Analysis of the effect of different types of additives added to a low viscosity polyalphaolefin base on micropitting

Wear 322-323 (2015) 238–250

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Analysis of the effect of different types of additives added to a low viscosity polyalphaolefin base on micropitting E. de la Guerra Ochoa a,n, J. Echávarri Otero b, E. Chacón Tanarro b, J.M. Munoz-Guijosa b, B. del Río López c, Cristina Alén Cordero d a

Talgo R&D, Patentes Talgo, Paseo del Tren Talgo 2, Las Rozas, 28290 Madrid, Spain Grupo de Investigación en Ingeniería de Máquinas, Universidad Politécnica de Madrid, José Gutiérrez Abascal 2, 28006 Madrid, Spain c Física Aplicada e Ingeniería de los Materiales, Universidad Politécnica de Madrid, José Gutiérrez Abascal 2, 28006 Madrid, Spain d Teoría de la Señal y Comunicaciones, Escuela Politécnica Alcalá de Henares, Plaza de San Diego s/n, Alcalá de Henares, 28801 Madrid, Spain b

art ic l e i nf o

a b s t r a c t

Article history: Received 11 August 2014 Received in revised form 10 November 2014 Accepted 13 November 2014 Available online 21 November 2014

This article explores the influence of the appearance of micropitting in different types of additives added to a polyalphaolefinic low-viscosity base oil, namely a polyalphaolefin PAO6. Additives commonly used in mechanical transmissions by gears were used, i.e., extreme pressure, anti-wear and friction modifier. For the study, we have conducted a series of tests on a disc machine under various operating conditions. The temperature and surface roughness of the test specimens have been changed in order to study different lubricant specific film thicknesses, as the last is one of the most influential parameters on the appearance of micropitting. Tests have shown the important influence of the additives, and their concentration, in the development of micropitting and their associated effects, depending on the operating conditions of the contact. Along with the study of the effect of additives, the influence of specific film thickness on the friction coefficient, surface appearance, conditions of lubrication, and wear was also studied. & 2014 Elsevier B.V. All rights reserved.

Keywords: PAO Rolling contact fatigue Sliding wear Sliding friction Lubricant additives Gears

1. Introduction Surface fatigue is a phenomenon associated with mechanical contacts under cyclic loading. There are different scales depending on the size of the defect: micropitting, (macro)pitting and spalling [1]. This article studies micropitting in lubricated line contacts. Micropitting is defined by ASTM [2] as a form of surface damage in rolling contacts consisting of numerous pits and associated cracks on a scale smaller than that of the Hertz elastic contact half-width. An example of this phenomenon is shown in Fig. 1, after a test with an approximated value of 100 μm of the Hertz half-width. The appearance of micropitting in mechanical contacts is influenced by multiple factors [3,4], such as contacting materials, their roughness, types of lubricants (both base oil and additives) and factors associated with operating conditions, namely load, temperature, average velocity, and sliding velocity. Of all these possible causes, specific film thickness is the key parameter [5,6]. It compares the lubricant film thickness to the n

Corresponding author. E-mail addresses: [email protected] (E. de la Guerra Ochoa), [email protected] (J.E. Otero), [email protected] (B. del Río López). http://dx.doi.org/10.1016/j.wear.2014.11.014 0043-1648/& 2014 Elsevier B.V. All rights reserved.

combined surface roughness of the contact surfaces and serves to find the lubrication regime in which the contact operates. In the case of low specific film thicknesses, high stresses are concentrated on the surface of the material. This leads to an early, and localized, failure on the surface. Therefore, all parameters that improve film formation improve resistance to micropitting. These parameters can be related to the lubricant (high viscosity, high viscosity–pressure coefficient, etc.), operating conditions (low temperature of the lubricant, high average velocity, low slide-toroll ratios, reduced loads, etc.) or materials and manufacturing process (reduced roughness, etc.). Once fluid film lubrication is reached the risk of micropitting is reduced. However, the rheological behavior of the lubricant plays an important role in the onset of macropitting [7]. On one hand, through the coefficient of friction attained: increasing the coefficient of friction rises the risk of surface fatigue by increased stresses, as well as the undesired effect of reducing the machine's performance. On the other hand, through the pressure distribution due to the elastohydrodynamic contact: the occurrence of pressure peak at the outlet accelerates the onset of pits. Besides operating conditions, pressure–viscosity coefficient and compressibility are some parameters which influence the pressure peak, friction coefficient and film thickness.

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239

MPR

32μm

100μm

Rings

Temperature probe Roller

67μm

Fig. 1. An example of surface with micropitting damage.

Therefore, the behavior of surface fatigue is closely related with the base oil. In this sense it seems synthetic lubricants perform better than mineral lubricants with similar characteristics [8,9], because they achieve higher film thickness, have better stability, lower operating temperature, lower coefficient of friction and lower viscosity–pressure coefficient. These aspects have been analyzed in [7] developing predictive behavior models to predict macropitting and micropitting. However, the effect of additives, although complex to model, is crucial to operation in regard to wear, friction and surface fatigue, especially when the contacts are working under mixed or boundary regimes, where the properties of additives become more important [10,11]. The base oil used in this article is a PAO6, usually used in gears, with a kinematic viscosity of 6 cSt at 100 1C. The effect of temperature, surface roughness and several types of additives were studied: a friction modifier (FM), an anti-wear (AW), and an extreme-pressure (EP). For the FM, we chose olein, for the AW a ZnDTP and for the EP a mixture of amine phosphates. These kinds of additives are very common in mechanical transmissions. In general, locally induced high temperatures in the contact asperities activate the EP additives to promote chemical etching at the points of contact resulting in a less rough surface. However, the AW generates a protective layer which adheres to the surfaces. The asperities slide over the layer without metallic contact, and therefore, the roughness is not reduced by chemical etching as in the case of EP. The crucial difference between the EP/AW and FM additives are their mechanical properties. The layer formed by the EP/AW additives is semi-plastic, and therefore offers certain resistance to shearing so that the coefficient of friction is generally high. However, the FM additive films are composed of multiple layers of ordered and closely packed molecules. These layers are loosely adhered to one another, with the polar part of the molecule remaining anchored to the surface. The outer layers can be easily sheared which allow low friction [12,13]. The danger of these additives, especially EP and AW, is that if chemical activity is uncontrolled, its effects may not be localized to the asperities of the contact surface and may extensively attack the surface in high temperature and high pressure conditions. In addition to studying the effect on surface fatigue, the influence of additives on friction and wear of the contact were also analyzed. For this analysis, a commercial test machine Micro Pitting Rig (MPR) by PCS Instruments was used (www.pcs-instruments.com).

