1

Probing the tribochemical degradation of hydrogenated

amorphous carbon using mechanically stimulated gas

emission spectroscopy

Anton Rusanov a,b, Roman Nevshupa b,c*, Julien Fontaine a, Jean-Michel Martin a, Thierry Le

Mogne a, Vera Elinson d, Andrey Lyamin d, and Elisa Roman e

a Laboratoire de Tribologie et Dynamique des Systèmes, Ecole Centrale de Lyon, UMR 5513,

69134 Ecully Cedex, France

b Bauman Moscow State Technical University, 2-Baumanskaia 5, Moscow 105005, Russia

c IETCC-CSIC, C/ Serrano Galvache 4, Madrid 28033, Spain

d Russian State Technological University MATI, Orshanskaya 3, Moscow 121552, Russia

e Institute of Materials Science of Madrid, ICMM-CSIC, C/ Sor Juana Inés de la Cruz 3,

Madrid 28049, Spain

Abstract

Mechanically Stimulated Gas Emission (MSGE) spectroscopy was used for investigation into

tribochemical reactions and gas emission for four types of amorphous hydrogenated carbon

(a-C:H) coatings, which were obtained by either ion beam deposition (IBD) or plasma

enhanced chemical vapour deposition (PECVD). The results of statistical analysis, which was

employed to identify the components of the emitted gases from the mass-spectrometry data,

argue against the hypothesis that considerable amount of CH3 could be present in the emitted

* Corresponding author. E-mail: r.nevshupa@csic.es (Roman Nevshupa)

*ManuscriptClick here to view linked References

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gases. For the IBD coatings the main components of the emitted gases were methane and/or

argon, whereas for the PECVD coatings they were mainly methane and hydrogen. Noticeable

emission of ethane, propane, carbon mono- and dioxides was also detected under sliding of

PECVD coatings deposited with the lowest ion energy. While frictional heating has been

definitely ruled out as the driving mechanism for MSGE, there are experimental evidences

that MSGE has to be associated with structural degradation of the coating.

1. Introduction

Hydrogenated amorphous carbon (a-C:H) is one of the most promising solid lubricant coating

for ultrahigh vacuum (UHV) and aerospace applications because of its super-low friction,

high wear and chemical resistance in vacuum [1, 2]. The structure of a-C:H film represents a

multiphase heterogeneous saturated/unsaturated twisted network including polymeric carbon

materials and hydrogenated carbon structures with sp, sp2 and sp3 carbon centres and

alternating local order and random local distortions [3-5]. Presently, there is a consensus that

both the ratio of two hybridized forms of carbon sp2 and sp3 and the total content of

incorporated hydrogen are responsible for the large variety of types of hydrogenated

amorphous carbon and its properties [6-9].

Hydrogen is essential for achieving super-low friction (SLF) of a-C:H with friction

coefficient below 0.02 in inert conditions [10-14]. In fact, thermal desorption of hydrogen

usually led to the increase in friction coefficient in vacuum [2, 13, 15]. SLF regime is

associated with a build-up of a transfer film and easy shear between hydrocarbon chains on

the mating surfaces due to coulombic repulsion between positively charged hydrogen atoms

[14, 16-18].

Nevertheless, a-C:H coatings have a serious weak point in that SLF regime is usually

unstable in vacuum: after certain, in many cases quite large, number of friction cycles or

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sliding distance, it spontaneously switches to solid friction regime with coefficient of friction

much higher than 0.1 [11, 12, 16, 19]. Most of the researchers agree that the instability of SLF

has to be attributed to spontaneous hydrogen desorption from the coating during sliding [1, 2,

15]. Here, a question arises about the molecular mechanisms of this process. Recently Tight-

Binding Quantum Chemical Molecular Dynamics (TB-QCMD) simulation was used to

elucidate the chemical reactions in the sliding interface of a-C:H [18, 20]. These studies have

demonstrated that terminal H atoms can be easily transferred between neighbouring C centres

contributing either to cleavage of C-C bonds or generation of molecular hydrogen. Hayashi et

al. [20] concluded that because of these reactions hydrogen termination cannot lubricate

carbon for contact pressures above 7 GPa. Furthermore, they suggested that the molecular

hydrogen could lubricate because of the steric effect due to larger separation of the mating

surfaces. Bearing in mind that most of molecular hydrogen has to rapidly escape from the

contact zone because of large activation energy for adsorption on both hydrogenated and non-

hydrogenated carbon [21], emission of hydrogen (and other gases) from the sliding interface

must correlate with near-surface chemistry and, therefore, can serve as virtually real-time

probe for a-C:H decomposition in UHV under mechanical loading [22, 23]. Numerous

investigations into chemical mechanisms of thermal decomposition of a-C:H support this

hypothesis [24-26].

