*3.2.1. Measurements of cyclic potentiodynamic polarization curves*

associated with these parameters or settings *in-situ*. They need also to be instrumented with electrochemical techniques, which enables the management and the recording of applied electrochemical parameters and/or their responses (e.g., polarization of the contacting materials, charge density, etc.). The ultimate goal is to obtain promptly any information on the evolution of the input and output chemical-mechanical measurements during the test. These *in-situ* data outcome combined with *ex-situ* surface characterization techniques (e.g., high resolution SEM imaging, TEM, XRD, EDAX, FIB, XPS, Auger spectroscopy, FT-IR, roughness surface profilometry, micro- or nanoindentation hardness, etc.) and chemical analyses of post-test solutions (e.g., ICP-AES, ICP-MS) allow for the disclosure of the wear-corrosion mode, and thereby contributing to a better understanding of the tribocorrosion mechanisms involved

The typical configuration of a triboelectrochemical cell experiment involves an inert material (e.g., corundum counter-body) sliding against the investigated material (i.e., working electrode) under mechanoelectrochemical well-controlled conditions. Although metal-onmetal contact configurations are possible, but the use of an inert counter-body simplifies the interpretation of the electrochemical results of the corrosion-wear process. In addition, it is most advising to perform the electrochemical measurements under stationary regime conditions, at least prior to starting up of the measurements. The development of relevant models for the interpretation of the tribocorrosion mechanism concurrently depends, essentially, on the choice of electrochemical techniques to be implemented in a tribocorrosion test and the mechanical contact conditions (e.g., relative motion). The state-of-the art and reviews on tri-

boelectrochemical techniques and experiments are available elsewhere [8–13].

**3.2. The use of electrochemical techniques in the study of** *in-situ* **tribocorrosion** 

Corrosion processes are of electrochemical nature and therefore electrochemistry likely interferes with the tribological behavior of tribocorrosion systems. Attempts were made in recent years to control wear by electrochemical methods in engineering and biomedical systems. These electrochemical techniques provide a very convenient way to measure the rate of corrosion processes in either the laboratory or the field. Such methods can also be used in many different ways to assess either their efficiency of monitoring corrosion material degradation (e.g., concrete steels, and marine alloy structures) or their capability for material protection (e.g., inhibitors, protective layers, coatings, and appropriate metals and alloys). The theoretical bases and the practical implementation of the electrochemical techniques have been published in books and review articles [8, 9, 11, 14] and this material is compendiously repeated here only to the extent that it is needed to define the terminology as well as the utility of the topic for the chapter. In mechanical lubricated contact systems where surface interactions do occur, these techniques offer the possibility to simulate different corrosion conditions under well-controlled electrochemical environments. This can be done for instance by the measurements of open-circuit potential (*E*oc or *E*cor) during a corrosion-wear test, measurements of cyclic potentiodynamic polarization curves under wear and pure corrosion conditions, records of the current-induced

(corrosion mechanism, wear regime, friction process, etc.).

82 Metallic Glasses - Properties and Processing

**processes**

The susceptibility of metals to localized corrosion is usually expressed by the breakdown potential, *E*<sup>b</sup> , or designated to, as the pitting potential, *E*pit, and the repassivation potential, *E*<sup>r</sup> or termed the protection potential, *E*p. At the breakdown potential, localized corrosion starts. The *E*pit of a metal is often associated with the potential at which the current density suddenly increases and with the breakdown of its passive surface film. The higher the potential (more noble), the less likely the alloy is to cause the initiation of localized corrosion. At the repassivation potential, pitting stops.

A cyclic potentiodynamic polarization technique can be employed to determine both *E*<sup>b</sup> and *E*r of a corrosion system under wear (i.e., tribocorrosion) and pure corrosion conditions.

Many researchers [15–21] have successfully used the potentiodynamic anodic polarization technique to study the corrosion, the wear-corrosion synergism, and the tribocorrosion behavior of metallic alloy systems including the BMGs in various electrolytes [17–21].

ASTM G61-86 provides a procedure for conducting cyclic potentiodynamic polarization measurements [22]. By convention a cathodic current is negative whereas an anodic current is positive [23].

This technique uses a typical three-electrode system (WE working electrode or metal being investigated, RE (SHE, reference electrode), and CE (Platinum or Graphite, counter-electrode)) controlled by a potentiostat as shown in **Figure 2**. The potential, which applies to the WE, is usually swept from the active (cathodic) direction to the noble (anodic) one, while tracking the current density continuously until it reaches a selected value of current density, where after, the scan is inverted in the active direction, until the hysteresis loop closes or until the corrosion potential is reached. The potential of the WE can be considered as the "driving force" of the corrosion system, while, the anodic current density can be regarded as proportional to the corrosion rate of the WE.

