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The friction and wear behaviour of polyamide 6 sliding against steel at low velocity under very high contact pressures
WEAR ELSEVIER
wear 209 (1997) 106-114
The friction and wear behaviour of polyamide 6 sliding against steel at low velocity under very high contact pressures F. Van De Velde *, P. De Baets Universi~. of Gent, Department of Mechanical alut Thermal Engineering, Sint-Pietersnieuwstraat4 I, B-9000 Gent, Belgium
Received 17 July 1996: accepted4 December1996
Al~ract The friction and wear behaviour of polyamide 6 sliding against steel under huge contact pressures and at very low sliding velocities was examined using large scale test specimens. The applied contact pressures and PV factors exceed the maximum allowable values cited in the literature. It is found that polyamide 6 resists huge contact pressures and allows relative sliding against steel without damaging the steel surface. The wear (material loss) is moderate, even at the huge contact pressures for which plastic flow (creep) is considerable. Polyamide 6 sliding against steel is found to be sensitive to stick-slip motion, which complicates accurate positioning, and the resulting vibrations can also lead to damage of the bearing material or even of the construction in which it is used. © 1997 Elsevier Science S.A. Keywords: Polyamide6: Friction:Wear; Overload; Stick-slip
1. Introduction The friction and wear behaviour of polyamide (nylon) 6 (PA 6) sliding against steel under huge contact pressures and at very low sliding velocities was examined. The purpose of the tests was to investigate the suitability ofPA 6 as a bearing material for the manipulation and positioning of huge construcfions (e.g. in off-sbore applications ). These applications are characterized by huge normal loads, small sliding velocities, relatively short total sik~:,ngdistance (e.g. a few metres ) and difficult or nearly impossible ~,,:brication. It is of great importance that the constructions, which have to be positioned, are not damaged during manipulation (e.g. by wear or vibrations). The bearing blocks have to resist the huge normal loads, but considerable wear is acceptable as they are /a'eated as waste material. As the contact pressures and the PV-factors of these applications exceed the maximum allowable values cited in the literature, tests are necessary to evaluate suitable bearing materials. Large scale specimens were used for the wear tests in orde[ to minimize edge effects (creep, contact stress concentrations,...). Friction force, thickness reduction of the PA specimens and temperatures of the rubbing specimens were measured on-line during ~ e tests. The behaviour of PA 6 during positioning (under stop-restart conditions) and its * Con'esoondingauthor. Fax: + 32 9 264 35 86. 0043-1648/97/$17,00 © 1997ElsevierScience S.A. All rightsreserved Pll S0043-1648 ( 96 ) 07500-X
sensitivity to stick-slip was evaluated. The possibility of keeping the bearing blocks in their place by means of a rough surface instead of fixing them was also investigated.
2, Experimental conditions 2.1. Test rig
The friction and wear of PA 6 against steel was investigated with large-scale specimens tested with an appropriate test rig already presented in previous papers [ 1,2]. The PA specimens are placed into special specimen holders. Tile steel counterplates are mounted on both sides of the central sliding block, which slides between the vertically loaded PA specimens. The PA specimens and their holders are shown in Fig. I. The specimens ( 1) are placed in their holder (2) with their bevelled edges against the steel counterplates, perpendicular to the sliding direction. They can move in their holders in the sliding direction over a distance equal to the gap between specimens and holders. In this way relative sliding can occur between the PA blocks and their counterplates or between the PA blocks and their holders, depending on the friction force at each interface. Two types of specimen (long and short, see Fig. 1) are tested. For the long specimens there is no clearance in the holder and sliding can only occur between the specimen and the steel counterplate.
F. Van De Velde, P. De Baets / Wear 209 (1997) 106-114
I. c I
B
~ ¢
107
f
®
: PA 6 block
dimensions [mm] C
/
\
I
,
%
i
i i
T
®
I 1 I
i
type I
140 ~o,5
100 eo,~5
20
type 2
+0,5 H&5 o
100 ~o,25
24,75
(®
i
i i
[
l I
4
i
i slt~).9
direction
5
ii J
\\
L
/ Fig. I. Dimensionsof the PA specimensand holders.
The vertical loading force, horizontal frictional forces, thickness reduction of' the PA specimens, bulk temperatures of the specimens or the mating plates and ambient temperature and humidity are recorded during the tests by means of a PC equipped with a data-acquisition card. The measurements are synchronized with the horizontal movement, i.e. every time the sliding block goes through its symmetrical position.
Table I Apparentcontact pressu.,'esat the frictioninterfaces Normal load (kN)
300 825 1440 2250
Apparentcontact pressure (MFa) Plastic-matingplate
Plastic-holder
20 55 96 150
15.5 43 74 116
2.2. Test conditions Tests are conducted with three different normal loads: 300 kN, 825 kN and 2250 kN. However, in the latter case the normal load was reduced during the test to 1440 kN. because of the inability of the hydraulic power station to generate the required pressure and flow for the horizontal jacks necessary to overcome the friction force and to keep a constant sliding velocity of 5 mm s-~. The apparent contact pressures between PA blocks-mating plates and PA blocks-holders for each applied load are gathered in Table I. The stroke is 240 mm.
The sliding velocity is 5 mm s-~, except during the first 13 cycles of the test with the highest normal load. Beside a reduction in the normal load during this test, a decrease in the (absolute) pressure in the horizontal counter jack (opposing motion of the sliding block ) was necessary to obtain a sliding speed of 5 mm s - ~. This pressure was reduced manually after the 13th cycle of the test, resulting in an increase in the sliding velocity from 2.7 to 5 mm s - i. The ambient conditions are a temperature of 20-23 °(2, ai~; humidity of 40%-50%.
108
F. Van De Velde, P. De Baetsl Wear209 (1997) 106-114
Table 2 Surface roughness of the steel mating plates
Normalload (kN)
300 825 2 2 5 0 ~ 1440
/
Surfaceroughnessof matingplatesRa (tLm) Top plate
Bottomplate
5.3 4.6 4.6
4.6 5.8 5.8
I
0.40
G
p .t440 kN
P . ~2S0 k ~
The PA 6 specimens were produced by moulding. The mechanical and thermal properties (in a standard atmosphere of 23 °C and 50% relative humidity) are as follows: tensile yield stress 55 N mm-2; elasticity modulus 1900 N ram-2; melting point 220 °C; thermal conductivity 0.29 W m - ' °C-J; mass density (measured according to ISO 1183) 1.15× !03 kg m-3; water absorption (at saturation in standard atmosphere of 23 °C and 50% relative humidity) 2.2%. The dimensions of the two shapes tested are shown in Fig. 1. Both shapes result in an apparent contact area of 14 950 mm 2 and 19 360 mm2 between test specimens-mating plates and test specimens-holders respectively. The initial clearance between the PA specimens and the holders equals 10 mm for the sbo~ specimens (type I) and 0-0.5 mm for the long specimens (type 2). The counter plates are cold rolled steel (German Standard No. i.0037): hardness 175 HB; dimensions 420× 180×20 ram3; surface roughness, ground roughly to a value as shown in Table 2, with the grinding grooves parallel to the sliding direction. The test with a normal load of 300 kN is performed with a different surface preparation (grinding) from the two other tests, because the plates were reground after being damaged by accidental contact between the bottom plate and holder. The specimen holders are construction steel (German Standard No. 1.0570): hardness 190 HB; dimensions as shown in Fig. 1; roughness of the bottom surface (against which sliding of the PA specimens is possible) Ra = 9 to 10 mm, obtained by spark erosion (resulting in an isotropic surface roughness). The same counter plates and the same specimen holders were used for all performed tests. The maximum operating conditions of PA 6 as a bearing material sliding against steel according to the literature [ 3-7 ] are as follows: maximum pressure 7 × I 0 - 3to I 0 MPa, maximum operating temperature 90-110 °C; PV-limit 10 -4 to0.7 MPa m s - L As the pressure and PV-valnes cited in the literature are greatly exceeded in the applications in question, experiments were necessary. Scale effects were minimized by using large test specimens. 3, Test results Fig. 2 shows the kinetic friction force between the plastic specimens and their mating plates during three tests with
o. Number of cycMc
Fig. 2. Friction coefficient vs. number of cycles: steeI-PA 6; s = 240 mm; c = 5 mm s - I. 40 cycles.
f'
,o"/""
-50O 0
. 5
.
I--
10
.