So far, only limited studies of mechanically stimulated gas emission (MSGE) from

amorphous carbon coatings were done, while the available literature data is rather

inconsistent. The group of Zaidi [12, 27] was, probably, the first who measured the

composition of the emitted gases during friction of various amorphous carbon coatings in

vacuum (10-4 Pa). They identified ions with mass-to-charge ratios, m/z, 1, 2, 17 and 19, which

were attributed to H+, H2+, OH+, and H3O

+. The reasons for the occurrence of hydroxyl and

hydronium ions were not evident since the coatings contained neither water nor oxygen. Lack

4

of hydrocarbons was also surprising since methane has typically been observed in MSGE

spectra for most technical materials and coatings [22, 28]. Though metals and ceramics may

have no carbon in their chemical compositions, a nanometre-thick surface layer of air-borne

carbon contaminants can act as a precursor for methane emission [22]. Furthermore, the

results in [12, 27] notably contrasted with thermal desorption spectra of various amorphous

carbon films (this subject is discussed in detail in Section 4.1).

Recently, Rusanov et al. [28] reported initial results on the composition of the emitted gases

from highly hydrogenated a-C:H coatings in UHV (10-8 Pa) obtained by cycling the mass-to-

charge ratio of a mass-spectrometer through the range from 1 to 99. The only ions detected in

this study were m/z 2, 12-16. While the ions with m/z 2 were undoubtedly corresponded to

H2+, the origin of a group of ions with m/z 12-16 was ambiguous since these ions can be

assigned to fragments of both methane molecules and methyl radicals. An analysis of thermal

desorption performed in the work [24] suggested a common rate determining step for methane

and methyl formation, i.e., methane was formed in a consecutive recombination step, after

liberation of methyl radical from the a-C:H network. So, analysis of the composition of

emitted hydrocarbons may shed light onto the near-surface chemistry that is extremely

important for the elucidation of friction and wear mechanisms of a-C:H [29]. In view of

mostly qualitative studies of MSGE in the past, obtaining the quantitative data on the MSGE

for both hydrogen and hydrocarbons is another challenging problem.

In the present work mass-spectra of MSGE from four different a-C:H coatings were

analysed in order to understand the effect of the coating chemical composition and structure

on the composition and behaviour of MSGE. On the basis of the experimental findings

possible mechanisms of MSGE from a-C:H are discussed.

2. Experimental technique and procedure

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2.1. Coatings deposition

The present exploratory study pursue the goal of identification of the main characteristics of

the phenomenon for several coatings, which have different hydrogen content and may have

Ar molecules incorporated in the coatings. For this purpose, four different a-C:H coatings

(D1-D4) were selected. The coatings were deposited on silicon substrates using one of the two

methods: ion beam deposition (IBD) and plasma-enhanced chemical vapour deposition

(PECVD). The parameters of the deposition are shown in Table 1. IBD with higher ion energy

was used to deposit the coatings with relatively low hydrogen content, whereas PECVD was

used to obtain the coatings with higher hydrogen content [25, 30, 31]. Deposition of the

coatings D-2 was assisted by Ar+ bombardment in order to incorporate Ar atoms into the

coating. Acetylene was used as a precursor gas for deposition of the PECVD D-3 coatings and

cyclohexane was used for other three coatings. The coatings D-1, D-2 and D-4 were

developed in this work, whereas D-3, identical to AC8 [1, 28], were provided by IBM.

Table 1. Coatings deposition conditions

sample designation

deposition method

precursor gas pressure (Pa)

plasma parameters

Thickness, nm

D-1 IBDa C6H12 0.056 U= 3 kV 100±10 D-2 IBDb 1:C6H12

2:Ar 0.1-0.16 0.05-0.08

U1= 2 kV; U2= 2 kV

100±10

D-3c PECVD C2H2 13 DC, Ubias= -800 V

922±17

D-4 PECVD C6H12 8 RF, P=100 W 750±150 a One ion source b Two ion sources c From ref. [1]

The base pressure in the deposition system before introducing the reactive gases was 8×10-3

Pa. All silicon substrates were cleaned by ion sputtering during 5 min using argon r.f. plasma

at Ar pressure of 2.4 Pa and a power of 80 W.

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2.2. Characterization of hydrogen content

Hydrogen content in the coatings was characterized by Elastic Recoil Detection Analysis

(ERDA) using incident He2+ ions with energy of 2 MeV, ion current in the range between 12

and 15 nA and the total dose between 2.5 and 10 µC. The forward recoiled atoms were

filtered by a 13 µm thick Mylar film.

2.3. Friction test and gas emission measurement

Sliding friction tests were carried out using an original ultrahigh vacuum tribometer with a

“ball-on-flat” configuration and reciprocating motion of the pin. Undesirable and harmful for

desorption measurements outgassing from the mechanical parts was largely avoided by

placing all the motors and guides of the tribometer outside the vacuum chamber. The details

of the experimental system were previously reported [32]. The pins were made of AISI 52100

steel and had a spherical mirror-polished tip with a radius of 8 mm. The pins were

ultrasonically cleaned first with heptane and then with ethanol. Experiments were conducted

at ambient temperature under ultrahigh vacuum in the 10-8 Pa pressure range. The main series

of the tests was carried out with the following sliding conditions: normal load 1 N and sliding

velocity, V, 0.5 mm s-1. Additional tests at severer sliding conditions (normal load 3 to 4 N

and sliding velocity 2 mm s-1) were conducted for the coatings D-3 and D-4 having very

stable SFL regime, in order to study the transition from SLF to solid friction regimes.

Gas emission from the coatings was studied using a quadrupole mass-spectrometer (QMS)

operated in one of two modes: Multiple Ion Determination (MID) or Mass Scanning (MS).

