On the effect of starved lubrication on elastohydrodynamic (EHL) line contacts

Accepted Manuscript On the effect of starved lubrication on elastohydrodynamic (EHL) line contacts M. Ebner, M. Yilmaz, T. Lohner, K. Michaelis, B.-R. Höhn, K. Stahl PII:

S0301-679X(17)30292-X

DOI:

10.1016/j.triboint.2017.06.004

Reference:

JTRI 4764

To appear in:

Tribology International

Received Date: 2 September 2016 Revised Date:

31 May 2017

Accepted Date: 2 June 2017

Please cite this article as: Ebner M, Yilmaz M, Lohner T, Michaelis K, Höhn B-R, Stahl K, On the effect of starved lubrication on elastohydrodynamic (EHL) line contacts, Tribology International (2017), doi: 10.1016/j.triboint.2017.06.004. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

ACCEPTED MANUSCRIPT On the Effect of Starved Lubrication on Elastohydrodynamic (EHL) Line Contacts M. Ebner1, M. Yilmaz1*, T. Lohner1, K. Michaelis1, B.-R. Höhn1, K. Stahl1 1 Gear Research Centre (FZG), Technical University of Munich (TUM), Boltzmannstraße 15, 85748 Garching, Germany.

Abstract

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Highly-loaded EHL contacts in gears often operate under starved lubrication, albeit unintentionally. This is frequently correlated to mixed lubrication regimes resulting in reduced lifetime. Consequently, the aim of this study is to investigate the effect of starved lubrication on the operating behaviour of EHL contacts. Different amounts of initial oil volumes and surface structures were investigated by measuring the frictional behaviour at the FZG twin-disk test rig. Analytically calculated gap fill factors for initial oil volumes were assigned to the experiments. Results show that a very small amount of initial oil is sufficient for lubrication and different operating behaviours for peripherally ground, axially ground and polished surfaces. All experiments were accompanied by surface photos and roughness measurements.

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Keywords: minimised lubrication, starved lubrication, rolling-sliding contacts, elastohydrodynamic lubrication (EHL) *Corresponding author: Mustafa Yilmaz ([email protected])

1. INTRODUCTION

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Highly-loaded rolling-sliding contacts in industrial and automotive transmission gears, as well as roller bearings, usually require spray or dip lubrication in order to avoid severe mixed and boundary lubrication regimes and to dissipate frictional heat. When minimizing the amount of circulating lubricant, the no load losses (e.g. splash losses) can be reduced significantly. However, with an insufficient supply of oil, the inlet zone (see Figure 1) of elastohydrodynamically lubricated (EHL) contacts is not fully flooded. Consequently, the EHL contact is under starved lubrication.

inlet zone

solid body 2

lubricant solid body 1

pressurized zone

outlet zone

Figure 1: Highly-loaded EHL line contact under starved lubrication

Figure 1 shows schematically characteristic distributions of pressure and temperature of a highly-loaded EHL line contact with smooth surfaces under starved lubrication. The contact zone is divided into an inlet, pressurized and outlet zone. The inlet zone and outlet zone are partially flooded and the pressurized zone is 1

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fully flooded. The inlet distance xm (also called the “meniscus”) defines the point where the lubricant film height of the surfaces flow together due to geometry and elastic deformation of the EHL contact (Hamrock and Dowson [1], Cann et al. [2]). Starved lubrication exists if the inlet zone is partially flooded (see Figure 1) and the film thickness in the pressurized zone has decreased compared to the achievable film thickness in the pressurized zone in the case of a fully flooded inlet zone (Hamrock and Dowson [1], Cann et al. [2], Goksem and Hargreaves [3]).

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Wolveridge et al. [4] analyse the influence of the inlet distance on film thickness and demonstrates a decrease in the film thickness under starved lubrication. Analytical equations for the minimum and central film thickness under starved lubrication are available from Hamrock and Dowson [1, 2] for EHL point contacts and from Goksem and Hargreaves [3] for EHL line contacts. Both quantify the decrease in film thickness in the pressurized zone (see Figure 1) due to starved lubrication. Prexler [5] shows that a minimum amount of initial oil volume is enough to lubricate EHL contacts, and suggests the presence of hydrostatic stresses. Otto [6] investigates the influence of minimised dip lubrication on the flank load carrying capacity of gears. His results show an increased risk of scuffing, as well as higher bulk temperatures with decreasing oil volume. Masjedi and Khonsari [7] quantify the decrease in film thickness due to starved lubrication by EHL simulations of point and line contacts based on reducing the lubricant mass flow rate. Chevalier et al. [8] numerically analyse the influence of different gap fill factors of EHL point contacts on the film thickness and show a decrease in film thickness under starved lubrication. Kostal et al. [9] confirm the results of Chevalier et al. [8] by film thickness measurements.

