Effect of tempering temperature on the wear behaviour of martensitic ductile iron

Abstract

Ductile cast irons (DCIs) with different microstructures have been widely investigated and applied. The present work aims to illustrate the effect of different tempering temperatures (150–600°C) on the wear behaviour of martensitic ductile cast iron. The results indicate that the wear rate increases with tempering temperatures above 300°C. The lowest wear rate is found after tempering at 150–300°C, where the smearing of graphite forms a solid lubricating film and maybe an oxide layer on the wear surfaces during the sliding process. With further increases in the tempering temperature, the main wear mechanism changes from abrasive wear to adhesive wear. It shows that martensitic ductile cast iron can be used as a self-lubricated bearing material in the future.

Keywords

Ductile cast iron tempering temperature wear rate wear mechanism

Introduction

Ductile cast iron (DCI) has been used increasingly in applications involving tribological components such as gears, piston skirts, engine cylinders, crankshafts, and camshafts owing to its higher strength and toughness than carbon steel, high fatigue endurance, machinability, and relatively low cost. DCI is considered to be an appropriate substitute for steel widely used in various industries. The so-called ‘iron instead of steel’ mainly refers to DCI [1–7]. Bearings are the basic components in the field of equipment manufacturing, but bearing steel is expensive, with poor wear resistance and low fatigue strength. The main failure modes of bearings are wear and fatigue under the service condition [8]. Lubricating oil can be added to some bearings, but it is troublesome to add lubricating oil regularly. A good seal is required, and some lubricating oils will increase wear [9]. Because the graphite phase in DCI decreases the friction in sliding contacts against other solids, graphite can act as a lubricant which can reduce wear under certain conditions [10,11]. The microstructure of DCI is composed of matrix and nodular graphite. However, the shape, size, and distribution of nodular graphite cannot be changed by heat treatments due to its thermodynamic stability [12]. Therefore, the mechanical properties of DCI are mainly dependent on the matrix microstructure, and the matrix microstructure can be modified by heat treatments to obtain the properties required for specific applications.

In recent years, the wear resistance of DCI with different matrix microstructures has been studied by many researchers. Abedi [13] studied the effect of graphite nodule count on the sliding wear behaviour of ferritic-pearlite ductile iron under different loads. The study showed that under a lower applied load, the specimens with high nodule count exhibited a lower wear rate than those with low nodule count; at higher load, the wear resistance decreased with the increase of nodule count. Sugisghita [14] investigated the role of nodular graphite under rolling-sliding contacts with a ferrite-pearlite matrix microstructure, and found that a graphite film was formed due to the extrusion of graphite during rolling-sliding, played a certain lubricating effect and improved wear resistance. In addition, graphite can also be considered as a void in DCI, due to the severe plastic deformation of subsurface layers, the void formed by the connection of small graphite pull-out to surface, which is the preferential location for crack nucleation and propagation [15]. Hardness can affect the stress needed to deform during the sliding wear process, toughness is the amount of deformation a material can withstand before failure, so wear resistance is largely determined by the combination of hardness and toughness [16]. Apart from the hardness and nodular graphite, the surface characteristics, including the surface roughness and the type of matrix microstructure, play important roles in the wear resistance of DCI [17–19]. The wear behaviour of a material is also affected by the shape and size of the abrasive grit, roughness, attack angle, normal load applied, sliding speed, and fracture toughness of the material as well [20–22]. For the DCI with different matrix as shortly reviewed above, the predominant damage phenomena are adhesive wear and surface fatigue [23–26]. As mentioned above, the wear behaviours have been well studied on the ductile iron with a matrix of ausferrite, pearlite, and ferrite, and all these types of the matrix are relatively soft, compared with the martensitic matrix. However, the wear behaviour of ductile iron with martensitic matrix at different tempering temperatures has seldom been reported in the literature.

