channel
( furnace
Carbon
blacks
Charcoal
234
silver
ptfe
MoS 2
resins
Fig t7 gives some values for the coefficients of fricuon
and rates of wear of a variety of types of carbons°graphites
during sliding against hardened steel. The coesideraNe
variations may be noted. In genera] the friction is lowest
for the most graphitic materiais, but the wear rates do
not appear to depend so much on the grap_hite content
The iowest wear rare of N1, in fact, is that ef an amorphous
carbon. The ciassical explanation for the low friction of
graphite relates friction to the cG~stal s~mcture. Graphite
possesses a layer-lattice structure in which networks of"
hexagonally arranged carbon atoms are separated from
each other by a distance much !argot than the interatom~e
spacb.g within the layers. The binding forces between
layers are comparatively weak; little energy is therefore
Table 12
Properties of carbons-graphites
,,,•
Properties
Specific gravity
Tensile strength
Compressive strength
Elastic modulus
Expansion coefficient
Thermal conductivity
Oxidation rate at 500°C
Maximum temperature for
COlltinuous
Metallized carbon
Carbon-graphite
Amorphous
carbon
of carbon
MN/m 2
MN/m 2
GN/m 2
10 5/°C
W/m°C
g/m 2h
°C
Low
graphite
High
graphite
Electrographite
1.6
1.6
25
180
11
0.5
7.1
200
300
23
160
1I
0.2
q.2
80
350
I .65
21
83
9.5
0.3
35
15
350
14
65
7
0.4
55
2
500
0.2
0.35
0.4
1.65
White
metal
Pb-Cu
2.8
45
230
16
0.35
-
55
260
18
0.5
25
130
300
Mild
Steel
7.85
560
410
183
1.1
46
2.7
use
Limiting PV at low speeds
MN/m 2 X m/s
Specific wear rate
10 7 mm_~/N m
0.7
0.7
Generally within the range 5 - 5 0
needed to induce cleavage or shear, and so the coefficient
of friction is low. ttowever, the cleavage energy is low only
when water or other condensable vapours are present in
the environment, and this observation correlates with tile
fact that the coefficient of friction is low only when the
water vapour content exceeds a critical pressure (of the
order of 3 inmHg partial pressure in air). The presence
of water vapour also influences the magnitude of the rate
of wear, and catastrophic increases can occur, by factors
of up to 103 or 104, when the water vapour content falls
below the critical value. In sliding conditions it should be
realised that the relevant concentration of water vaponr
is not that in the environment as a whole, but that in the
vicinity o f the actual sliding surfaces. If there is a significant rise in surface temperatures, the concentration of
water vapour relative to the saturation concentration
decreases and the carbon surface may begin to lose its
physically adsorbed water vapour, leading to an increase
in both the coefficient of friction and the rate of wear.
As the temperature increases further the rate of adsorption
of water no longer suffices to satisfy the fresh carbon
surfaces produced as a result of mechanical wear, and both
the friction and wear rate increase dramatically. These
effects are illustrated schematically in Fig 18 which shows
typical variations of friction and wear with temperature
for carbons-graphites sliding on metals. High friction and
wear may also be induced at ambient room temperatures
if the combination of load and speed becomes sufticient
to raise the surface temperatures to tire order of 500°C
or greater.
Apart from the water vapour concentration, there are
two other factors which can influence the magnitude of
the coefficient of friction. Finely divided debris from tile
wear process may become consolidated on the surface of
the carbon or its counter-face and exhibit a preferred
crystallographic orientation and direction of easy shear
thus reducing friction. Alternatively these layers may fill
the surface irregularities and increase tile real area o f
J
O-4
~.r o.2
0
u
O-I
I
I
IOO
2OO
I
300
I
4OO
5OO
Temperature (°C)
IO2
'E
,o
_u .~
.-=_~
Specific wear rate {mm3/Nm)
10-7
Compacted natural gCaphlt¢
~ eiatin-e - ~ o nde~ n aTu~T 9r ~pNt ~
Copper - ~ (a ~ - ~ - ~ 9 t ~ b ~
Copper - g r a phite (low copper)
Spectroscopic st andard ~sg~hLt e'
Am~r Phous carbon
]E/ectrogr~h~te (brush grade)
~ b o n 2 ~ b J ~ l T o w~ _ ~
~
a
~
~
~
I0 6
io-S
Friction coefficient
10-4 10-3
0.4 0.8
1:2
16
g.o
or)
.
"
m
12
C a r b o n - rag5~hire ~
C a r b on - copp¢ r - ~ _ 1 5 ¢ 5 ] r ,n 91~a87 !
N a t u r a l qraphite - h i g h temp treated
£Sr-bdn- white m e t a l
- ~
Carbon lead bronze
E J ¢ ¢ t r ~ r ~ h i t l t (bearing qrade)
Amorphous c a r b o n (electrode grade~
.....
~
.
~
.
.
.
.
.
.
I
IOO
I
m
Fig 17 Friction and wear of various carbon/graphites sliding
on hardened 1% carbon steel of 0.025/lm cla roughness.
Load = ION speed = 18 m/s
2OO
3OO
400
Temperature (°C)
500
Fig 18 Schematic variation of friction and wear rate of
carbons/graphites with temperature during sliding against
steel
TRIBOLOGY
December 1 9 7 3
235
×/
d
initial
~O'ml
wear
z gG:~E
E
Q
fd ~_
P
a Natural graphite
x ESectrographit¢
e Hard carbon
u
¢J
/
h
Steady- state
wear
~
/ ~~'x~ uransTer
{~
Transfer
--xNo~-----~-------...~e
o
abrasion
Abrasion
lC?6L
0-0t
O-i
I
Io
t
IOO
Surface roughness (gin cla)
Fig ! 9 Variation of wear rate with counterface roughness
for carbons/graphites sliding on copper
contact which tends to increase the coefficient of friction°
Nddgley and co-workers ~ have observed the latter effect
occurring in a cyclic manner during the sliding o f carbon
thrust washers. When the friction rises to a high value, the
surface stresses become sufficient to disrupt the layer of
consolidated debris, the surface is roughened, the real area
of contact decreases and the friction again decreases.
Because the moduli of elasticity of carbons and graphites
are relatively low and comparable with those of reinforced
polymers, elastic deformation and fatigue play an important
role in the wear process. As with polymers, the wear rates
in the early stages of sliding are very dependent on the
surface roughness of the counter-face and Fig 19 shows
typical variations for three types of carbons. In the final
stages of sliding, when steady-state conditions are attained,
there is much less dependence of wear rate on initial
counter-face roughness, see Fig 19, for the same reasons
as discussed earlier for filled polymers; the counter-face
surface is modified either by abrasion or by transfer.
Abrasion of a counter-face by carbons-graphites arises
from two causes. Firstly, since graphite is anisotropic in
structure, it is also anisotropic in properties, and the maximum hardness in the direction parallel to the basal planes
is of the order o f 1500VPN. Secondly, carbons and
graphites are seldom pure and contain small quantities of
abrasive materials such as metal oxides. These impufites
tend to volatilize during heat-treatment, so that the degree
of purity depends not only on the type of starting materiai
but also on the heat-treatment temperature. In very general
terms the degree of abrasiveness from both causes, intrinsic
and impurities, increases in the order: electrographite,
carbon-graphite (high graphite), carbon-graphite (low
graphite), natural graphite and amorphous carbon.
236
The formation of transfer films on the counter-face
revolves structural breakdown of the 5carbon into units
of the order of 5 50 nm m size. Breakdown of the more
graphitic materials to t.his level is relatively easy and most
etectrographites, for example~ wN readily form transfer
Films ever, on counter-face metals as soft as copper. For
the stronger, non-graphitic carbons, however, locNized
stresses sufficiently high to achieve breakdown are oPJy
possible when sliding against very hard counter-face
materials. Thus amorphous carbons or low graphitecarbons will generate a tra~sfer ~irn on ceramics or tungsten carbide but not on steels or copper. As a genera]
conclusion, each carbon-grapNte generates its own characteristic topography on the counter-face by abrasion or
transfer, or a combination o f the ~wo It is this factor
which is primarily responsible for most of the difficulty
in predicting the wear rates of carbons-graphites directl]
from a knowledge of their structure, composition and
mechanical properties.