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The literature review indicates a shortage of systematic experimental investigations on the operating behaviour of highly-loaded EHL contacts under starved lubrication. Therefore, the aim of this study is to investigate the effect of starved lubrication of different surface structures on the operating behaviour of highly-loaded EHL line contacts by measuring coefficients of friction at the FZG twin-disk test rig and consequent surface analyses. Furthermore, gap fill factors are calculated analytically and assigned to the experiments.

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2. GAP FILL FACTOR AND MENISCUS

The gap fill factor θin(x) in the inlet zone and the meniscus xm depend directly on the initial oil volume. θin is defined after Bartel [10], Damiens et al. [11], Popovici [12], Venner et al. [13] and according to Figure 1: θin(x) = (hlub,in,1(x) + hlub,in,2(x)) / hin(x)

(Eq 1)

The initial oil volume is distributed evenly over the circumferential running surface of the disk. Lubricant does not flow off sidewards in the axial direction of disk. Film thickness height in the pressurized zone is neglected.

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The gap fill factor and meniscus are derived from geometrical considerations based on Hertz’ theory [14]. Two cylindrical disks, as used in the subsequent experimental investigations, are considered (see Figure 2, right). The calculation of θin(x) is based on the following assumptions:

Based on these assumptions, Figure 2 shows the calculated gap fill factor θin at x = -4.5·bH and meniscus xm for different amounts of initial oil volume Voil and a Hertzian pressure of pH = 1200 N/mm2. The disk material considered is 16MnCr5E and bH is the Hertzian contact half-width. The gap fill factor is calculated at θin(x = -4.5·bH), because this position is often used as the left-hand border of the lubricant domain in TEHL simulations (Lohner et al. [15], Habchi et al. [16]).

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Figure 2: Calculated gap fill factor θin(x = -4.5·bH) and meniscus xm in relation to the initial oil volume Voil for pH = 1200N/mm² (left) and disk as used at the FZG twin-disk test rig (right)

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Figure 2 shows a linear dependence of θin(x = -4.5·bH) and Voil. For Voil > 0.11 ml, the gap fill factor is θin(x = -4.5·bH) = 1, so that the inlet zone is fully flooded at xm = -4.5·bH. The meniscus xm increases degressively with increasing Voil. For Voil = 0.11 ml, the meniscus is xm = -4.5·bH.

3. EXPERIMENTAL SETUP AND PARAMETERS

FZG Twin-Disk Test Rig

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The FZG twin-disk test rig according to Michaelis [17] (with some modifications) is used for all experimental investigations. The following description of the test rig is based on the work and formulations of Ebner et al. [18] and Lohner et al. [19].

The FZG twin-disk test rig is widely used for basic investigations on rolling-sliding contacts of machine elements e.g. gears or rolling bearings. The mechanical layout is shown in Figure 3. upper disk v2, ϑM2

pivot arm

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load cell (normal force FN)

oil injection ϑoil

pivot

steel sheet

mount pneumatic cylinder

lower skid

lower disk v1, ϑM1

load cell (friction force FR)

Figure 3: FZG twin-disk test rig

The upper and lower disk are press-fitted onto shafts, which are separately moved by two speed-controlled electric motors. This allows continuous variation of the sliding velocity vg (see equation (2)), meaning the

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difference between the surface velocities (v1, v2), and of the sum velocity vΣ (see equation (3)), meaning the sum of the surface velocities of the disks.

vg = v1 – v2

(Eq 2)

vΣ = v1 + v2

(Eq 3)

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The normal force FN in the contact is continuously applied by a pneumatic cylinder via a pivot arm where the upper disk is mounted. The lower disk is mounted in a skid which is attached to the frame by thin steel sheets. The skid is supported laterally by a load cell so that the friction force FR in the disk contact for sliding velocities vg ≠ 0 m/s can be measured as reaction force with hardly any displacement of the skid. The coefficient of friction µ is calculated according to equation (4):