In this study, the DCI was processed by austenitising and quenching to obtain a multiphase microstructure including martensitic matrix, retained austenite, and nodular graphite, followed by tempering at different temperatures to change its microstructure, hardness, and wear behaviours. The effect of different tempering temperatures on the sliding wear behaviour of DCI tested by a block-on-ring testing apparatus will be analysed and discussed.

Experimental

Sample preparation

The material was obtained by melting molten iron (pig iron, scrap steel, and cast iron alloy) in an intermediate frequency induction furnace and casting it into a crystalliser with an inner diameter of 60 mm through vertical continuous casting. FeSi75 and FeSiMg were used as inoculants and nodulariser, respectively. The chemical composition of the ductile cast iron is shown in Table 1. In order to obtain the martensitic matrix with different hardness, the ductile iron was heat treated by varied heat treatment processes, as shown in Figure 1. These treatments include: (i) holding the material at an austenitising temperature of 840°C for 1 h to obtain an austenite matrix; (ii) quenching into an oil bath to get a martensitic matrix and retained austenite; (iii) tempering at 150–600°C for 2 h with steps of 50°C to obtain different matrix hardness. Samples for the tensile test were prepared according to ASTM E-8 with a dimension of ϕ5 mm × 65 mm [27]. Tensile tests were carried out at room temperature on a WE-300 experimental machine. The tensile properties, including tensile strength σ_b (MPa) and elongation δ (%), were calculated as the averages of three tests. The hardness test was carried out by using a Rockwell hardness tester with a load of 150 kg, and the average value of 5 points for each specimen was calculated. Table 2 is a list of abbreviations to be used in the following sections.

Table 1

Chemical composition of the samples (mass fraction, %).

Element	C	Si	Mn	S	P	Mg	RE
Fraction	3.51	2.8	0.32	0.0085	0.023	0.04	0.03–0.05

Table 2

List of abbreviations.

Ductile cast iron	DCI
Austempered ductile iron	ADI
DCI tempered at different temperatures	TDCI
DCI tempered at low temperature (150°C < Temp. ≤ 250°C)	LTDCI
DCI tempered at medium temperature (250°C < Temp. ≤ 450°C)	MTDCI
DCI tempered at high temperature (450°C < Temp. ≤ 600°C)	HTDCI

Wear test

Testing machine and samples

The wear experiment was performed on an OAT-U high temperature tribometer, and the schematic of the wear test equipment is shown in Figure 2. The sample was cut into a standard size of 30 mm × 17 mm × 10 mm, washed with alcohol, and dried. The prepared samples were demagnetised. In this experiment, the TC-1 demagnetiser was used, which is a small platform type demagnetiser. The operating voltage is 127 V, and the energisation duration is 60%. A U-shaped iron core is installed in the middle of the bottom of the stainless steel platform. A set of coils are wound around its iron core. When demagnetising, place the specimen in the middle of the line, move to the other side through the middle line, and then move away. This method belongs to the distance method. After the demagnetisation treatment, the samples were completely non-magnetic, did not attract other ferromagnetic substances, and were easy to clean. In addition, demagnetisation treatment can reduce the residual stress caused by machining. Wear test was conducted at room temperature, and in a dust-free environment without vibration and corrosive gas. The counterbody material (ring) was GB GCr15 (also called ISO 100 Cr6 or AISI 52100) bearing steel, with a hardness of 60 HRC.

The friction force was recorded as soon as the test machine turned on. All the tests were conducted under a contact load of 200 N, the final pressure was about 69 MPa, and a sliding speed of 0.323 m/s with a maximum sliding distance of 200 m was applied. According to the standard ASTM G77-2017, each specimen was tested three times to obtain the wear rate and friction coefficient [28]. When the sliding distance reached 200 m, testing was automatically stopped by the travel switch. The wear volume of the test block is calculated from the track length.