2. Testing plan 2.1. Testing bench and test specimens A MPR disc machine, shown in Fig. 2, was used for the tests. This machine is for testing surface fatigue and also allows the measurement of the friction coefficient and wear in lubricated line

Initial contact width=1mm Fig. 2. Overview of MPR and a detail of the test zone; contact width of the roller.

Table 1 Geometrical dimensions and surface finish of specimens.

Roller Rings

Radius r (mm)

Contact width L (mm)

Roughness Ra (μm)

6 27.075

1 8

0.1 0.1 or 0.4

contacts between discs. The MPR is currently used by the Industry to look at both macro- and micro-pitting resistance of lubricating oils on gears and bearings [14]. The equipment consists of three rings which rotate at the same velocity and in contact with a central roller connected to a separate axis, which permits rotating at different velocity from the rings. This way you can set the average (rolling) velocity um and the slide-to-roll ratio (SRR) of the contact. SRR is defined as the ratio of the sliding velocity Δu to the average velocity um, expressed as a percentage. The zone of interest is the rolling track of the testing roller, due to the higher strength of the outer rings [15]. The rolling track of the roller has an initial contact width of 1 mm, as can be seen in Fig. 2. This geometry of the roller produces widening of the rolling track in case of wear. The normal load W of the contact is applied on the upper ring against the roller and lower rings. The arrangement is such that the load is evenly distributed among them. The contact area is lubricated by a thermostatic bath with controllable temperature. A temperature probe is located into the test chamber with the tip of the probe close to the contact region, as shown in Fig. 2. In addition, it is equipped with a commercial ICP accelerometer (PCB, model M353B16) mounted on the test head, which is used to measure the vibration in the contact and provide real time information on the condition of the specimen's surface. The rings and roller are made of steel 16MnCr5, equivalent to F1516, which is used in the construction of gears, pinions, and cemented parts. It has undergone carburizing and quenching treatments to achieve hardness of around 780 HV in the rings and 680 HV in the roller. This material is the same as that of the gears in the FZG gear testing machine [16,17]. The geometry and surface finish of the test specimens are shown in Table 1. To observe the effect of specific film thickness, and study the behavior of the base and additives, rings of two roughnesses (Ra) are used: 0.1 and 0.4 μm. This geometry of the specimens achieves a reduced contact radius of 4.91 mm, similar to that found in the contact inlet in the standard type C gears for testing micropitting in the FZG machine [18–20] as seen in Fig. 3. The FZG is used in standard surface

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 EP-mixture of amine phosphates: secondary amine salt of a

Reduced contact radius (Gear Type C) 9

C11–14-branched alkyl, monohexyl and dihexyl phosphates.

8

In this way, the effect on surface fatigue of the additives and their concentrations can be studied individually.

R(mm)

7

2.3. Test profile

6 MPR Reduced contact radius

5

4

3

0

50

100

150

200

250

300

Points in line of action

Fig. 3. Evolution of the reduced contact radius of the standard gear type C of the FZG. Table 2 Physical properties of the PAO6 at different temperatures. T (1C)

η0 (mPa s)

α (GPa  1)

30 40 60 80 100

37.95 25.00 12.57 7.36 4.78

12.3 11.5 10.1 9.0 8.2

fatigue testing, with the disadvantage of requiring a long time to achieve a high number of cycles. Keeping in mind that a gear can be modelled by a contact between rolling and sliding discs [21] the MPR reproduces these test conditions much more quickly and cheaply. The central roller has three contacts per cycle and a small radius shortens the cycle duration and therefore more fatigue cycles are achieved in a shorter time. 2.2. Lubricant The lubricants used are as follows: a PAO6 oil base without additives; the base with an FM additive at 2 wt%; the base with an AW additive at 1 and 0.3 wt%; the base with an EP additive at 1 wt% and 0.3 wt%. The highest percentages are achieved at base saturation, and in the case of AW and EP, have also been used at a lower percentage for lower reactivity as these additives are consumed during chemical etching. Viscous properties of the PAO6 oil base were measured in laboratory and are given in Table 2, namely the viscosity at ambient pressure η0 and the pressure–viscosity coefficient α. The addition of the different types of additives has shown nonsignificant variations in the results. The non-Newtonian behavior of the PAO6 has been modelled by means of the Carreau equation [22], since references [23–25] reveal that the Carreau model properly characterizes the behavior of shear-thinning lubricants like the PAO6. The Carreau equation is defined by two additional parameters of lubricant, namely Carreau's exponent n and the shear modulus G. In the case of PAO6, the values taken at 40 1C are n¼0.53, G¼ 1.59 MPa and at 90 1C are n¼0.81, G¼ 0.1 MPa, in line with references [7,26]. The types of additives used in the experiments are as follows:

 FM-olein: Glyceride of oleic acid (C17H33COO)3C3H5.  AW-ZnDTP: Zinc dialkyldithiophosphate primary alkyl groups with 8 carbon atoms.