Despite "chair long history o f use, the P - V properties
of carbons-graphites do no~ appear ,~o be as well categorized
as those o f polymer-based bearing materials. Some approximate values of the limiting P V factors at tow speeds are
given in Table t2, and estimates o f the P - Y relationships
for continuous operation are shown in Fig 20. As for rei~>
forced &errnosetting resin bearings, the !LmRing PV factors
are greatly increased if fluids are available to increase the
rate of heat dissipation. Carbons-graphites are particularly
suitable for operation in fluids because not only are t~hey
chemically compatible with most types, other than strong
oxidizing agents, but they are not prone ~o dimensional
changes ~o the'sm~ne extent as polymers. However, fluids
may often result in increased wear o f the more graphitic
grades which in dry conditions exhibit low wear as a consequence of transfer fitm formation. Grades contNning
little or no graphite are less sensitive in this respect~ Addi..
!o2[
tO
jBronz
' ~
e -graphite MoS2
E
Z
O_
Carbongraphite
4-Sn-Pb-Sb
Carbon Cu/Pb
low g r,~phit~\~N~
carb°n
04High greph~
-carbon
O.Oi
O!O~
O.I
v [ra/s)
Fig 20
P-V relationships for soma carbon/graphite bearings at a depth wear rate of 25 #m/100 m (from Ref 6)
tions of ptfe to carbon-graphites may be helpful in maintalning transfer films in wet, or alternate wet and dry,
conditions. Ptfe/MoS2 additions also tend to reduce the
dependence of friction and wear on the environmental
humidity during dry operation.
Design points
The counter-face roughness usually recommended for
carbon-graphite bearings is 0.2-- 0.4 ~um cla, but as for polymers, a further reduction in roughness will generally lead
to lower wear rates, particularly in the early stages of sliding. Mild steel is not a particularly good choice of counterface metal for operation against carbons-graphites because
of its lack of corrosion resistance in humid environments.
Corrosion may also be enhanced electrochemically
because carbon is strongly electro-negative with respect
to iron. If a relatively soft and inexpensive shaft material
is essential austenitic cast iron is probably the best. However, harder materials are generally preferred to minimize
serious abrasion by the carbons-graphites and stainless
steels and stellites with hardnesses within the range 4 0 0 600 VPN are most suitable. Very hard counter-faces, such
as chromium plate or ceramics and cermets are also satisfactory but because they cannot easily be polished by
carbons during sliding, care must be taken to minimize
the initial surface roughness.
Carbon bearings are readily available in finished form
suitable for installation into housings. Machining, however,
presents few difficulties provided that cemented carbide
tools are used to offset the abrasiveness of carbon debris.
Because of poor impact resistance and tensile properties,
the wall thicknesses of carbon bearings need to be somewhat greater than those used with polymer-based bearings.
Typical values range from about 3 mm for a 12 mm bore
bush to 12 mm for 100 mm bore. Flanges on bushes
intended to support axial loads should also be avoided
because of the risk of tensile stress concentrations at the
neck of the flange. Bearing aspect ratios are preferably
kept below 2, again to minimize the effects of any tensile
stresses arising from distortion or misalignment.
It is acceptable to press-fit bushes into housings with an
interference of about 2 ~m/mm maximum, provided that
the outer edges of the bearings are chamfered. This technique, however, is mainly restricted to relatively small
bushes, of less than about 25 mm in bore, intended for
operation over a restricted temperature range. For larger
bushes, or when wide temperature fluctuations are expected.
shrink-fitting is to be preferred, the relative dimensions
of the bush and housing being chosen so that interference
will still be maintained at the maximum temperature
encountered. Typical values are 4 6 #m/ram. If the
housing has a very large expansion coefficient relative to
the carbon bush, eg bronze or aluminium alloy, it may
be advantageous to shrink-fit the bush into a steel sleeve
which is then attached to the housing by any conventional
technique. Preformed metal-backed carbon bearings are
available commercially. Because carbon bearings are
much more dimensionally stable than those based on
polymers, running clearances can be significantly lower.
Allowance is mainly required to offset the build-up of
consolidated wear debris between the sliding surfaces,
and typical clearances are about 3 # m / m m with a minimum of 50/am. Finally, it should be mentioned that
because carbons do not exhibit any plasticity and are weak
in tension, their tolerance towards abrasive contamination
is generally lower than that of polymers. To avoid high
wear rates in dirty environments, therefore, some form of
shielding for carbon bearings is essential.
Solid-film lubricants
The advantages of solid lubricants as additives to polymers
have already been mentioned. Lubricants such as MoS 2
or ptfe function essentially by forming a film over the
surface of the composite or counter-face during sliding
so that the friction becomes characteristic of this film
rather than of the matrix material. In principle, the lubricating film is self-replenishing and should provide lubrication throughout the life of the bearing. It is obvious that
an alternative method to provide lubrication is to preform
the lubricating film on the surfaces before assembly By
adopting this technique, the properties of the film can
be optimized to achieve the lowest coefficient of friction,
or the longest life, or any suitable compromize, and the
substrate materials can be chosen to obtain maximum
load-carrying-capacity. These preformed fdms, however,
inevitably have a finite life and it is the difficulty of predicting this life in practical conditions which is responsible
for most of the uncertainties associated with this type
of lubrication.
Table 13 gives a list of the more common solid-film lubricants, together with estimates of their temperature limit
in air. This limit is determined primarily by the oxidation
characteristics, and the values quoted are clearly arbitrary
because oxidation is time as well as temperature dependent.
Thermal stabilities of lamellar solid lubricants in vacuum
or inert gas are much higher than the oxidative stabilities,
and for MoS2 and similar materials are of the order of
1000°C. MoS 2 is by far file most widely used lubricant.
The simplest method of applying a film of a lamellar
solid to a substrate is by burnishing of dry powder on to
a clean surface with a soft cloth. Particles adhere locally
to the surface, mainly by mechanical interlocking with
surface defects or depressions, and a coherent film builds
up via the cohesive forces between the crystallites within
the particles. There is some evidence to suggest that softer
substrates give more adherent films whilst harder ones
produce more coherent films. As might be expected in
these circumstances, therefore, the effects of substrate
hardness on friction and wear-life of rubbed films tends to
be ill-defined. Film thicknesses for MoS 2 are usually less
than about 1 #m, and increase with relative humidity of
the environment and with time of rubbing. Practical applications of rubbed films, however, are few and confined
mainly to assembly of components or for lubrication of
precision parts when the amount of sliding involved is
relatively small.
The most common type of preformed solid film lubricant is a 'bonded-coating' in which particles of the lubricant are cemented together and to the substrate by some
type of binder, usually an organic resin. Some of the most
widely used binders are listed in Table 13, together with an
estimate of their temperature limit. Because the temperature stabilities o f organic resins are generally less than those
of the solid lubricants themselves, the temperature limits
of the coating are determined primarily by the resin properties. For temperatures in excess of about 400°C, inorganic binders have been developed based on sodium silicate,
aluminium phosphate, or various mixed oxides. The choice
of binder is also influenced by a number of other factors
in addition to temperature, such as mechanical properties,
compatibility with the environment, ease of processing,
and cost. The last two can often be of major importance
237
TaNe 13 Types of solid lubricants and surface pretreatme~qts
Maximum
Temperature
(°C)
Lubricants
Lamd|ar solids
MoS2
WS 2
GrapNte
TaS2
CaF 2
Other solids
ptfe
phthalocyanine
B203-PbS
SiO2oPbO
Na2WO4
MoO 3
350
400
500
550
1000
300
400
550
650
700
800
Binders
Maximum
Temperature
(°C)
acry~cs
cellulose
any&
phonetics
epoxies
silicones
65
65
95
160
200
300
grit -blast
add-etch
phosphate (steels)
anodize (At, Ti)
dichromate (Mg)
phosphate-fluoride (Ti)
oxalate (Cu)
polyimides
silicates
phosphates
vitreous
350
450
500
550
porous sintered or sprayed iavers
and there is an increasing interest in the use of air-curing
resins, such as acrylics and cellulose-based materia/s, which
are applied, together with the lubricant itself, from pressurized aerosol containers.
The ratio of lubricant to binder varies with the materials
invoNed, but is usually within the range 1 : I - 4 : t.
Tge higher ratios generally minimize ~ e coefficient of
friction, whilst the lower ones maxLmize the wear-life.