A typical plot of polarization curve generated by this method as *E*–log (*i*) is shown in **Figure 3**. If a specified material is susceptible to localized corrosion, a hysteresis loop as shown in **Figure 3** will be observed as the potential scan is reversed. Otherwise, a uniform corrosion takes place in the transpassive or oxygen evolution region. Note that the larger the area of the hysteresis loop, the lower the ability of the metal to repassivate.

The approach to be considered for getting a quick picture of the electrochemical behavior of a metal under polarization is to identify some of the key corrosion parameters directly from the

• The process is designated as anodic polarization when the potential is above the *E*cor, and

• The corrosion potential, *E*cor, corresponds to the potential at which the current density

• The extrapolation of the experimental anodic and cathodic branches of the polarization

• The primary passivation potential (*E*pp) corresponds to a potential positive to which passive surface layers are formed, and at which a maximum corrosion current density is reached

• The region, where the current density remains approximately 10 mA.m−2 (i.e., passive cur-

This corresponds to the stabilization of a passive film. A material that exhibits a high resistance to localized corrosion in a specified environment form an adherent, and dense, nonporous, thin passive oxide/hydroxide film, which protects it from high corrosion rates.

− *E*oc) may indicate a strong resistance of the metal to pitting at *E*oc. High values of

continued driven by a rapid active dissolution (e.g., case of some stainless steels immersed in chloride solutions), and usually at more elevated potentials, oxygen evolution may occur

• In the reverse scan, as in case of path (1), the localized corrosion will not be triggered at *E*oc; in this case, the working electrode will not suffer pitting under free *E*oc conditions. If the path (2) is observed, where *E*pp is below *E*oc, the working electrode will undergo pitting

evolution or

85

cor.

Metallic Glasses for Triboelectrochemistry Systems http://dx.doi.org/10.5772/intechopen.78233

cc, current requisite prior to the formation of

) as the potential increased, is called the passive region.

(or *E*p) is the potential at which passive layers are stable and

), pits are initiated and propagated. In fact, the material

and *E*<sup>r</sup>

, transpassive region starts. The local breakdown of the passive film

.

− *E*oc), and protection overpotential,

relative to the *E*oc.

• The corrosion behavior of a metal is assessed at *E*cor (or *E*oc) and other potentials.

• At cathodic potentials lower than *E*cor, the cathodic reactions takes place (H<sup>2</sup>

curves, as shown in **Figure 3**, indicate the corrosion current density, *i*

cyclic potentiodynamic scan mode (**Figure 3**), mainly:

approaches or theoretically equal zero (*i* ≈ 0).

coinciding with the critical current density (*i*

<sup>p</sup> of the electrode at *E*<sup>r</sup>

• High values of both pitting overpotential (ηpit = *E*<sup>b</sup>

both (ηpit and ηp) are desirable to reflect high values of *E*<sup>b</sup>

corrosion on the surface defects or after *E*oc incubation periods.

− *E*<sup>r</sup>

may suffer from localized corrosion when the potential is higher than *E*<sup>r</sup>

O2

reduction prevails).

surface layers).

rent density, *i*

protective.

(η<sup>p</sup> = *E*<sup>r</sup>

• The protection potential *E*<sup>r</sup>

• At the potential interval, (*E*<sup>b</sup>

• At potential above *E*<sup>b</sup>

(*cfr*. Pourbaix diagram, *E*-pH).

cathodic polarization when the potential is below the *E*cor.

**Figure 2.** Schematic view of a tribocorrosion experimental set-up under potentiodynamic polarization conditions. Potential (*E*), and current density (*i*) measurements are performed on a working electrode (*WE*) sliding against a counterbody ball (unidirectional reciprocating sliding, sphere-on-flat) with respect to a *RE* reference electrode (e.g. Ag/ AgCl (3 M KCl)) *via* a *V* voltmeter and *CE* counter-electrode (platinum or graphite) *via* an *A* ammeter respectively. *F*<sup>N</sup> normal force, *F*T tangential force, *f* sliding frequency, *D* displacement amplitude.

**Figure 3.** Schematic view of a hypothetical cyclic potentiodynamic polarization (cathodic and anodic) plot for determining localized corrosion parameters.

The approach to be considered for getting a quick picture of the electrochemical behavior of a metal under polarization is to identify some of the key corrosion parameters directly from the cyclic potentiodynamic scan mode (**Figure 3**), mainly:


**Figure 2.** Schematic view of a tribocorrosion experimental set-up under potentiodynamic polarization conditions. Potential (*E*), and current density (*i*) measurements are performed on a working electrode (*WE*) sliding against a counterbody ball (unidirectional reciprocating sliding, sphere-on-flat) with respect to a *RE* reference electrode (e.g. Ag/ AgCl (3 M KCl)) *via* a *V* voltmeter and *CE* counter-electrode (platinum or graphite) *via* an *A* ammeter respectively. *F*<sup>N</sup>

**Figure 3.** Schematic view of a hypothetical cyclic potentiodynamic polarization (cathodic and anodic) plot for

normal force, *F*T tangential force, *f* sliding frequency, *D* displacement amplitude.