Several test methods and profiles have been developed for analyzing micropitting in the MPR, taking into account the experience of companies like PCS Instruments (www.pcs-instru ments.com), Powertrib (www.powetrib.com) or the Society SAE International (www.sae.org). Bearing in mind this know-how, the test profile performed consists of an initial running-in step, followed by four steps, as reflected in Table 3. Until step 3 included, this test profile corresponds with the DGMK GFKT-C/8.3/90 Micropitting short test on the FZG machine (www.dgmk.de), although it is more severe in load because it works with higher pressures (0.85 GPa to 1.7 GPa in respect to the standard 0.51 GPa to 1.4 GPa), with less cycles (3E6 in the test in respect to 4.33E6 which is the standard) and with a lower SRR than that achieved in the profile of the type C gear, which reaches up to 120% in the beginning of the line of action. The duration is 200 min, whereas the duration of the DGMK GFKT-C/8.3/90 Micropitting short test is 2000 min approximately. Additionally, we have introduced step 4 to evaluate the endurance of the test specimen and the effect of additives in the long run. Therefore, the total duration of the test is 21 million cycles (1400 min) with stops to inspect the roller at the end of each step. During the test, the friction coefficient μ, the bath temperature T, the specimen wear level δ and the acceleration/vibration of the contact are continuously measured. The vibration level displayed by this rig is a dimensionless result which compares the instantaneous acceleration with an initial reference value. The control system was set to automatically stop the test in case of reaching a peak-to-peak vibration level of 1000, which corresponds to an acceleration of 90 m/s2, since this indicates severe worsening of the roller surface by wear or the appearance of pits. Moreover, after each step a micrograph is taken of the condition of the roller surface and the contact width is measured. This test profile is performed for two different temperatures (40 and 90 1C) in the case of the base without additives. The effect of the additives in the proportions indicated is only analyzed in the worst case, i.e., at 90 1C. The tests are performed for two initial ring roughnesses, 0.1 and 0.4 μm.

2.4. Estimation of specific film thickness The specific film thickness λ is defined [1] as the ratio of the lubricant film thickness hc to the combined surface roughness Table 3 Test profile: step, average velocity um, slide-to-roll ratio (SRR), load W and Hertz pressure p0, time t and number of cycles. Step

um (m/s)

SRR (%)

W (N) and p0

t(min) and cycles

Running-in Step 1 Step 2 Step 3 Step 4

3.15 3.15 3.15 3.15 3.15

1-20 20 20 20 20

100 (0.85 GPa)-230 230(1.3 GPa) 305(1.5 GPa) 390(1.7 GPa) 390(1.7 GPa)

20 (3E5 cycles) 60 (9E5 cycles) 60 (9E5 cycles) 60 (9E5 cycles) 1200(18E6 cycles)

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σ (RMS) of the contact specimens. λ¼

hc

σ

hc ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi σ 21 þ σ 22

ð1Þ

where σ1 and σ2 are the surface roughnesses (RMS) of the roller and the rings, respectively. Lubricant film thickness is approached as constant and equal to the central value [27]. Shear thinning effects in film thickness are considered by means of the following equation obtained by Bair [28]: "    1 #3:6ð1  nÞ1:7 hNc SRR um η0 1 þ 0:002SRR ¼ 1 þ0:79 1þ 100 hNc G hc

ð2Þ

Eq. (2) takes into account the lubricant parameters n and G of the Carreau rheological model [22], the low-shear viscosity at ambient pressure η0 and the slide-to-roll ratio SRR. Finally, hNc represents the film thickness for the Newtonian approach, which is determined with the Hamrock equation [29] for line contacts.  0:692 0:110 0:308  0:332 E' R p0 ð3Þ hNc ¼ 2:154α0:47 η0 um where α is the viscosity–pressure coefficient, E0 Young's reduced modulus, R the reduced contact radius and p0 the maximum Hertz pressure. The results of specific film thickness λ calculated according to Eqs. (1)–(3) are presented in Table 4, together with the Hertz contact half-width a. Thermal effects are neglected [30] due to the low values of slide-to-roll ratio considered in the test profile. Otherwise, a thermal correction factor should be used for film thickness estimation, e.g., the formulas proposed in references [29,31].

3. Results obtained 3.1. Effect of temperature In Fig. 4, the micropitting test micrographs are shown for PAO6 without additives and using rings of lowest roughness (0.1 μm). The results at 40 and 90 1C are compared at the end of each step of the test. In both cases, micropitting has begun to appear in step 2, where the flakes caused by surface fatigue are appreciable. The results of the test at 90 1C, as expected, were more severe than those at 40 1C due to the decrease in the viscous properties of the lubricant and consequent reduction of film thickness. As shown in Table 4, for step 1 the specific film thickness is approximately 1.17 for 40 1C and 0.33 for 90 1C. In steps 3 and 4, specific film thicknesses were reduced to 1.06 and 0.29 for 40 1C and 90 1C, respectively. The severity of the 90 1C test is revealed in the more affected surface, mainly by the increase in pit size. In addition, greater pit