. . . . 15 20 25 N~m',berot c y c ~
Topblock
---
J
e / ~
30
35
--
40
B~domblock I
Fig, 3. Thickness reduction of PA blocks vs. number of cycles: steel-PA 6; s = 240 nun; v = 5 mm s - i: 40 cycles.
different normal loads. The friction force is measured each time the sliding block goes through its symmetrical position; the average values of two subsequent strokes (I cycle) are presented in Fig. 2. The friction force increases sharply during the running-in period for all tested normal loads. Thereafter, it stabilizes for the smallest load (300 kN) and it decreases again for the higher loads (825 and 1440 kN). Fig. 2 also shows that the friction coefficient decreases with increasing normal load. Fig. 3 shows the reduction (as a result of wear and plastic flow) of the thickness of the PA specimens during the tests. The thickness reduction is accelerated when the load is increased. The thickness reduction of the top specimen is higher than that of the bottom specimen for the smallest load, and smaller for the higher loads. The temperatures of the bottom mating plate during the tests are shown in Fig. 4. The temperatures of the bottom PA specimen are also recorded, but this measurement mostly failed owing to the deformation of the plastic material. However, the temperature evolution of the PA specimen is found to be comparable with that of its mating plate (lower temperatures are measured because of the smaller thermal conductivity of PA in comparison with steel). Fig. 4 shows that higher temperatures are reached for the higher loads.
F, Van De Velde, P. De Baets / Wear 209 (1997) 106-114
15
2O
25
Number ol cycles ~ g . 4. Temperature of the bottom plate vs, number of cyclcs: steel-PA 6; s = 240 mm. v = 5 mm s - I; 40 cycles.
During the test with the highest load, the normal load and the pressure in the horizontal counter jack were decreased manually owing to insufficient power of the hydraulic system to overcome the friction force. The test was stopped after 23 cycles because the thickness reduction of the PA blocks was too high and contact between the steel plates and the specimen holders occurred. The other two tests were stopped after 37 cycles ( 17.76 m sliding distance). Thereafter, the sliding block is restarted again alter a dwell-time of 3 rain in order to investigate the behaviour during stop--restart.
4. Discussion
4.1. Friction force 4.1.1. lnfluence of temperature From the literature [3] it is known that the friction force between a thermoplastic and a steel surface decreases during the running-in period of sliding owing to the increasing molecular orientation of the thermoplastic parallel to the frictional direction. Such an evolution is not observed here (on the contrary the friction coefficient rises during running-in), which indicates that other mechanisms are active. During the running-in phase, both the friction force and the temperature of the sliding specimens increase, compare Fig. 2 and Fig. 4. Different authors [ 3,8 ] found that the friction force between PA and steel increases with growing temperature. The actual relation between both follows from the combined dependence of the elasticity modulus E and the shear strength ~'on the temperature T. Two mechanisms contribute to the friction force Fbetween a thermoplastic and steel: adhesion and deformation of the plastic [ 7,9]. The friction component resulting from adhesion equals the product of the real contact area and the shear strength of the plastic (softest) material [ 10]. An increase in Tabove the glass transition temperature (which equals 4050 0(2 for PA 6) leads to a gradual decrease in the elasticity modulus E [ 7 ]. This decrease in its turn results in an increase
109
in the real contact area. However, the shear strength of thermoplastics decreases when they are warmed up. Similar considerations apply for the defermatien component of the friction force. The decrease in the elasticity modulus with growing temperature leads to deeper indentation of the steel roughness asperities in the plastic material. Consequently, more deformation of the plastic material is required for relative sliding between the plastic and the steel surface. However, the necessary deformation occurs more easily (a lower force is needed) at higher temperature. The temperature dependences of the elasticity modulus and the shear strength thus have an opposing effect on the friction force for both adhesion and deformation (an increase in the temperature leads to an increase in the friction force via the elasticity modulus and to a decrease in the friction force via the shear strength). Fig. 2 shows that the friction coefficient reaches a maximum at the 13th cycle for a normal load equal to 1440 kN. It can be seen from Fig. 4 that the reduction in the friction force after the 13th cycle is accompanied by a sharp rise in the bottom plate temperature, which is caused by an increase in the relative sliding velocity from 2.7 to 5 mm s - J (see Section 2.2). Hence, the reduction of the elasticity modulus with increasing temperature dominates at relatively low temperatures leading to an increase in the friction force with temperature. At higher temperatures however, the reduction of the shear strength becomes dominant and the friction force decreases with growing temperature. For the normal load of 825 kN, the friction coefficient also goes through a maximum. In this case the maximum is less pronounced, but it is remarkable that both maxima correspond to a bottom plate temperature between 75 °C and 80 °C (compare Fig. 2 and Fig. 4). During the test with normal load equal to 300 kN: no maximum friction force can he observed, because the warming up of the sliding specimens is insufficient. Watanabe [8] found a maximum friction force for PA 6 at 150°C. In our tests tbe maximaoccur at abulk temperature of the bottom plate of about 75 0(2, but the actual temperature of the sliding contact surface is surely much higher.
4.1.2. Influence of apparent contactnressure It is clear from Fig. 2 that the friction coefficient between PA 6 and steel decreases with increasing normal load (or apparent contact pressure). However, the influence of the load is less pronounced for the higher apparent contact pressures (above 55 MPa). For apparent contact pressures which are sufficiently low to neglect the interaction of the individual contact spots between rough surfaces, the real contact area (or the indentation of the roughness asperities of the harder surface into the softer) is proportional to the normal load for both elastic and plastic deformation [ 10-12]. Such a situation results in a friction coefficient independent of the normal load, if the sheer strength of the softer material is assumed to be constant. Hewever, when the apparent contact pressure is so high that individual contact spots interact, the real contact area A,
110
F. Van De Velde. P. De Baets / Wear 209 (1997) 106-114
increases less than proportionally with the load W. Consequently the friction coefficient decreases with growing normal loads at high apparent contact pressures. It is remarkable that the friction coefficients for the normal loads equal to 825 and 1440 kN are almost identical, although the interaction of individual contact spots ( and thus the influence of normal load on the friction coefficient) is expected to he more important at these high loads. The fact that the friction coefficient is more or less independent of the mean contact pressure at huge normal loads results from the influence of another mechanism. The shear strength of polymers depends on the hydrostatic pressure in the following way [91:
,wa
(a)
laa
f
y
Jo
dlsDl~emem Imml
with ~"the shear strength, p the hydrostatic pressure, ~'oand a constants depending on the polymer. At very high normal loads, the real contact area is equal to the apparent contact area. An increase in the normal load then no longer enlarges the real contact area, but increases the real (hydrostatic) contact pressure. The friction coefficient then can be expressed as
t I
/.L=-r = "r°+ ot P P which tends to a (a constant) for very high pressures. From this discussion it follows that the friction coefficient decreases with growing contact pressure for moderate contact pressures owing to the less than proportional increase in the real contact area (or indentation of the roughness asperities of the harder surface into the softer) with normal load (this relation is found during our tests for apparent contact pressures between 20 and 55 MPa). At huge apparent contact pressures though, the increase in the shear strength with growing contact pressure compensates for the less than proportional increase in the real contact area and the friction coefficient is found to he more or less independent of the apparent contact pressure (this relation applies for our tests with appment contact pressure above about 50 MPa).