84 Metallic Glasses - Properties and Processing

determining localized corrosion parameters.


Many of the foregoing determined corrosion key parameters are based on empirical observations. As with any empirical method, it is perplex with many questions about the extent of its validity. The power of this technique should not be over-estimated, since the values of *E*<sup>b</sup> and *E*<sup>r</sup> are subject to change by a number of factors. Typically, for instance, environmental changes (e.g., temperature, pH, reagent as chloride ions) will influence these values drastically. Moreover, this method can be function of scan rate, pit size or depth, polarization curve shape, and specimen geometry, which can affect the accuracy of these electrochemical parameters (*E*<sup>b</sup> and *E*<sup>r</sup> ) [24]. For example, if the scanning rate is too high, the *E*<sup>b</sup> usually has a higher value than the correct value. This is due to the potential dependence of the induction period (time required for localized corrosion to initiate) which is long at low potentials and short at high potentials. To overcome this issue, the *E*<sup>b</sup> can be measured by using a stationary potentiostatic method, in which the electrode is polarized at a constant potential and the time dependence of the current is measured. If the current starts to increase with time at a particular potential, then this potential is the correct *E*<sup>b</sup> . Due to these uncertainties, this technique should be considered as qualitative and should never be used alone. In fact, it is often desirable to use at least two different types of corrosion monitoring devices whenever possible to weed out spurious or inaccurate readings.

where, *W*<sup>v</sup>

, and *W*<sup>r</sup>

the dimensionless Archard wear coefficient, *A*<sup>r</sup>

load, and *H* the hardness of the worn material.

occurrence of triboluminescence and triboelectricity.

This remains so far valid only as part of the case-by-case study.

represent the volumetric loss (assigned as total volume of wear debris

the real area of contact, *F*N the applied normal

Metallic Glasses for Triboelectrochemistry Systems http://dx.doi.org/10.5772/intechopen.78233 87

produced), and the wear rate (usually expressed per unit sliding distance) respectively. *k* is

This equation was originally used for the case of adhesive wear [27, 28], then it was extended to more cases including that of tribocorrosion. This is because the *k* parameter in the Eq. (1) exclusively remained the only flexible parameter consistent with the case to which the wear may originate. For example, unidirectional sliding of mild steel against mild steel without any lubricant has a *k* of 10−2, whereas, for stellite sliding against tool steel, *k* is 10−5 [29]. Even more confusing is that, according to literature, the Archard wear coefficient can vary by two orders of magnitude for the same couple of materials just due to a slight change in load or speed [30]. These findings should be taken with precautions in view of the number of empirical error cases reported with respect to the wear reproducibility and validation of test methods. Friction and wear properties are often considered as subjects of poor accuracy in comparison with materials intrinsic properties. Indeed, comparative round robin studies on the topic have shown that the reproducibility of wear derived from different inter-laboratories with the same material pairing was often very poor [31]. Using the wear track width, the scattering was roughly 50% whereas the scatter in the wear coefficient was over three orders of magnitude. No clear correlation was found between a single and constant parameter (type of tribometer, normal force, and sliding velocity) and the wear rates measured in inter-laboratories. Interestingly, a good convergence was found between the wear volume loss and the energy dissipated in the tribocontact zone [32–36]. This can readily be explained by the fact that the dissipation of frictional energy is one, among others, of the main causes of triboelectrochemistry, playing an essential contributing role in wear mechanism, in this case entailing an acceleration (e.g. chemical wear rates) or modification of tribochemical reactions. The yielded frictional heat between interacting surfaces leads to a stationary rise in temperature at surface contact asperities and flashes. Furthermore, such frictional energy can take the form of high quantum excitations with short lifetime of surface and bulk sites due to the mechanochemical forces involved during the sliding process. Those excitations are also responsible for the

Tribologists nowadays are seeking for an agreement due to the fact that there is an unavoidably need to address more fundamental research towards the establishment of an original formulation or a universal methodology to define a "wear criterion" in order to better understand the complexity of the wear process in a tribocorrosion test. Although, this aim has not yet been achieved, a fair amount of progress has been made on this matter-oriented approach.

To conclude, research must focus on establishing an approach that emphasizes the nature of the dependence of the mechanochemical wear rate (output) on the energetic aspect of sliding friction, the electrochemical aspect of the exposure of bare metal surface, and the transformation of the subsurface material (input). This usually should incorporate materials properties, and behavior. If it does, this could be very useful to solve most the issues and difficulties encountered in the specific field (e.g., cases involving the failure of mechanical systems related matter at any given time), and thereby leading to a better improvement of the reliability life of

A complete discussion regarding DC and AC electrochemical techniques for the measurement of corrosion rates is beyond the scope of this chapter. A more comprehensive treatment of this subject area may be found elsewhere [9, 11, 22, 65].