µ = FR / FN

(Eq 4)

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Normal force FN, friction force FR, surface velocities v1 and v2, as well as bulk temperatures of the upper disk ϑM1 and lower disk ϑM2 are measured. A spray lubrication unit with heating and cooling possibilities enables to provide lubricant at the desired temperature (ϑoil) to be injected directly into the inlet region of the disk contact. An oil filter with a filtration degree of 10 µm is included in the main circuit of the spray lubrication unit. ϑM1 and ϑM2 are recorded by a Pt100 resistance temperature sensor 5 mm below the surface of the disk. All experiments are carried out under line contact conditions with cylindrical disks having a diameter of 80 mm and width of 5 mm. The disks are made of case hardened steel (16MnCr5E). The disk geometry is shown in Figure 2 (right).

Test Parameters

Table 1 shows the peripherally ground, axially ground and polished surface structures used for the experimental investigations. Besides for one reference test run under dry lubrication, same types of surface structures are paired. Thereby, a new pair of disks is used for each test run.

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Table 1: Light microscope pictures of peripherally ground, axially ground and polished surfaces before test run Axially ground

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Peripherally ground

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1 mm

Ra = 0.10 – 0.15 µm

Polished

1 mm

Ra = 0.2 – 0.3 µm

1 mm

Ra < 0.01 µm

For surface roughness measurements, the profile method is used with a measurement length of Lt = 4 mm for peripherally ground and polished surfaces and Lt = 4.8 mm for axially ground surfaces. Correspondingly, the cut-off wavelength is λc = 0.08 mm for polished surfaces and λc = 0.80 mm for axially ground and peripherally ground surfaces. For ground surfaces, roughness measurements are performed perpendicular to the grinding direction. In order to apply defined amounts of initial oil volumes to the running surfaces of the lower disks, a mechanical dosing unit with a precision of 0.002 ml (in the case of water) is used. Thereby, the initial oil volume is applied piecewise at multiple positions around the running surface of the lower disk. An ISO VG 100 mineral oil without additives (FVA3 according to Laukotka [20]) is used as lubricant. The initial oil volume is applied only one-time before the oil distribution run for all experiments. Table 2 shows the test program.

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Table 2: Test program

Hertzian pressure pH in N/mm2

Sum velocity vΣ in m/s

Sliding velocity vg in m/s

Running time t

Oil distribution run

325

1

0.01

60 s

Test run

1200

4

0.4

3 h if µ < 0.13

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The oil distribution run lasts 60 seconds and distributes the applied initial oil volume evenly on the circumference of the running surfaces of the lower and upper disk. Thereby, a very low Hertzian pressure (pH = 325 N/mm2) and sliding velocity (vg = 0.01 m/s) are set. Preliminary tests have not revealed any changes to the disks’ surface structures.

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Following the oil distribution run, the two disks are separated and the sum velocity of vΣ = 4 m/s is set. The load of pH = 1200 N/mm2 is then applied. The maximum running time of each test run is three hours, where test runs with stationary operating behaviour have shown almost constant values for coefficients of friction and bulk temperatures. The test run is aborted if the coefficient of friction is higher than µ > 0.13. Due to the back coupling of coefficient of friction and bulk temperature, stationary values are only observed when the heat balance of the FZG twin-disk test rig is achieved. This is when the frictional heat in the highly-loaded rolling-sliding contact (heat source) is equal to the dissipated heat due to heat radiation, heat conduction and convection (heat sinks). The test program in Table 2 is conducted for the following initial oil volumes Voil. Thereby, it is assumed that a value of θin(x = -4.5·bH) < 1 results in starved lubricated EHL contacts.

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0 ml (θin(x = -4.5·bH) = 0, dry lubrication) with solvent-cleaned running surfaces (no oil distribution run) 0.02 ml (θin(x = -4.5·bH) = 0.19), 0.04 ml (θin(x = -4.5·bH) = 0.38), 0.06 ml (θin(x = -4.5·bH) = 0.57), 0.08 ml (θin(x = -4.5·bH) = 0.76), 1.6 ml (θin(x = -4.5·bH) = 1) Spray lubrication

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4. EXPERIMENTAL RESULTS Dry Lubrication

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Table 3 shows light microscope pictures and arithmetic mean roughness Ra values for the peripherally ground and polished surfaces before and after a four-minute test run under dry lubrication. The test was aborted due to measured coefficients of friction of µ > 0.13.