4.1.3. lnfluence of the surface roughness of the steel counter specimens One of the objectives of this work was to investigate the possibility of applying PAL6 bearing blocks without the use of a special fixation, but with different roughnesses of the steel surfaces separated by the PA blocks. Tests were performed with PA specimens which are shorter than the width of their holder ( 1 and 2 in Fig. I respectively), resulting in a 10 mm wide gap between specimens and holders in the sliding direction. For all tests the bottom surface roughness of the specimen holders is taken to be considerably larger than the roughness of the steel plates, see Section 2.2. The tests were designed to show whether relative sliding can he prevented at one interface by applying a rougher steel counter surface compared with the steel surface at the actual sliding interface. During our tests, only the interface between the PA specimens
t .~
.leo
.5o
o
50
~oo
~5o
dis~err~r~ lmml Fig. 5. (a) Friction force vs. displacement during running-in: steel-PA 6; P = 825 kN; s = 240 ram; r = 5 m m s - J ; cycle I-5. ( b ) Friction force vs. displacement under steady-state conditions: steel-PA 6; P = 825 kN; s = 240 mm; l , = 5 m m s - ~; cycle 6-10.
and the steel plates was meant as a sliding contact. The surface roughnesses of the holders and the steel plates were chosen in order to approximate the practical situation, which was simulated by tiie tests, as closely as possible. Fig. 5(a) and (b) show friction logs for such a test (with normal load equal to •25 kN) during running-in (the first 5 cycles ) and steady state respectively. During the first 3 cycles, relative sliding can only be observed between the PA specimens and the steel mating plates. No relative sliding occurs between the PA blocks and their holders, because the friction at this interface is larger. This larger friction results from the higher surface roughness of the specimen holders and from the larger apparent contact area of this interface compared with the PA specimens-steel plates contact (see Section 2.2). The higher surface roughness of the specimen holders results in a larger friction, because the indentation of the steel roughness asperities into the PA surface is deeper when the steel surface is rougher. Consequently the deformation component of the friction force and also the total friction is higher. From Section 4.1.2 it follows that the friction coefficient between PA and steel decreases with growing apparent contact pressure. A larger apparent contact area results in a smaller apparent contact pressure and thus in a
F. Van De Velde, P. De Baets / Wear 209 (1997) 106-.I 14
larger friction. However, this influence is relatively small for the high apparent contac~ pressures considered here. Fig. 5(a) shows that the friction force increases monotonically during running-in owing to the rise in the contact temperature (see Section 4. I. I ). From the 4th cycle on, relative sliding is also observed between the PA specimens and their holder. This can be seen on Fig. 5(a) and (b) as the stepwise increase in the friction force after each velocity reversal. During I0 mm (equal to the gap width between the PA specimens and their holders) a decreasing coefficient of friction is noted. Suddenly it rises to a higher value for the remaining 230 mm. This means that the first step corresponds to the friction force between the PA specimens and their holders and the second to the friction force between the PA blocks and their mating steel plates. After the 3rd cycle, the friction between specimens-mating plates exceeds the friction between specimens-holders although the surface roughness of the holders is larger and the apparent contact area is greater at the latter interface. This results from the higher contact temperature at the specimens-mating plates interface compared with the specimens-holders interface (the friction force increases with growing temperature, see Section 4.1. I ). The higher temperatures at the contact zone between PA specimens and mating plates are caused by .'.helarger amount of dissipated energy during each cycle at this interface compared with the interface between PA specimens and holders. More energy is dissipated at the former contact because the relative sliding distance during each cycle is larger at this contact compared with the specimens-holders interface (460 mm compared with 20 mm). Thus at the beginning of relative sliding, the PA blocks remain stationary relative to the steel surface with the highest roughness (which also forms the interface with the largest apparent contact area). However, the relative sliding of the plastic specimens against the steel mating plates leads to an increase in the temperature and consequently in the friction force at this interface. Hence in the simulated practical situation where PA blocks are pressed between two surfaces, one surface with high surface roughness designed as the stationary surface, the other with lower roughness designed as sliding surface, after approximately 1.5 m sliding distance unwanted sliding at the stationary surface can occur. Holders are therefore necessary to fix the PA bearing blocks on one side. 4.1.4. The behaviour of PA 6 during positioning A proper bearing material tor the positioning of very large constructions has to exclude stick-slip and should behave well during stop-restart conditions. in order to investigate the behaviour of PA 6 during stoprestart, tests were performed with a restart of relative sliding after a dwell-time of 3 min. Fig. 6 shows the friction force for a normal load equal to 300 kN during 2 cycles following a standstill of 3 rain after 37 cycles of relative sliding. During the whole first stroke and at the beginning of the following
ill
L v~
t
o
.200
./o
6
z,
,/~
~so
¢i*OId~,rr~m tmml Fig. 6. Friction force vs. displacement during 2 cycles following a stand~:.;ll o f 3 rain after 37 cycles o f relative sliding: s t e e l - P A 6 ; P = 3 0 0 klq; s = 2 4 0 ram: c = 5 m m s - i: cycles 38-39.
strokes stick-slip is observed. Tests with a different number of cycles preceding the stop-restart were performed for a normal load of 825 kN. At restart after 5 cycles, heavy stickslip occurs resulting in brittle fracture of the PA blocks, at restart after I 0 cycles, stick-slip only occurs at the beginning of the first stroke (thereafter smooth sliding is observed) while no stick-slip at all is observed at restart after 37 cycles. No stick-slip is found for a normal load equal to 1440 Idq. Five test results with different stick-slip behaviour are gathered in Table 3. This table clearly demonstrates that stickslip is more pronounced when the temperature of the PA specimens is lower. It is thought that stick-slip disappears at higher temperatures owing to the increase in frictional and internal material damping with temperature (with frictional damping is meant the slope of the friction force-relative sliding velocity relation). Stick-slip is caused by a negative slope oftbe friction force-relative sliding velocity dependence. A decrease in this negative friction damping can eliminate stick-slip [ 1 3 ] . As PA 6 is a thermoplastic, it deforms more easily at interactions with the harder steel roughness asperities of the mating surface when its temperature is higher. Consequently, at higher temperatures the roughness asperities of the steel counter surface plough through the PA, rather than normal separation of the rubbing surfaces due to the forced tangential motion of inclinations of roughness peaks over one another ( asperity interlocking) occurs. The friction component resulting from the ploughing mechanism increases with growing relative sliding velocity because PA 6 is viscoelastic. The normal vibrations caused by asperity interlocking though, lead to a reduction in the friction force at higher velocities [ 14]. As the relative importance of the ploughing mechanism in the total friction force increases with growing temperature, a ~emperature rise results in larger frictional damping ( the slope of the friction force-relative sliding velocity relation becomes less negative or more positive) and therefore in a smaller risk of stick-slip. Moreover, the internal material damping of thermoplastics (PA 6) increases with the temperature above
F. Van De Velde. P. De Baets / Wear 200 (1997) 106-114
!12
Table 3 Con'elationbetweencontacttemperatureand stick-slipbehaviour Test specifications Load (kN)
Numberof cyclespreceedingstop--restart
Temperatureof bottomplate at restant(°C)
S:ick-slip
825 300
5 37
45 53
h
825 825 1440
10 37 -
60 85 -
no stick-sllp no stick-slip
"Very heavy stick-slip resulting in brit,.le fracture of the PA specimens. b Stick-slip during several strokes. c Stick-slip at the beginning of the first stroke.
the glass transition temperature [ 7 ]. As mechanical damping has a stabilizing effect on intermittent motion [ 15], a temperature rise of the PA blocks reduces the risk of stick-slip also by way of its influence on the material daraping. Fig. 6 shows that the mean friction force during stick-slip is lower than during smooth sliding. During stick-slip serious vibrations of the entire test rig are observed. From the literattire, it is known that both normal and tangential vibrations lead to a reduction in the friction force [ 14,16]. Therefore, the decrease in the mean friction force during intermittent motion can he attributed to the appearance of vibrations induced by stick-slip. For most tests, the friction force at restart is found to be higher than before stopping (see e.g. Fig. 2 for W= 825 kN ). This increase in the friction force during relative standstill results from creep of the PA, which leads to growth of the real contact area ( greater friction by adhesion ) and to deeper indentation of the steel roughness asperities into the plastic surface (greater friction by deformation). However, Fig. 2 shows different behaviour at restart for a normal load equal to 300 kN. In this case, the friction force is decreased owing to the immediate appearance of stick-slip.