Wear rate and friction coefficient determination

The test block after the wear test was shown in Figure 3 [28]. The wear scar length b was measured at the middle and both ends of the wear scar observed for the block sample. A minimum of three tests were conducted, and the average of the three measurements was taken as the experimental result. According to the standard ASTM G77-2017, the wear volume was calculated via Equation (1), and Equation (2) was used to calculate the wear rate [28]: (1) (2)

where: V_k is the wear volume (mm³), w is the wear rate (mm³/m), D is the ring diameter (mm), which is 30 mm, b is the average length of wear mark (mm), t is the test block width (mm), which is 3 mm, L is the sliding distance 200 m. The friction coefficient (μ) varied with the material in contact with the ring and the relative velocity of movement. It can be recorded directly during the experiment without calculation.

Microstructural characterisation

After the wear test, Figure 4 shows the wear mark length of the sample after tempering at 250°C. The average of seven wear scar lengths at equal intervals was taken as the value of b. The microstructure and wear surface morphologies were characterised by optical microscope (OM, Zeiss 40MAT) and scanning electron microscopy (SEM, ZEISS Microscope EVO25), in combination with energy dispersive spectroscopy (EDS) to determine wear mechanisms. The volume fraction of retained austenite fractions (V _γ) was determined by X-ray diffraction (XRD) in an X-ray diffractometer of D8 ADVANCE with Co target, a tube current of 40 mA, a tube voltage of 35 kV and a Lynxeye XE detector [29]. A Rigaku rotating head anode diffractometer was used to scan the angular 2θ range from 45° to 115° using an angular speed of 1.2° per minute.

Results and discussion

Eﬀect of tempering temperature on microstructure

The as-cast microstructure of the ductile iron is the typical ‘bull's eye’ microstructure which consists of graphite nodules surrounded by ferrite (white) in the optical micrograph, as shown in Figure 5. The dark grey microstructure is pearlite. According to standard ASTM-A247-2016, the nodularity rate is 100% (with the highest nodularity rate) and the average diameter is about 10 μm, measured by Photoshop and Image-pro plus (IPP) software.

Figure 6 shows metallographic microstructures of ductile cast iron tempered at different temperatures. Due to the high carbon content of DCI, the microstructure of DCI can be regarded as composed of high carbon steel and nodular graphite, and the microstructure evolution of the matrix at different tempering temperatures is similar to the martensitic steel [30]. At low temperatures of 150–250°C, few carbides precipitate from carbon-supersaturated martensite, the final microstructure consists of tempered martensite and nodular graphite, as shown in Figure 6(a). With the increase of tempering temperatures to the medium range (300–450°C), more carbides precipitate from carbon-supersaturated martensite, and the final microstructure consists of tempered trooper and nodular graphite, as shown in Figure 6(b). When tempering at high temperatures (500–600°C), almost all carbides precipitate from carbon-supersaturated martensite, the final microstructure is carbides, ferrite, and nodular graphite, as shown in Figure 6(c). Figure 7(a) shows the XRD pattern after quenching and Figure 7(b) shows the XRD patterns at three tempering temperatures of 200, 400, and 600°C. The volume fractions of retained austenite (V _γ) at three tempering temperatures of 200, 400, and 600°C are shown in Figure 7(c). The V _γ after quenching is 11.3%. As the tempering temperature increases from 200°C to 600°C, the V _γ decreases from 6.75% to 0.34%, as shown in Figure 7(c). In Figure 7(b), the α (110) peak rises significantly from 200°C to 600°C, which is a result of the phase transformation of V _γ → martensite.

Eﬀect of tempering temperature on mechanical properties

Figure 8 shows the tensile strength, hardness, and elongation of ductile cast irons after tempering at different temperatures. Tensile strength values higher than 1600 MPa, 900–1600 MPa, and 400–900 MPa, corresponding to hardness higher than 53 HRC, 42–52 HRC, and 24–35HRC, are obtained LTDCI, MTDCI, and HTDCI, respectively. It can be seen that when the tempering temperature increase, the hardness, and strength of TDCI gradually decrease, but the elongation gradually increases. At the low tempering temperature, the strength is similar and the hardness decreases, but the hardness is still quite close to the test ring (60 HRC). At the medium tempering temperature, the hardness and strength both drastically decrease, and the hardness is much lower than the test ring. At the high tempering temperature, the strength and hardness decrease to the lowest values and the elongation reaches the highest.