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Table 3: Light microscope pictures and arithmetic mean roughness Ra of running surfaces before and after the test run at vΣ = 4 m/s, vg = 0.4 m/s, FN = 3920 N (pH = 1200 N/mm²), Voil= 0 ml (dry lubrication)

Upper disk

Peripherally ground – polished Before test run After test run

1 mm

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1 mm

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Lower disk

Ra = 0.15 µm Ra = 1.01 µm Running direction

1 mm

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1 mm

Ra = 0.01 µm Ra = 0.27 µm Running direction

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The light microscope pictures after the test run show a clear scuffing failure with black and shattered upper and lower disk surfaces. Ra increases from Ra = 0.15 µm to Ra = 1.01 µm for the upper disk and from Ra = 0.01 µm to Ra = 0.27 µm for the lower disk.

Starved Lubrication

Figure 4 displays the measured coefficients of friction µ (top) and related bulk temperatures ϑM2 (bottom) over running time for test runs with initial oil volumes of Voil = 0.04 ml and Voil = 0.08 ml for peripherally ground, axially ground and polished surfaces. 0.08

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Voil = 0.04 ml
θin(x = - 4.5·bH) = 0.38 Voil = 0.08 ml
θin(x = - 4.5·bH) = 0.76

0.075

peripherally ground

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coefficient of friction µ

0.07 0.065

0.06 0.055

axially ground

0.05 0.045 polished 0.04 0 0

0.5

1

1.5 running time t in h

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2

2.5

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130

Voil = 0.04 ml
θin(x = - 4.5·bH) = 0.38 Voil = 0.08 ml
θin(x = - 4.5·bH) = 0.76

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peripherally ground

100 90 80 70

polished

axially ground

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bulk temperature ϑM2 in °C

110

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1

1.5

2

2.5

3

running time t in h

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Figure 4: Measured coefficients of friction (top) and related bulk temperatures (bottom) with FVA3 at vΣ = 4 m/s, vg = 0.4 m/s, FN = 3920 N (pH = 1200 N/mm2), Voil = 0.04 ml, Voil = 0.08 ml

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The coefficients of friction and related bulk temperatures show stationary values after an initial running time of approximately 0.5 hours for an initial oil volume of Voil = 0.04 ml and with axially ground and polished surfaces. After three hours’ running time, the coefficient of friction is µ = 0.049 (ϑM2 = 75 °C) for axially ground surfaces and µ = 0.043 (ϑM2 = 85 °C) for polished surfaces. Peripherally ground surfaces do not show any stationary values for the coefficient of friction and bulk temperature. The test run is aborted due to µ > 0.13 after 2.3 hours’ running time.

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For an initial oil volume of Voil = 0.08 ml, very similar coefficients of friction and bulk temperatures are measured for axially ground and polished surfaces. The deviation of the coefficient of friction between Voil = 0.04 ml and Voil = 0.08 ml is ∆µ < 0.001 and therefore within the measurement error of the FZG twindisk test rig (Prexler [5]). For the peripherally ground surfaces, the trend of the coefficient of friction and its related bulk temperature differs from Voil = 0.04 ml. For Voil = 0.08 ml, the test run succeeds in reaching a running time of three hours, although the coefficient of friction increases from 0.052 to 0.068 before decreasing again from 0.068 to 0.064. Consequently, no stationary operating behaviour is observed, as it is the case for Voil = 0.04 ml. All coefficients of friction and bulk temperatures in Figure 4 are representative. Table 4 shows light microscope pictures and arithmetic mean roughness Ra values for peripherally ground, axially ground, and polished surface structures before and after the test run with an initial oil volume Voil = 0.04 ml. Table 4: Light microscope pictures and arithmetic mean roughness Ra values of running surfaces before and after test runs at vΣ = 4 m/s, vg = 0.4 m/s, FN = 3920 N (pH = 1200 N/mm²), Voil = 0.04 ml (0: before test, e: after test)