The first mode was used to measure the kinetics of gas desorption for selected gases with

higher time resolution. In this mode, the selected range of m/z was scanned continuously

before, in the course of rubbing and after the sliding test end.

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The MS mode was used to determine the general composition of desorbed gases. For this

purpose the mass-spectra of residual gases were acquired before and during sliding. At least

15 measurements of each mass-spectrum were acquired to reduce random errors.

Furthermore, to avoid data scattering resulting from intrinsic variation of the MSGE rate, the

mass-spectra were measured in short time intervals, typically in the range of 10 to 20 s. The

time intervals for data acquisition were selected to have the desorption rate most stable

without loss of representativeness of the data samples. After checking the distribution of the

measured data samples for normality using Shapiro-Wilk method [33], the statistical

parameters – mean value and standard error of the mean – were determined for each ion

component of the mass-spectra. Finally, mean differential mass-spectrum (MDMS) was

calculated as a difference between the mean mass-spectrum during sliding and the

corresponding reference mass-spectrum acquired before sliding.

The partial pressures of different gas components emitted during sliding of the a-C:H

coatings were determined from the mass-spectrometry data using the matrix regression

method and the reference fragmentation patterns of gas molecules (see Supplementary

materials SI and Table S1). Though no information on the fragmentation pattern of CH3 was

found in literature, it has been reasonably assumed that the main ion fragments have to be in

the range of m/z from 12 to 15. Bearing in mind natural abundance of heavier isotopes of

hydrogen and carbon, ions with m/z 16 can also be found in the mass-spectrum of CH3, but

their contribution has to be quite small and can be neglected. If the emitted gases contain the

mixture of CH4 and CH3, the resulting MDMS should have an excess of ions with m/z 12-15

as compared with the reference mass-spectrum of pure methane. Thus, the method of analysis

of variances (ANOVA) [34] was applied in this work for statistical comparison of the MDMS

and the standard mass-spectrum of methane. To this end, mean experimental fragmentation

coefficients, , were determined from the MDMS for i=m/z in the range from 12 to 15:

8

/, (1)

where is the mean ion current of the principal component of methane mass-spectrum at

m/z 16.

By assuming that the measurement errors of j i and j16 are independent the standard error (se)

of the mean of can be determined from the Gauss’ formula:

."

, (2)

where #$ and # are the standard errors of and , correspondingly.

The kinetic parameters of MSGE – mean total gas desorption rate of the l-th gas

component, Ql, and the specific amount of the l-th gas component per total mechanically

affected volume, NVW,l, – were determined from the MID mass-spectrometry data using the

following formulae [22, 35, 36]:

( )∫ −−

=kt

tll

k

ll dtptp

tt

SQ

0

00

)( , (3)

w

kllVW V

ttQN

)( 0,

−= , (4)

where Sl is the pumping speed for the l-th gas components, tk – t0 is the period of time during

which the MSGE occurred; pl(t) is the pressure of the l-th gas component, p0l is the

equilibrium pressure for the l-th gas component before sliding, Vw is the total mechanically

affected volume of material.

3. Experimental results

3.1. ERDA

Hydrogen content in the topmost layers of the coatings D-2, D-1, D-3, and D-4 scaled as 15

at.% : 22 at.% : 31 at.% : 46 at.%, correspondingly. These results are consistent with literature

data for a-C:H obtained under similar deposition conditions [31, 37]. For D-3 there is also a

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satisfactory agreement with the results of Donnet et al. [30] who reported H concentration of

34 at. %. in a coating AC8 identical to D-3. The difference of 3 at. % can be considered rather

insignificant bearing in mind the measurement uncertainty (standard error 1.2 at. %). Our

results suggest that the IBD and PECVD coatings can be classified as moderately

hydrogenated and highly hydrogenated amorphous carbon, correspondingly. The coatings D-4

can probably be considered as transitional between hydrogenated DLC and PLC. Relatively

low hydrogen content in D-1 and D-2 has to be responsible for poor tribological

characteristics of these coatings in vacuum that is consistent with literature [1, 15, 38].

With exception of D-2, the increase in H concentration correlated with the decrease in a

substrate bias during deposition or equivalent impinging ion energy. This tendency was

generally observed in literature [25, 30, 31, 37, 39-41]. Nevertheless, the tendency was

opposite for the IBD coatings: for D-2 both the ion energy and H content were lower than for

D-1. This finding can be related to the effect of Ar+ during deposition due to surface

sputtering and H re-emission from the bulk [31, 42]. Comparing the PECVD coatings, larger

H content in D-4 can, in part, be attributed to the difference in precursor gases: the coatings

grown from linear and cyclic alkanes, i.e. cyclohexane, tend to have more polymer-like

structure with higher content of hydrogen than the coatings obtained from acetylene [30].