4.2. Wear The total thickness reduction of the PA blocks during the tests with different normal load is shown in Fig. 3. The total thickness reduction results from the combined action of the following mechanisms: elastic and plastic deformation, thermal expansion and wear (material loss ). The initial deformation of the specimens on application of the normal load is excluded from the results of Fig. 4 by setting the output of the probes to zero after applying the load and before the start of relative sliding. The normal deformation of the specimens due to application of the test load is shown in Table 4. The deformation at 2250 kN could not be measured accurately because of the important creep of file plastic material under the high normal load. Fig. 4 shows that the thickness reduction of the plastic blocks occurs faster for higher normal loads. The maxima of the curves for the highest load follow from the elastic mate-
Table 4 Normal deformation of the PA specimens due to loading Normal deformation (itm)
Top specimen Bottom specimen
Initial load (kN)
Test load (kN)
30
300
825
2250
0 0
237 192
430 430
1660 1040
riai recovery at the reduction of the load from 2250 kN to 1440 kN. In order to distinguish real wear (material loss) from deformation, the weight loss of the plastic blocks during the different tests was also determined and wear rates are calculated from the weight loss and the total thickness reduction (including material loss, deformation and thermal expansion), The weight of the PA blocks is not stable in time as PA can absorb water from the air [ 17]. The influence of water absorption by the PA specimens on the measured material loss is minimized by weighing the specimens immediately before and immediately after the friction and wear tests (the duration of all tests was less than 70 min and water absorption by PA occurs very slowly [ 17 ] ). The wear results determined by both methods (weighing and thickness measurement ) are gathered in Table 5. The relative part of material loss in the total thickness reduction is also shown. Table 5 indicates that the specific material loss (true wear) does not change significantly with growing normal load. Consequently, the material loss is more or less proportional to the normal load. The wear rates calculated from the total thickness reduction though, increase more than proportionally with the normal load. Hence, the part of material loss in the total thickness reduction decreases with increasing normal load. Fig. 4 shows that the temperature of the PA blocks is lower for the test with a normal load of 300 kN than for the other two tests. It can be seen in Table 5 that the material loss during this test is larger than the measured thickness reduction suggests. For the tests with the higher normal loads (and the higher temperatures of the plastic specimen) though, the
F. Van De Velde, P. De Boers/Wear 209 (1997) 106-114
I 13
Table 5 Wear rates of the PA specimens Normal load (kN)
300 825 1440
Specimen Wearrate from total weightloss
Top Bottom Top Bottom Top Bottom
Wear rate from total thicknessreductioni
Weight loss Specificwear (g) (mm3N -' m-')
Thicknessreduction Specificwear (Izm) (mm~N -i m -I)
3.8 2.8 4.7 9.9 6.3
5.89 x I0 - 4 4.34x 10-4 2.58 X In -~ 5.44X 10-4 3.22 X l0 -4
?
?
190 55 1183 2424 1818 2649
5.06X 1.46× II.17 X 22.88 X 15.93 X 23.21X
10 -4 10-'* I0-'* I0 - 4 10 -4 10 -4
Part of material Wear rate during steady state loss in the total fromtotal thicknessreduction b thickness reduction (%) (mmkm -I ) (rams N-t m-I) 116.4 297.3 23.1 23.8 20.2
16A 8.4 50.5 115.9 126.8
8.2X 10 -4 4.2X 10 -4 9.18X 10 -4 21.1X 10 -4 13.2X 10 -4
?
171.9
17.9× 10-4
"Only the last measuredpoint is considered. b From linear regressionthroughthe last 'linear' part of each curve. material loss is found to be considerably smaller than expected based on the measured thickness reductions. It can therefore be concluded that the thermal expansion of the plastic specimens dominates at moderate temperatures (for P = 300 kN), leading to a smaller thickness redaction compared with the real material loss, while plastic deformation (creep) of the PA blocks under the huge normal loads results in a more important increase in the thickness reduction than in the material loss at higher temperatures (for P = 8 2 5 kN and P = 1440 kN). The PA blocks are considerably (plastically) deformed after the tests with the higher normal loads. They are about 5 mm shorter in the sliding direction, 10 mm longer crosswise and warped owing to thermal stresses. The actual wear (material loss) of the plastic blocks is caused by abrasion, which could be expected for such rough counterplates [ 18]. No material transfer duc to adhesion of the plastic material to the steel counter surfaces is observed. Fig. 3 and Table 2 and Table 5 indicate that the largest wear has always occurred at the interface with the roughest mating plate ( for the test with normal load equal to 300 kN the largest wear is found at the top specimen, while the bottom specimens have worn more during the other two tests). The literature supports the finding that abrasive wear of polymers increases considerably with growing surface roughness of the steel mating plates [ 7,18 ]. However, the differences between the thickness reduction of top and bottom specimens shown in Fig. 3 cannot result from different material loss by abrasion only. A greater deformation of the PA blocks has occurred at the interface with the roughest mating plate. This indicates that the temperature, and thus also the friction force which causes the temperature rise, is larger at the roughest mating plate (see also Section 4.1.3). The surface roughness of the steel counter plates and holders was measured before and after the tests. No significant change in the surface roughness due to wear is observed.
$. Conclusions !. PA 6 resists huge contact pressures and allows relative sliding against steel without damaging the steel surface.
2. The sliding velocity should be sufficiently low in order to restrict the temperature rise of the PA 6 bearing blocks. This temperature rise f~:~,altsin a considerable increase in the friction force, excluding the possibility of using PA 6 bearing blocks without fixation. 3. The wear (material loss) of PA 6 sliding against steel is moderate, even at very high contact pressures, and is more or less proportional to the normal load. 4. The thickness reduction of the PA 6 blocks due to deformation (creep) becomes dominant for huge contact pressures (and large temperatures), while it is not clearly observable at relatively small normal loads. 5. PA 6 sliding against steel is sensitive to stick-slip motion, which complicates accurate positioning, but the resulting vibrations can also damage the bearing blocks or even the construction which has to be manipulated. The use of PA 6 in combination with proper lubricants (e.g. grease, PTFE) could eliminate the stick-slip problem and is a topic of further research.
References [ I ] P. De Baets. Wear simulationof a sliding system by means of largescale specimentesting, We, r, 173 (1994) 65-74. [21C. Dekouinck. P. De Boers and F. Van De Velde, A highlight on tribologicalresearch: friction and wear devices. Enr. J. Mech. Eng. M..40(3) (1995) 154--156. 13[ Y. Yamaguchi.Tribology of Plastic Materials, TfibologySeries, Vol. 16, Elsevier.Amsterdam, 1990. [4] MJ. Neale. Tribology Handbook, Butterworth,London. 1973. [5] M.B. Peterson and W.O. Wirier, Wear Control Handbook. ASMEo New York, 1980. [6] B. Bushan and B.K. Gapta, Handbook of Tribology. Materials. Coatings and Surface Treatments, McGraw-Hill.New Ymk, 1991. [7l H. Uetz and J. Wiedemeyer,Tribologie der Polyraere. Carl Hanser Verlag, Munich. 1985. [8] M. Watanabe, The frictional propertiesof nylon. Wear. 12 (1968) 185. 191 I.M. Hutchings, Trlbology. Friction and Wear of Engineering Materials.Edward Arnold, London, 1992. [ 10l F.P. Bowden and D. Tabor. The Frictionand Ialbricationof Solids, Clarendon Press,Oxford, 1950. l II ] J.F. Archard. Elasticdeformation and the laws of friction.Proc. R. Soc. London. Ser.A. 243 (1957) 190.
114
F. Van De Velde. P. De Baets l Wear 209 (1997) 106-114
[12] J.A. Greenwood and J.B.P. Williamson, Contact of nominally fiat surfaces,Proc R. Sac. London. Set. A, 295 (1966) .'~0--319. [13] G. Cockerhamand G.R. Synunous, Stability criterion for stick-slip motion using a discontinuous dynamic friction model, Wear, 40 (1976) 113--120. [14] D.M.Tolstoi,Significanceof die uormaldegrecof freedomand natural normal vibrationsin contact friction, Wear, I0 (1967) 199-213. [ 15] B.V. Derjaguin,V.E. Push and D.M. Tolstoi, A ~ of stick-slip sliding of solids, Proc. Conf. on Lubrication and Wear, Inst. Mech. Eng., London, 1957,pp. 255-268. [ 16] W. Lenkiewicz, The sliding friction process.effect of external vibrations,Wear, 13 (1969) 99-108. [ 17] G. Schreyer, Konstruieren mit Kunststoffen, Vol. 2, Carl Hauser Verlag,Munich, 1972. 118l D. Dawson. J.M. Challonand K. Holmes,The wear of non-metallic materials, Proc. 3rd Leeds-Lyon Syrup. on Tribology, Mech. Eng. Publ., 1976,pp. 99-102.
mograp~es Frederik Van De Velde graduated in mechanical engineering in 1993 at the University of Gent, Belgium. The same year
he became an assistant at the Laboratory of Machines and Machine Construction at the same university, where he is involved in lecturing on production methods and machining. His research is concentrated in the area of dry and boundary lubricated friction and wear, with special reference to instabilities during machining, for example stick-slip at machine tool slideways. He is also involved in several industrial research projects concerning bearing material and lubricant testing. Patrick De Baets graduated in mechanical engineering in 1989 at the University of Gent, Belgium. That year he joined the Laboratory of Machines and Machine Construction at the same university, and in 1995 received his PhD degree. As doctor-assistant he is involved in lecturing and research on the dimensioning of machine elements and tribology. His research interests are mainly in the area of dry friction and wear, especially fretting wear, large-scale wear testing and industrial research. He is a member of the Mechanical Engineering and Transport Mechanics Group of the Flemish Organization of Engineers.