Before test run

Peripherally ground

Axially ground

1 mm

Running direction

1 mm

Running direction

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1 mm

Running direction

1 mm

Ra0 = 0.27 µm Rae = 0.22 µm

1 mm

1 mm

Ra0 = 0.21 µm Rae = 0.19 µm

1 mm

Ra0 = 0.01 µm Rae = 0.01 µm

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Ra0 = 0.11 µm Rae = 0.12 µm

Ra0 = 0.01 µm Rae = 0.01 µm

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Lower disk

Ra0 = 0.10 µm Rae = 0.10 µm

1 mm

1 mm

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Upper disk

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The light microscope pictures of peripherally ground surfaces after the test run with Voil = 0.04 ml show a scuffing stripe with a width of approximately 0.5 mm on the running surface of the upper and lower disk. Thereby, Ra values of the upper and lower disk before and after the test run are approximately the same. The running surfaces of axially ground disks show light circumferential marks in the centre of the surface (perpendicular to ground direction). The corresponding Ra values of the upper and lower disk decrease slightly. The light microscope pictures of polished disks in Table 4 show circumferential marks on the running surface of the upper and lower disk after test run, although the Ra values of the upper and lower disk do not change.

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From Starved to Spray Lubrication

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Figure 5 shows an overview of the measurement results for the coefficient of friction µ and its related bulk temperature ϑM2 at the beginning of test runs and after three hours of running time for different surface structures in relation to the initial oil volume. For completeness, the calculated gap fill factors (see chapter 2) are added. Furthermore, the measured coefficients of friction and related bulk temperatures for spray lubrication with fully flooded conditions in the inlet zone (see Figure 1) are added to Figure 5. In order to do this, the bulk temperature has been adjusted to the bulk temperature for Voil = 0.08 ml by adjusting the oil inlet temperature.

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3h

area of constant coefficients of friction (linear regression)

peripherally ground axially ground polished

0.075

area of damage

0.065 0.06 0.055 0.05

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coefficient of friction µ

0.07

area of damage

0.045

area of damage

0 Voil = 0 θin = 0

Voil = 0.02 θin = 0.19

Voil = 0.04 θin = 0.38

Voil = 0.06 θin = 0.57

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0.04

Voil = 0.08 θin = 0.76

Voil = 1.6 θin = 1.0

spray lubrication

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test begin

3h peripherally ground axially ground polished

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100 90

70

area of damage area of damage

60 50 40 30

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20 Voil = 0 θin = 0

area of damage

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area of constant coefficients of friction (linear regression)

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bulk temperature ϑM2 in °C

110

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initial oil volume VOil in ml / calculated gap fill factor θin(x = - 4.5·bH)

Voil = 0.02 θin = 0.19

Voil = 0.04 θin = 0.38

Voil = 0.06 θin = 0.57

Voil = 0.08 θin = 0.76

Voil = 1.6 θin = 1.0

spray lubrication

initial oil volume VOil in ml / calculated gap fill factor θin(x = - 4.5·bH)

Figure 5: Overview of measured coefficients of friction (top) and related bulk temperatures (bottom) with FVA3 at vΣ = 4 m/s, vg = 0.4 m/s, FN = 3920 N (pH = 1200 N/mm2)

Below, the results are first discussed according to the type of surface structure and then they are compared with each other. Polished: The coefficients of friction and bulk temperatures after three hours’ running time are unchanged from Voil = 0.04 ml to Voil = 1.6 ml (µ = 0.043, ϑM2 = 83 °C). This area is denoted by area of constant coefficient of friction and bulk temperature. Below Voil = 0.04 ml, the test run does not reach three hours of running time. This area is called area of damage. The coefficient of friction under spray lubrication is µ = 0.042 at the same bulk temperature compared to that in the area of constant coefficients of friction and bulk temperatures. Hence, almost the same coefficient of friction is measured for both starved and spray

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lubrication. The difference between the coefficients of friction under starved lubrication at the beginning of the test run compared to after a running time of three hours is higher the lower the initial oil volumes are. Axially ground: The area of constant coefficient of friction and bulk temperature after three hours’ of running time ranges from Voil = 0.02 ml to Voil = 0.08 ml (µ = 0.049, ϑM2 = 75 °C). The area of damage starts below Voil = 0.02 ml. The coefficient of friction under spray lubrication is µ = 0.045 at the same bulk temperature compared to that in the area of constant coefficients of friction and bulk temperatures and subsequently lower than under starved lubrication. The trend of the difference between the coefficients of friction under starved lubrication at the beginning of the test compared to a running time of three hours is comparable to that with polished surfaces. Peripherally ground: No area of constant coefficient of friction and bulk temperature is found. The area of damage starts below Voil = 0.06 ml. The coefficient of friction under spray lubrication is µ = 0.055 at the same bulk temperature compared to that at Voil = 0.06 ml and Voil = 0.08 ml (ϑM2 = 92 °C). The coefficients of friction under starved lubrication at the beginning of the test run are lower compared to the coefficients of friction after three hours of running time.