241

depth is seen in steps 2 and 3 in respect to the test at 40 1C, see Fig. 4. In step 4 of the test at 90 1C, a coalescence phenomenon has occurred and micropitting has resulted in incipient macropitting given the size of the indentations formed (approximately 100 μm in diameter). The larger size of the indentations corresponds to a higher recorded acceleration signal, as can be seen in Fig. 5a. In both cases, the 1 mm contact width has remained in the original dimensions of the specimen, without significant expansion of the original contact width, despite the low specific film thickness, in contrast to reference [5]. This may be due to polishing during running-in and first step, so that the original machining marks are not appreciable, with the consequent increase of specific film thickness. Fig. 5b shows wear quickly grows to step 4 where it stabilizes to a level of approximately 140 μm. The blue areas on the surface of the contact, seen in Fig. 4, are the result of tempering steel by high temperatures reached locally at the contact between asperities, which could be in a typical range of 250–290 1C. Fig. 5c and its measures of friction coefficient at 40 1C, demonstrates friction increases as steps progress: in step 1 a value of 0.046 is reached, in step 2 it increases to 0.055, step 3 continues to 0.063, and finally reaches 0.070 in step 4. Steps 1 and 2 present a low specific film thickness which diminishes when gradually increasing the load and operating temperature during successive steps, leading to a higher coefficient of friction. For the test performed at 90 1C, the friction coefficient slightly decreases at the beginning of step 4, whereas lower fluctuations are found for the tests at 40 1C. Fig. 5d shows a drop in temperature at the beginning of this step 4, from 100 1C (approximately) to the commanded value of 90 1C. The reduction in temperature is associated with a rise in film thickness and explains the decrease observed in the friction coefficient according to Stribeck's curve [1]. This point is supported by the results presented in Table 4, concerning very low values of specific film thickness in the test at 90 1C. Once working conditions are stabilized in step 4, the average friction coefficient is about 0.065. A continuous increase in temperature is observed in the test at 40 1C because of the heat generated by friction. Thus, a refrigeration system would be required to keep the temperature in a controlled range. In contrast, a simpler layout is required at 90 1C, i.e. the temperature selected for the following tests, because the heat generated in the contact is lower than is required to attain the commanded temperature. Then, the use of an electric cartridge that heats up the lubricant bath in a controlled way is enough to control the temperature in the tests at 90 1C. 3.2. Effect of roughness To study the effect of increased roughness, the previously used rings of Ra ¼ 0.1 μm are exchanged for rings of Ra ¼ 0.4 μm. This implies specific film thicknesses of the previous case are approximately divided by three, taking into account the results of Table 4.

Table 4 Specific film thickness calculation for the test profile considered. Case

Step

W (N)

p0 (GPa)

a (lm)

hc (lm)

λ

Temperature: 40 1C, Ra1 ¼0.1 mm; Ra2 ¼ 0.1 mm, σ(RMS) ¼ 0.156 mm

Step 1 Step 2 Steps 3 and 4 Step 1 Step 2 Steps 3 and 4 Step 1 Step 2 Steps 3 and 4

230 305 390 230 305 390 230 305 390

1.3 1.5 1.7 1.3 1.5 1.7 1.3 1.5 1.7

112 129 146 112 129 146 112 129 146

0.183 0.174 0.165 0.051 0.049 0.046 0.051 0.049 0.046

1.17 1.12 1.06 0.33 0.31 0.29 0.11 0.11 0.10

Temperature: 90 1C, Ra1 ¼0.1 mm; Ra2 ¼ 0.1 mm, σ(RMS) ¼ 0.156 mm

Temperature: 90 1C, Ra1 ¼0.1 mm; Ra2 ¼ 0.4 mm, σ(RMS) ¼0.454 mm

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Running in+step 1

Step 2

Step 3

Step 4

40ºC 200μm

200μm

100μm

100μm

200μm

200μm

100μm

50μm

90ºC

Fig. 4. Appearance of the roller at the end of each step. Experiment without additives at 40 and 90 1C with 0.1 μm rings. (For interpretation of the references to color in this figure, the reader is referred to the web version of this article.)

Comparison of wear

Comparison of vibration level (Peak-to-Peak) 200

700

180 600

160 140

400 PAO6-40C PAO6-90C

300

δ (μm)

500

120 100

PAO6-40C PAO6-90C

80 60

200

40 100

20

0

0 0

0

10000 20000 30000 40000 50000 60000 70000 80000 90000 t(s)

t(s)

Comparison of temperature 120 100 80 PAO6-40C PAO6-90C

T(º)

μ

Comparison of friction coefficient 0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0

10000 20000 30000 40000 50000 60000 70000 80000 90000

60

PAO6-40C PAO6-90C

40 20 0

0

10000 20000 30000 40000 50000 60000 70000 80000 90000 t (s)

0

10000 20000 30000 40000 50000 60000 70000 80000 90000 t(s)

Fig. 5. Comparison of friction coefficient, wear, vibration level and temperature between PAO6 at 40 1C and 90 1C.

Fig. 6 shows the micrographs at the end of each testing step for the PAO6 at 90 1C, without additives and roughest rings, compared with the results of rings with a lower roughness. As seen in Fig. 6, step 1 shows considerable wear and polishing of the surface of the roller which has grown from an initial width of 1 to 1.2 mm. For these conditions, the specific film thickness is around 0.11, according to Table 4. Fig. 7 sheds light on the pits formation process, since subsurface cracks are clearly shown under the worn surface of the roller. A main crack is observed in the subsurface, propagating through the material in the sliding direction. Secondary cracks originated close to the surface and near this main crack reach the surface. In this way, pits are formed which can be easily removed by wear, visually eliminating the appearance of a fatigued surface.

Thus, wear has probably masked the appearance of micropitting during the test, unlike the previous case with the rings of 0.1 μm, Fig. 4. However, primary cracks can produce larger and deeper pits (macropitting) more difficult to be masked by wear. In authors' opinion, despite the loss of material at the surface level, the inside continues to accumulate high stresses and contribute to originate primary cracks (step 2). This phenomenon maintained for a high number of cycles can lead to severe macropitting [7] with pits from 300 to 500 μm in diameter, like those seen in step 4. Testing was stopped without finishing step 4 because maximum acceleration was reached after only 250,000 cycles in step 4 (or 13,000 s from the start of the test), see Fig. 8a. Also, being specimens with higher roughness, the friction coefficient in the first steps (running-in and step 1) is approximately 0.08

E. de la Guerra Ochoa et al. / Wear 322-323 (2015) 238–250

Step 1

Ra=0.1μm

Step 2

200μm

Step 3

Ra=0.1μm

200μm

243

Ra=0.1μm

200μm Ra=0.1μm

Step 4

0.25E6 cycles

200μm

Fig. 6. Appearance of rollers at the end of each step. Test without additives at 90 1C with 0.4 μm rings. In miniature, image under the same conditions with 0.1 μm rings.