Eﬀect of tempering temperature on wear resistance

The wear rates of ductile cast iron tempered at different temperatures are shown in Figure 9(a). The most important finding from this graph is that tempering temperatures between 150°C and 300°C have no significant effect on the wear rate. The TDCI after tempering at 150–300°C, the value of the wear rate is about 0.25 × 10⁻³–0.36 × 10⁻³ mm³/m. It is much smaller than the wear rate of GCr15 under the load of 10–50 N (0.49 × 10⁻³–0.61 × 10⁻³ mm³/m) and the load of 50–250 N (3.43 × 10⁻³–5.13 × 10⁻³ mm³/m), which may be related to the smearing of nodular graphite [31,32]. After tempering at 300–450°C, the wear rate of MTDCI increases gradually. The wear rate increases significantly for MTDCI tempered at a temperature higher than 450°C. The real-time friction coeﬃcient of TDCI determined by the wear test is shown in Figure 9(b). During the running-in period, μ _MTDCI > μ _HTDCI > μ _LTDCI, and μ increases rapidly to a high value at the beginning of the 100 s interval. After the running-in period, μ enters a stable period. This process may reveal wear mechanisms and the contact status during the wear test. In the steady state, μ _HTDCI > μ _MTDCI > μ _LTDCI, the fluctuation of μ after tempering at 400°C is the largest and smallest after tempering at 150°C. DCI is a heterogeneous material [33]. During the steady stage, the μ of all specimens shows fluctuation with time. This may be related to the detachment of graphite and surface unevenness during the wear process. This phenomenon will be explained in a subsequent section.

Dry sliding worn surface of TCDI

Worn surface of LTDCI

Figure 10(a,b) shows the worn surface of LTDCI after tempering at 200°C. For the TDCI after tempering at 200°C, the hardness measured at a distance of 3 μm from the surface increases from the initial 57 HRC to 58.8 HRC (672.3 HV) after wearing. The increase in hardness is due to the retained austenite transformation into martensite during the wear test, and this transformation only occurs at the surface layers. It is interesting that the worn surfaces are covered with deformation layers, and it is hard to observe nodular graphite. The worn surfaces in Figure 10 show some dark patches which may be a mixture of oxide layers and graphite films, and the number of dark patches is related to the distribution of matrix microstructure and nodular graphite, which results in delamination on the surface. The surface has only a small amount of material peeling off. The materials after tempering at low temperatures have the lowest wear rate, which exhibits high wear resistance, as shown in Figure 9(a).

Worn surface of MTDCI

Figure 11(a,b) shows the worn surface of TDCI after tempering at 300°C. The topographies of the surface are completely different from previous ones. It can be seen that the round black particles are graphite according to the EDS in Figure 11(c). As seen in Figure 11(a,b), some of the graphite nodules on the worn surface are covered by the matrix, whereas some of the graphite nodules are bare on the surface. The voids originate from the nodular graphite; while debris particles may be hard tempered martensite/carbide particles that coat or intermix with nearby graphite particles during the sliding process, exfoliating the surface material.

Figure 12 shows the worn surfaces of MTDCIs after tempering at 350°C and 400°C. Compared with the worn surface of the sample tempered at 300°C, the nodular graphite areas in these samples show many cracks. These cracks are observed to be connected to the nodular graphite. The reasons for nucleation and propagation of cracks include (i) increased surface deformation due to the decrease of strength; (ii) the debonding of graphite from the matrix during the surface deformation due to the weak bonding between graphite and matrix [34], and (iii) the propagation of cracks by tearing introduced by surface wear/sliding. This abrasive wear also leaves behind deep grooves on the surface from hardened abrasive debris, as shown in Figure 12(b). After the wear test, the morphology formed on the surface after the surface material is lost is called a pit, as shown in Figure 12(c,d). The TDCI after tempering at 400°C, the wear rate is greatly increased and the standard deviation for the three samples tempering at 400°C is the largest, as shown in Figure 9. This may be due to the deformation of the matrix increases and adhesion between the friction pair and the material surface forming pits where the surface is uneven (Figure 12(c)).