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To sum up, the following main points can be drawn for the starved lubrication regime: • Polished surfaces show the lowest coefficients of friction of µ ≈ 0.043 (ϑM2 ≈ 83 °C). • Peripherally ground surfaces feature the highest coefficients of friction (0.064 < µ < 0.07, 92 °C < ϑM2 < 114 °C). • Axially ground surfaces have lower coefficients of friction (µ ≈ 0.049, ϑM2 ≈ 75 °C) compared to peripherally ground surfaces, although the Ra values of the axially ground surfaces are higher (see Table 1). • The coefficients of friction of polished surfaces at the beginning of the test run are higher than the corresponding coefficients of friction of axially ground surfaces. • The bulk temperatures of axially ground surfaces in the area of constant bulk temperature after three hours’ running time are lower than the corresponding bulk temperatures of polished surfaces.

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Investigations on the repeatability of the experiments in the area of damage and in the area of constant coefficients of friction in case of polished and axially ground surfaces confirm the results presented. For purposes of illustration, the results are not included in Figure 5. A new pair of disks is used for each repetition test run. Four additional test runs for polished surfaces with Voil = {0.005; 0.01; 2x 0.02} ml, one additional test run for axially ground surfaces with Voil = 0.01 ml and two additional test runs for peripherally ground surfaces at Voil = {2x 0.04} ml confirm the area of damage. An additional test run each for polished surfaces at Voil = 1.6 ml and axially ground surfaces at Voil = 0.02 ml show at three hours running time measured coefficients of friction of µ = 0.043 and µ = 0.049, respectively and bulk temperatures of ϑM2 = 83 °C and ϑM2 = 75 °C, respectively. As in case of peripherally ground surfaces no area of constant coefficients of friction is found, the characteristic trend during the three hours’ running time as well as the absolute values at three hours running time vary slightly more. For example, an additional test with peripherally ground surfaces at Voil = 0.06 ml results in a coefficient of friction of µ = 0.065 and a bulk temperature of ϑM2 = 105 °C.

5. DISCUSSION

Doleschel [21] suggests a correlation between the relative film thickness λrel and the fluid load portion ξ with the surface structure factor Xos.

ξ = 1 - (1 - Xos · λrel)2 ξ=1

for λrel < 2 for λrel ≥ 2

(Eq 5a) (Eq 5b)

Here, the relative film thickness λrel is the quotient of the minimum lubricant film thickness and the average arithmetic mean roughness Ra of the surfaces:

λrel = hm / (0.5 · (Ra1 + Ra2))

(Eq 6)

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As the polished surfaces feature a very low Ra value, λrel is significantly higher than for ground surfaces so that the fluid load portion ξ is also higher than for ground surfaces. Hence, polished surfaces operate in a much less severe lubrication regime. The coefficient of friction measured is the result of a fluid and solid coefficient of friction and correlated by the fluid load portion ξ:

µ = ξ · µ f + (1- ξ) · µ s

(Eq 7)

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As for lubricant FVA3, the solid coefficient of friction µ s is higher than the fluid coefficient of friction µ f (Lohner et al. [18]), the measured coefficient of friction decreases with increasing λrel. This explains the lower coefficients of friction of polished surfaces compared to axially and peripherally ground surfaces. Note that the light microscope pictures at the end of the test run indicate mixed to boundary lubrication regime for all surface structures (see Table 4).

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The observed coefficients of friction in the starved lubrication regime are higher than for spray lubrication. In this context, Goksem and Hargreaves [3] show a reduced lubricant film thickness under starved lubrication. Hence, according to (Eq 5) λrel and ξ decrease. Based on the comments above, the coefficient of friction is higher in the starved lubrication regime.

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In comparison of axially and peripherally ground surfaces, axially ground surfaces should have a better ability to transport lubricant to the disk contact due to grooves perpendicular to the running direction. In this context, Kreil [22] finds better film formation capabilities of axially ground surfaces compared to peripherally ground surfaces. This correlates to the start of the area of damage at smaller initial oil volumes for axially ground surfaces compared to peripherally ground surfaces and comes along with considerable higher coefficients of friction for peripherally ground surfaces (cf. Figure 5), which can be again traced back to low fluid load portions.