(Fig. 8b), which is slightly higher than those obtained previously. This increased friction coefficient also increases the risk of macropitting [7]. The friction and wear (Fig. 8b and c) have not clearly increased from step 2 due to the decrease in effective pressure by the widening of the rolling track during the test. Using Hertzian formulae for line contact [32], there is approximately a 10% reduction of the maximum Hertz pressure at the end of the test when the track width is 1.2 mm. On the other hand, lubricant temperature remains stable during the test, with slight oscillations around the commanded value of 90 1C (Fig. 8d). 3.3. Effect of additives The next phase is the introduction of additives up until the maximum weight ratio: FM at 2%, and AW and EP at 1%, i.e., reaching the base saturation at the temperature of 90 1C. Test rings of 0.1 μm are used. Fig. 9 shows the micrographs obtained at the end of each test for the base without additives and for each of the additives. Fig. 10 includes the results of the parameters measured during the new tests, where a stable temperature of 90 1C is observed in the lubricant bath (Fig. 10a). 3.3.1. Results for FM additive at 2 wt% In the case of FM, significant differences in the appearance of the surface are observed when compared to the base without additives. In the base without additives, surface pits are larger because surface fatigue has led to macropitting. Meanwhile, for the base with FM additive, the surface has more micropitting marks, but they have not increased as much in size. If you examine Fig. 10b, the graphic showing wear with FM additive, it is somewhat higher than that obtained with the base alone. Wear has also given the surface a polished appearance, which has made micropitting less noticeable. This is justified [7] since the reduction of roughness increases specific film thickness and reduces the degree of micropitting. Furthermore, Fig. 10c shows that the effect of FM in reducing the friction coefficient is clear, from 0.060 in the base without additives to 0.045 with FM.

Thus helping to improve the surface appearance of the specimen in line with the lower level of peak-to-peak vibration level measured, as shown in Fig. 10d. 3.3.2. Results for AW additive at 1 wt% As for the AW, the ZnDTP additive forms a protective reaction layer on the surfaces leading to a reduction of mild wear, as described in reference [10]. This slows the running-in process and causes persistent roughness. A very low specific film thickness and a high concentration of AW promote an uncontrolled continuous chemical reaction on the surface (Fig. 9). This leads to an increasing material loss, as observed in Fig. 10b, in agreement with reference [12]. Due to chemical etching, signs of surface fatigue in step 4 can be masked and not clearly visible. Thus, if the time evolution of wear and friction in Fig. 10b and c is analyzed for AW additive, it has achieved higher levels than PAO6 without additives where a more appropriate running-in process may occur. It can be seen that wear is almost double than in the case of the base without additives, and also does not stabilize over time but rather follows a rising trend. The increased friction is due to deterioration of the surface by etching. The persistent roughness induces high local stresses that increase the degree of micropitting, in line with the results of reference [10], which explains the high vibration level observed. The acceleration measurement increases in the first steps to later descend and stabilize in the first quarter of step 4, where the surface state of the roller does not continue to worsen. The decrease in acceleration may be due to the combination of two effects: the decrease of pressure (due to the increase of the contact width) and mild wear (which eliminates roughness). 3.3.3. Results for EP additive at 1 wt% As with AW, the EP additive has produced widespread attack of the surface as shown in Fig. 9. The EP additive can promote corrosion in the contact [12,13] and has resulted in early failure of the roller, which did not finish step 4. It reached the maximum allowed acceleration at 4.22 million cycles of step 4 due to the damage in the surface of the specimen. Wear levels are maintained at the same values as in the case without additives until step 2, where they begin to increase and

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25μm

50μm

10μm

50μm

25μm

Fig. 7. Detail of the subsurface of the roller at the end of the test without additives at 90 1C with 0.4 μm rings.

equal wear obtained with the AW additive. The friction coefficient is slightly higher than without additives, maintaining a value close to 0.070. Unlike the AW additive, mild wear has not been able to compensate for the surface deterioration of the roller and the signal of acceleration has increased significantly in step 4. Due to the poor behavior of the AW and EP additives, and motivated by their high reactivity, the proportion is reduced to 0.3 wt% in both cases, yielding at the end of the test the results shown in Figs. 11 and 12. The lubricant temperature remains stable in the new tests and close to the commanded value of 90 1C (Fig. 12a and e).

3.3.4. Results for AW additive at 0.3 wt% If Fig. 11 is compared to Fig. 9 for the cases with AW, further deterioration of the specimen is observed for lower proportion of additives. This deterioration begins at step 1, where pits begin to appear. Due to the scarce concentration of AW, an insufficient running-in process may have caused higher roughness than in the

previous case of higher proportion of additive. This higher roughness has led to an earlier and more prominent appearance of micropitting. The worse appearance and widespread presence of pits have caused the roller to fail in step 3, at 600,000 load cycles when the maximum acceleration allowed was reached (Fig. 12b), unlike the concentration of 1% which survived to step 4. The base without additives shows faster increase of wear (Fig. 12c), but remains constant after achieving a certain level; this causes the specimens to better withstand the test. However, additives which cause chemical etching of the surface tend to increase wear throughout the whole test. Furthermore, in some cases, is has been observed a worsening of the surface of the roller, i.e. increase of acceleration, that leads to an early failure, as has been the case of AW additive at 0.3% by weight. Regarding friction (Fig. 12d), it is slightly higher than results obtained with the higher concentration due to the worse specimen surface appearance.

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Vibration level (Peak-to-Peak)

1200 1000

μ

800 600 400 200 0 0

2000

4000

6000

8000

10000

12000

Friction Coefficient

0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0 0

14000

245

2000

4000

6000

Wear

300 250

T(º)

δ(μm)

200 150 100 50 0 0

2000

4000

6000

8000

8000

10000

12000

14000

10000

12000

14000

t(s)

t(s)

10000

12000

14000

Temperature

100 90 80 70 60 50 40 30 20 10 0 0

2000

t(s)

4000

6000

8000 t(s)

Fig. 8. Signals of friction, wear, vibration level and temperature for the test without additives at 90 1C with 0.4 μm rings.