wear 209 (1997) 106-114
The friction and wear behaviour of polyamide 6 sliding against steel at low velocity under very high contact pressures F. Van De Velde *, P. De Baets Universi~. of Gent, Department of Mechanical alut Thermal Engineering, Sint-Pietersnieuwstraat4 I, B-9000 Gent, Belgium
Received 17 July 1996: accepted4 December1996
Al~ract The friction and wear behaviour of polyamide 6 sliding against steel under huge contact pressures and at very low sliding velocities was examined using large scale test specimens. The applied contact pressures and PV factors exceed the maximum allowable values cited in the literature. It is found that polyamide 6 resists huge contact pressures and allows relative sliding against steel without damaging the steel surface. The wear (material loss) is moderate, even at the huge contact pressures for which plastic flow (creep) is considerable. Polyamide 6 sliding against steel is found to be sensitive to stick-slip motion, which complicates accurate positioning, and the resulting vibrations can also lead to damage of the bearing material or even of the construction in which it is used. © 1997 Elsevier Science S.A. Keywords: Polyamide6: Friction:Wear; Overload; Stick-slip
1. Introduction The friction and wear behaviour of polyamide (nylon) 6 (PA 6) sliding against steel under huge contact pressures and at very low sliding velocities was examined. The purpose of the tests was to investigate the suitability ofPA 6 as a bearing material for the manipulation and positioning of huge construcfions (e.g. in off-sbore applications ). These applications are characterized by huge normal loads, small sliding velocities, relatively short total sik~:,ngdistance (e.g. a few metres ) and difficult or nearly impossible ~,,:brication. It is of great importance that the constructions, which have to be positioned, are not damaged during manipulation (e.g. by wear or vibrations). The bearing blocks have to resist the huge normal loads, but considerable wear is acceptable as they are /a'eated as waste material. As the contact pressures and the PV-factors of these applications exceed the maximum allowable values cited in the literature, tests are necessary to evaluate suitable bearing materials. Large scale specimens were used for the wear tests in orde[ to minimize edge effects (creep, contact stress concentrations,...). Friction force, thickness reduction of the PA specimens and temperatures of the rubbing specimens were measured on-line during ~ e tests. The behaviour of PA 6 during positioning (under stop-restart conditions) and its * Con'esoondingauthor. Fax: + 32 9 264 35 86. 0043-1648/97/$17,00 © 1997ElsevierScience S.A. All rightsreserved Pll S0043-1648 ( 96 ) 07500-X
sensitivity to stick-slip was evaluated. The possibility of keeping the bearing blocks in their place by means of a rough surface instead of fixing them was also investigated.
2, Experimental conditions 2.1. Test rig
The friction and wear of PA 6 against steel was investigated with large-scale specimens tested with an appropriate test rig already presented in previous papers [ 1,2]. The PA specimens are placed into special specimen holders. Tile steel counterplates are mounted on both sides of the central sliding block, which slides between the vertically loaded PA specimens. The PA specimens and their holders are shown in Fig. I. The specimens ( 1) are placed in their holder (2) with their bevelled edges against the steel counterplates, perpendicular to the sliding direction. They can move in their holders in the sliding direction over a distance equal to the gap between specimens and holders. In this way relative sliding can occur between the PA blocks and their counterplates or between the PA blocks and their holders, depending on the friction force at each interface. Two types of specimen (long and short, see Fig. 1) are tested. For the long specimens there is no clearance in the holder and sliding can only occur between the specimen and the steel counterplate.
F. Van De Velde, P. De Baets / Wear 209 (1997) 106-114
I. c I
B
~ ¢
107
f
®
: PA 6 block
dimensions [mm] C
/
\
I
,
%
i
i i
T
®
I 1 I
i
type I
140 ~o,5
100 eo,~5
20
type 2
+0,5 H&5 o
100 ~o,25
24,75
(®
i
i i
[
l I
4
i
i slt~).9
direction
5
ii J
\\
L
/ Fig. I. Dimensionsof the PA specimensand holders.
The vertical loading force, horizontal frictional forces, thickness reduction of' the PA specimens, bulk temperatures of the specimens or the mating plates and ambient temperature and humidity are recorded during the tests by means of a PC equipped with a data-acquisition card. The measurements are synchronized with the horizontal movement, i.e. every time the sliding block goes through its symmetrical position.
Table I Apparentcontact pressu.,'esat the frictioninterfaces Normal load (kN)
300 825 1440 2250
Apparentcontact pressure (MFa) Plastic-matingplate
Plastic-holder
20 55 96 150
15.5 43 74 116
2.2. Test conditions Tests are conducted with three different normal loads: 300 kN, 825 kN and 2250 kN. However, in the latter case the normal load was reduced during the test to 1440 kN. because of the inability of the hydraulic power station to generate the required pressure and flow for the horizontal jacks necessary to overcome the friction force and to keep a constant sliding velocity of 5 mm s-~. The apparent contact pressures between PA blocks-mating plates and PA blocks-holders for each applied load are gathered in Table I. The stroke is 240 mm.
The sliding velocity is 5 mm s-~, except during the first 13 cycles of the test with the highest normal load. Beside a reduction in the normal load during this test, a decrease in the (absolute) pressure in the horizontal counter jack (opposing motion of the sliding block ) was necessary to obtain a sliding speed of 5 mm s - ~. This pressure was reduced manually after the 13th cycle of the test, resulting in an increase in the sliding velocity from 2.7 to 5 mm s - i. The ambient conditions are a temperature of 20-23 °(2, ai~; humidity of 40%-50%.
108
F. Van De Velde, P. De Baetsl Wear209 (1997) 106-114
Table 2 Surface roughness of the steel mating plates
Normalload (kN)
300 825 2 2 5 0 ~ 1440
/
Surfaceroughnessof matingplatesRa (tLm) Top plate
Bottomplate
5.3 4.6 4.6
4.6 5.8 5.8
I
0.40
G
p .t440 kN
P . ~2S0 k ~
The PA 6 specimens were produced by moulding. The mechanical and thermal properties (in a standard atmosphere of 23 °C and 50% relative humidity) are as follows: tensile yield stress 55 N mm-2; elasticity modulus 1900 N ram-2; melting point 220 °C; thermal conductivity 0.29 W m - ' °C-J; mass density (measured according to ISO 1183) 1.15× !03 kg m-3; water absorption (at saturation in standard atmosphere of 23 °C and 50% relative humidity) 2.2%. The dimensions of the two shapes tested are shown in Fig. 1. Both shapes result in an apparent contact area of 14 950 mm 2 and 19 360 mm2 between test specimens-mating plates and test specimens-holders respectively. The initial clearance between the PA specimens and the holders equals 10 mm for the sbo~ specimens (type I) and 0-0.5 mm for the long specimens (type 2). The counter plates are cold rolled steel (German Standard No. i.0037): hardness 175 HB; dimensions 420× 180×20 ram3; surface roughness, ground roughly to a value as shown in Table 2, with the grinding grooves parallel to the sliding direction. The test with a normal load of 300 kN is performed with a different surface preparation (grinding) from the two other tests, because the plates were reground after being damaged by accidental contact between the bottom plate and holder. The specimen holders are construction steel (German Standard No. 1.0570): hardness 190 HB; dimensions as shown in Fig. 1; roughness of the bottom surface (against which sliding of the PA specimens is possible) Ra = 9 to 10 mm, obtained by spark erosion (resulting in an isotropic surface roughness). The same counter plates and the same specimen holders were used for all performed tests. The maximum operating conditions of PA 6 as a bearing material sliding against steel according to the literature [ 3-7 ] are as follows: maximum pressure 7 × I 0 - 3to I 0 MPa, maximum operating temperature 90-110 °C; PV-limit 10 -4 to0.7 MPa m s - L As the pressure and PV-valnes cited in the literature are greatly exceeded in the applications in question, experiments were necessary. Scale effects were minimized by using large test specimens. 3, Test results Fig. 2 shows the kinetic friction force between the plastic specimens and their mating plates during three tests with
o. Number of cycMc
Fig. 2. Friction coefficient vs. number of cycles: steeI-PA 6; s = 240 mm; c = 5 mm s - I. 40 cycles.
f'
,o"/""
-50O 0
. 5
.
I--
10
.