#### **3.3. Aspects of tribocorrosion**

Wear is an unavoidable and a potentially serious problem in all areas of engineering. In particular, wear due to tribocorrosion is reflected by a loss of material from the exposure to corrosion of contacting solid surfaces and in relative motion. Designers and engineers who have to make optimal decisions in situations where tribocorrosion considerations are significant, need to know "how long will a component last?" To solve this question, numerous models have been developed so far to distinguish this material loss due to tribocorrosion. These models usually correlate a wear volume or a wear rate with physical and geometrical quantities. Various expressions have since been attributed to this material loss, of which the material loss can be defined in terms of weight, volume, surface, depth, width or even charge density or current density, per unit hardness, per unit frictional dissipated energy (work due to the tangential force), per unit input energy (work due to the normal force), or even per unit sliding distance, or sliding time, sliding frequency, contact frequency, etc. It becomes readily understandable of the complexity of comparing results between the various wear data published so far. It is expected then that the terminology in this field is rather uncertain, and it will remain so for a certain time, hence the need for a specific standardization, despite some recent progress made in this area [25].

One of the earlier attempts to predict the wear rate or wear volume loss of a material in sliding contact is the commonly Archard wear criterion [26] used during the second half of the twentieth century. That criterion is usually expressed as follows,

$$\mathcal{W}\_v = k \frac{F\_\text{\textquotedblleft}}{H} \quad \text{or} \quad \mathcal{W}\_r = k A\_r \tag{1}$$

where, *W*<sup>v</sup> , and *W*<sup>r</sup> represent the volumetric loss (assigned as total volume of wear debris produced), and the wear rate (usually expressed per unit sliding distance) respectively. *k* is the dimensionless Archard wear coefficient, *A*<sup>r</sup> the real area of contact, *F*N the applied normal load, and *H* the hardness of the worn material.

Many of the foregoing determined corrosion key parameters are based on empirical observations. As with any empirical method, it is perplex with many questions about the extent of its validity. The power of this technique should not be over-estimated, since the values of *E*<sup>b</sup>

 are subject to change by a number of factors. Typically, for instance, environmental changes (e.g., temperature, pH, reagent as chloride ions) will influence these values drastically. Moreover, this method can be function of scan rate, pit size or depth, polarization curve shape, and specimen geometry, which can affect the accuracy of these electrochemical

) [24]. For example, if the scanning rate is too high, the *E*<sup>b</sup>

higher value than the correct value. This is due to the potential dependence of the induction period (time required for localized corrosion to initiate) which is long at low potentials and

ary potentiostatic method, in which the electrode is polarized at a constant potential and the time dependence of the current is measured. If the current starts to increase with time

technique should be considered as qualitative and should never be used alone. In fact, it is often desirable to use at least two different types of corrosion monitoring devices whenever

A complete discussion regarding DC and AC electrochemical techniques for the measurement of corrosion rates is beyond the scope of this chapter. A more comprehensive treatment

Wear is an unavoidable and a potentially serious problem in all areas of engineering. In particular, wear due to tribocorrosion is reflected by a loss of material from the exposure to corrosion of contacting solid surfaces and in relative motion. Designers and engineers who have to make optimal decisions in situations where tribocorrosion considerations are significant, need to know "how long will a component last?" To solve this question, numerous models have been developed so far to distinguish this material loss due to tribocorrosion. These models usually correlate a wear volume or a wear rate with physical and geometrical quantities. Various expressions have since been attributed to this material loss, of which the material loss can be defined in terms of weight, volume, surface, depth, width or even charge density or current density, per unit hardness, per unit frictional dissipated energy (work due to the tangential force), per unit input energy (work due to the normal force), or even per unit sliding distance, or sliding time, sliding frequency, contact frequency, etc. It becomes readily understandable of the complexity of comparing results between the various wear data published so far. It is expected then that the terminology in this field is rather uncertain, and it will remain so for a certain time, hence the

need for a specific standardization, despite some recent progress made in this area [25].

*F*\_\_\_*N*

twentieth century. That criterion is usually expressed as follows,

*Wv* = *k*

One of the earlier attempts to predict the wear rate or wear volume loss of a material in sliding contact is the commonly Archard wear criterion [26] used during the second half of the

*<sup>H</sup>* or *Wr* = *kAr* (1)

usually has a

can be measured by using a station-

. Due to these uncertainties, this

and *E*<sup>r</sup>

parameters (*E*<sup>b</sup>

and *E*<sup>r</sup>

86 Metallic Glasses - Properties and Processing

**3.3. Aspects of tribocorrosion**

short at high potentials. To overcome this issue, the *E*<sup>b</sup>

possible to weed out spurious or inaccurate readings.

of this subject area may be found elsewhere [9, 11, 22, 65].

