A comprehensive review on synergistic and individual effects of erosion–corrosion in ferrous piping materials

2.1 Uniform corrosion

Uniform corrosion or general corrosion is an electrochemical process in which material corrodes uniformly throughout the surfaces (King 2012). It was considered less dangerous than other forms of corrosion and was more common on unprotected surfaces exposed to corrosive environments. Uniform thinning followed by failure takes place in general corrosion. Materials like steel, aluminum, and copper undergo uniform corrosion when exposed to a corrosive environment. It can be prevented by painting or coating the surface or by galvanizing.

2.2 Flow-assisted corrosion (FAC)

The flow-assisted corrosion occurs in the metallic materials that transport aqueous solution. The accelerated loss of material occurred as a result of the fluid velocity and the corrosive action of the aqueous solution. Moreover, the mechanism of flow-assisted corrosion is entirely different from erosion–corrosion because the aqueous solution does not contain any of the solid particles that cause mechanical erosion. Furthermore, low-carbon steels and low-alloy steels exposed to flowing water are more susceptible to FAC. However, materials with the passive film have good resistance against FAC. For instance, in flow-assisted corrosion, the material degradation is due to electrochemical dissolution and the corrosion products are washed away by the flow. The increased damage to a material is due to the dissolution rate of metals and the condition of an aqueous solution (Kain 2014).

2.3 Erosion–corrosion

Erosion–corrosion is a form of corrosion where mechanical erosion and electrochemical corrosion come into the picture. It is more prevalent in pipes carrying solid entrained fluids like gas and oil field pipelines. Factors like slurry concentration, flow velocity, temperature, particle size, shape, and concentration significantly influence erosion–corrosion (Aribo et al. 2013; Selvam et al. 2019). Erosion corrosion has attracted the attention of researchers nowadays. As a result, detailed research is being carried out to explore the mechanism, characteristics, and prevention methods of erosion–corrosion. In this review, an attempt has been made to present a complete overview of erosion–corrosion.

3.1 Finnie’s erosion model for ductile material

Finnie was the first to formulate a mathematical model for single-particle erosion. The author correlated the erosion in flow direction and particle velocity (Finnie 1960). The author assumed that the target material surface was rigid plastic and hard erodent particle impinges the surface with an impact angle α. Moreover, two mathematical models were developed for low-angle and high-angle impacts.

In this model, the path traveled by the erodent on the surface is used to derive the particle’s motion equation. The material removal volume is calculated by multiplying the area and width of cutting faces. Equations (1) and (2) represent loss of material due to cutting action at low and high impact angles, respectively (Finnie 1960).

where Q is the material removal volume, M is the erodent mass, α is the impingement angle, σ_f is the plastic flow stress, V is the velocity of the erodent particle, Ψ is the ratio of the depth of contact to the depth of cut, and K is the ratio of normal force to shear force on the particle surface (Finnie 1960).

Although the Finnie erosion model is considered simplified and older, it is valid only for ductile materials. It does not include the erosive behavior of brittle materials (ElTobgy et al. 2005). Finnie and Mcfadden (1978) re-examined the single particle erosion model and pointed out that the erosion rate prediction complies with the experimental data up to 45° incident angle. However, at 90° incident angle, the model predicts no erosion rate. Shewmon and Sundararajan (1983) reviewed Finnie’s erosion model and concluded that this model does not predict erosion rate for 0° incident angle of erodent. Keating and Nešić (2001) evaluated Finnie’s model and found that the prediction was inconsistent for an axisymmetric expansion. They have also modified the erosion model by introducing Bitter’s critical velocity phenomena. Aslam Noon and Kim (2017) considered Finnie’s erosion model to study erosion wear of Francis turbine components.

3.2 Bitter’s erosion model for ductile and brittle material

According to Bitter (1963a), two types of wear occur on fluid transport lines, one is due to repeated impingement of abrasive particles (W_D), and the other is caused by erosion of loose abrasive particles (W_c1 & W_c2). He has derived an expression by considering the erodent’s mass and flow velocity, impact angle, and properties of the erodent and target material. For instance, the erosive wear due to the repeated deformation would be calculated by eq. (3). Equations (4) and (5) depict the cutting when particles leave the surface with a horizontal velocity component, and the horizontal component becomes zero after the collision respectively (Bitter 1963a). Equation (3) was developed as material removed in terms of energy absorbed by the impinging particles (Bitter 1963a).

where W_D is the material loss, M is the erodent particle mass, V is the particle velocity, α is the impingement angle, ξ is the peak erodent velocity at which the particle–target surface interaction is still purely elastic, χ is the cutting wear factor, and δ is the deformation wear factor.