3.2. MSGE and friction behaviour

The coatings obtained by IBD and PECVD had very different friction and MSGE

behaviour. In agreement with the previously reported results [28], both the PECVD coatings –

D-3 and D-4 – reached the SLF friction regime. The typical behaviour of friction coefficient

for these coatings was characterized by four stages: I – the initial friction decrease during run-

in, II – the SLF regime, III – the transitional friction increase, and IV – the steady “solid”

friction regime (Figs. 1 and 2). For all the coatings MSGE vanished when the coating had

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been fully wore out. MSGE varied between the friction stages, but not synchronously with the

friction coefficient. Actually, during the stages I and II no appreciable pressure increase could

be observed for both PECVD coatings. MSGE manifested itself only on the stages III and IV,

although differently for D-3 and D-4 coatings. For the coatings D-3, the rate of MSGE rose

just at the beginning of the stage III, then slightly decreased and stabilized on the stage IV

(Fig. 1 b and c). After prolonged sliding, the rate of MSGE further decreased and, finally, died

away. In contrast, for the coatings D-4, MSGE occurred as a series of sharp spikes on the

stage III and at the beginning of the stage IV (Fig. 2 b and c).

Figure 1. Friction coefficient (a), residual pressures of hydrogen (b) and methane (c)

measured in the sliding test of D-3 coatings. Vertical dashed lines define the characteristic

friction stages denoted by Roman numerals. Sliding initiated at the beginning of the stage I

and finished at the end of the stage IV.

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Figure 2. Friction coefficient (a), residual pressures of hydrogen (b) and methane (c)

measured in the sliding test of D-4 coatings. Vertical dashed lines define the characteristic

friction regions denoted by Roman numerals. Sliding initiated at the beginning of the stage I

and finished at the end of the stage IV.

The transient pressure decay on the falling edges of the spikes was much slower than the

pumping process: the time constants were >100 s and 0.94±0.14 s, correspondingly.

Therefore, the pressure decay seems to be an intrinsic property of the MSGE characteristic for

this type of a-C:H. The pressure spikes generally coincided with the peaks of friction

coefficient. Certain correlation in amplitude of the pressure spikes and the friction peaks can

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also be observed. Similar spike MSGE behaviour was recently reported for the coatings

obtained by PECVD from acetylene gas precursor [28].

For the IBD coatings – D-1 and D-2 – the SLF regime did not occur (Fig. 3). After short

run-in, friction behaviour switched directly to the stage III with friction coefficient gradually

increasing with time. Hydrogen emission was very weak for these coatings, so that in many

experiments the signal of the ion current at m/z 2 could hardly be distinguished from the

background noise. Nevertheless, the emission of light hydrocarbons was noticeable and

appeared just from very beginning of sliding. After the first peak, the emission rate decreased

and stabilized. This MSGE behaviour was somewhat similar to that of D-3 coatings.

Figure 3. Friction coefficient (a), residual pressures of hydrogen (b) and methane (c)

measured in the sliding test of D-1 coatings. Vertical dashed lines define the characteristic

friction regions denoted by Roman numerals. Sliding initiated at the beginning of the stage I

and finished at the end of the stage IV.

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3.3. Composition of the emitted gases

Figure 4 shows the MDMS of the emitted gases for four types of the coatings. For D-3 the

MDMSs were determined on the stage III during the pressure spikes only, whereas for other

coatings they were determined on the stage IV when the MSGE was nearly uniform. Two-fold

standard errors of the mean are shown by vertical error bars.

Figure 4. Mean differential mass-spectra of gases emitted during sliding of the a-C:H

coatings: D-1 (a), D-2 (b), D-3 (c) and D-4 (d). Inset (e) shows the enlarged portion of the

graph (d) in the range of m/z from 24 to 45.

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The ion components present in all MDMSs were CHn+ with n from 0 to 4. Intensive peaks

of Ar+ and Ar2+ were identified in the MDMSs for D-2. In some tests weak hydrogen

desorption could also be observed, however, the desorption rate had large variation between

different tests even on the same sample. This behaviour was consistent with previous results

and can be related with non-uniform degradation of the coatings under tribological activation

[28].

For the PECVD coatings, H2+ and CHn

+ with n from 0 to 4 were the main components of

the MDMSs. In addition, two groups of ions with m/z from 25 to 29 and from 41 to 45 with

main peaks at m/z 28 and 44 were found for the coatings D-4. Though the main peaks could

be attributed to carbon mono- and dioxide, presence of smaller peaks nearby indicated

possible contribution from ethane, propane, cyclopropane and/or propene. The MDMSs were

fitted by the reference mass-spectra of these possible gas components, using linear regression

method. Then, backward elimination procedure with a sequence of F-tests was employed in

order to discard the gas components which contribution might be statistically insignificant

[33]. Adjusted coefficient of determination, R2, is shown on Table 2 for various combinations

of gas components. From this analysis, cyclopropane and propene were discarded at

significance level of 0.05. When the model included the following gases: H2, CH4, C2H6,

C3H8, CO and CO2, the best regression quality was achieved as indicated by the highest value

of R2, which was very close to unity.

Table 3 shows mean values and standard error of the mean for partial pressures determined

using linear regression analysis of the mass-spectra. In this analysis, standard values of the

ionization probability of each gas, fractional abundance of ion fragments, sensitivity,

transmission and detection factors of the mass-spectrometer were taken in accordance to the

QMS manufacturer’s guide [43]. In terms of pressure, methane was the dominant component

followed by hydrogen. For H2 and CH4 mean partial pressures were statistically significant at

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α = 0.05. Though C2H6, C3H8, CO and CO2 gave statistically meaningful contribution to the

MDMS (at significance levels α=0.2, 0.15, 0.15 and 0.15, correspondingly), the residuals

were quite large. This could be due to various reasons including possible presence of

unidentified components in the emitted gases or measuring errors related to fast-varying

pressure signals. The latter is of special concern since the time required for the measurement

of 5 channels with m/z 12-16 had to be about 625 ms, provided the multiplexor operation

mode of the QMS with the dwell time per channel about 125 ms. This time is large enough

compared with the rate of the pressure variation.