at a particular potential, then this potential is the correct *E*<sup>b</sup>

This equation was originally used for the case of adhesive wear [27, 28], then it was extended to more cases including that of tribocorrosion. This is because the *k* parameter in the Eq. (1) exclusively remained the only flexible parameter consistent with the case to which the wear may originate. For example, unidirectional sliding of mild steel against mild steel without any lubricant has a *k* of 10−2, whereas, for stellite sliding against tool steel, *k* is 10−5 [29]. Even more confusing is that, according to literature, the Archard wear coefficient can vary by two orders of magnitude for the same couple of materials just due to a slight change in load or speed [30]. These findings should be taken with precautions in view of the number of empirical error cases reported with respect to the wear reproducibility and validation of test methods. Friction and wear properties are often considered as subjects of poor accuracy in comparison with materials intrinsic properties. Indeed, comparative round robin studies on the topic have shown that the reproducibility of wear derived from different inter-laboratories with the same material pairing was often very poor [31]. Using the wear track width, the scattering was roughly 50% whereas the scatter in the wear coefficient was over three orders of magnitude. No clear correlation was found between a single and constant parameter (type of tribometer, normal force, and sliding velocity) and the wear rates measured in inter-laboratories. Interestingly, a good convergence was found between the wear volume loss and the energy dissipated in the tribocontact zone [32–36]. This can readily be explained by the fact that the dissipation of frictional energy is one, among others, of the main causes of triboelectrochemistry, playing an essential contributing role in wear mechanism, in this case entailing an acceleration (e.g. chemical wear rates) or modification of tribochemical reactions. The yielded frictional heat between interacting surfaces leads to a stationary rise in temperature at surface contact asperities and flashes. Furthermore, such frictional energy can take the form of high quantum excitations with short lifetime of surface and bulk sites due to the mechanochemical forces involved during the sliding process. Those excitations are also responsible for the occurrence of triboluminescence and triboelectricity.

Tribologists nowadays are seeking for an agreement due to the fact that there is an unavoidably need to address more fundamental research towards the establishment of an original formulation or a universal methodology to define a "wear criterion" in order to better understand the complexity of the wear process in a tribocorrosion test. Although, this aim has not yet been achieved, a fair amount of progress has been made on this matter-oriented approach. This remains so far valid only as part of the case-by-case study.

To conclude, research must focus on establishing an approach that emphasizes the nature of the dependence of the mechanochemical wear rate (output) on the energetic aspect of sliding friction, the electrochemical aspect of the exposure of bare metal surface, and the transformation of the subsurface material (input). This usually should incorporate materials properties, and behavior. If it does, this could be very useful to solve most the issues and difficulties encountered in the specific field (e.g., cases involving the failure of mechanical systems related matter at any given time), and thereby leading to a better improvement of the reliability life of selected material and design technologies when adopted in mechanical articulations. Further, this could predict materials performance in an environment where tribocorrosion plays a significant role.

In tribocorrosion phenomena, where tribological contacts are exposed to corrosive environments, such as aqueous lubricants, the contact materials are subject to both mechanical, and chemical/electrochemical solicitations, which contribute to material removal from sliding surfaces. The rate of material degradation/removal cannot be predicted simply by adding the wear rate in absence of corrosion to the corrosion rate in absence of wear. The reason is that corrosion and wear do not proceed independently and synergistic effects usually (but not always) result in accelerated material degradation (tribocorrosion). In that respect, theoretical models have been developed so far with respect to mechanical, chemical, and electrochemical factors and their mutual interactions, and which can be tested under well-controlled experimental conditions. In general, modeling has followed either an empirical or a mechanistic approach. The empirical approach is based on the independent measurement of material loss due to wear and corrosion. These parameters are summed up and compared to the material loss due to tribocorrosion. The difference between the two is termed synergy (Δ*W*syn). A general equation for this approach is of the form [37–39],

$$\text{IV}\_{\text{tot}} = \text{IV}\_{\text{mac}} + \text{IV}\_{\text{cor}} + \Delta\text{IV}\_{\text{sym}} \tag{2}$$

effect of repeated sliding may cause the removal of metal particles by asperities burrowing

*W*tot = *W*che(wac) + *W*mec (3)

where, *W*che(wac) is the electrochemical contribution to wear; it is termed wear accelerated corrosion and it reflects the material loss due to corrosion in the presence of wear. *W*mec is the mechanical wear, and it reveals the material loss due to wear in the presence of corrosion, and which can be related to processes as that for the formation-ejection of oxide debris, oxide

*W*tot can be determined by measuring the volume of the wear scar post-experiment using, for instance, a laser non-contact profilometry or by on-line measurement of the rate of moving down of the counter-body (e.g., a pin) on the surface wear track during sliding. The latter method has the advantage of recording an instantaneous wear rate, but it would only be applicable if no significant amount of solid reaction products (such as third body particles) accumulate in the contact zone during the tribocorrosion experiment. Under potentiostatic control, the electrochemical term (*W*che(wac)) can barely be related to the anodic corrosion cur-

a,tribocor) measured under mechanical sliding wear (occasionally by subtracting the background current) using Faraday's law. The amount of anodically oxidized metal under such

where, *W*che(wac) is the volume of the metal transformed by anodic oxidation in a triboelec-

process, which is obtained by integrating the measured current *I*a,tribocor over the time of the triboelectrochemical experiment, *M* is the atomic mass of the metal, *z* is the valence for oxidation reaction, *F* is the Faraday constant (96,480 C/mol) and *ρ* is the density of the metal.