The Bitter’s erosion model for brittle material requires that the wear factor χ and deformation wear factor δ be known. However, these factors depend on erodent size, speed, and other parameters. Bitter’s erosion model considers more material properties than Finnie’s erosion model. However, as like Finnie’s model, bitter’s model also predicts no erosion rate at 0° impact angle (Lyczkowski and Bouillard 2002).

3.3 Hutchings erosion model

Hutchings (1981) has developed an erosion model by considering the erodent as spherical particles with normal impingement. The earlier work of Finnie (1960) had revealed that erosion occurs at only a low impact angle, and there was no erosion at a higher incidence angle. According to the Bitter model (Bitter 1963a), deformation erosive wear was considered the dominant wear mechanism, and little justification was done for the low-angle cutting wear mechanism. Hutchings formulated two erosion models, one is for low impact angle, and the other is for high impact angle. The erosion rate due to the cutting of an individual erodent is expressed in eq. (6). Although the impingement of rounded particles produces no cutting, the erosion model can be done in two approaches. (1) Assuming that the material removal occurs when the surface plastic strain is greater than the critical value. (2) Considering this phenomenon as low-cycle fatigue due to the repeated impingement of erodents. The erosion wear related to the low-cycle is given in eq. (7).

where H_s is the target material hardness, K₁ & K₂ are the sections of material removed from the indentation as small debris, ρ_s is the erodent density, ρ_m is the target material density, ε_c is the critical plastic strain, and E_m is the ratio mass loss of target material to mass of erodent (Hutchings 1981).

The main limitation of this model is that hutching proposed the model for spherical shape particles and is not suitable for angular particles. Bingley and O’Flynn (2005) validated Hutching’s erosion model and reported that the erosion process was strongly influenced by the surface hardness of the target material. In addition, the orientation and shape of the particle determine the wear mechanism. They also reported that the erosion model showed an agreement with the experimental results.

3.4 Hashish’s erosion model

Hashish (1989) has modified Finnie’s erosion model by incorporating the particle shape, and this model is suitable for predicting ductile material wear at shallow angles. In addition, the author has introduced a constant C_k in eq. (8), which considers erodent and target materials properties.

where M is the mass of the abrasive particle, ρ_p is the erodent particle density, V is the particle velocity, α is the impingement angle, and C_k is the constant (Hashish 1989).

The C_k can be computed from eq. (9)

where R_f is the roundness factor and σ_f is the flow stress (Hashish 1989).

The main advantage of this erosion model is that it does not have any experimental constant and is the only model that considers the particle shape. However, this model is only suitable for low impact angles for ductile materials (ElTobgy et al. 2005).

5.1 Jet impingement rig

From the literature, many researchers have examined the erosion–corrosion behavior of the target material with slurry jet apparatus. The main advantage of this rig is that many parameters can be changed to evaluate the individual effects of each parameter on erosion–corrosion. The basic configuration of the jet impingement apparatus is illustrated in Figure 1. This apparatus has a centrifugal pump that circulates the abrasive-laden slurry through the nozzle to impinge the target material. The target material can be impinged with different angles by tilting the sample holder. The corrosion can be measured by incorporating the reference and auxiliary electrodes. The jet impingement apparatus with corresponding target materials chosen by various authors listed in Table 1. The disadvantage is the erodent concentration is inconsistent (Aribo et al. 2013; Giourntas et al. 2014; Hu and Neville 2009; Karafyllias et al. 2019; Owen et al. 2018; Stack and Abdulrahman 2010, 2012; Wongpanya et al. 2020; Yang and Cheng 2012; Yi et al. 2018).