Table 2. Adjusted coefficient of determination for various models of gas composition for

samples D-4

Model of gas composition m/z range Adj. R2

H2, CH4, C2H6, C3H8, CO, CO2 2-45 0.9979

C2H6, C3H8, CO, CO2 22-45 0.9779

C2H6, C3H8 22-45 0.8115

CO, CO2 22-45 0.8743

Table 3. Mean pressure increase and standard error of the mean for various gases emitted

from the coatings D-4 during a spike

Gas H2 CH4 C2H6 C3H8 CO CO2

∆&, 10-10 Pa 4.83 35.2 1.42 1.31 2.17 2.16

se, 10-10 Pa 0.050 0.410 1.73 1.12 1.69 1.24

Table 4 shows mean experimental fragmentation coefficients with corresponding standard

errors for the group of ions hypothetically related with methane and methyl radical. In order

to compare the experimental fragmentation coefficients with the reference ones, a significance

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test with null-hypothesis H0: ae = aref and two-side alternative hypothesis H1a: ae ≠ aref was

done. For the most tested coefficients, the test failed to reject the null-hypothesis at the critical

significance level α=0.05. When H0 was rejected, two different single-side alternative

hypotheses H1b: ae < aref and H1c: ae > aref were additionally considered. The hypothesis H1b

was accepted at α=0.05 in all cases. These results have indicated that none of the examined

experimental fragmentation coefficients exceeded the standard reference value; therefore, no

statistically significant amount of methyl radicals could be identified in the mass-spectra of

the emitted gases. It can be suggested that the ions with m/z 12-16 originated mainly from

methane with a small contribution from ethane.

Table 4. Statistical comparison of experimental and reference fragmentation coefficients for

methane

m/z ,() sample parameters accept. hypothes.

Nr

D-1

12 0.038 - a 13 0.107 - a 14 0.204 - a

15 0.888 0.697 0.0882 15 H1b

D-2

12 0.038 - a 13 0.107 0.0813 0.0550 15 H0 14 0.204 - a 15 0.888 0.951 0.105 15 H0

D-3

12 0.038 0.035 0.0040 28 H0 13 0.107 0.077 0.0037 27 H1b 14 0.204 0.144 0.0067 28 H1b 15 0.888 0.710 0.0420 28 H1b

D-4

12 0.038 0.0523 0.0133 15 H0 13 0.107 0.101 0.0182 15 H0 14 0.204 0.197 0.0421 15 H0 15 0.888 0.931 0.119 15 H0

a Statistically insignificant at α=0.05

3.4. Quantitative characteristics of MSGE

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Statistical parameters of experimental MSGE rates for hydrogen, methane and argon are

shown in Table 5. In order to reduce the measuring error related with the pressure variation,

the parameters for D-1 – D-3 were determined during the whole stage IV, whereas, for D-4

they were evaluated during the spikes only.

Table 5. Total gas desorption rate, Ql, standard error of the mean, seQl, maximum desorption

rate, maxQl, and specific amount of desorbed gases per unit volume of worn coating for H2,

CH4 and Ar

Ql / seQl / maxQl (10-12 Pa m3 s-1) l H2 CH4 Ar D-1 1.95 / 1.02 / 3.95 D-2 0.73 / 0.17 / 0.94 18.3 / 6.5 / 29.4 D-3 8.95 / 6.01 / 13.2 17.2 / 7.4 /22.4 D-4 48.0 / 21.0 /110.0 31.0 / 14.0 / 59.0 NVW,l (mol cm-3) D-1 1.04×1019 D-2 1.33×1018 9.9×1019 D-3 7.0×1017 1.64×1018 D-4 1.14×1020 5.8×1019

For the IBD coatings the rate of hydrogen emission was statistically insignificant and is not

shown. For D-2 Argon was the dominant species comprising 96% of the total emission yield,

whereas the rate of methane emission was the lowest among all the coatings. With respect to

the methane emission rate, the coatings can be ranged in the following order: D-2 < D-1 < D-

3 < D-4, while the emission rate between D-2 and D-4 increased almost forty-fold. This order

is the same as found for the H concentration in the coatings. Comparing D-3 and D-4, it was

observed that the proportion between methane and hydrogen changed from about 2 for D-3 to

0.65 for D-4. Furthermore, quite surprising was the finding that the instant ratio CH4:H2

considerably varied in course of sliding for the same coating. Figure 5 shows cross-

correlations between the instantaneous values of the emission rates for methane and hydrogen

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for D-3 (a) and D-4 (b). Solid lines are the linear fits to the experimental data, while the

dashed lines show the upper and the lower bounds. In addition, for the coating D-3, the data

corresponding to the stages III (open dots) and IV (solid dots) were fitted separately (the lines

bIII and bIV). These results show a clear trend of gradual increase in the ratio CH4:H2 along the

stages III and IV. The total span of the ratio was almost one order of magnitude. For the

coating D-4 the ratio also had large dispersion, but without some definite trend. These finding

can conceivably be related with different H content in the PECVD coatings. Indeed, the

depletion of hydrogen in the coating has to be faster when the initial H concentration is lower.