Figure 13 shows the worn surfaces of MTDCI after tempering at 450°C. Under low magnification, it can be seen that the worn surface is uneven, as shown in Figure 13(a). High magnification images reveal that many protrusions are present on the surface of the material (see Figure 13(b)), which causes an uneven surface. Furthermore, EDS analysis as shown in Figure 13(d) of point 2 (see Figure 13(c)) indicates that the convex part is nodular graphite wrapped by the matrix. After tempering at 450°C, the plasticity of the matrix increases, and these protrusions lead to surface unevenness. The surface of the material undergoes plastic deformation due to compressive stress and sliding friction, and the wear mechanism changes from abrasive wear to adhesive wear. Massive plastic yields at the onset of sliding and subsequent delamination appear to mask the lubricating effect of graphite on the surface, as shown in Figure 13(b). The relatively loud noise from the machine during the test is attributed to the unevenness of thesurface.

Worn surface of HTDCI

Figure 14 shows the worn surfaces of HTDCIs at different temperatures. Along the sliding direction, the shape of the groove in the middle is regular and relatively smooth and no nodular graphite is observed (Figure 14(b)). This can be attributed to the fact that the matrix's plasticity has increased, and the matrix has suffered severe deformation under adhesive wear. The name of this sub-mechanism is micro-ploughing. The material is pressed out of the grooves under the load and accumulates on both sides of the track forming the adhering layer. Figure 14(c,d) and Figure 14(e,f) show the worn surfaces of the samples tempered at 550°C and 600°C, respectively. As shown in Figure 9(a), the wear rate increases significantly at 550°C, resulting in severe wear. Due to the compressive stress, numerous deep grooves generate on the surface of the material and arc cracks appear in the middle of the grooves (see Figure 14(c)). Figure 14(d) reveals that many voids form due to the nodular graphite being pulled-out. The plasticity increases obviously after tempering at 600°C (Figure 14(e,f)) during the adhesive wear.

For the TDCI after high tempering temperature, the matrix hardness decreases and plasticity increases, and the wear rate is the highest. Several cracks are found to propagate along the direction of plastic deformation and surround the squeezed nodular graphite owing to smearing, tearing of the matrix, and the weak bonding between the matrix and nodular graphite [34,35]. As a result, the wear resistance of HTDCI is reduced, which results in more cavities, and large voids on the surface (see Figure 14(c–f)), and the wear resistance is lower than that of LTDCI and MTDCI.

SEM study of cross section of the worn samples

To reveal the wear failure mechanisms of TDCI, cross sections of TDCI after wear tests are observed, as shown in Figure 15. Due to the 200 N load being applied on the sample surface, the matrix underwent diﬀerent levels of deformation at different tempered temperatures, which caused nodular graphite to play various roles in the wear process. Sometimes the rough wear surface makes nodular graphite cannot be distinguished, but can be clearly seen in the sub-surface microstructure. The nodular graphite in the subsurface layer makes the matrix microstructure more rheological and is a potential cause of subsurface layer cracks. It exhibits low plasticity in the LTDCI matrix, a little deformation under the load, and a relatively smooth cross section as shown in Figure 15(a). As illustrated in Figure 15(b), nodular graphite can act as a lubricant, and the wear rate does not increase compared to the LTDCI, as shown in Figure 9(a).