Figure 1:

Schematic representation of the slurry jet impingement test rig. (Reprinted from Aribo et al. (2013), Copyright (2012), with permission from Elsevier B.V)

5.2 Slurry pot erosion apparatus

The lack of standards related to erosion–corrosion has led researchers to develop the apparatus to meet their specific requirements. Although having different configurations but work on the same principle. The slurry pot apparatus is classified as a rotating type apparatus. The researchers utilized the slurry pot erosion apparatus to assess erosion–corrosion by rotating the target material against the slurry kept in a pot. The erosion occurs as the rotary motion of the target material in the slurry pot tends to remove the material from the surface and the corrosive medium will corrode the material. This combined action is responsible for the increased material loss. As far as the configuration is concerned, this apparatus has a cylindrical container that accommodates the test samples and the three electrodes for corrosion testing: the working electrode, counter electrode, and reference electrode. The container is filled with slurry, made of either normal tap water or a corrosive medium. The samples are held by a holder, as shown in Figure 2, connected to an electric motor to rotate the samples in the slurry. Furthermore, the heating system is also used to control the temperature of the slurry. Despite having the advantage of handling slurry and erodent easily, the impact angle cannot be varied (Jones and Llewellyn 2009; Lindgren and Perolainen 2014a, b; Rajahram et al. 2009, 2011; Yu et al. 2013).

Figure 2:

Illustration of slurry pot erosion–corrosion test rig. (Reprinted from Rajahram et al. (2011), Copyright (2010), with permission from Elsevier Ltd.)

5.3 Coriolis erosion tester

To simulate the actual working condition of pumps and other hydraulic components, the researchers have developed an apparatus utilizing the Coriolis acceleration and centrifugal force. Figure 3 depicts the schematic diagram of the Coriolis erosion tester. It is a rotating type apparatus that has a rotor with a radial passage and sample holders on either side, equidistant from the center. The samples are fixed on either side of the holder. The slurry containing erodent particles is supplied to the central hole of the rotor, and it flows through the radial passage and over the sample surface as the rotor rotates. Thereby, erosion occurs in the surface of the sample. The rotor speed, solid concentration, and erodent size greatly control erosion (Clark et al. 1999; Hawthorne 2002; Hawthorne et al. 2002; Hawthorne and Xie 2005; McI Clark et al. 2000; Xie et al. 1999).

Figure 3:

Diagram showing the Coriolis erosion test rig. (Reprinted from Xie et al. (1999), Copyright (1999), with permission from Elsevier B.V.)

5.4 Flow loop apparatus

Besides turbines, pumps, and valves, erosion–corrosion is the major failure mechanism in oil and gas field pipelines and pipes used in desalination plants. It can be assessed by many test rigs, as mentioned in the previous sections. However, the assessment made by the laboratory test rigs cannot assimilate the actual erosion–corrosion. To overcome these problems, researchers have developed a setup like the actual pipeline shown in Figure 4, which contains joints, bends, and elbows. This system has a slurry tank and pumps to create a flow, and the slurry is recirculated through the pipes and other joints for a particular time. Corrosion studies can be made by installing the electrodes at different locations. The temperature control systems are often integrated with the loop system to control the thermal gradient of the slurry. Then the elbows and joints are removed from the loop system for further analysis of erosion–corrosion (Amara et al. 2018; Elemuren et al. 2018, 2019, 2020; Liu et al. 2017; Zhou et al. 2022).

Figure 4:

Illustration of the loop system employed for the assessment of erosion–corrosion in elbows and other fittings used in the piping system. (Reprinted from Elemuren et al. (2020), Copyright (2019), with permission from Elsevier Ltd.)

6.1 Effect of erodent hardness

The hardness of the erodent is an important property that must be considered when investigating the erosion–corrosion synergy of metallic materials. Hardness is the surface property often defined as the resistance to indentation. Many research studies have considered the erodent hardness in pure erosion. However, only a few papers considered the erodent hardness in the erosion–corrosion synergy. It has been found that the harder erodent resulted in an increased material loss in the pure erosion tests. However, further studies are needed to conclude the contribution of erodent hardness on synergy. Roy et al. (1993) pointed out that particle hardness does not always result in an increase in mass loss. They also established the combined effect of particle hardness and shape. Hutchings and Shipway (2017) have reported that the ratio of erodent hardness H_a to material’s surface hardness H_s has a significant role in erosion wear. It was concluded that the plastic flow occurs when the H_a/H_s > 1.2. The interaction of the particle and the metallic surface is shown in Figure 6. The micro-cutting or plowing occurs when the impinging pressure is higher than the yield strength of target material.

Figure 6:

Interaction of erodent and the surface under normal load (A) H_a > 1.2 H_s, indentation occurs on the surface (B) H_a <1.2 H_s, plastic deformation occurs in the erodent. (Reprinted from Hutchings and Shipway (2017), Copyright (2017), with permission from Elsevier Ltd.)