. . . . 15 20 25 N~m',berot c y c ~
Topblock
---
J
e / ~
30
35
--
40
B~domblock I
Fig, 3. Thickness reduction of PA blocks vs. number of cycles: steel-PA 6; s = 240 nun; v = 5 mm s - i: 40 cycles.
different normal loads. The friction force is measured each time the sliding block goes through its symmetrical position; the average values of two subsequent strokes (I cycle) are presented in Fig. 2. The friction force increases sharply during the running-in period for all tested normal loads. Thereafter, it stabilizes for the smallest load (300 kN) and it decreases again for the higher loads (825 and 1440 kN). Fig. 2 also shows that the friction coefficient decreases with increasing normal load. Fig. 3 shows the reduction (as a result of wear and plastic flow) of the thickness of the PA specimens during the tests. The thickness reduction is accelerated when the load is increased. The thickness reduction of the top specimen is higher than that of the bottom specimen for the smallest load, and smaller for the higher loads. The temperatures of the bottom mating plate during the tests are shown in Fig. 4. The temperatures of the bottom PA specimen are also recorded, but this measurement mostly failed owing to the deformation of the plastic material. However, the temperature evolution of the PA specimen is found to be comparable with that of its mating plate (lower temperatures are measured because of the smaller thermal conductivity of PA in comparison with steel). Fig. 4 shows that higher temperatures are reached for the higher loads.
F, Van De Velde, P. De Baets / Wear 209 (1997) 106-114
15
2O
25
Number ol cycles ~ g . 4. Temperature of the bottom plate vs, number of cyclcs: steel-PA 6; s = 240 mm. v = 5 mm s - I; 40 cycles.
During the test with the highest load, the normal load and the pressure in the horizontal counter jack were decreased manually owing to insufficient power of the hydraulic system to overcome the friction force. The test was stopped after 23 cycles because the thickness reduction of the PA blocks was too high and contact between the steel plates and the specimen holders occurred. The other two tests were stopped after 37 cycles ( 17.76 m sliding distance). Thereafter, the sliding block is restarted again alter a dwell-time of 3 rain in order to investigate the behaviour during stop--restart.
4. Discussion
4.1. Friction force 4.1.1. lnfluence of temperature From the literature [3] it is known that the friction force between a thermoplastic and a steel surface decreases during the running-in period of sliding owing to the increasing molecular orientation of the thermoplastic parallel to the frictional direction. Such an evolution is not observed here (on the contrary the friction coefficient rises during running-in), which indicates that other mechanisms are active. During the running-in phase, both the friction force and the temperature of the sliding specimens increase, compare Fig. 2 and Fig. 4. Different authors [ 3,8 ] found that the friction force between PA and steel increases with growing temperature. The actual relation between both follows from the combined dependence of the elasticity modulus E and the shear strength ~'on the temperature T. Two mechanisms contribute to the friction force Fbetween a thermoplastic and steel: adhesion and deformation of the plastic [ 7,9]. The friction component resulting from adhesion equals the product of the real contact area and the shear strength of the plastic (softest) material [ 10]. An increase in Tabove the glass transition temperature (which equals 4050 0(2 for PA 6) leads to a gradual decrease in the elasticity modulus E [ 7 ]. This decrease in its turn results in an increase
109
in the real contact area. However, the shear strength of thermoplastics decreases when they are warmed up. Similar considerations apply for the defermatien component of the friction force. The decrease in the elasticity modulus with growing temperature leads to deeper indentation of the steel roughness asperities in the plastic material. Consequently, more deformation of the plastic material is required for relative sliding between the plastic and the steel surface. However, the necessary deformation occurs more easily (a lower force is needed) at higher temperature. The temperature dependences of the elasticity modulus and the shear strength thus have an opposing effect on the friction force for both adhesion and deformation (an increase in the temperature leads to an increase in the friction force via the elasticity modulus and to a decrease in the friction force via the shear strength). Fig. 2 shows that the friction coefficient reaches a maximum at the 13th cycle for a normal load equal to 1440 kN. It can be seen from Fig. 4 that the reduction in the friction force after the 13th cycle is accompanied by a sharp rise in the bottom plate temperature, which is caused by an increase in the relative sliding velocity from 2.7 to 5 mm s - J (see Section 2.2). Hence, the reduction of the elasticity modulus with increasing temperature dominates at relatively low temperatures leading to an increase in the friction force with temperature. At higher temperatures however, the reduction of the shear strength becomes dominant and the friction force decreases with growing temperature. For the normal load of 825 kN, the friction coefficient also goes through a maximum. In this case the maximum is less pronounced, but it is remarkable that both maxima correspond to a bottom plate temperature between 75 °C and 80 °C (compare Fig. 2 and Fig. 4). During the test with normal load equal to 300 kN: no maximum friction force can he observed, because the warming up of the sliding specimens is insufficient. Watanabe [8] found a maximum friction force for PA 6 at 150°C. In our tests tbe maximaoccur at abulk temperature of the bottom plate of about 75 0(2, but the actual temperature of the sliding contact surface is surely much higher.
4.1.2. Influence of apparent contactnressure It is clear from Fig. 2 that the friction coefficient between PA 6 and steel decreases with increasing normal load (or apparent contact pressure). However, the influence of the load is less pronounced for the higher apparent contact pressures (above 55 MPa). For apparent contact pressures which are sufficiently low to neglect the interaction of the individual contact spots between rough surfaces, the real contact area (or the indentation of the roughness asperities of the harder surface into the softer) is proportional to the normal load for both elastic and plastic deformation [ 10-12]. Such a situation results in a friction coefficient independent of the normal load, if the sheer strength of the softer material is assumed to be constant. Hewever, when the apparent contact pressure is so high that individual contact spots interact, the real contact area A,
110
F. Van De Velde. P. De Baets / Wear 209 (1997) 106-114
increases less than proportionally with the load W. Consequently the friction coefficient decreases with growing normal loads at high apparent contact pressures. It is remarkable that the friction coefficients for the normal loads equal to 825 and 1440 kN are almost identical, although the interaction of individual contact spots ( and thus the influence of normal load on the friction coefficient) is expected to he more important at these high loads. The fact that the friction coefficient is more or less independent of the mean contact pressure at huge normal loads results from the influence of another mechanism. The shear strength of polymers depends on the hydrostatic pressure in the following way [91:
,wa
(a)
laa
f
y
Jo
dlsDl~emem Imml
with ~"the shear strength, p the hydrostatic pressure, ~'oand a constants depending on the polymer. At very high normal loads, the real contact area is equal to the apparent contact area. An increase in the normal load then no longer enlarges the real contact area, but increases the real (hydrostatic) contact pressure. The friction coefficient then can be expressed as
t I
/.L=-r = "r°+ ot P P which tends to a (a constant) for very high pressures. From this discussion it follows that the friction coefficient decreases with growing contact pressure for moderate contact pressures owing to the less than proportional increase in the real contact area (or indentation of the roughness asperities of the harder surface into the softer) with normal load (this relation is found during our tests for apparent contact pressures between 20 and 55 MPa). At huge apparent contact pressures though, the increase in the shear strength with growing contact pressure compensates for the less than proportional increase in the real contact area and the friction coefficient is found to he more or less independent of the apparent contact pressure (this relation applies for our tests with appment contact pressure above about 50 MPa).
4.1.3. lnfluence of the surface roughness of the steel counter specimens One of the objectives of this work was to investigate the possibility of applying PAL6 bearing blocks without the use of a special fixation, but with different roughnesses of the steel surfaces separated by the PA blocks. Tests were performed with PA specimens which are shorter than the width of their holder ( 1 and 2 in Fig. I respectively), resulting in a 10 mm wide gap between specimens and holders in the sliding direction. For all tests the bottom surface roughness of the specimen holders is taken to be considerably larger than the roughness of the steel plates, see Section 2.2. The tests were designed to show whether relative sliding can he prevented at one interface by applying a rougher steel counter surface compared with the steel surface at the actual sliding interface. During our tests, only the interface between the PA specimens
t .~
.leo
.5o
o
50
~oo
~5o
dis~err~r~ lmml Fig. 5. (a) Friction force vs. displacement during running-in: steel-PA 6; P = 825 kN; s = 240 ram; r = 5 m m s - J ; cycle I-5. ( b ) Friction force vs. displacement under steady-state conditions: steel-PA 6; P = 825 kN; s = 240 mm; l , = 5 m m s - ~; cycle 6-10.