Step 4

18E6 cycles

Step 4 FM 2% wt.

50μm

Step 4 AW 1% wt.

18E6 cycles

18E6 cycles

200μm

Step 4 EP 1% wt.

200μm

4.22E6 cycles

200μm

Fig. 9. Comparison of the surface appearance of the 0.1 μm rings in step 4 with different additives.

3.3.5. Results for EP additive at 0.3 wt% Comparing both tested EP additive concentrations (Figs. 11 and 9) it can be seen that for step 3 the concentration of 0.3% is not able to mask flakes that appear, showing significant deterioration. However, for the concentration of 1%, beginning from step 3 the appearance of etching on the roller has mitigated signs of fatigue. For the concentration of 0.3% in weight, the specimen failed at 760,000 cycles in step 3 because of the acceleration peak (Fig. 12f).

With regard to wear (Fig. 12g), the lower proportion of the EP additive caused a decrease in the wear rate. As in the case of AW, friction is higher than that obtained with the highest concentration because of surface deterioration (Fig. 12h).

3.3.6. Summary of results for rings with 0.1 μm roughness Therefore, for low roughness it seems the use of very reactive additives does not improve (or even worsens, as in the case of EP and

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Comparison of wear 300

100

250

80

200 δ(μm)

T(º)

Comparison of temperature 120

60

100

40

50

20 0

150

0

0

0 10000

20000

30000

40000

50000

60000

70000

80000

90000

0

10000

20000

30000

40000

μ

PAO6-90C

PAO6+FM-90C

PAO6+EP-90C

PAO6-90C

PAO6+AW-90C

60000

70000

80000

90000

PAO6+FM-90C

PAO6+EP-90C

PAO6+AW-90C

Comparison of vibration level (Peak-to-Peak)

Comparison of friction coefficient

0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0

50000 t(s)

t (s)

1200 1000 800 600 400 200 0

0

10000

20000

30000

40000

50000

60000

70000

80000

90000

0

10000

20000

30000

40000

PAO6-90C

PAO6+FM-90C

50000

60000

70000

80000

90000

t(s)

t(s) PAO6+EP-90C

PAO6+AW-90C

PAO6-90C

PAO6+FM-90C

PAO6+EP-90C

PAO6+AW-90C

Fig. 10. Comparison of the signals of wear, friction, vibration level and temperature for PAO 6, PAO6 þFM at 2%, PAO6 þ AW at 1%, and PAO6 þ EP at 1%.

Step 3 EP 0.3% wt.

Step 3 AW 0.3% wt.

0.59E6 cycles

200μm

0.76E6 cycles

200μm

Fig. 11. Surface of roller for test using PAO6 with AW and EP additives at 0.3 wt% with 0.1 μm rings at 90 1C.

AW) the behavior of the base without additives, in regard to surface fatigue strength. In addition, lower concentrations of additive give even worse results: a combination of wear and deterioration of surface appearance. On using less reactive additives, the FM additive has delivered improved operating conditions due to the reduction of vibration level and significant improvement of friction in the contact, as exposed above.

3.4. Combined effect of roughness and additives After checking the effects of additives with the less rough discs, rings of Ra ¼ 0.4 μm, instead of Ra ¼0.1 μm previously used, are selected to analyze the effect when roughness is increased. In this case the additives are introduced in the proportion of saturation. The results are shown in Figs. 13 and 14. A stable temperature of 90 1C is also attained in the new test (Fig. 14a).

3.4.1. Results for FM additive at 2 wt% For the FM additive, as with no additives, the high roughness and associated wear of the specimens has concealed the appearance of micropitting, although in this case flakes are seen on the surface (Fig. 13). With the Ra ¼0.4 μm rings the friction modifier has been ineffective and has not caused the friction coefficient to decrease as was observed with the 0.1 μm discs. With such low specific film thicknesses, the shearing effect of the FM layers is negligible compared to the metal-on-metal contact. Thus, friction coefficients come close to 0.075 (Fig. 14b), which are almost double those of the same test with 0.1 μm rings. As for wear (Fig. 14c), it slightly decreased compared to the base without additive, but obviously it was far superior to the similar test with 0.1 μm rings. The acceleration signal reflects the poor state of the roller surface (Fig. 14d). The test stopped after 11,000 load cycles in step 2 because the resulting pits caused peak of acceleration higher than the fixed maximum.

E. de la Guerra Ochoa et al. / Wear 322-323 (2015) 238–250

Comparison of Temperature

100

Comparison of vibration level (Peak -to-Peak)

1200

90

247

1000

80 T(º)

800 70 600

60 50

400

40

200

30 0

10000

20000

30000

40000

50000

60000

70000

80000

0

90000

0

10000

20000

30000

40000

t(s) PAO6-90C

PAO6+AW 1%

PAO6+AW 0.3%

PAO6-90C

Comparison of wear

300 250

150

μ

δ(μm)

200

100 50 0 0

10000

20000

30000

40000

50000

60000

70000

80000

90000

0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0

PAO6+AW 1%

0

PAO6+AW 1%

10000

PAO6+AW 0.3%

20000

30000

PAO6-90C

90000

PAO6+AW 0.3%

40000

50000

60000

70000

80000

90000

PAO6+AW 1%

PAO6+AW 0.3%

Comparison of vibrartion level (Peak -to-Peak) 1000

80 T(º)

80000

1200

90

800

70 600

60 50

400

40

200

30

0 0

10000

20000

30000

40000

50000

60000

70000

80000

90000

0

10000

20000

30000

40000

t(s) PAO6-90C

PAO6+EP-90C

PAO6+EP 0.3%-90C

PAO6-90C

μ

150 100 50 0 10000

20000

30000

40000

50000

60000

70000

80000

90000

0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0

0

10000

20000

t(s) PAO6-90C

PAO6+EP-90C

60000

70000

80000

90000

PAO6+EP-90C

PAO6+EP 0.3%-90C

Comparison of friction coefficient

200

0

50000 t(s)

Comparison of wear

250

δ(μm)

70000

t (s)

Comparison of temperature

100

60000

Comparison of friction coefficient

t(s) PAO6-90C

50000

t(s)

30000

40000

50000

60000

70000

80000

90000

t (s)

PAO6+EP 0.3%-90C

PAO6-90C

PAO6+EP-90C

PAO6+EP 0.3%-90C

Fig. 12. Comparasion of wear, friction, vibration level and temperature of PAO6, PAO6þ EP and PAO6 þAW in concentrations of both 1 and 0.3%.