Apart from lubricant and binder, however, other additives
are often incorporated to enhance one or more aspects of
performance. Soft metals may facilitate re-adhesion of
debris to the substrate during sliding, smalt additions of
graphite and other metal sulphides to MoS2 films enhance
wear-life, and inhibitors prevent corrosion of the substrate
;!
by the oxidation products of MoS2' in humid eaviro~ments
eg H2SO4. Sb203 is a widely-used additive with MoS~
films to increase wear-Ere, but its mode of action still
remains obscure. ~n general, the precise folxnulations G~
commercially-available coatings are no~ avNlable to the
user. Most coatings have been developed in the USA under
the stimulus of military and aerospace requirements° anc~
because of this a greater degree of standardization has
been achieved in the US than elsewhere. Four specifications
are relevanu
a air-drying lubricant MIL L-23398B.
b general purpose, heat-cured, bonded solid fiim lubricant
M i L L 8937A,
c corrosion resisting, heat-cured, bonaed solid film iubacan~ MIL-46010A,
d extreme enviromnem 300°F 750°F. bonded solid
f t m h b r i c a n t M I L L 81329ASG.
A partial list of some of the commerciN oroducts safisfyin~
these specifications is given in Table 14. There are~ of course
numerous other products which, whilst not satisfying the
specifications in one or more respects, may nevertheless
be quite suitable for particular applications.
i
App]ica tfon
¢-
%
d~
.w
d
0-25
0.5
O'75 !.0
~'25
1.5
Substrate roughr~ess {gin de)
I 75
Ficj 21
Life-surface roughness relationships for Falex tests
on one particular bonded sblid film lubricant {from Peterson
and Finkin 6 )
238
Substrate pretreatments
The most effective method of application of bonded solid
£im lubricant coatings is by spraying on to a carefully
cleaned and roughened metal substrate. For small nurabers
of components, brushing or dipping is sometimes used_
but it is then much more difficult to control film thickness
and quality to the required standard. Even spraying is
best carried out, if possible, on an automated or semiautomated basis to ensure consistency. In addition ~o
cleaNng of the substrate, various types of pre-~reatment
can also be "used to enhance the wear-life of bonded coatings, and some of these are ~isted in Table i3. The mos~
knportant facet of surface pre~reatment is roughening ~e
increase the~egree of mechanical 'keying' of ~ie N m to
the surface. Fig 21 shows how roughening produced by
grit-blasting is more effective than grinding to the same
numerical cla roughness. Wet git-hlasting is also more
effective than dry. The essential feature reG;aired in ~ e
Table 14 Partial list of some commercial solid lubricant coating formulations satisfying US specifications
(From Peterson and Finkin 8)
MI L-L-8937A
MIL-23398B
MIL-L-46010 (A)
Acheson Colloids, Dag 254
Electrofitm, Lubri-Bond A
Dow ('orning 3400A
Dow Coming, Molykote X106
Hohman, Surfkote A1625
Everlube, Ecolube 642
Electrofihn, Lub-Lok 5306
Lubrifthn, 600A
Fel-Pro C651 A
Everlube 620
Electrofihn, Lub-Lok 2109
Fel-Pro (%40
Lubrifihn LF710A
Hohman, Surfkote M1284
Sandstrom 9A
M! L-L-8132q
(ASG)
Dow Corning,
Molykote X1 5
Lubrifihn, LF700
Table15 Effect of coating different parts in Falextests
with a bonded MoS 2 coating to MIL-L-8937
(From McCain 9)
Coating applied to
Wear life
(min)
V-blocks only
Pin only
Pin and V-blocks
10
958
965
roughening process is the production of as uniform a distribution of surface depressions as possible, see Fig 22a.
Techniques other than grit-blasting either give anisotropic
topographies, non-uniform depths of depressions, or both,
see Figs 22b and c. Further improvements in wear-life
can usually be obtained by phosphating of steel surfaces,
as shown in Fig 21, but in practice the advantages may
not always be considered to be sufficiently great to justify
the additional processing cost. In addition, phosphate
coatings break down thermally above about 300°C, and
are therefore unsuitable for use with the higher temperature
lubricant formulations.
After spraying, the coating should be carefully examined
to check uniformity and then cured for a time and temperature appropriate to the binder; these values are usually
specified by the manufacturer. The coating thickness may
be estimated by weighing or, on steel surfaces, measured
with a magnetic gauge. The thickness normally recommended is from 7 17/am, and the solid content of tire
lubricant-resin dispersion is, in fact, often adjusted to give
fihn thicknesses of this order during a single spraying
operation. Thicker coatings (up to about 50/am) may
sometimes give improved life in low stress conditions and
for these it is preferable to build-up the coating gradually
from successive spraying/curing cycles.
a variant thereof (LFW 1, one block; Macmillan, one block;
'.tohman, two blocks, Dual Rub Shoe, two blocks).
Both of these tests are used to obtain wear-lives during
continuous running, or load-carrying-capacity by increasing the load in stages as shown previously (Fig 4) When
time and opportunity permit, these accelerated tests are
supplemented by plain bearing assessments, either with
continuous rotation or more usually, with oscillatory
motion, see Fig 23c. In all three tests it is usual to coat
both of the rubbing surfaces with solid film lubricant, but
the layer on the rotating surface is by far the most critical
as is shown in Table 15 for Falex tests. Similar experiments
with plain bearings have shown that coating of the staaft
surface is responsible for about two-thirds of the wear-life
and coating of the bush for about one-third. A major
difficulty with all testing of solid film lubricants, both
in accelerated or service conditions, is lack of reproducibility of wear-life. Even under very carefully controlled
conditions a scatter of +-50% is common for Falex tests
and for Timken-type tests the scatter is even worse,
-+100%. Fig 24 illustrates the variations observed in the
wear-life of one particular film between different laboratories using identical testing conditions and apparatus.
a G r i t - blasted
Across grinding marks
Along grinding marks
b Ground
Testing
The development of solid film lubricant formulations has
been, and still is largely an empirical process in which
friction and wear testing plays the dominant role. Because
the possible combinations of materials are numerous, tests
are frequently made in apparatus specifically designed
to produce data in short time intervals. Fig 23 shows two
of the most common tests; the Falex, and the Timken, or
C Randomly abraded
Fig 22 Profiles of mild steel surfaces prepared in different
ways to a roughness of 0.5--0.62/am cla ~ X 5000 4+ X 100
239
Load
a Faiex
Lood
Load
b Timken -type
Fi~ 23
C Oscillating plain b¢(lrirEj
Tests for bonded solid film lubricants
Falex
V-blocks: free-cutting stee~, 250 VPN 12.7 mm diameter
10 mm long, 91 ° angle.
Pin: Ni-Cr steel, 150 VPN 6.35 mm diameter, 31.8 mm long
Typical operating conditions: V = 290 rev/min (~0.1 m/s),
W=upto13MN
T~mken-type
Ring: C-Cr steel, 750 VPN 35 mm diameter, 8.9 mm long
Block: Mo-steel, 620 VPN 15.4 mm X6.35 mm X 10 mm
TvpicN operating conditions: V = 72 rev/min (~0.13 m/s},
W=3MN
Oscillating plain bearing
Shaft: Cr-ptated Mo steel, 1000 VPN 15.8 mm diameter
Bush: C-Cr steel, 750 VPN 22 mm od 16 mm wide, 0.1 mm
clearance
Typical Operating conditions: 64 ° arc, 10 cycles/rain,
W = 70 MN
Because each type of test uses a different geometrical
sliding arrangement as welt as different loads and speeds,
it is obvious that there is unlikely to be any correlation
o f weardives when these are expressed in terms o f time or
distance o f sliding. However, a reasonable correlation
has been shown to exist if the life o f the coating is
expressed in terms of the number o f cycles o f compression/
flexure to which each element of the film has been subjected 9. [n a Timken-type test with a coated ring, there
is one such cycle per revolution o f the ring, whereas in
the Falex test there are 4 cycles/revolution. There are
complications involved in extending this concept to the
case of oscillating plain bearings because in these conditions the load m a y never be removed from part of the
coated shaft. In these situations o f conforming geometry,
there is also some evidence to suggest that wear-life is
lower than with nominal point or line contact geometries.
The reduced life has been attributed to the fact that debris
cannot readily escape from the contact areas, and this
explanation appears to be confirmed by the observation
that the wear-life increases if grooves or depressions are
provided in one or both of the sliding surfaces.