Depletion of H can be the reason for gradual decrease in relative H2 concentration in the

emitted gases. Though under given experimental conditions MSGE is not driven by frictional

heating (see Section 4.2), there is some analogy with thermal effusion experiments [44, 45],

where H2 emission usually slightly anticipated the emission of hydrocarbons. These similarity

may indicate that the same reaction pathways underpin both thermal and mechanochemical

degradation of a-C:H.

Apart from the chemical composition both the coating’s thickness and wear rates could

influence MSGE [46]. In fact, the thicker the coatings the larger should be the total amount of

emitted gases. The same applies for the wear rate. To account for these factors, the total

amount of the emitted gases was divided by the total volume of the material in the

mechanically affected zone. The resulting parameter - specific amount of desorbed gases per

unit volume of worn coating, NVW, – allows comparison of the MSGE for the coatings of

different thickness (Table 5). For methane, NVW increased in the same order as both the H

concentration and the emission rate: D-2<D-1<D-3<D-4. Again, the difference in NVW for

CH4 between D-3 and D-4 is nearly forty-fold. Even higher is the difference in NVW for H2

between D-3 and D-4: more than 160-fold.

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Figure 5. Cross-correlations between the rates of gas emission for methane and hydrogen for

the coatings D-3 (a) and D-4 (b). Solid lines are linear fits, dashed line are the lower and

upper bounds. Open and solid dots on (a) correspond to the stages III and IV, respectively.

The lines with the slope bm are the linear fits to the total data sample.

4. Modelling and discussion

4.1. Gas composition and rate of MSGE

Though MSGE of hydrogen and methane for a-C:H was previously reported [28], this is the

first time when ethane and propane emission have been observed during sliding of PECVD a-

C:H coatings. Most likely, the hydrocarbons were recombinative products as the coatings

hardly contained free C1 to C3 alkanes. Emission of hydroxyl and hydronium ions has not

been confirmed in our experiments. The gas composition was very different from that

20

reported in [12], but resembled the spectra of thermal decomposition which typically had

hydrogen [2, 24-26, 45, 47-50], methyl radical [24, 26, 45], methane [24, 45, 48, 49],

acetylene [24, 26], ethylene [24, 26, 45, 49], ethane [24, 48, 49], propene [49], propane [25,

49], and higher mostly unsaturated volatile hydrocarbon species [49]. Similar results were

also obtained in thermal desorption studies of graphite subject to atomic hydrogen [13], where

the most important products were hydrocarbons with the general formulae CnH2n and CnH2n+2

and the most abundant species with n=2 and 3.

Argon emission from D-2 has to be a natural consequence of Ar+ bombardment of the

coating during its deposition. Generation of carbon oxides is still puzzling since no oxygen

has to be incorporated into the coatings. Probably, carbon oxides could stem from the

reactions between a-C:H and oxide phases on the steel pin.

An apparent contradiction between, on the one hand, the widespread belief that sliding

leads to hydrogen depletion in a-C:H and, on the other hand, failure to detect any appreciable

H2 desorption during run-in and SLF in these experiments can be plausibly explained if one

considers the detection limit of the QMS and the expected rate of hydrogen desorption. The

smallest emission rate, Qmin, which can be measured by a QMS, is a function of the minimal

statistically significant pressure increase due to MSGE, ∆pmin, and the pumping speed, S,:

*+, ∆-./012 34, (5)

where Na is Avogadro’s number, R is the gas constant, and T is the temperature. On the

reasonable assumptions of the independence of the pressure measurements, ∆pmin can be

determined from the following formula:

∆&+, 5√,√289,,:, (6)

where t2n-2,α is the Student’s statistics at the significance α and 2n-2 degrees of freedom. The

data samples of the pressure must be of the same size, n, and have equal standard deviation,

21

sd. For the typical size of a data sample 10 to 20 and α = 0.05 t2n-2,α is ranged between 1.734

and 1.684 [51]. Then, for given S = 0.159±0.017 m3s-1 and the standard deviation of p(H2)

sd=2.30×10-11 Pa we have ∆pmin = 1.50×10-11 Pa and Qmin = 5.77×108 mol s-1.

Furthermore, the rate of sweeping of the surface area by the pin is nearly the product of the

contact diameter, d, and the sliding velocity, V. Then, minimal measurable number of H2

emitted from the unit surface area is determined from the following expression:

3;,+, <./=5 . (7)

For V=3 mm s-1 and d=100 µm, NA,min=1.92×1011 mol cm-2 that is approximately 0.01% of

the number of adsorbed atoms in a monolayer. In the hypothesized case that 0.01% of

adsorbed H2 is desorbed from the friction zone during each cycle and that the readsorption of

H2 from the gas phase is negligible, one can assume that the adsorbed layer would be fully

removed in 104 strokes. So, the expected duration of SLF regime, which critically depends on

the saturation of the surface dangling bonds with H, would be much shorter than 104 strokes.