6.2 Effect of erodent size

The size of the erodent also influences erosion wear. Many researchers have investigated the effect of erodent size concerning pure erosion. Still, considerable work is needed to comprehend the impact of erodent size on erosion–corrosion synergy. The evaluation of erosion characteristics of stainless steel with three different erodent sizes (75 µm, 150 µm, & 300 µm) was studied by Lin et al. (2018). It was found that the erosion ratio increases when the erodent size is above 150 µm. The sharpness of the erodent also influences the erosion rate. However, smaller erodents with high angularity resulted in high erosion rate. The SEM image of quartz with different size ranges (Lindgren and Perolainen 2014a) is shown in Figure 7. Rajahram et al. (2009) performed the erosion–corrosion test on carbon steel, aluminum bronze, and stainless steel with a slurry pot tester under various test conditions. They found that the higher erosion rate resulted in the medium sand size followed by coarse sand size. The fine size sand yields a minimum erosion rate because of the lower collision energy. The small particles are affected by particle retardation, leading to lower kinetic energy dissipation and decreased material loss. It was also evident that because of the squeeze film effect, the particles with a size less than 100 µm could not rebound from the surface, and particles were not able to penetrate the squeeze film (Lynn et al. 1991). According to Bahadur and Badruddin (1990), the width (w) to length (l) ratio w/l and perimeter squared to area P²/A were considered an indicator of particle size. They have experimentally proved that an increase in erodent size significantly increases mass loss. Venkatraman Krishnan and Lim (2021) examined the slurry erosion of S275JR mild steel material by considering individual and coupled effects of flow velocity, erodent size, impingement angle, and stress exerted by an external load. It has been reported that the increase in erodent size in the slurry decreases the erosion wear rate. This statement was contrary to the previous research findings mentioned in this section. The author has argued that increasing the erodent size reduces the impinging particles on the target surface, resulting in less wear rate. Figure 8 illustrates the impact of particle size on the erosion wear rate of P110 steel in 1.2 wt% of SiC powders at 18.7 m/s nominal test speed.

Figure 7:

SEM image of quartz with different size ranges: (A) quartz (75–100 µm), (B) quartz (125–180 µm), and (C) quartz (100–600 µm). (Reprinted from Lindgren and Perolainen (2014a), Copyright (2014), with permission from Elsevier B.V.)

Figure 8:

Variation of erosion rate (g m⁻² min⁻¹) as a function of particle size (µm) with 18.7 m/s particle velocity and 1.2 wt% SiC concentration in P110 steel. (Reprinted from Lynn et al. (1991), Copyright (1991), with permission from Elsevier B.V.)

6.3 Effect of erodent shape

Like many other properties, erodent shapes also affect erosion wear (Bahadur and Badruddin 1990; Lin et al. 2018; Roy et al. 1993). However, most of the research work pertains to pure erosion. Moreover, the impact of particle shape on erosion–corrosion is rarely discussed in the literature and is still not fully understood. Metal removal mechanism has been significantly influenced by the shape of erodent particles. Three mechanisms of metal removal occur when the solid particle and surface interact: cutting, plowing, and indentation. The different shapes of particles chosen by the authors are listed in Table 2. The angular erodent and low impact angle results in micro-cutting and the formation of microchips. In contrast, spherical erodent and high impact angle resulted in either plowing with extruded lips or indentation of erodent on the surface. The angular abrasive particles significantly increase erosion wear. Walker and Hambe (2015) used two methods: the circularity factor and spike parameter to characterize the particle shape. Singh et al. (2021) determined the effect of erodent shape on the erosion–corrosion of mild steel. They have found that erosion wear increases with particle angularity. Sundararajan and Roy (1997) assessed the effect of angular and circular particles on metallic materials. Figure 9 illustrates the impact of an angular and spherical particle on the target surface. For the same impact energy upon impingement on the target material, the angular particle concentrated on a smaller area resulted in increased wear by a high strain rate. For instance, the spherical particles caused less wear due to the impact energy distribution over a larger area (Sundararajan and Roy 1997). Levy and Chik (1983) conducted the erosion wear test on carbon steel to determine the influence of erodent composition and shape. The authors used six different compositions of erodent (calcite, apatite, alumina, sand, silicon carbide, & steel) and two different particle shapes (angular and circular) to examine the influence of erodent shape on wear. The results revealed that the angular grits caused sharper craters on the metallic surface and produced more efficient extruded platelets. The rounded or spherical grits caused shallow craters and did not produce platelets effectively. They have also found that the angular erodents produced four times more erosion than the circular erodents.