F r i c t i o n a n d wear
Wear volume-time relationships for b o n d e d solid film
lubricants are generally o f the type already shown in Fig t 1.
In the early stages o f sliding, loose material is rapidly
removed and the film is consolidated, the thickness being
reduced b y as m u c h as 50% in the first few hundred revo-
240
lutions. The wear rate then decreases to an extremely iow
value, sometimes virtually to zero, as a consequence of
sintering of the particles together ~ d the formation of a
smooth, c~stallographicaliy orientated surface layer~
Failure ultimately occurs either as a resui~ of continuous
wear down to the substrate or to the gradual build-up o f
compressive stresses within the coating leading ~o 'bl~stering' and a loss o f adhesion to the substrate, i n practice It
is oRen difficult, if not impossible, to differentiate between
these ~wo failure modes. Although the wear-life o f a
bonded coating thus comprises two par~s, there is little
or no information available about the effect of the conditions of sliding on each part separately. It is consequently
difficult to apply the concept of a specific wear rate to
solid fitm lubricants except as a mean value over the whole
wear-life. WbJtst such values lack precision, they may nevertheless be useful for preliminary design purposes, and
Tabie t 6 shows estimates for several types of solid f~.m
lubricants based on accelerated tests, tt m a y be noted
that these mean wear-rates are not, in general, partic~Aarly
iow in comparison ro the steady-state wear-rates of fitled
ptfe's (Table 5) and some carbons (Fig 17).
Both the wear tile and the coefficient o f friction of
bonded coatings are svrongly affected by the conditions
of sliding. The general trends are shown in Fig 25, and
it can be seen that some parameters produce opposing
effects on friction and on wear F o r example, h e a ~ loads
and !ow fi!m thicknesses reduce the coefficient of fr/ction,
but they also tend ~o reduce the wear life. The magnitude
of "the coefficient o f friction obviously depends on the
particular coating formulation used and the conditions
of sliding. Values as iow as 0.03 are not u n c o m m o n for
MoS2 f:flms o f the order o f 5 g m thick a{ heavy loads m
conditions o f nominal point or !ine contact. With decreas=
ing load, increasing Film thickness, or an increase Ln_the
apparent area of the surfaces towards a conforming geometry, the coe~]cient of friction m a y rise to values of ~ e
order of 0.15 0.2. It is often observed that the coefficient
of friction o f MoS2 films also tends to decrease with
increasing time of sliding, ~ypically by factors of abo~t
1.5 2. This is partly attributable to the reduction of
fitrn thickness which occurs during the early stages o f
sliding, together with the development o f a preferred
orientation o f the crystallites, and partly to an increase
in surface temperature which leads to loss of physicaily
adsorbed water vapour and a weakening of the [nterparticie
bonds°
q
x
R
u
i
e
I;
ooo ]
.,".:o : ....
fi
O
3
5
7
Fdex
Fig 24
me~ me
em
U
t
v
I
-
9
A C E G ~ K
Im i3
Different laborGtori~s
LFW-~
Reproducibility of wear Hfe of one bonded solid
film lubricant in two tests in different laboratories {from
McCain 9)
Table 16 Order of magnitude values of the mean specific wear rate for various types of MoS 2 bonded coatings at 20°C
(From Finkin 1°)
Mean specific
wear rate
(mm3/N m)
Coating
Ceramic
MoS 2
MoS2
MoS2MoS2
MoS2 -
10 4
10-5
10-5
10 6
10-6
10 7
bonded oxides
ceramic bonded (no metals)
thermosetting resin binders (except polyimide)
Sb203 - polyimide binders
graphite - sodium silicate binders (above 200°C)
graphite - metals - ceramic/glass binders
The effects of substrate hardness on the wear life of
solid film lubricants are conflicting, but do not, in general,
appear to be very great. Softer substrate metals have the
advantage that the film may sometimes be able to repair
itself following a localized penetration, and failure is not
then so abrupt as on the harder substrates. Greater effects
on the wear life of MoS2 coatings have been observed by
changing the type of substrate metal from steel to molybdenum, or a high molybdenum alloy such as TZM (Mo +
small proportions of Ti and Zr). The increase in life obtained with Mo substrates is greatest in conditions of sliding
where the localized flash temperatures become sufficiently
high (>1000°C) to dissociate the MoS 2 thermally. Sulphur
may then react with the Mo to reform a lubricating film.
Summary
The main advantages of solid film lubricants for dry bearings,
compared to polymer- and carbon-based materials are that:
a it is possible to use very high loadings, up to the yield
stress of the substrate metal,
b high speeds can be tolerated because the films are comparatively thin and of thermal conductivity comparable
to that of steels,
c formulations are available for use at temperatures up
to 1000°C (CaF2/BaF 2 eutectic coatings)
d very stiff bearing assemblies can be obtained by using
thin films, and little or no back-lash is introduced as
a result of wear,
e coatings are able to provide the lowest coefficients of
friction of any sliding dry-bearing system, except
perhaps those based on ptfe fibre at heavy loads and
low speeds.
Against these advantages must be offset the facts that:
a the prediction of the wear life of a coating is only
possible in order of magnitude terms and tests are
virtually mandatory,
b the performance of the great majority of coatings is
extremely sensitive to fluid contamination during
operation, leading to greatly reduced life,
c careful attention to detail is essential at all stages during
processing, and in particular all forms of fluid contamination must be avoided,
d even with die processing conditions optimized it is
not usually possible to prepare coatings with wear lives
reproducible to better than +50%,
e the ultimate failure of solid film lubricants at heavy
loads may be rapid and catastrophic, leading to seizure.
Increase
Mo substrotes
High speeds
Substrate pretreotments
Increased film thickness
I
High humidities
I Friction
coefficient
Wear
life
Heavy loads
High temperatures
High humidities
Rough counterfaces
Abrasive contamination
Fluid contamination
Heavy loads
High substrate hardness
Reduced film thickness
High temperatures
Line or point contracts
Decrease
Fig 25 Factors affecting the wear life of bonded solid film
lubricants
Little or no mention has been made of the various hightemperature solid film lubricants in Table 13. Apart from
formulations using MoS 2 sodium silicate (Molykote
X15) and glass-bonded 1~oS2 (Vitrolube), these are not
generally available commercially and their use for any given
application normally involves a development and test
programme by the user.
Pecause of the major part played by processing in the
performance of solid film lubricants, there is a growing
trend towards custom-coating of parts by specialized
processors. Most coatings can be obtained in this way,
sometimes from the manufacturer of the coating itself.
There are also a number of specialized coating treatments
which are only available on a custom-coating basis.
These include aluminium or titanium alloys with anodized
surfaces impregnated with ptfe ('Canadizing' "Tufram')
and coatings of MoS2, graphite etc, produced by particle
impingement, electrostatic spraying or sputtering. At
the present time there is insufficient data available on these
specialized coatings to evaluate their performance generally
in comparison to the more conventional resin-bonded
types.
241
TaMe 17
Some commercially availabfe composites of metals-!ametlar solid ~ubricants
Material
[steaded. application
Tra de name/S~pp~iez
Ag-ptfe
Ag-NbSe 2
Ag-Cu-MoS2
Small bearings
Polymer AG~ Polypenco
~
Ag-ptfe-NbSe 2 I
Ag-WSe 2
Ni.WS 2
a
Ni-CaF 2
l
Ag-WS2-polyi~Jde
Metats-polyimide-lameliar solids
Ta-MoS2
Ta-Mo-MoS 2
Ni-NiO-CaF 2
Bronzes 1
ken
- Graphite
Ni
Bronze-soIid lubricant inserts
Bronze-MoS2
1
Bronze-ptfe-MoS2
Ag-porous carbon
vacuum, electrical contacts
Slfding electricalcontacts
Self-lubricating bearing cages
Self-lubricating
alloys
Berne!