and the steel plates was meant as a sliding contact. The surface roughnesses of the holders and the steel plates were chosen in order to approximate the practical situation, which was simulated by tiie tests, as closely as possible. Fig. 5(a) and (b) show friction logs for such a test (with normal load equal to •25 kN) during running-in (the first 5 cycles ) and steady state respectively. During the first 3 cycles, relative sliding can only be observed between the PA specimens and the steel mating plates. No relative sliding occurs between the PA blocks and their holders, because the friction at this interface is larger. This larger friction results from the higher surface roughness of the specimen holders and from the larger apparent contact area of this interface compared with the PA specimens-steel plates contact (see Section 2.2). The higher surface roughness of the specimen holders results in a larger friction, because the indentation of the steel roughness asperities into the PA surface is deeper when the steel surface is rougher. Consequently the deformation component of the friction force and also the total friction is higher. From Section 4.1.2 it follows that the friction coefficient between PA and steel decreases with growing apparent contact pressure. A larger apparent contact area results in a smaller apparent contact pressure and thus in a
F. Van De Velde, P. De Baets / Wear 209 (1997) 106-.I 14
larger friction. However, this influence is relatively small for the high apparent contac~ pressures considered here. Fig. 5(a) shows that the friction force increases monotonically during running-in owing to the rise in the contact temperature (see Section 4. I. I ). From the 4th cycle on, relative sliding is also observed between the PA specimens and their holder. This can be seen on Fig. 5(a) and (b) as the stepwise increase in the friction force after each velocity reversal. During I0 mm (equal to the gap width between the PA specimens and their holders) a decreasing coefficient of friction is noted. Suddenly it rises to a higher value for the remaining 230 mm. This means that the first step corresponds to the friction force between the PA specimens and their holders and the second to the friction force between the PA blocks and their mating steel plates. After the 3rd cycle, the friction between specimens-mating plates exceeds the friction between specimens-holders although the surface roughness of the holders is larger and the apparent contact area is greater at the latter interface. This results from the higher contact temperature at the specimens-mating plates interface compared with the specimens-holders interface (the friction force increases with growing temperature, see Section 4.1. I ). The higher temperatures at the contact zone between PA specimens and mating plates are caused by .'.helarger amount of dissipated energy during each cycle at this interface compared with the interface between PA specimens and holders. More energy is dissipated at the former contact because the relative sliding distance during each cycle is larger at this contact compared with the specimens-holders interface (460 mm compared with 20 mm). Thus at the beginning of relative sliding, the PA blocks remain stationary relative to the steel surface with the highest roughness (which also forms the interface with the largest apparent contact area). However, the relative sliding of the plastic specimens against the steel mating plates leads to an increase in the temperature and consequently in the friction force at this interface. Hence in the simulated practical situation where PA blocks are pressed between two surfaces, one surface with high surface roughness designed as the stationary surface, the other with lower roughness designed as sliding surface, after approximately 1.5 m sliding distance unwanted sliding at the stationary surface can occur. Holders are therefore necessary to fix the PA bearing blocks on one side. 4.1.4. The behaviour of PA 6 during positioning A proper bearing material tor the positioning of very large constructions has to exclude stick-slip and should behave well during stop-restart conditions. in order to investigate the behaviour of PA 6 during stoprestart, tests were performed with a restart of relative sliding after a dwell-time of 3 min. Fig. 6 shows the friction force for a normal load equal to 300 kN during 2 cycles following a standstill of 3 rain after 37 cycles of relative sliding. During the whole first stroke and at the beginning of the following
ill
L v~
t
o
.200
./o
6
z,
,/~
~so
¢i*OId~,rr~m tmml Fig. 6. Friction force vs. displacement during 2 cycles following a stand~:.;ll o f 3 rain after 37 cycles o f relative sliding: s t e e l - P A 6 ; P = 3 0 0 klq; s = 2 4 0 ram: c = 5 m m s - i: cycles 38-39.
strokes stick-slip is observed. Tests with a different number of cycles preceding the stop-restart were performed for a normal load of 825 kN. At restart after 5 cycles, heavy stickslip occurs resulting in brittle fracture of the PA blocks, at restart after I 0 cycles, stick-slip only occurs at the beginning of the first stroke (thereafter smooth sliding is observed) while no stick-slip at all is observed at restart after 37 cycles. No stick-slip is found for a normal load equal to 1440 Idq. Five test results with different stick-slip behaviour are gathered in Table 3. This table clearly demonstrates that stickslip is more pronounced when the temperature of the PA specimens is lower. It is thought that stick-slip disappears at higher temperatures owing to the increase in frictional and internal material damping with temperature (with frictional damping is meant the slope of the friction force-relative sliding velocity relation). Stick-slip is caused by a negative slope oftbe friction force-relative sliding velocity dependence. A decrease in this negative friction damping can eliminate stick-slip [ 1 3 ] . As PA 6 is a thermoplastic, it deforms more easily at interactions with the harder steel roughness asperities of the mating surface when its temperature is higher. Consequently, at higher temperatures the roughness asperities of the steel counter surface plough through the PA, rather than normal separation of the rubbing surfaces due to the forced tangential motion of inclinations of roughness peaks over one another ( asperity interlocking) occurs. The friction component resulting from the ploughing mechanism increases with growing relative sliding velocity because PA 6 is viscoelastic. The normal vibrations caused by asperity interlocking though, lead to a reduction in the friction force at higher velocities [ 14]. As the relative importance of the ploughing mechanism in the total friction force increases with growing temperature, a ~emperature rise results in larger frictional damping ( the slope of the friction force-relative sliding velocity relation becomes less negative or more positive) and therefore in a smaller risk of stick-slip. Moreover, the internal material damping of thermoplastics (PA 6) increases with the temperature above
F. Van De Velde. P. De Baets / Wear 200 (1997) 106-114
!12
Table 3 Con'elationbetweencontacttemperatureand stick-slipbehaviour Test specifications Load (kN)
Numberof cyclespreceedingstop--restart
Temperatureof bottomplate at restant(°C)
S:ick-slip
825 300
5 37
45 53
h
825 825 1440
10 37 -
60 85 -
no stick-sllp no stick-slip
"Very heavy stick-slip resulting in brit,.le fracture of the PA specimens. b Stick-slip during several strokes. c Stick-slip at the beginning of the first stroke.
the glass transition temperature [ 7 ]. As mechanical damping has a stabilizing effect on intermittent motion [ 15], a temperature rise of the PA blocks reduces the risk of stick-slip also by way of its influence on the material daraping. Fig. 6 shows that the mean friction force during stick-slip is lower than during smooth sliding. During stick-slip serious vibrations of the entire test rig are observed. From the literattire, it is known that both normal and tangential vibrations lead to a reduction in the friction force [ 14,16]. Therefore, the decrease in the mean friction force during intermittent motion can he attributed to the appearance of vibrations induced by stick-slip. For most tests, the friction force at restart is found to be higher than before stopping (see e.g. Fig. 2 for W= 825 kN ). This increase in the friction force during relative standstill results from creep of the PA, which leads to growth of the real contact area ( greater friction by adhesion ) and to deeper indentation of the steel roughness asperities into the plastic surface (greater friction by deformation). However, Fig. 2 shows different behaviour at restart for a normal load equal to 300 kN. In this case, the friction force is decreased owing to the immediate appearance of stick-slip.