Figure 9:

Diagram showing the impact of (A) angular and (B) spherical particle. (Reprinted from Venkatraman Krishnan and Lim (2021), Copyright (2021), with permission from Elsevier B.V.)

6.4 Effect of particle concentration

Particle concentration, or solid concentration or slurry concentration in some literature, is another factor that affects the erosion–corrosion wear behavior of engineering materials. Over the years, it has been found that many researchers have evaluated the impact of slurry concentration on the erosion–corrosion wear characteristics of metallic materials. All the research work concludes with the same result. The increase in solid concentration in the slurry facilitates the impingement of the maximum quantity of erodents on the material’s surface, resulting in a high wear rate related to erosion–corrosion (Elemuren et al. 2018, 2020; Rajahram et al. 2009; Ribu et al. 2022; Yang and Cheng 2012; Zhou et al. 2022. According to Yang and Cheng (2012), the increase in the erosion–corrosion rate of the steel corresponds to the increased slurry concentration and flow velocity. It has been shown in Figure 10. Rajahram et al. (2011) analyzed the impact of solid concentration on the erosion–corrosion behavior of austenitic stainless steel. It was found that increased solid concentration increases mass loss and passive oxide film removal rates. As a consequence, passive film removal enhances the material dissolution rate. In the study of Elemuren et al. (2020), the results showed that particle velocity was the most influencing parameter on erosion–corrosion wear rate, followed by slurry concentration. It has been found that the slurry velocity and solid concentration responsible for 87 % of the total mass loss of mild steel. In another research work by Rajahram et al. (2009), the SEM images showed that the increased solid concentration damaged the steel surface, and multiple extruded lips, platelets, craters, and shallow indentations were observed.

Figure 10:

Plot depicting the effect of sand concentration (wt%) on weight loss (mg cm⁻² h⁻¹) for erosion–corrosion, corrosion and erosion in carbon steel pipes. (Reprinted from Yang and Cheng (2012), Copyright (2011), with permission from Elsevier B.V.)

In contrast to the previous findings, some research work (Turenne et al. 1989) stated that the increase in solid concentration reduces the erosion–corrosion wear rate by rebounding the particles near the material’s surface, shielding the surface from successive impacts.

6.5 Effect of temperature of slurry

Apart from particle properties, the slurry’s temperature is considered an important parameter in studying erosion–corrosion. The researchers covered mainly medium carbon steels and alloy steels. Tian et al. (2009) have done the erosion–corrosion test on high-Cr cast iron alloys. The results exhibited that the temperature increase from 32 °C to 47.5 °C resulted in an increased mass loss rate of 19.6 % from 126.3 % at different particle sizes. Hu and Neville (2009) explained the relationship between temperature and total mass loss in X65 steel under CO₂ conditions. It indicates that the total mass loss and temperature have an exponential relationship, as seen in Figure 11.

Figure 11:

Total mass loss (mg) versus temperature with 200 mg/L solid concentration and three flow velocities. (Reprinted from Hu and Neville (2009), Copyright (2009), with permission from Elsevier B.V.)

From the literature (Hu and Neville 2009; Lindgren et al. 2016; Senatore et al. 2021; Tian et al. 2009), it is clear that temperature rise decreases the viscosity of the slurry; as a consequence, the erodent particles strike the target material surface with less screen effect, causing a high wear rate. Meng et al. (2007) have compared the impact of slurry temperature on the erosion–corrosion behavior of austenitic and duplex stainless steels. They found that temperature strongly influences the mass loss by pure corrosion than pure erosion. Aribo et al. (2013) tested different stainless steels and found that the total material loss increased with temperature increase.