High temperature bearings
Self-lubricating bearing cages vacuum
High toad bearings vacuum
Self-lubricating bearing cages - high temperature
High !oad beatings
High temperature bearir~gs
Feuralons
Feuralloys
1
Molalioys, P~re Carbon Co,
,TX4organiteo553
General applications. Moderate ~emperamre
Deva Metals
General applications, particularly at heavy loads
Lubrite, Fraaerlube
Sinite
Sinitex
Grapha!loy
Self-lubricating retainers
Metat-soJid lubricant mixtures
As an Nternative to preformed solid lubricant Nms on
metals, numerous attempts have been made to incorporate
the solid lubricant within the structure of the bearing
metal itself to provide lubrication continuously. One of
the earliest ways of achieving t_his was to machine grooves
or holes into conventional bearing altoys and frill the
recesses with a solid lubricant originally based on graphite.
Bearings of this type have now been available for many
years, eg 'Lubrite' and Franertube', and have become
increasingly sop~sticated with a wide range of different
alloys and types of lubricants. Unfortunately details
of the various lubricants are not divulged by the manufacturers. P V factors range from 0.5 MN/m 2 X m/s for
continuous operation to about 3 MN/m 2 X m/s for low
speed, intermittent service. The maximum speed is about
2.5 m/s. In contrast to all the materials so far discussed
it is claimed that optimum performance is obtained with
rough shafts of the order of 1.5 3 #mcta. Recommended
clearances are 2 #m/mrn with a minimum of 75 #m and
are thus very similar to those for carbon/graphite bearings
The coefficient of friction decreases with increasing load
from about 0.15-0.05. There is no information available
on wear.
More recent developments have concentrated on the
production of more uniform mixtures of metals and solid
lubricants, and one product of this type, porous bronze/
ptfe/Pb, has already been described. Mixtures of metals
and graphite are produced by powder-metallurgical techniques (Deva metals) and the graphite contents range
from 6-25% wt (~12 50% vot). Metal matrices include
bronze, leaded bronze, brass, iron and nickel, the particular
choice being dictated by the temperature of operation, tn
general, the composite strengths tend to decrease with
increasing graphite content, but the coefficients of friction
and wear rates also decrease. Friction coefficients of the
bronze-graphites during dry operation against steel are
typically 0.15-0.3. The particular advantages of these
242
)
TR~BOLOGY December 1973
materials over plain carbon/graphites and carbon/graphites
containing small proportions of metal are a grea~er resist.ance to impact and shock-ioadings, and a higher load
carrying capacity of !5 35 MN/m 2 for the bronzegraphites° The specific wear rates, however~ are somewhat
greater than those of carbon/grapNtes and are typically
in the range10 5 t0 4 m m 3 / N m .
Numerous types of other metai-lm~etlar solid rmxtures
have been developed for particular applicationso mostly
m aerospace, and some of those which have reached commercial exploitation are listed in Table 17 tn connection
with all these materials, it is worth making the genersJ
point that lubrication is provided v/a the development of
a transfer N m of the so!id lubricant o~ the counier-~ace
surface. Transfer of lameHar solids is an inefficiei~t process
and. in general, a considerable excess of wear debris _~.s
required to form a sufficiently coherent fi!m to reduce
friction and wear. There is consequently a tendency for
the wear rate to correlate inversely with the soefficient
of friction, and this is illustrated in Fig 26 for a number
of materials for which the larnei!ar so!id is a minor constb
men~. Ptfe, however, transfers readily to rnetals and no~
oNy functions as a lubricant in Rs own right but may
also facilitate transfer of any lamellar solid present° The
~mproved performance of the metea-ptfe-lametlar solid
mixtures over those without ptfe may be noted° V e ~
recently, a group of composites has been developed based
on Ta-Mo-MoS2 and containing >5£~o of MoS2, ie
'Molalloys'. These exhibit reasonably tow wear rates and
coefficients of fliction, as shown in Fig 26. and are suitable
for operation a~. temperatures up to ~500°C in air or
~1000°C in inert atmospheres or vacuum. Despite the higg~
content of MoS2, the compressive strengths can be as higiq
as 700 MN/m 2, but the tensile strengths are at least a factor
of t0 lower, and the materials are relatively brittleo Similar
assembly techniques are used to those already described
for carbon/graphites. FinNly, it may be noted that the
specific wear rates of all the metai/lameltar solid mixtures
in Fig 26 are significantly greater than the porous bronzeptfe-Pb composite already described. Their main area of
application as bearings, therefore, lies at a temperature in
excess of that pemrissible for ptfe, ie above about 275°C.
10-2
l\\ \
--Cu-lamellar solids
Solid lubrication of rolling-element bearings
All discussion so far has been concerned with materials
for plain bearings where the coefficients of friction are of
the order of 0.05 or greater. In some applications, the
magnitude of the coefficient of friction may be a critical
parameter because of restrictions on the power available
or because of difficulties in heat dissipation. The use of
rolling element bearings is then indicated. Such bearings
are "also preferred for many instrument applications where
precise location of shafts may be critical. Conventional
rolling bearings operated without any fornr of lubrication
whatsoever exhibit relatively short lives as a result of
retainer, ball and race wear. The problem of dry operation
thus becomes one of selecting the appropriate nrethod
of solid lubrication, and there are three possibilities:
continuous supply of solid lubricant powders (MoS>
graphite/CdO mixtures, phthalocyanine) in an inert
carrier gas
preformed fihns of lubricants on the surfaces of the
retainer, balls and races
self-lubricating retainers which provide lubrication by
transfer t~ the balls and/or races.
The first technique has so far been used only for feasibility
studies on the lubrication o f high temperature bearings
and has not achieved commercial exploitation. Attention
will therefore be concentrated on the remaining two
methods. II is extremely difficult to reach general conclusions about the performance of solid lubricants in rolling
bearings. The problem is not simply one of the correct
choice of solid lubricant because design, type and materials
of the bearing itself play an equal or even more important
role in performance. It is widely accepted that the optimum
lubricant/bearing combination for one application is seldom
the same as for another, and in the space available here it is
only possible to outline a few of the combinations which
have shown promise in particular applications.
Solid films
All the types of bonded solid film coatings described
earlier are potentially suitable for use in rolling element
bearings. Because of the limited amounts of sliding involved,
it is also possible to use thin films of lamellar solids produced by burnishing of dry powder, sputtering, or from
the conversion of Me coatings to MoS 2 in H2S. Most of
the interest in solid lubricant coatings has arisen from
space applications where finite, and relatively short, bearing lives can often be tolerated, provided that test programmes are able to establish these lives with a reasonable
degree of certainty. For vacuum use, MoS 2 is normally
the preferred lubricant over other types of lamellar solids:
as nrentioned earlier, graphite is ineffective in these conditions. In one particular example, a life o f > 1 0 8 cycles
has been obtained for MoS2-Na2SiO 3 coatings on the retainer and races of 440C stainless steel bearings of R4 size
(~(~.3 mm bore) o]~erating in a vacuum of l0 9 torr
(133 × 10 9 N/nrZ) at light loads (~500 g) and low speeds
(~400 rev/min) 11. Other work with similar sized bearings
has shown, however, that the lives of various MoS 2 coatings
m vacuunr are very irreproducible. By far the most success-
\
\
Mixtures of grephil¢
Ag
MO S2
Cu
WSe2
Ni &
NbS¢2
Co
MoSe 2
F¢
MoTe 2
Cu -Ni
CoF2/Bal
O
~.
\
°0
o
E" I O - 4
Z
0
71~ 0 0
~q
o\
,.g
\
o
iO-s
~a
•
o
"- .. \ ,
"=-
." . . . .
U~lO-6 t
o\ \
~
graphite
'\\
Bronze \graphite
o
......."
Ta_Mo_MoS2
~-,.\~.~
~Metals -ptfe lamellar solids
\o~
o
\
\
\
1(7-7
Porous
b r o n z e - p t f e -Pb
w o v e n p t f e f i b r e / g l a s s fibre
iO-8
0.2
0-4
O.6
0.8
I.O
1.2
Coefficient o f f r i c t i o n
Fig 26 Relationship between wear rate and coefficient of
friction for metal-lamellar solid composites sliding on metals
(from various sources)
ful vacuum lubricant was a thin electroplated lead coating
on the races. Lives in excess of 1010 cycles at 3000 rev/
rain were obtained with full-complement bearings and the
addition of a retainer reduced the life somewhat, although
the torque was "also reduced. The optimum retainer material
was a leaded bronze. Other soft metal fihns as lubricants
for rolling bearings in vacuum have also been examined
(Ba, Ga, In, Sn, Ag and Au) and promising results are
reported for Au and Ag in lightly loaded R2 bearings
(~3 mm bore) at 10000 rev/min t3. The deposition of
these films, however appears to be more critical than
for Pb; gold plated balls in conjunction with silver plated
races are much more effective than the opposite arrangement.