The contrast between this estimation and the experimentally observed duration of the SLF

regime, e.g., 2.5×104 cycles for D-4, implies that the real number of emitted hydrogen

molecules during each cycle has to be at least one order of magnitude smaller than NA,min and,

therefore, could not be detected.

On the other hand, the upper bound of H2 emission rate can be estimated from the results

reported by the group of Kubo who simulated a-C:H sliding interface using TB-QCMD [20].

In a period of time of 6 ps, they observed formation of 1 dihydrogen molecule. This yielded

specific number of emitted H2 per unit surface area 2.0×1013 mol cm-2, which value is two

orders of magnitude higher than NA,min. Thus, detection of hydrogen emission during SLF

regime is principally possible under extreme sliding conditions like those used in the

simulation (load pressure 1 GPa and sliding speed 100 m s-1) [20]. However, these results

22

must be interpreted with caution since H2 emission rate was determined from the only

emission event and is conceivably overestimated.

4.2. Effect of frictional heating

When observing appreciable correlation in the position and the amplitude between the

pressure spikes and friction peaks on the stages III and IV (Fig. 2) and certain similarity in the

gas composition between MSGE and thermal decomposition, the by now familiar model of

the gas desorption stimulated by frictional heating comes immediately to mind. This model is

based on the belief that the important increase in temperature, referred to as flash temperature,

is produced on the tops of protruding summits on the contacting rough surfaces due to

significant density of dissipated energy [52]. Drawing on the assumption of severe abrasive

wear of a-C:H various researchers estimated the range of flash temperatures: >600 K [53],

320 K [39], 150 K [27]. Nevertheless, we cannot be confident about the applicability of these

estimations in our case since the condition of severe wear is unrealistic for both run-in and

steady super-low friction (SLF) regimes. More reasonable value of only 15 K was suggested

by Paulmier et al. [27] for low-friction regime and mild wear.

In this work, an approximate solution of the heat problem proposed by Tian and Kennedy

[54] for circular contact zone and homogeneous half-space was used to calculate the lower

bound of flash temperature, ∆Tmax, assuming that the coating is infinitely thin and the upper

bound assuming that the coating is a half-space(the model is given in supplementary materials

SII). The initial parameters used in the calculation were: velocity of the pin V = 2 mm/s,

radius of the contact zone 0.1 mm, flat sample was stationary. The calculation results

(Peclet number, Pe, and ∆Tmax) are shown in Table 6 (for the material characteristics used in

the calculation see Table S2 in supplementary materials). For all regimes ∆Tmax was below 15

K. Such low values are not surprising when low sliding speed and moderate normal load are

23

considered. These results are consistent with the data of Katta et al. [55], who reported the

flash temperature for DLC coatings below 100 K even at the sliding velocity as high as 10

m/s.

Definitely, the calculated temperature increase is too small to produce any significant

hydrogen desorption, especially considering short duration of the temperature flashes [56] and

the threshold temperature for hydrogen evolution (see Table S3 in Supplementary materials).

This conclusion is in line with the works of Racine et al. [39, 57] who suggested that elastic

interaction rather than frictional heating (up to 320° C) is responsible for loss of hydrogen

from a-C:H. Using simulation, Rusanov et al. [28] drew the conclusion that under the same

sliding conditions as in this study the model of thermal desorption can explain only about 10-

5% of the total MSGE yield. Similar findings were reported for different materials under mild

sliding conditions [23, 56, 58, 59].

Table 6. Calculation results of the maximal temperature increase for the PECVD coatings

Flat sample

(half space)

Pe ∆Tmax (K)

Friction stage II

(µ=3×10-3)

Friction stage IV

(µ=0.52)

a-C:H 0.485 0.057 14.8

Si 1.3×10-3 2.8×10-3 0.47

In addition to microscopic temperature rise, the existence of nanometre-scale localized hot

spots of several hundreds and even thousands K that may result from quasi-adiabatic energy

release accompanying molecular interaction and bonds breaking has been argued by many

researches as the activation factor for various mechanochemical reactions including the gas

emission [52, 60, 61]. However, the effect of these hot spots seems to be overestimated. Using

24

molecular dynamic (MD) simulation of a-C:H Schall et al. [62] determined that under severe

sliding with normal stress of 7.6 GPa and sliding velocity of 90.2 m s-1 the temperature

increase in a 2.5 Å thick section containing interface was only 282 ± 146 K. Ma et al. [63]

carried out MD simulation under even severer sliding conditions and proofed that the effect of

the hot spots on hydrogen depletion and degradation of a-C:H could be neglected. These

conclusions are in line with the basic principles of mechanochemistry pointing that neither the

total heat evolution nor local hot spots have significant influence on the mechanochemical

processes [64, 65]. Another argument against the hot spots is their short life, which typically

does not exceed 10-12 – 10-11 s due to decay via phonon excitation [14]. Since this time is

several orders of magnitude smaller, than the time constant for desorption of hydrogen and

methane [66], the probability of a molecule desorption during the lifetime of a hot spot is

vanishing.