This equation is credible and independent of whether the anodic oxidation leads to the forma-

It is worthwhile to note that few assumptions must be met in order for the Eq. (4) to be used

• The measured current must be equal to the anodic partial current for metal oxidation, which means that cathodic partial currents due to the reaction of oxidizing agents must be negligible. This can be performed by anodic polarization into the passive potential region.

The mechanical wear (*W*mec) is taken as the difference between the total wear volume *W*tot and

tion of dissolved metal ions or solid reaction products, such as oxide films.

• The charge number *z* for the oxidation reaction must be known [12, 40, 42].

the chemical wear volume *W*che(wac) determined from the electric charge.

<sup>a</sup>,tribocor.*dt*, is the electric charge generated during that transformation

*<sup>ρ</sup>*.*z*.*<sup>F</sup>* (4)

Metallic Glasses for Triboelectrochemistry Systems http://dx.doi.org/10.5772/intechopen.78233 89

layers or any corrosion products, and plastically detached metal.

Therefore, the overall wear volume due to tribocorrosion, *W*tot, can be defined as follow:

beneath the surface [12, 42, 43].

conditions is calculated as follows [12]:

trochemical test. *q*(*t*) = ∫*I*

[12, 40], namely:

*<sup>W</sup>*che(wac) <sup>=</sup> *<sup>M</sup>*.*<sup>q</sup>* \_\_\_\_\_

rent (*I*

where, *W*mec represents the material loss due to wear measured in the absence of corrosion, and *W*cor is the material loss due to corrosion only without any influence of mechanical wear.

Although, the empirical approach is technically feasible which allows for the ranking and the performance of materials based on their resistance to tribocorrosion in engineering systems, it is still time-consuming, quite economically not justifiable in the long-term, and furthermore, it integrates a synergy term, which has no physical meaning.

The advantage of a mechanistic approach is that it leads for a better understanding of the physical processes involved in tribocorrosion by incorporating the notion of synergism into the mechanical and electrochemical terms. Many factors can be responsible for the mutual dependence of mechanical and chemical material removal in a tribocorrosion system. For example, local abrasion of the passive film can lead to wear accelerated corrosion due to rapid dissolution of the locally depassivated metal surface, followed by repassivation [40]. The abrasive action of hard oxide particles formed by corrosion can accelerate the mechanical metal removal by wear [41]. The plastic deformation of the surface layer of a rubbing metal can lead to a transfer of material to the opposite body resulting in a reduction of the corrosive wear rate [42].

Therefore, it is important to distinguish material loss due to chemical or electrochemical oxidation (i.e., wear accelerated corrosion) from material removed due to mechanical wear (i.e., mechanical material removal from the sliding contact). The former arises from the fact that an asperity sliding on a material surface produces a fresh wear track zone of clean bare material (i.e. metal), which is usually more susceptible to corrosion than the same surface subjected to free corrosion under no mechanical plastic contact or sliding conditions. The effect of repeated sliding may cause the removal of metal particles by asperities burrowing beneath the surface [12, 42, 43].

Therefore, the overall wear volume due to tribocorrosion, *W*tot, can be defined as follow:

selected material and design technologies when adopted in mechanical articulations. Further, this could predict materials performance in an environment where tribocorrosion plays a

In tribocorrosion phenomena, where tribological contacts are exposed to corrosive environments, such as aqueous lubricants, the contact materials are subject to both mechanical, and chemical/electrochemical solicitations, which contribute to material removal from sliding surfaces. The rate of material degradation/removal cannot be predicted simply by adding the wear rate in absence of corrosion to the corrosion rate in absence of wear. The reason is that corrosion and wear do not proceed independently and synergistic effects usually (but not always) result in accelerated material degradation (tribocorrosion). In that respect, theoretical models have been developed so far with respect to mechanical, chemical, and electrochemical factors and their mutual interactions, and which can be tested under well-controlled experimental conditions. In general, modeling has followed either an empirical or a mechanistic approach. The empirical approach is based on the independent measurement of material loss due to wear and corrosion. These parameters are summed up and compared to the material loss due to tribocorrosion. The difference between the two is termed synergy (Δ*W*syn). A gen-