Self-lubricating retainers
The solution most widely adopted for vacuum lubrication
of rolling bearings is to use a self-lubricating retainer
fabricated frona a composite of ptfe/MoS2/glass fibre
(Duroid 5813, Bartemp). Bearings of this type are available commercially in sizes up to 12.5 mm bore, together
with a limited amount of design data. Compared with
fluid-lubricated bearings, the ratio of the total dynamic
load to the total static load is very small and ranges from
about 4% with 2.36 mm bore to about 0.3% with 12.5 mln
bore. Within the recommended limits of maximum load
(~2 kg for 12.5 mm bore), lives in excess of 109 revolutions have repeatedly been obtained in vacuum. Some
success has also been reported in extending the use of
243
filled ptfe retainers to larger bearings ~4. Lives exceeding
108 revolutions were obtained with 204 bearings (20 mm
bore) operating with radial loads of about 5 kg in a vacuum
of 10- 5 tort at room temperature. The life decreased by
a factor of about 10, however, when the temperature
increased to 150°C.
One difficulty which arises in attempting to collate
information on the performance of different lubricants
in different bearings is the dependence on bearing design.
This has been demonstrated very forcibly in some recent
work where the performance o f 6 different types of 204
size bearings was compared in vacuum in the absence of
any form of lubrication ~s. The bearings involved different
degrees of precision, different retainer materials and
designs, and different numbers of balls. The average lives
showed a variation of 400 000 to i, ranging from 3 rain
to 20 000 h, the most successful being the bearing with
highest precision, a greater number of balls than usual
for this size (11 c f 7) and a leaded-bronze retainer. This
bearing also performed well in air without lubrication,
although the life was a factor of 1 0 - 2 0 times lower than
in vacuum.
Bearings with ptfe/MoS2/glass fibre retainers are also
effective in air over the temperature range - 185°C to
300°C, provided that there is no condensation of vapours
to interfere with the transfer process of ptfe to the balls
and races. A major area of application of these bearings
is, in fact, at temperatures above !50°C where the performance of miniature bearings lubricated by even the best
synthetic high temperature fluid lubricants begins to
deteriorate markedly. At 200°C, the life of fluid-lubricated
miniature bearings seldom exceeds 108 cycles, whereas dry
bearings with ptfe/MoS2/glass fibre cages wi!l give lives of
this order at 300°C.
As an alternative to fabricating retainers from filled
ptfe composites, and for operation at temperatures
exceeding 300°C, metallic retainers can be made selflubricating by suitable modifications, eg machining holes
or grooves in critical areas, which are then filled with solid
lubricant. Some examples of the various designs which
have been investigated are given in Fig 27. The most successful appears to have cylindrical reservoirs in the ball-pockets
together with rectangular reservoirs on the inner surface
of the retainer and on the lands of the inner race. MoS2graphite-sodium silicate has been the solid lubricant formulation most widely used with these designs and some results
obtained with 204 size bearings at 400°C are shown in
Table 18. The improvements obtained by providing lubrication in the lands as well as in the bali-pockets are clearly
evident, and it may also be noted that molybdenum is
the rnost effective retainer material.
In connection with the solid lubrication of relatively
large ball or roller bearings at high temperature, two recent
developments in retainer materials are worth noting. Boes
and co-workers iv have produced a series of self-lubricating
composites in which the lamellar sotid hibricant VTSe2 is
bonded together with a gallium-indium alloy. A typical
composition is 80% wt WSe2 and 20% Ga-In (75/25),
with a compressive strength of ~15 MN/m 3. Because the
tensile strength is tow, the most satisfactory design of
retainer is one in which the composite is shielded by a
metal of similar coefficient of thermal expansion, such
as titanium. With this retainer material, lives of up to 108
cycles have been obtained for 207 size bearings (35 mm
bore) at 420°C &qd loads of 45 kg and speeds of
244
10 000 rev/min. The second development is by Van Wyk
and co-workers ~s who have prepared a series of composites
based on MoS2 with minor proportions of refractory
metals, such as Ta or Mo. The compressive strengths are
similar to those of the WSe2/Ga4n composites, but the
tensile strengths are significant!y higher, which fac~ities
the fabrication of retainers. Some form of metallic support
for the larger sizes is still, however, desirab!e: One potential
application is for 'fail-safe' conditions in helicopter rotor
bearings in the event of failure o f the main oit suFpiy.
The target life is 30 rain at 12 000 roy/rain with a
radial load of 1350 kg, and conventional metalIic retainers
normatty ey&ibit lives of less than 5 rain. A retainer of
MoS 2-Ta-Mo, however, has given lives in excess of 60 rain,
and in the complete absence of oil, where transfer film
formation can become more uniform,, lives of up to 22 ir
have been obtained.
Summary
An attempt to summarize the above data in rerms o f a tileload reiationship is shown in Fig 28. The individual lines
refer to different sizes of bearings operating at different
speeds and represent the best result achieved with any one
particular system. For comparison, two lines are drawn for
the El0 Iives of conventional bearings lubricated by ~uids,
and the difference in slope may be noted. Failure under
lubricated conditions results from fatigue o f the bails or
races and in these conditions iife is inversely proportional
to the third power of the load. With self-lubricating bearings,
however, life is determined either by cage wear or by wear
of the lubricating film, and ~o a first approximation, iife
is therefore inversely proport!onal ro load !t is a~so instructive to compare the life of self-lubricated rolling bearings
with that of a porous bronze/ptfe/Pb sliding bearingo Data
for the latter has been derived from Fig i2, assuming a
speed of 1500 rev/min, and is shown by the dotted ?hqe
in Fig 28 The lives of roiling bearings with self-lubricating
cages are clearly of the same order of magnitude as those
of the ptfe-based sliding bearing, and the roiling bearings
have the additional advantages of a hig~her limiting speed
and temperature, and a lower coefficient: of friction. The
major factor Precluding the more widespread use of selflubricating rolling bearings a~ present is the cost.
Table 18 Effect of retainer material and position of soiid
lubricant reservoirs
(Lubricant 71% weight MoS2, 7 % weight graphite,
22 % weig,~t Na2SiO 3. Tool steel (M10) balls and races,
t 0 0 0 0 rev/min, 400°C, Radial load = 1.4 kg, Thrust load
2.3 kg)
{From Devine et at/%
Life(h~
Retainer material
Reservoirs
in cage
only
Reservoirs in
cage ane
lands
Fe-Si-bronze
Tt Tool steel ~W-Cr-V)
M10 Tool steei (Cr-Mo-V)
Mo-0.5% Ti-0.08% Zr
25
32
61
[ 07
! 50
~39
300
~~48
Fig 27
Reservoir designs and locations for self-lubricating retainers in rolling element bearings (from Devine et al, Ref 16)
245
Leed
film~
m}
R416-3mm)
207 (35ram)
IolO
Table 19 'Hard metals', and 'super a!~oys' suitabie for nigh
tern perature bearings
tO 4
IX'o',;;~lass
:~';~
MOS -Te-MO
I,;ooI~o,,:'\
,,~;,°o,
, r:-X
lub icat¢
Material
~
II09_
e0
w $¢2 -G~ - I n
I~°8
[42OOC)
p t f ¢ - MoS2-gless
iTs= , 50°<}
-.....
X
Iolr)
=107
Io
Limiting
:emperature
ea i g
,oo=
lo2
~1~3
184
Lood,IN)
Fig
28 Life-load relationships for rolling element bearings
w i t h solid lubricant films or retainers (from various
sources)
High temperature materials
The choice of bearing materials for operation at temperatures
in excess of 500°C is somewhat restricted and it is usually
necessary to compromise on the conflicting requirements
of low friction and low wear. Many hard metals, and superaltoys exhibit wear rates which tend to d rcrease with increasing temperature as a result of reaction with the environment
and the formation of protective oxide layers: however, the
coefficients of friction of these iayers seldom fall below
about 0.2-0.3. Some materials which have been used for
high temperature bearings are listed in Table t9, together
with an estimate of their limiting temperature. The latter
arises partly as a result of increasing oxidation and loss of
material during sliding from this cause, and partly because
of loss of strength and elastic modulus. For temperatures
in excess of about 800°C, or for applications at very high
speeds of sliding where the localized flash-temperatures
are sufficiently high to me!t most metals, a variety of
ceramics and cermets are available. Some of the more
Mo alloys (TZM)
Mo t o o l steels
Nitrided steels
500°C
HastetIoy C
(57% NL 17% Mo, 16% Cr, 5%Fe TMn,
Si, C)
Stetlite d
Stellite Star J
(43% Co, 32% Cr, 17% W, 3% Fe, +Ni, C.