3.4.2. Results for AW additive at 1 wt% With the roughest specimens, the AW additive can produce a running-in more appropriate than in the case without additives. This causes a reduction of the degree of micropitting (Fig. 13), as reported in reference [33]. Behavioral differences of 0.4 μm specimens in regards to those of 0.1 μm seem clear with the AW additive: apparently chemical etching has not been as severe, as pits have not appeared until step 4. For the 0.1 μm rings they appeared during step 1.

In this case, one can still see the flakes that indicate fatigue and with a smooth appearance. However, if the evolution of wear is observed (Fig. 14c), it increases continuously throughout the whole test, leading to a rolling track close to 1.8 mm at the end of the test. This represents a drop of 25% in maximum Hertzian pressure. As in the case with 0.1 μm rings, the friction coefficients are higher than for the base without additives due to deterioration of surface appearance, see Fig. 14b.

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200μm

Step 2 FM 2% wt.

11.000 cycles

Ra=0.1μm

4.5E6 cycles

Step 4 EP 1% wt.

Ra=0.1μm

7.2E6 cycles Step 4 AW 1% wt. 200μm

Step 4 PAO without additives

0.25E6 cycles Ra=0.1μm

200μm Fig. 13. State of roller surface after tests with PAO6 þFM, PAO6 þ EP, and PAO6 þ AW in proportion of saturation with 0.4 μm rings.

Comparison of temperature

Comparison of friction coefficient

0.1

80

0.08

60

0.06 μ

T(º)

100

40

0.04

20

0.02

0

0 0

5000

0

10000 15000 20000 25000 30000 35000 40000 45000 50000

10000

20000

PAO6-90C

PAO6+FM-90C

PAO6+AW-90C

40000

50000

PAO6+FM-90C

PAO6+AW-90C

PAO6+EP-90C

Comparison of vibration level (Peak-to Peak)

1200 1000

400 δ(μm)

PAO6-90C

PAO6+EP-90C

Comparison of wear

500

30000 t(s)

t(s)

800

300

600 200

400

100

200 0

0 0

10000

20000

30000

40000

50000

0

5000

10000

15000

20000

PAO6-90C

PAO6+FM-90C

PAO6+AW-90C

25000

30000

35000

40000

45000

50000

t(s)

t(s) PAO6+EP-90C

PAO6-90C

PAO6+FM-90C

PAO6+AW-90C

PAO6+EP-90C

Fig. 14. Comparasion of friction, vibration level, wear and temperature for the base with and without additives, using the roughest rings.

The AW test was stopped in step 4 when the maximum acceleration was reached after 7.2 million cycles (Fig. 14d). Therefore, the specimen outlasted the base without additive.

3.4.3. Results for EP additive at 1 wt% With the EP additive, the attack seems less severe when compared with the specimens with less roughness (Fig. 13). However, you can see widespread surface flaking and a clear increase in the width of the rolling track due to wear. At the end of step 4 the rolling track measurement is 1.4 mm, this translates into a reduction of 15% in maximum Hertzian pressure. Wear has a tendency to grow throughout all testing steps and with a slope greater than that observed for the test with 0.1 μm

rings, but similar to that in the base with no additives test (Fig. 14c). The friction coefficient follows the tendency of the test with 0.1 μm discs, and is superior over the base without additives with a value close to 0.085 (Fig. 14b). It has managed to prolong the life of the roller up until 4.5 million cycles in step 4, before reaching the maximum acceleration (Fig. 14d).

3.4.4. Summary of results for rings with 0.4 μm roughness Therefore, for conditions of more roughness, the EP and AW additives have proved effective to improve resistance to micropitting of the base without additives. However, due to the very low

E. de la Guerra Ochoa et al. / Wear 322-323 (2015) 238–250

249

Table 5 Summary table of the results for the PAO6 with roller roughness of 0.1 μm. Test

Additive (wt%)

T (1C)

Ring Ra roughness (μm)

Failure step

Cycles until failure (  1E6)

Wear track (mm)

Aspect

1 2 3 4 5 6 7 8 9 10 11

– – FM 2 AW 1 EP 1 AW 0.3 EP 0.3 FM 2 AW 1 EP 1

40 90 90 90 90 90 90 90 90 90 90

0.1 0.1 0.1 0.1 0.1 0.1 0.1 0.4 0.4 0.4 0.4

4 4 4 4 4 3 3 4 2 4 4

16.24 16.24 16.24 16.24 4.22 0.59 0.76 0.25 0.01 7.25 4.49

1 1.02 0.98 1.32 1.22 1.03 1.03 1.2 1.02 1.73 1.38

Light micropitting Large pits Micropitting Chemical etching Chemical etching Severe micropitting Severe micropitting Macropitting Severe micropitting Severe micropitting and chemical etching Severe micropitting and chemical etching

specific film thickness, a high wear level has been seen in all tests for the base containing additives due to etching on the metal surfaces. In the case of the FM additive, it proves less effective than in the case of higher specific film thickness.

Acknowledgments

3.5. Final summary

References

Table 5 briefly presents the results obtained in all the cases described above.