The cross section of the sample after tempering at 400°C was shown in Figure 15(c). There are many pits on the surface, and the cross section is uneven. The cross section of the sample after tempering at 450°C was shown in Figure 15(d). In accordance with the behaviour described in Figure 13, some of the graphite nodules on the worn surface are covered by the matrix, and some of the graphite nodules are pulled-out on the surface. For MTDCI, the thickness of the deformation layer is around 6 μm.

It can be seen from Figure 15(e) (the tempering temperature is 600°C) that both the nodular graphite and the matrix deformed after wear. The undulation of the boundary reflects the pit deformed on the wear surface. The matrix on the upper part of the nodular graphite on the left is about to be lost, and the nodular graphite has been severely deformed and shed off. This indicates that materials with low strength and high plasticity cannot provide strong support for nodular graphite. For HTDCI, the thickness of the deformation layer increases up to around 18 μm.

Normally speaking, what needs to be focused on with respect to the above should be the aspects concerning the formation and stability/adherence of the mechanically mixed transfer layer as well a solid lubricating film and oxide layer on to the wear surfaces during the sliding process that would actually control the overall wear response of the samples with different matrix hardness, ductility and strength levels. In this context, too soft (matrix) surfaces would make the transfer/oxide/lubricating layers/films unstable thereby allowing fresh metal-to-metal contact to be established during the wear process and thus leading to high material adhesion and inferior wear response. On the other hand, too hard a surface would experience severe spalling leading once again to the non-formation and/or premature removal of the mechanically mixed transfer layer and hence poor wear behaviour. Accordingly, an optimum level of material (matrix) mechanical properties (hardness, strength, and ductility) would lead to good wear characteristics. However, in the present case, the spalling of graphite of the LTDCI forms a solid lubricating film and maybe plus an oxide layer on the wear surfaces during the sliding process, which guarantees the low wear rate of the LTDCI.

The failure of materials generally initiates from the surface. Hence, improving the wear resistance of these materials via surface heat treatment is helpful to reduce such failures. However, the wear resistance increases non-monotonically with an increasing amount of nodular graphite, it is also related to the matrix deformation and hardness. In future work, a detailed microstructural investigation, the microstructure-strength/hardness relationship [36],the heterogeneous stress and strain distribution [37,38], ductility [39], and systematic wear mechanisms will be carried out, as well as the effect of different amount of nodular graphite on the wear resistance subjected to a given set of heat treatments and the possible advanced mechanical testing in microscopes [40].

Conclusions

In this work, the eﬀects of tempering temperature on the microstructure, mechanical properties, and wear resistance of TDCI were investigated. The conclusions are summarised as follows:

With an increase in tempering temperature, the V _γ, tensile strength, and hardness of the TDCI gradually decreases, and the tensile elongation increases.

The wear mechanisms of the samples after tempering at different temperatures are different, and the fundamental reasons are the decrease in matrix strength and hardness and its different roles in the block-on-ring sliding process.

For the LTDCI (150–300°C), as the retained austenite transformed mechanically intomartensite during the wear test, the hardness is similar to the counterbody. The spalling of graphite of the LTDCI forms a solid lubricating film and maybe plus an oxide layer on to the wear surfaces during the sliding process, which guarantees a similar low wear rate of the LTDCI. The wear mechanism is delamination with the smeared graphite as a lubricant.

For the MTDCI after tempering at 300–450°C, the matrix hardness decrease to the zone of 42–52 HRC, many pits and protrusions form on the surface, and the wear rate increases with the matrix hardness/strength decrease, and the main wear mechanism gradually transforms from abrasive wear to adhesive wear.

For the HTDCI after tempering at 500–600°C, the wear rate increases sharply with the matrix hardness/strength decrease, and the main wear mechanism is the adhesive wear. And the main failure mechanisms are smearing, tearing, and pitting.

In the block-on-ring wear test, the best wear resistance is found in the LTDCI after tempering during 150–300°C and lower than the GCr15 steel under the load of 200 N. It shows that ductile iron may be used as a self-lubricated bearing material in the future.

Footnotes

Disclosure statement

No potential conflict of interest was reported by the author(s).

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