6.6 Effect of pH of the slurry

The pH is an important factor that can influence the erosion–corrosion of engineering materials. However, only a few papers have considered the influence of pH as a parameter (Moloto et al. 2019). Tian et al. (2009) investigated the erosion–corrosion behavior of high-Cr white cast iron with three pH values and three chloride concentrations. The authors reported that the corrosion was stable at pH 7 and absence of (0 ppm) chloride concentration, and the highest mass loss corresponds to the lowest pH value and maximum chloride concentration. Figure 12 depicts the mass loss rate with corrosion conditions. The recent work by Laukkanen et al. (2020) proved that an acidic environment increases the oxide layer dissolution rate. The maximum mass loss was observed in 316L stainless steel at the lowest pH values. Dalbert et al. (2020) investigated the effect of sulfate solution (pH 1.5, 6.5, & 12.5) on ferritic stainless steels. The authors confirm that the pH is the most influencing parameter of wear rate. They also found that pH strongly influenced the kinetics of passive film growth, which determines the repassivation rate. The ferritic stainless steels showed faster repassivation in an acidic solution than in a neutral solution. Although stainless steels were considered high corrosion resistance alloys, in acidic conditions, the passive layer becomes more susceptible to dissolution. Karafyllias et al. (2019) have investigated the effect of pH on erosion–corrosion resistance of high chromium cast irons and stainless steels. The authors have found some interesting results. At a neutral pH level, the high chromium cast irons showed superior corrosion resistance than stainless steels. However, the austenitic structured material in acidic environments showed higher resistance than martensitic stainless steels. It is attributed to the highest chromium content in the high chromium cast irons and austenitic stainless steels. In addition, the total mass loss and synergy of erosion–corrosion were intact for the austenitic stainless steels when lowering the pH from 7 to 3.

Figure 12:

Plot showing the mass loss rate of high chromium white irons with corrosion conditions (different pH levels and Cl⁻ concentrations) at 10 µm solids particle size at 32 °C. (Reprinted from Tian et al. (2009), Copyright (2009), with permission from Elsevier B.V.)

6.7 Effect of flow velocity

Flow velocity or particle velocity, or impact velocity, is the most studied parameter concerned with the erosion–corrosion analysis of materials. Researchers have conducted extensive research to establish the relationship between particle velocity and erosion–corrosion synergy. Many researchers with different materials have reported the impact of particle velocity on erosion–corrosion: X65 steel (Aguirre et al. 2019; Hu and Neville 2009; Owen et al. 2018; Rameshk et al. 2020; Zhao et al. 2016), 2205 duplex stainless steel (Aribo et al. 2013; Kim et al. 2019; Lindgren and Perolainen 2014a; Loto 2017), AISI 1018 steel (Elemuren et al. 2018, 2019, 2020), austenitic stainless steels (Hussain et al. 2013; Lindgren and Perolainen 2014a; López et al. 2011; Tian et al. 2022), and titanium (Khayatan et al. 2017; Lindgren and Perolainen 2014a; Wang et al. 2019). Yang and Cheng (2012) have proved that in carbon steels the erosion rate increases with the particle velocity. They have also highlighted that the increased slurry velocity has changed the mechanism of material removal from erosion–corrosion to pure erosion mechanism. Hu and Neville (2009) analyzed the interaction of solid concentration and particle velocity on erosion–corrosion. The findings showed the dependency of solid concentration of velocity. The total mass loss increased with impact velocity and slurry concentration. In most literature, it was found that the change in mass loss after crossing a threshold limit of slurry velocity is called critical flow velocity. Beyond that, there is a change in the material removal mechanism from erosion–corrosion to pure corrosion or erosion. The nozzle diameter and flow velocities mentioned in the literature are listed in Table 2.

The research work of Yi et al. (2018) found that beyond the critical flow velocity, the surface roughness and weight loss tend to rise in stainless steels. Yu et al. (2013) experimented with carbon steel to evaluate the influence of slurry velocity on erosion–corrosion. It revealed that erosion dominates material removal more than corrosion at high velocities.

It has been found that only a little literature is available regarding the impact of particle velocity on the erosion–corrosion synergy of stainless steels. Furthermore, Nguyen et al. (2014) studied flow velocity’s effect on austenitic stainless steel. It was determined that erosion rate and surface roughness increased with increasing slurry velocities. Meng et al. (2007) conducted an erosion–corrosion study on duplex and austenitic stainless steels and discovered that the overall mass loss and erosion component are considerably affected by the flow velocity and solid concentration. Furthermore, the authors also reported a strong relationship between velocity and solid concentration.

Following the previous findings, Zadeh and Rashidi (2020) also performed an erosion–corrosion test on CK45 steel with different velocities and reported that the flow velocity tends to increase the synergism rate. Furthermore, the recent work by Venkatraman Krishnan and Lim (2021) revealed that the erosion rate increases when the slurry velocity increases. Moloto et al. (2019) investigated the erosion–corrosion of standard duplex steel in mine water using Raman spectroscopy and X-ray diffraction. The authors correlated the presence of corrosion products found by Raman spectroscopy and X-ray diffraction with the erosion–corrosion rate. Figure 13 displays the variation mass loss rate with flow velocity in duplex stainless steel.