[1] B. Bhushan, Introduction to Tribology, John Wiley & Sons, USA, 2002. [2] ASTM International. ASTM G-40-13. Standard Terminology Relating to Wear and Erosion, ASTM Book of Standards, vol. 03.02, 2013. [3] J.F. Brandão, Gear micropitting prediction using the Dang Van high-cycle fatigue criterion (Master's thesis), Faculdade de Engenharia da Universidade do Porto, Oporto, Portugal, 2007. [4] W.J. Bartz, Tribologische Aspekte bei Zahnradgetrieben—Speziell für Fahrzeuge, 5. Internationales CTI Symposium. [5] E. Lainé, A.V. Olver, T.A. Beveridge, Effect of lubricants on micropitting and wear, Tribol. Int. 41 (2008) 1049–1055. [6] ISO (The International Organization for Standardization). ISO/TR 15144-1: calculation of micropitting load capacity of cylindrical spur and helical gear —Part 1: introduction and basic principles, 2010. [7] J. Echávarri, E. de la Guerra, E. Chacón, P. Lafont, A. Díaz, J.M. Munoz-Guijosa, J.L. Muñoz, Influence of the rheological behaviour of the lubricant on the appearance of pitting in elastohydrodynamic regime, Fatigue Fract. Eng. Mater. Struct. 35 (2012) 1047–1057. [8] J.E. Fernandez Rico, A. Hernandez Battez, D. Garcia Cuervo, Rolling contact fatigue in lubricated contacts, Tribol. Int. 36 (1) (2003) 35–40. [9] R. Errichello, Selecting and applying lubricants to avoid micropitting of gear teeth, Mach. Lubr. 2 (6) (2002) 30–36. [10] C. Benyajati, A.V. Olver, The Effect of a ZnDTP Anti-wear Additive on Micropitting Resistance of Carburised Steel Rollers, AGMA, Alexandria, USA, 2004. [11] T. Ahlroos, H. Ronkainen, A. Helle, R. Parikka, J. Virta, S. Varjus, Twin disc micropitting tests, Tribol. Int. 42 (10) (2009) 1460–1466. [12] R.M. Mortier, M.F. Fox, S.T. Orszulik, Chemistry and Technology of Lubricants, Springer, Dordrecht, Netherlands, 2010. [13] L.R. Rudnick, Lubricant Additives, Chemistry and Applications, CRC Press, Boca Raton, USA, 2009. [14] T. Mang, K. Bobzin, T. Bartels, Industrial Tribology: Tribosystems, Friction, Wear and Surface Engineering, Lubrication, Wiley-VCH, Germany, 2011. [15] H.A. Spikes, A.V. Olver, P.B. Macpherson, Wear in rolling contact, Wear 112 (2) (1986) 121–144. [16] L. Winkelmann, O. El-Saeed, M. Bell, The effect of superfinishing on gear micropitting, Gear Technol. (2009) 60–65. [17] C. Benyajati, A.V. Olver, C.J. Hamer, An experimental study of micropitting, using a new miniature test-rig, Tribol. Ser. 43 (2003) 601–610. [18] G. Niemann, H. Winter, Maschinen elemente Band II Getriebe allgemein, Zahnradgetriebe-Grundlagen, Stirnradgetriebe, Springer-Verlag, Berlin, 1985. [19] Deutsches Institut für Normung (DIN), DIN 3990 Teil 1. Tragfähigkeitsberechnung von Stirnrädern, Einführung und allgemeine Einflusfaktoren, Germany, 1987. [20] International Organization for Standardization (ISO), ISO/DIS 6336: Calculation of Load Capacity of Spur and Helical Gears, 2006. [21] P. Lafont Morgado, A. Díaz Lantada, J. Echávarri Otero, Diseño y cálculo de transmisiones por engranajes, Sección de Publicaciones de la ETSI Industriales de Madrid, Madrid, 2009. [22] P.J. Carreau, Rheological equations from molecular network theories, Trans. Soc. Rheol. 16 (1) (1972) 99–127. [23] A.D. Chapkov, S. Bair, P. Cann, A.A. Lubrecht, Film thickness in point contacts under generalized Newtonian EHL conditions: numerical and experimental analysis, Tribol. Int. 40 (2007) 1474–1478. [24] Y. Liu, Q.J. Wang, S. Bair, P. Vergne, A quantitative solution for the full shearthinning EHL point contact problem including traction, Tribol. Lett. 28 (2007) 171–181.

4. Conclusions Through the tests, differences have been detected in resistance to surface fatigue, wear, and friction with the base without additives, as well as, the base with different types of additives. The conditions set for the test have shown themselves effective to study the behavior of the contact from the viewpoint of the influence of the lubricant, using a test profile which presents a high degree of equivalence with the standard micropitting short test on the FZG gear testing machine. As for the results of less severe contact conditions (lower roughness, lower supply temperatures, etc.) it has been found that signs of surface fatigue are reduced and wear is low, even with low specific film thickness. As operating conditions toughen, by increasing the roughness and temperature, surface fatigue becomes clearer and the endurance of specimens is reduced. This confirms the enormous influence specific film thickness has in the appearance of micropitting. The use of very reactive additives, such as those for extreme pressure (EP) or anti-wear (AW), can either have no effect or even worsen surface fatigue when the system is operating under the less severe conditions associated with low roughness. For low concentrations of additive, there is a combination of wear and surface deterioration due to chemical etching. Besides not improving fatigue life, the tests with EP and AW have shown a significant increase in the friction coefficient. On using less reactive additives, the FM additive has improved operating conditions by reducing acceleration and significantly improving friction in the contact. However, for more severe conditions, i.e. higher roughness, the EP and AW additives have been able to improve the duration of the specimen, in contrast to when the base without additive is used. In return for longer duration, more wear is seen on the roller. This increase of wear on the rolling track leads to significant reductions in contact pressure because the initial contact width increases as the wear rises during the test. The combined effect of lower pressure, reducing surface roughness, and removal of the layer of surface fatigue through wear helped prolong the life of the specimen.

The authors would like to thank the valuable collaboration of the Lubricants Laboratory of Repsol.

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