Mn. Si)
750°C
[ncone! X
(73% Ni, ! 5% Cr, 7% Fe, 2½% Ti + Mm Si.
Nb,0
Stellite ] 9
Rend 4!
(55% Ni, t9% Cr, 10% Co, t0% No, 3% Ti
+ At, Fe. Si, Mn, C, B)
850°C
widely used compositions are shown m Table 20. The
advantages of cermets (metal-bonded ceramics) over ceramics alone are increased toughness, ductility and resistance
to shock loads. However, with increasing metai content
the overatl hardness decreases and the wear rate tends to
increase, as shown in Fig 29a for tungsten carbide-cobalt
mixtures. The two cermets containing A1203 have been
found parficu!arly suitable for high temperature bearing
applications: LTIB
19% A1203.59% Cr, 20% No and
2% TiO2; LT2 --- 15% A1203. 60% W and 25% Cr
There are few general guide lines from which to predict
1
--.~ZE
I0-'~
I
Z
0
IO-S[
E
0
~9
0
E?
t_
IO~L-
OI
u
u
u
u
a_ IO-~
o_
uO
U3
I
4
8
I
8
I
I
12
16
O/o C o b a l t
o-81
R - -
20
24
Fig 29 a Variation of wear rate w i t h cobalt content for
tungsten carbide-cobalt m ixtu res sliding on hard 18% W tool
steel. W = l O 0 - 5 0 0 N ,
V=O.2-3.2m/s
246
T R I B O L O G Y December t 9 7 3
i0_71
b
0
lC
!O2
03
L o a d tNi
b Variation of wear rate w i t h toad for tun£sten carbide/12%
cobalt sliding on tself. V = 0,7 2,6 m/s
Table 20
Types and properties of some ceramics and cermets
Ceramics
Cermets
a-A12° 3
B4C
TiC I
Cr3C2
SiC
WC
Si3N4
A1203
Cr -- Mo
ZrO 2 (MgO-stabilized)
A1203
W
Ni
Coatings
A1203 + TiO2
Ni
Co
Mo
+
Cr2C 3 + Ni/Cr/Co
Cr
WC + Ni/Fe/Co/Cr
Cr
Cr203 + Cr/A1203
NiO
Typical properties
Thermal
stress
resistance
factor]" (°C)
Materials
UTS
(MN/m 2)
E
(GN/m 2)
Thermal
diffusivity*
(10 4 m2/s)
a-Al20 3
240
360
0.08
60
B4C
240
450
0.3
100
170
70
0.1
50
LTIB (19% A1203, 59% Cr, 20% Mo)
280
260
0.1
150
K162B (64% TiC, 25% Ni, 5% Mo}
800
380
0.15
250
20
8
Si3N 4
SiC
Graphite
1
>1000
* Thermal diffusivity = thermal conductivity/density x specific heat
t l'hermat stress resistance factor = o(1 v)/o.E.
IO -~
the wear of ceramics and cermets from a knowledge of
their properties and composition. One major difficulty
is the marked dependence of properties on minor changes
ira composition or in methods of manufacture. The wear
process at high speeds has been attributed to thermal
fatigue on an asperity scale due to repeated cycles of
localized heating and cooling. This concept is supported
by tile fact that there appears to be a significant correlation between the wear rate and a thermal stress resistance
factor, as shown in Fig 30. The scatter is too great, however, for the results to be of much value for design or
prediction purposes. It may be noted from Table 20 that
graphite has by l;ar the, highest themral stress resistance
factor of any material. Its use for high temperature or
high speed bearings is limited, however, by oxidation above
about 5000( ` and by the fact that adhesive wear and high
friction become important in the absence of condensable
vapours for adsorption.
Various techniques for producing ceramic and cermet
coatings up to 0.5 nrm thick on metal substrates offer a
convenient way of utilizing the wear resistance of these
materials with a minimum of processing cost. Plasma
spraying is the most widely used of these techniques but
more recent developments include impingement coatings
from a detonation gun ('Linde flame plating') and electrolyric co-deposition from an electrolyte containing ceramic
particles ('Tribomet coatings'). The advantages o f ' f l a m e
plating" are an improved adhesion to the substrate metal
and a lower porosity, whereas the advantage of electrolytic
co-deposition lies in its ability to coat small internal surfaces
inaccessible by any olher technique. In each case the
c
C'
O
E
O
E
=L
O
O
O
-6
IO
o
o
E
i-
o
o°
o
o
0
o_
i
0
O
oO
0 0
O
°o°
o
o
o
O
o©
IO
O O
I
I O -I
O
O
I
I
I
IO
I
IO 2
Wear rote {t~m/min) - theoreticol, derived
from W = 1,5 p.R -p25 D -O'75
Fig 30
C o r r e l a t i o n of the wear rates o f ceramics and
cermets at high speeds w i t h an empirical r e l a t i o n s h i p involving their thermal and mechanical properties ( f r o m Sibley
and A l l e n 6 }
TRIBOLOGY
D e c e m b e r 1973
247
Z
+
@
Z
~
6
8
I
9
Sp¢c.ific wcee
F=%¢ ( ~ m ' / N t ~ )
0"~
LWt
Coatin9~
QgainsC
themselves
540oC
~0 - ~
]
0o~
LWI N30
LC4
LCRH
©.?_~
O,Z7
LA2
Coatinqs
against
~hcmselv¢~
7~0oC
0.~I
LW5
LClC
O.l@
Lr.,5 - -
O.ZZ
0.@ - 0 . 7
LA7
0.2.¢
LCC;A
0"f9
LCIC
C o ~ i ncjs
__
LW 5
Qgainst
Haynes c111oy
0'~9
O-Z3
LA7 ~
LC5
Z5
?(~OOC
0-17
LAZ
C o e f f i c i e n t of frictJom
i At. -tO0 ° c
Composition and properties of coatings
Hardness
VPN
LW1
LW1N30
LC4
LC9H
LA2
LW5
LC1C
LC5
LA7
LC9A
WC - 9% Co
W C - 13%Co
Cr203
Cr203 2O% Cr (heat-treated)
A1203
WC 5%Ni
Cr2C 3 - 15% Ni-Cr
Cr203 20% A1203
A1203 40% Ti 02
Cr203 40% Cr
1300
1150
1300
i !00
1075
800
925
950
T max
(°C)
540
540
540
870
000
760
000
870
700
870
Expansion
coefficient
( t 0 -6 °C-1)
218
218
8.1
8.1
!75
!75
84
85
56
127
56
77
6.8
8.3
-
70
140
~25
70
63
Apart from the coatings themselves, other suitable counterfaces are:
Haynes alloy 25; 20% Cr, 15% W, 10% Ni, 3% Fe-Co, 1 °5% Mn, 0.1% C
Hardened stainless steels (EN 59, 440 C)
Haynes LT-1B; 19% A1203, 59%, Cr, 20% Mo, 2% TiO2
TiC cermets (K161, 162 and 163)
and for light loads, carbon/graphites
Thrust bearing tests, P = 3.5 MN/m 2, V = 0.05 m]s,
surfaces ground and lapped to 0.025 0.05 #m cla
Fult fines - rotating member, dotted lines stationary member
(Data from Union Carbide 2°)
Fig 31
248
Specific wear rates and friction of various ceramic and cermet coatings at high temperatures
Bond
strength to
substrate
Elastic
modulus
(GN/m 2)
(MN!m2)
properties of a coating depend at least as much on the
processing conditions as on its nominal composition.
Generalizations about specific types of materials are therefore very difficult to make; each coating is unique.
The friction and wear properties of a series of ceramic
and cermet coatings from one particular manufacturer
are shown in Fig 31, based on thrust washer tests. The
most suitable mating surface is usually another coating of
either the same or a similar composition, but where this
is not possible hardened stainless steels may be satisfactory.