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As very small oil volumes are applied only initially to the disks, also other effects influence the observed experimental characteristics. The wetting surface and therefore the adhesive forces of rough surfaces are reported to be higher compared to smooth surfaces (Bobe [23]). Hence, the ability of the lubricant to maintain contact with the surface is superior for ground surfaces compared to polished surfaces. This suggests that the possible loss of lubricant due to centrifugal forces is lower for ground surfaces. Hence, a correlation to the start of the area of damage at smaller initial oil volumes for axially ground surfaces compared to polished surfaces may be proposed (cf. Figure 5). Note that beside centrifugal forces, lubricant can also be lost from the disk surface by flow in axial direction of the disks. Differences between the surface structures could not be quantified experimentally.

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The profile of coefficient of friction over the time in Figure 4 shows similar profiles for axially ground and polished surfaces: an initial decrease of the coefficient of friction is followed by an almost stationary operating behaviour. In case of polished surfaces, the initial decrease of the coefficient of friction can be mainly traced back to the increasing bulk temperature, whereas for axially ground surfaces the initial decrease is additionally related to the fluid load portion in mixed lubrication regime. The profiles of the coefficient of friction over time of peripherally ground surfaces are different. The initial increase of the coefficient of friction can be related to the poor lubricant transport ability and film formation capability of peripherally ground surfaces, which immediately results in very low fluid load portions and superior influence of the solid coefficient of friction of FVA3. This is associated with a not clearly pronounced stationary operating behaviour after three hours running time. Furthermore, the coefficient of friction and the bulk temperature after three hours running time is found to depend on the profiles during the three hours running time, which vary more for peripherally ground surfaces. The bulk temperatures of axially ground surfaces in the area of constant bulk temperature after three hours’ running time are lower than the corresponding bulk temperatures of polished surfaces. At first sight, this is not in accordance with the measured coefficient of friction as it is lower for polished surfaces, resulting in less frictional heat. However, at second sight two essential differences between experiments with polished and axially ground surfaces are present. First, the coefficient of friction at the beginning of test runs is higher for polished surfaces than for axially ground surfaces. Subsequently, the trend of the bulk temperature of 11

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polished surfaces increases more steeply in the first operating period and may remain at higher level. Second, heat transfer in the EHL contact differs due to the different lubrication regimes for polished and ground surfaces. As in the experiments frictional heat is mainly removed by conduction, this may lead to different bulk temperatures below the surface. Further studies are required to fully understand the mainly phenomenological investigations. This may include an extended experimental programme as well as tribosimulations and experiments at tribometers based on optical interferometry.

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6. CONCLUSION

This study focuses on the operational behaviour of highly-loaded EHL contacts under starved lubrication considering different surface structures and initial oil volumes. Thereby, the friction coefficient is considered as the relevant measurand. The main conclusions drawn from this work are:

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Under dry lubrication, the experiment confirms, as expected, that a minimal amount of lubricant is necessary to ensure the functionality of highly-loaded EHL contacts. Very small amounts of initial oil volume are sufficient for lubrication of highly-loaded EHL contacts. Varying the initial oil volume has no influence on the coefficient of friction of polished and axially ground surfaces over the parameter range investigated. The coefficient of friction in starved lubrication regime at the beginning of the test run depends on the initial oil volume. Peripherally ground surfaces do not achieve stationary operating behaviour under starved lubrication regime within a running time of three hours. For the surface structures investigated, the area of damage starts at different amounts of initial oil volume. The coefficient of friction for ground surfaces under starved lubrication is higher than for spray lubrication.

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7. ACKNOWLEDGEMENT

8. REFERENCES

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The FZG would like to thank the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation, HO 1339/49-1), for their kind sponsorship of this research.

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Bullet Points Operational behaviour of EHL contacts under starved lubrication was investigated

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Coefficient of friction was measured for different surface structures

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Very small initial oil volumes were found to be sufficient for lubrication

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Required minimal initial oil volume was found to differ with surface structures

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Highest coefficient of friction was found for peripherally ground surfaces

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Lehrstuhl für Maschinenelemente - Forschungsstelle für Zahnräder und Getriebebau Technische Universität München, Boltzmannstr. 15, 85748 Garching www.fzg.mw.tum.de