4.2. Wear The total thickness reduction of the PA blocks during the tests with different normal load is shown in Fig. 3. The total thickness reduction results from the combined action of the following mechanisms: elastic and plastic deformation, thermal expansion and wear (material loss ). The initial deformation of the specimens on application of the normal load is excluded from the results of Fig. 4 by setting the output of the probes to zero after applying the load and before the start of relative sliding. The normal deformation of the specimens due to application of the test load is shown in Table 4. The deformation at 2250 kN could not be measured accurately because of the important creep of file plastic material under the high normal load. Fig. 4 shows that the thickness reduction of the plastic blocks occurs faster for higher normal loads. The maxima of the curves for the highest load follow from the elastic mate-
Table 4 Normal deformation of the PA specimens due to loading Normal deformation (itm)
Top specimen Bottom specimen
Initial load (kN)
Test load (kN)
30
300
825
2250
0 0
237 192
430 430
1660 1040
riai recovery at the reduction of the load from 2250 kN to 1440 kN. In order to distinguish real wear (material loss) from deformation, the weight loss of the plastic blocks during the different tests was also determined and wear rates are calculated from the weight loss and the total thickness reduction (including material loss, deformation and thermal expansion), The weight of the PA blocks is not stable in time as PA can absorb water from the air [ 17]. The influence of water absorption by the PA specimens on the measured material loss is minimized by weighing the specimens immediately before and immediately after the friction and wear tests (the duration of all tests was less than 70 min and water absorption by PA occurs very slowly [ 17 ] ). The wear results determined by both methods (weighing and thickness measurement ) are gathered in Table 5. The relative part of material loss in the total thickness reduction is also shown. Table 5 indicates that the specific material loss (true wear) does not change significantly with growing normal load. Consequently, the material loss is more or less proportional to the normal load. The wear rates calculated from the total thickness reduction though, increase more than proportionally with the normal load. Hence, the part of material loss in the total thickness reduction decreases with increasing normal load. Fig. 4 shows that the temperature of the PA blocks is lower for the test with a normal load of 300 kN than for the other two tests. It can be seen in Table 5 that the material loss during this test is larger than the measured thickness reduction suggests. For the tests with the higher normal loads (and the higher temperatures of the plastic specimen) though, the
F. Van De Velde, P. De Boers/Wear 209 (1997) 106-114
I 13
Table 5 Wear rates of the PA specimens Normal load (kN)
300 825 1440
Specimen Wearrate from total weightloss
Top Bottom Top Bottom Top Bottom
Wear rate from total thicknessreductioni
Weight loss Specificwear (g) (mm3N -' m-')
Thicknessreduction Specificwear (Izm) (mm~N -i m -I)
3.8 2.8 4.7 9.9 6.3
5.89 x I0 - 4 4.34x 10-4 2.58 X In -~ 5.44X 10-4 3.22 X l0 -4
?
?
190 55 1183 2424 1818 2649
5.06X 1.46× II.17 X 22.88 X 15.93 X 23.21X
10 -4 10-'* I0-'* I0 - 4 10 -4 10 -4
Part of material Wear rate during steady state loss in the total fromtotal thicknessreduction b thickness reduction (%) (mmkm -I ) (rams N-t m-I) 116.4 297.3 23.1 23.8 20.2
16A 8.4 50.5 115.9 126.8
8.2X 10 -4 4.2X 10 -4 9.18X 10 -4 21.1X 10 -4 13.2X 10 -4
?
171.9
17.9× 10-4
"Only the last measuredpoint is considered. b From linear regressionthroughthe last 'linear' part of each curve. material loss is found to be considerably smaller than expected based on the measured thickness reductions. It can therefore be concluded that the thermal expansion of the plastic specimens dominates at moderate temperatures (for P = 300 kN), leading to a smaller thickness redaction compared with the real material loss, while plastic deformation (creep) of the PA blocks under the huge normal loads results in a more important increase in the thickness reduction than in the material loss at higher temperatures (for P = 8 2 5 kN and P = 1440 kN). The PA blocks are considerably (plastically) deformed after the tests with the higher normal loads. They are about 5 mm shorter in the sliding direction, 10 mm longer crosswise and warped owing to thermal stresses. The actual wear (material loss) of the plastic blocks is caused by abrasion, which could be expected for such rough counterplates [ 18]. No material transfer duc to adhesion of the plastic material to the steel counter surfaces is observed. Fig. 3 and Table 2 and Table 5 indicate that the largest wear has always occurred at the interface with the roughest mating plate ( for the test with normal load equal to 300 kN the largest wear is found at the top specimen, while the bottom specimens have worn more during the other two tests). The literature supports the finding that abrasive wear of polymers increases considerably with growing surface roughness of the steel mating plates [ 7,18 ]. However, the differences between the thickness reduction of top and bottom specimens shown in Fig. 3 cannot result from different material loss by abrasion only. A greater deformation of the PA blocks has occurred at the interface with the roughest mating plate. This indicates that the temperature, and thus also the friction force which causes the temperature rise, is larger at the roughest mating plate (see also Section 4.1.3). The surface roughness of the steel counter plates and holders was measured before and after the tests. No significant change in the surface roughness due to wear is observed.
$. Conclusions !. PA 6 resists huge contact pressures and allows relative sliding against steel without damaging the steel surface.
2. The sliding velocity should be sufficiently low in order to restrict the temperature rise of the PA 6 bearing blocks. This temperature rise f~:~,altsin a considerable increase in the friction force, excluding the possibility of using PA 6 bearing blocks without fixation. 3. The wear (material loss) of PA 6 sliding against steel is moderate, even at very high contact pressures, and is more or less proportional to the normal load. 4. The thickness reduction of the PA 6 blocks due to deformation (creep) becomes dominant for huge contact pressures (and large temperatures), while it is not clearly observable at relatively small normal loads. 5. PA 6 sliding against steel is sensitive to stick-slip motion, which complicates accurate positioning, but the resulting vibrations can also damage the bearing blocks or even the construction which has to be manipulated. The use of PA 6 in combination with proper lubricants (e.g. grease, PTFE) could eliminate the stick-slip problem and is a topic of further research.
References [ I ] P. De Baets. Wear simulationof a sliding system by means of largescale specimentesting, We, r, 173 (1994) 65-74. [21C. Dekouinck. P. De Boers and F. Van De Velde, A highlight on tribologicalresearch: friction and wear devices. Enr. J. Mech. Eng. M..40(3) (1995) 154--156. 13[ Y. Yamaguchi.Tribology of Plastic Materials, TfibologySeries, Vol. 16, Elsevier.Amsterdam, 1990. [4] MJ. Neale. Tribology Handbook, Butterworth,London. 1973. [5] M.B. Peterson and W.O. Wirier, Wear Control Handbook. ASMEo New York, 1980. [6] B. Bushan and B.K. Gapta, Handbook of Tribology. Materials. Coatings and Surface Treatments, McGraw-Hill.New Ymk, 1991. [7l H. Uetz and J. Wiedemeyer,Tribologie der Polyraere. Carl Hanser Verlag, Munich. 1985. [8] M. Watanabe, The frictional propertiesof nylon. Wear. 12 (1968) 185. 191 I.M. Hutchings, Trlbology. Friction and Wear of Engineering Materials.Edward Arnold, London, 1992. [ 10l F.P. Bowden and D. Tabor. The Frictionand Ialbricationof Solids, Clarendon Press,Oxford, 1950. l II ] J.F. Archard. Elasticdeformation and the laws of friction.Proc. R. Soc. London. Ser.A. 243 (1957) 190.
114
F. Van De Velde. P. De Baets l Wear 209 (1997) 106-114
[12] J.A. Greenwood and J.B.P. Williamson, Contact of nominally fiat surfaces,Proc R. Sac. London. Set. A, 295 (1966) .'~0--319. [13] G. Cockerhamand G.R. Synunous, Stability criterion for stick-slip motion using a discontinuous dynamic friction model, Wear, 40 (1976) 113--120. [14] D.M.Tolstoi,Significanceof die uormaldegrecof freedomand natural normal vibrationsin contact friction, Wear, I0 (1967) 199-213. [ 15] B.V. Derjaguin,V.E. Push and D.M. Tolstoi, A ~ of stick-slip sliding of solids, Proc. Conf. on Lubrication and Wear, Inst. Mech. Eng., London, 1957,pp. 255-268. [ 16] W. Lenkiewicz, The sliding friction process.effect of external vibrations,Wear, 13 (1969) 99-108. [ 17] G. Schreyer, Konstruieren mit Kunststoffen, Vol. 2, Carl Hauser Verlag,Munich, 1972. 118l D. Dawson. J.M. Challonand K. Holmes,The wear of non-metallic materials, Proc. 3rd Leeds-Lyon Syrup. on Tribology, Mech. Eng. Publ., 1976,pp. 99-102.
mograp~es Frederik Van De Velde graduated in mechanical engineering in 1993 at the University of Gent, Belgium. The same year
he became an assistant at the Laboratory of Machines and Machine Construction at the same university, where he is involved in lecturing on production methods and machining. His research is concentrated in the area of dry and boundary lubricated friction and wear, with special reference to instabilities during machining, for example stick-slip at machine tool slideways. He is also involved in several industrial research projects concerning bearing material and lubricant testing. Patrick De Baets graduated in mechanical engineering in 1989 at the University of Gent, Belgium. That year he joined the Laboratory of Machines and Machine Construction at the same university, and in 1995 received his PhD degree. As doctor-assistant he is involved in lecturing and research on the dimensioning of machine elements and tribology. His research interests are mainly in the area of dry friction and wear, especially fretting wear, large-scale wear testing and industrial research. He is a member of the Mechanical Engineering and Transport Mechanics Group of the Flemish Organization of Engineers.