The wear process of ceramics and cermets often exhibits
discontinuities analogous to the transitions between mild
and severe wear found with many metals, and an example
is shown in Fig 29b. Above the critical load, the localized
stresses within the cermet are sufficiently high to cause failure of the ceramic-matrix bond, and the size of the wear
debris is therefore comparable to the size of the ceramic
particles in the composite. In escaping from within tile
contact zone these particles cause further stress concentrations and damage leading to roughened surfaces which
help to maintain high stresses and a high rate of wear.
On the other hand at light loads, and with surfaces which
are initially smooth, the wear process is one of gradual
attrition of the ceramic particles rather than complete
detachment. The surfaces then remain relatively smooth,
the localized stresses are low, and the wear rate, in turn.
also remains low. It follows that the avoidance of localized
stress concentrations is a major requirement in maintaining
low wear of ceramics and cermets, and for this reason, the
initial surface finish should always be as smooth as possible.
In obtaining these smooth surfaces, however, care must
be taken to ensure that the finishing process itself does not
introduce surface or sub-surface defects which weaken the
bonding of the ceramic particles. The wear resistance
of many ceramics and cennets can be ruined irretrievably
by excessively severe grinding operations. Taking into
account all the complicating factors discussed above, it is
clear that attempts to generalize the specific wear rates
of all the different hard metals, ceramics and cermets are
fraught with even more uncertainty than for other materials.
However, Table 21 shows tile orders of magnitude to be
¢-
~oo
u
expected in conditions o f relatively high stress at 500°C.
In addition to their use for high temperature plain
bearings, some hard metals and cermets are suitable for
rolling element bearings and gas bearings. The main advantages for these applications are a high hot-hardness and
dimensional stability. Fig 32 shows how the load-carrying
capacity of some materials fabricated into rolling element
bearings varies with temperature in comparison to a standard carbon steel bearing. The temperature limit is largely
the result of thermal softening. For even higher temperatures, ~1100°C, rolling bearings have been fabricated from
sintered a-A1203 and ZrO2. The difficulty with most
ceramic or cermet rolling elements, however, is their
tendency to failure by chipping under the high stress concentrations at the ball-race contacts. ZrO2 is reported to
be less prone to this defect than At203, possibly because
its lower elastic modulus may help to relieve the contact
stress via elastic deformation. The lives of all these bearings tend to be very short, however, unless some form of
solid lubrication is provided, either by coatings or by selflubricating retainers as discussed earlier.
Conclusions
Despite the very large number of dry bearing materials
discussed in this survey, the choice of the most suitable
material type for a particular bearing application is usually
not too difficult. As already mentioned, prediction of the
wear rates to be expected often gives rise to most
uncertainty, and it is therefore useful to condense in a
single chart all the data on specific wear rates given earlier.
Fig 33 shows this summary. Once again, it must be strongly
emphasized that the specific wear rate cannot be regarded
as a unique material property because its value will depend
on the particular conditions of sliding involved. No precise
predictions are therefore possible, and tests under service
conditions, or close laboratory simulations thereof, should
always be made wherever feasible.
If the only major material requirement for a bearing
application were a low rate of wear, the choice of material
would be obvious from Fig 33. However, this is seldom
the case because numerous other factors are normally
involved. By identifying these factors separately, it is
possible to draw up short lists of the most suitable types
of materials and from comparison of these lists, the appropriate compromise choice can then be made. A selection
chart o f this type is given in Table 22. Within a given group
of materials, the final choice can be more difficult because
80
o_
i -Mo)
6o
Table 21 Order of magnitude of the specific wear rates
for various high temperature materials sliding against themselves at 500°C. Based on pin/disc type tests
40
Specific
wear rate
(mm3/N m)
u
I io/~ r- ~.~,
2o
\
O
o
3-
2dO
.........
400
\ ~A~
\'~st¢~'~'
\
\
6dO
8bO
cast
N stellit¢
I000
Temperature (°C)
Fig 32 Limiting load capacity of various rolling-element
bearing materials at high temperatures (from Glaeser 21 )
Material
Ceramics (A1203, ZrO2, SiC)
Nickel-base alloys
Tool steels
Co-base alloys
Cermets (WC-Co, TiC-Ni-Mo;
Cr3C2-Ni-Cr; A1203-Cr-Mo)
10
10
10
10
3
3
4
5
10--5
10 5
10 5
10.-6
10-5
10-7
249
Table 22 Selection of bearing materials for various conditions
Operating requirement
Decreasing suitability
>
Low wear/long life
Low friction
High temperatures
Low temperatures
High loads
High speeds
High'stiffness
Dimensional stability
Compatibility with fluid lubricants
Corrosive environments
Compatibility with abrasives
Tolerance to soft counter-faces
Compatibility with radiation
Space/vacuum
Minimum cost
5
11
10
3
9
11
11
10
7
10
1
1
7
11
1
3
9
9
tt
10
9
9
11
10
7
3
9
4
9
2
KEY
1 Unfilled thermoplastics
2 Filled/reinforced thermoplastics
3 Filled/reinforced ptfe
4 Filled/reinforced thermosetting resins
5 Ptfe impregnated porous metals
6 Woven ptfe/glass fibre
7
8
9
10
11
6
5
11
9
6
8
5
9
4
3
2
2
9
4
3
7
3
?
4
5
5
4
7
8
4
4
3
10
3
4
do
J
g
~'
5
g
2
2
<
/~
2
2
Carbons-graphites
Metal-grap.hite mixtures
Solid film lubricants
Ceramics, cermets, hard metats
Rolling bearings with self-lubricating cages
the detailed properties o f individual materials are either
not always k n o w n or are not available to the user. This
situation applies particularly to carbons-graphites, ceramiccermet coatings, and some o f the recently developed polymermetalhamellar solid composites. In such cases, final selection is often possible on the basis of past experience in
similar applications, and most manufacturers o f dry bearings
and materials are able to provide information of this type.
Midland, Michigan, and Fig 27 by the American Society
of Lubrication Engineers. This paper is Crown Copyright
and is reproduced b y permission o f the Controller, Her
.Majesty's Stationery Office.
Acknowledgements
The follovAng short ~ist of review papers and book chap ~ers maybe
helpful in providing further detaited information on some of the
materia!s described in this survey.
A number o f figures have been redrafted from published
data and in each such case, the original source is quoted.
Figs 3 and 27, however, are direct reproductions by kind
permission o f the copyright holders; Fig 3 by Dow Coming,
Further reading
Polymer-based materials
Pratt, G. C. , 'Plastic-basedbearings', Lubrication and Lubricm-~ts,
edited by E. R. Braithwaite, Elsevier. Amsterdam (1967)
Lancaster° J. K., 'Friction and wear (of polymers)', Polymer Science:
edited by A. D. Jenkins, North Holland Publishing Co, Amsterdam
(1972)
Carbons and graphites
Badami D. V. and Wiggs, P. K. C. 'Friction and wear (of carbons
and graphites)*, Modem Aspects of Graphite Technology, edited by
L. C. F. Btaekman, Academic Press~ London (1970)
Mild steel
Metals - lametlar
solid l u b r i c a n t s
I
Ceramics
ptfe
Solid lubricants
Unfilled t h e r m o p l a s t i c s
Reinforced t h e r m o s e t s
+ SOlid l u b r i c a n t s
F i l l e d and reinforced
t hermopIostk:s
Metais-lam¢lfor solids-ptfe
I-
B o n d e d solid film ~bricants
CampbeU, M. E., Loser, J: B. and Sneegas, E. "Solid Lubricants'.
NASA SP 5059 (1966)
Benzing, R. J. 'Solid lubricants', Modem Material~ Vol 5, edited
by B. W. Gonser and H. H. Hausner, Academic Press, London
(1964)
1
C a r b o n s - gcaphites
k
Cermets
Filled p t f ¢
POROUS b r o n z e - P b - p t fe
woven p t f ¢/91ass f i b r e
Against t hemselves
"I
High temperature materials"
I
IO-8
t
IO-7
I
I
I
[O-6
IO-5
$O-4
Specific wear rate ( mmS/Nm )
I
IO-S
I0-2
Fig 33 Order of, magnitude values of wear rates for various
groups of materials during sliding against steel at room
temperature
250
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251

Dry bearings: a survey of materials and factors affecting their

Transcription

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