*W*tot = *W*mec + *W*cor + Δ*W*syn (2)

where, *W*mec represents the material loss due to wear measured in the absence of corrosion, and *W*cor is the material loss due to corrosion only without any influence of mechanical wear. Although, the empirical approach is technically feasible which allows for the ranking and the performance of materials based on their resistance to tribocorrosion in engineering systems, it is still time-consuming, quite economically not justifiable in the long-term, and furthermore,

The advantage of a mechanistic approach is that it leads for a better understanding of the physical processes involved in tribocorrosion by incorporating the notion of synergism into the mechanical and electrochemical terms. Many factors can be responsible for the mutual dependence of mechanical and chemical material removal in a tribocorrosion system. For example, local abrasion of the passive film can lead to wear accelerated corrosion due to rapid dissolution of the locally depassivated metal surface, followed by repassivation [40]. The abrasive action of hard oxide particles formed by corrosion can accelerate the mechanical metal removal by wear [41]. The plastic deformation of the surface layer of a rubbing metal can lead to a transfer of material to the opposite body resulting in a reduction of the corrosive

Therefore, it is important to distinguish material loss due to chemical or electrochemical oxidation (i.e., wear accelerated corrosion) from material removed due to mechanical wear (i.e., mechanical material removal from the sliding contact). The former arises from the fact that an asperity sliding on a material surface produces a fresh wear track zone of clean bare material (i.e. metal), which is usually more susceptible to corrosion than the same surface subjected to free corrosion under no mechanical plastic contact or sliding conditions. The

significant role.

88 Metallic Glasses - Properties and Processing

wear rate [42].

eral equation for this approach is of the form [37–39],

it integrates a synergy term, which has no physical meaning.

$$\mathcal{W}\_{\text{tot}} = \mathcal{W}\_{\text{chq(vac)}} + \mathcal{W}\_{\text{mac}} \tag{3}$$

where, *W*che(wac) is the electrochemical contribution to wear; it is termed wear accelerated corrosion and it reflects the material loss due to corrosion in the presence of wear. *W*mec is the mechanical wear, and it reveals the material loss due to wear in the presence of corrosion, and which can be related to processes as that for the formation-ejection of oxide debris, oxide layers or any corrosion products, and plastically detached metal.

*W*tot can be determined by measuring the volume of the wear scar post-experiment using, for instance, a laser non-contact profilometry or by on-line measurement of the rate of moving down of the counter-body (e.g., a pin) on the surface wear track during sliding. The latter method has the advantage of recording an instantaneous wear rate, but it would only be applicable if no significant amount of solid reaction products (such as third body particles) accumulate in the contact zone during the tribocorrosion experiment. Under potentiostatic control, the electrochemical term (*W*che(wac)) can barely be related to the anodic corrosion current (*I* a,tribocor) measured under mechanical sliding wear (occasionally by subtracting the background current) using Faraday's law. The amount of anodically oxidized metal under such conditions is calculated as follows [12]:

$$\mathcal{W}\_{\text{choyancy}} = \frac{M.q}{\rho.z.F} \tag{4}$$

where, *W*che(wac) is the volume of the metal transformed by anodic oxidation in a triboelectrochemical test. *q*(*t*) = ∫*I* <sup>a</sup>,tribocor.*dt*, is the electric charge generated during that transformation process, which is obtained by integrating the measured current *I*a,tribocor over the time of the triboelectrochemical experiment, *M* is the atomic mass of the metal, *z* is the valence for oxidation reaction, *F* is the Faraday constant (96,480 C/mol) and *ρ* is the density of the metal.

This equation is credible and independent of whether the anodic oxidation leads to the formation of dissolved metal ions or solid reaction products, such as oxide films.

It is worthwhile to note that few assumptions must be met in order for the Eq. (4) to be used [12, 40], namely:


The mechanical wear (*W*mec) is taken as the difference between the total wear volume *W*tot and the chemical wear volume *W*che(wac) determined from the electric charge.

#### **3.4. The causality of wear-corrosion synergism**

The synergy between wear and corrosion has recently attracted increasing attention to improve materials used in systems where tribocorrosion plays a role. However, wear-corrosion synergy still seems to be a developing topic of discussion as no convincing expression is available yet.

Most amorphous alloys, such as bulk metallic glasses (BMGs), usually do not depend on the presence of a protective surface film to exert a corrosion resistance. Therefore they can be expected to claim showing a negative synergistic effect, hence their opportunity to be selected as potential candidates over other passive alloys (such as stainless steels) in systems where

Metallic Glasses for Triboelectrochemistry Systems http://dx.doi.org/10.5772/intechopen.78233 91

Celis et al. [10, 13] identified five mechanisms in tribocorrosion, which could explain the synergism noticed between mechanical and chemical factors acting on contacting materials,

**1.** The debris can speed up or reduce wear compared to what happens in the same environment where debris does not exist like e.g., in sliding contacts polarized at a large cathodic