Apart from the above theoretical considerations, the evidences against thermal mechanism

of MSGE can be found in our experiments. Bearing in mind that friction coefficient is a

measure of the mechanical energy dissipation, the temperature of the contact surfaces under

constant sliding conditions has to linearly depend on friction coefficient. On the other hand,

the rate of gas desorption is an exponential function of the surface temperature. Taken

together, these results suggest that, if MSGE is a temperature-driven process, the rate of

MSGE has to be an exponential function of friction coefficient. However, that was not

generally true. Actually, when comparing the stages I and III in Fig. 2, one can note only little

variation in friction coefficient while the MSGE rate and behaviour significantly varied. In a

rough-and-ready approach, the upper bound of the possible pressure increase on the stage I

can be assumed equal to the standard deviation of the background pressure: 2.30×10-11 Pa for

hydrogen and 5.51×10-11 Pa for methane. Then, the difference in the MSGE rate between the

stages I and III is nearly two orders of magnitude. Furthermore, the rate of gas desorption on

25

the stage III was noticeable even though friction coefficient was smaller than on the stage I

and below 0.1 (see inserts in Fig 2 a and b). Likewise, for all coatings there is a lack of

correlation between the MSGE and friction coefficient on the stages III and IV. So, the

experimental results have provided evidence that MSGE is almost independent on friction

under given sliding conditions.

4.3. Dependence of the MSGE on the coatings damage

While thermal desorption can be confidently ruled out, there are firm evidences [56, 67-72]

that the MSGE can be related to fracture, abrasion, plastic deformation and other forms of

material damage. In fact, certain conformity of the MSGE and the wear behaviour can be

easily observed. For the coatings with the SLF regime, both the MSGE and the wear rate were

minimal on the stages I and II and drastically increased on the stage III. For the IBD coatings,

intensive wear occurred from the commencement of sliding and, again, it was accompanied

by MSGE. In this context, the pressure spikes can be attributed to individual damage events,

e.g., cracking or detachment of debris from the coating, which can be observed in Fig. 6.

Further grinding and attrition of the debris in the contact zone can lead to continuous

decrement in the emission rate, thus, explaining slow transient pressure decay. Actually, only

fine wear debris could be found on sides of the wear track after friction test.

The run-in and SLF stages are not completely weariless since a thin transfer layer must be

formed on the pin surface in order to achieve the SLF. Nevertheless, the degree of the coating

damage during SLF is much less than on the stages III and IV (Fig. 6). This corroborates the

hypothesis of MSGE as the process driven by the coating structural degradation. When the

coating is fully gone (at the end of the stage IV) the desorption rate completely vanished. This

occurred because on this stage the pin slid on the silicon substrate, which usually gives no

26

MSGE due to high hardness preventing from wear and very low content of dissolved and

occluded gases.

Figure 6. Images of the wear track for D-4 samples: after 5510 cycles in super-low friction

regime (a) and after the coating failure (b).

Though a minor part of the emitted hydrogen can proceed from unbound H2 occluded in the

coating, the results of this work have indicated that mechanochemical processes play an

important role in MSGE. Thus, MSGE has to depend on the chemical composition and

structure of the coatings. In the future the mechanism of MSGE from a-C:H will be

investigated in detail on the base of exhaustive mechanical, physical and chemical

characterization of the coatings.

5. Conclusions

27

For all studied coatings MSGE was observed only during the friction stages III and IV,

when both friction coefficient and wear rates were high. During run-in and SLF the rate of

MSGE was below the measurement limit being approximately 5.77×108 mol s-1 for hydrogen.

For the IBD coatings the emitted gases included methane and/or argon depending whether

or not IBD was assisted by Ar+ bombardment. For the PECVD coatings, in addition to

methane intensive hydrogen emission was observed. Noticeable emission of ethane, propane,

carbon mono- and dioxide was also found during rubbing of D-4 coatings. The following

three parameters: the emission rates of methane, the hydrogen concentration in the coatings

and the specific amount emitted methane per total volume of mechanically affected a-C:H on

the friction zone increased in the following order D-2<D-1<D-3<D-4. The lowest hydrogen

concentration in D-2 can be related to the effect of Ar+ bombardment.

Statistical analysis of the mass-spectrometry data failed to confirm presence of methyl

radicals in the emitted gases.

For PECVD coatings, the composition of the emitted gases significantly varied in the

course of sliding. For D-3 coatings, a definite trend was observed which consisted in gradual

increase in the ratio CH4 : H2. This behaviour can be attributed to continuous hydrogen

depletion in the coating during intensive wear.

The experimental data gave no evidences of the dependence of MSGE on friction

coefficient. The calculated temperature increase on the contact zone was below 10 K.

Therefore, frictional heating has to be discarded as the reason for MSGE under given sliding

conditions. MSGE from a-C:H can be related to fracture, plastic deformation and other forms

of material damage. In fact, there is close correlation between the degree of the coating wear

out and the rate of MSGE: on the stages I and II when the wear rate was very low the MSGE

was below the measurement limit. In contrast, on the stages III and IV, which are

characterized by intensive wear, MSGE was significant. These findings have important

28

implications for developing of advanced tribological coatings with tailored or controlled

MSGE properties.

Acknowledgements

Rusanov acknowledges the Embassy of France in Moscow and Ecole Central de Lyon in

France for financial support. Roman Nevshupa acknowledges the financial support from the

Ministry of Economy and Competitiveness of Spain through the grants RYC-2009-0412,

BIA-2011-25653 and IPT-2012-1167-120000 with participation of the European Regional

Development Fund (FEDER).

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