Figure 13:

Illustration of total mass loss (mg) as a function of impact velocity (rpm) for erosion–corrosion and pure erosion of standard duplex stainless steel in mine water. (Reprinted from Moloto et al. (2019), Copyright (2019), with permission from Elsevier Ltd.)

Though it is a proven fact that the wear rate increases in accordance with the flow velocity, some contradictions have been reported by the researchers; López et al. (2005) have compared the erosion–corrosion behavior of SS304 and SS420 stainless steel. They discovered that the mass loss of SS304 was shown to be increased until reaching a critical velocity, after which it decreased. In addition, Lindgren et al. (2016) asserted that in 904 stainless steels, the mass loss decreased when increasing the flow velocity.

6.8 Effect of impact angle

Researchers often considered impact angle or impingement angle, as a primary parameter in the study of erosion/erosion–corrosion. Moreover, the erosive wear rate can be effectively controlled by identifying which angle makes the higher and lowest wear rate. In addition to that, the impact angle determines the material removal mechanism (Andrews et al. 2014; Burstein and Sasaki 2000; Khayatan et al. 2017; Liu et al. 2021). Furthermore, the impingement of particles at various impingement angles changes the morphology of the target material surface. Therefore, SEM and TEM have been used to characterize the metal surface (Patel et al. 2020).

The influence of this parameter on erosion–corrosion has been extensively investigated in previous years for titanium, carbon steels, low alloy steels, and different grades of stainless steels. Most of the studies were carried out on ductile materials with low-impact angles. Researchers used 3.5 % NaCl-containing slurry to study the influence of impingement angle on the erosion–corrosion behavior of metallic materials (Aribo et al. 2013; Basumatary and Wood 2020; Yi et al. 2018). The investigation of erosion–corrosion on pure titanium revealed that the maximum mass loss due to erosion and erosion–corrosion rate was observed at a 40° impact angle and a further increase in impact angle decreased the mass loss (Khayatan et al. 2017). Liu et al. (2021) compared the significance of impingement angle on the erosion–corrosion behavior of X80 steel with corrosive and noncorrosive media. The weight loss of samples in pure water was observed to increase with decreasing impact angle, whereas in corrosive media, the mass loss of samples initially followed a decreasing trend after that it was increased with increasing impact angle.

Meanwhile, Burstein and Sasaki (2000) found that an oblique impact angle increases the synergism of erosion–corrosion. More specifically, the maximum mass loss by erosion–corrosion and the erosion for the SS304 were obtained at impingement angles between 40° and 50°. Andrews et al. (2014) highlighted the correlation between the scar depth and angle of impingement for Stellite 6 and SS316. The Stellite 6 exhibited brittle behavior with the highest mass loss observed for a 60° impact angle. On the other hand, due to the ductile nature of SS316, the maximum weight loss and considerable surface damage occurred at 45° impact angle. The mild steel S275JR (Venkatraman Krishnan and Lim 2021) and aluminum brass (Abedini and Ghasemi 2017) followed the same trend, i.e., the maximum weight loss was observed at low oblique angles. Venkatraman Krishnan and Lim (2021) reported that the erosion rate decreased with increasing impact angle in ductile materials. The authors also presented the material removal mechanism for different impingement angles, as given in Figure 14. The authors have reported three material removal mechanisms namely: cutting, plowing, and indentation. In cutting, the erodent particle removes the material in the form of microchips, which are physically separated from the surface of the target material. The impact angles between 5° and 20° resulted in chip formation as shown in Figure 14A. It is because of the sufficient shear component responsible for the material removal. In the case of plowing, the erodent particle impinges the surface and plastic deformation results in the extruded lip on either side of the particle as can be seen in Figure 14B. The plowing mechanism was observed for angles 25° to 85°, in that deformed lips do not physically separate from the target surface, and a 90° impact angle made an indentation (Figure 14C) on the surface and no chip was formed. Some researchers reported the highest weight loss observed at high-impingement angles (Hussain et al. 2013; Mesa et al. 2003).

Figure 14:

Diagram showing erosion mechanism: (A) cutting, (B) plowing, and (C) indentation. (Reprinted from Venkatraman Krishnan and Lim (2021), Copyright (2021), with permission from Elsevier B.V.)