**2.** A galvanic coupling is established between the worn (active) and unworn areas (passive).

**4.** An accumulation of dissolved species may take place in the liquid surrounding the contact. This may render the medium chemically or electrochemically more aggressive;

**5.** The mechanical loading in the contact area and its nearby zone may cause a work hardening of the materials. This work hardening can alter the kinetics of corrosion and/or repas-

Most of engineering and biomedical metals and alloys oxidize and frequently passivate "spontaneously" in contact with the ambient air, and with suitable aqueous media to form "natural" thin passive surface oxide films. Alternatively, alloying is also considered to be, inter alia, the most extensively used method for enhancing the passivity of base metals [45]. Passive films that growth on most surface alloys are found to be of two types: discontinuous, and continuous [46]. Discontinuous films are porous and are formed at the metal/solution interface from the reaction of metal cations with species in solution. They have a thickness of up to 1 mm and are less protective. Continuous films, on the other hand, are tenacious and

a high electrical field. They are called barrier protective layers, which serve to prevent current flow, and corrosion (i.e., dissolution). Commonly, the passive film on most metals and alloys consist both continuous, and discontinuous layers, with discontinuous film forming the outer layer, and the continuous film forming the inner barrier protective layer. A typical example of a bi-layer passive film formed on Fe-based amorphous alloy and interacting with anions

The approach that links the improved resistance of amorphous alloys to their ability to promote amorphous passive oxide formation is well accepted by the scientific community. In high-temperature gas working conditions, vitreous or amorphous oxides offer improved

ohm.cm−2) and support

thin (nm's to μm's thickness range). They have high resistances (≥10<sup>6</sup>

adsorption in various aqueous solutions [47] is depicted in **Figure 4**.

It accelerates the anodic dissolution in the area where the metal is depassivated; **3.** A galvanic coupling may be established between the two contacting counterparts;

tribocorrosion plays a role.

mainly:

potential;

sivation processes.

**3.5. Passivity breakdown**

A positive or a negative (antagonism) synergistic effect can occur in most cases where surface interactions interfere with tribocorrosion phenomena. It intervenes especially (in a positive way) by increasing the wear (volume) when the mechanical process affects the electrochemical process and *vice versa*. In these situations, the total wear (volume loss) will be very different and greater than the sum of the mechanical wear in the absence of corrosive environment and the loss of material by corrosion in the absence of any mechanical stress. A negative effect of synergy, however, will occur when the total wear is less than the sum of the two protagonists taken individually and independently, namely wear and corrosion.

Madsen et al. [37, 38] critically reviewed the measurement of wear-corrosion synergism, and proposed a group of penetration rate equations to quantify the wear and corrosion processes and the wear-corrosion synergism. Their results showed that the wear-corrosion synergism is of great extent for alloys, such as AISI 316 stainless steels, which depends on the formation of a film of passive layers for their corrosion resistance, sometimes only a few atom layers thick, resulting from an interaction between the material and the surrounding environment [37, 38]. On the contrary, the synergism was limited for alloys, such as low alloy steels (e.g. amorphous steels), which do not depend on the presence of a passive film for their resistance to a corrosive environment. The causality of this synergistic effect has been explained in part for some passive materials. The presence on their surface of passivation layers and the ability of their surface, even in a deformed or partially damaged state, due to the sliding contact by the counter-body or a third-body, to be rehabilitated (by forming reaction layers with a thickness of a few nm, such as oxides, solid precipitates, adsorbed layers or passive surface films) is at the origin of the increment of that synergy. Dense oxide layers, precipitates, or passive films may play a protective role in isolating the underlying metal from a direct contact with the surrounding corrosive environment, thereby protecting the metal from a corrosion increment, but not necessarily their mechanical wear. In particular, one of the possible explanations could be related to the mechanical or chemical shear strength of these formed layers. This can likely be one of the causes of the incremental or decline effect of the wear-corrosion synergism of passivating metals. This is mainly true in the case of stainless steels and other alloys containing chromium. Their passive surface film formed in the ambient air or in contact with an aqueous solution has a thickness of a few nanometers but gives them a high resistance to corrosion. The sliding of a hard counterbody material on such a surface is likely to damage that passive film what is known as a "depassivation" process by which the bare material is exposed to the corrosive environment [8, 10, 13, 32, 44]. Various but essentially electrochemical processes can then compete on these bare surfaces [8, 10, 13, 32, 44], namely:


Most amorphous alloys, such as bulk metallic glasses (BMGs), usually do not depend on the presence of a protective surface film to exert a corrosion resistance. Therefore they can be expected to claim showing a negative synergistic effect, hence their opportunity to be selected as potential candidates over other passive alloys (such as stainless steels) in systems where tribocorrosion plays a role.

Celis et al. [10, 13] identified five mechanisms in tribocorrosion, which could explain the synergism noticed between mechanical and chemical factors acting on contacting materials, mainly:

