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Pulsed laser synthesis of ceramic–metal composite coating on steel
Applied Surface Science 255 (2008) 3188–3194
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Pulsed laser synthesis of ceramic–metal composite coating on steel Baoshuai Du a,b, Anoop N. Samant a, Sameer R. Paital a, Narendra B. Dahotre a,* a b
Department of Materials Science and Engineering, The University of Tennessee, Knoxville, TN 37996, USA School of Materials Science and Engineering, Shandong University, Jinan 250061, China
A R T I C L E I N F O
A B S T R A C T
Article history: Received 9 July 2008 Received in revised form 9 August 2008 Accepted 4 September 2008 Available online 12 September 2008
A pulsed Nd:YAG laser was employed to modify the surface properties of AISI 1010 steel with precursor of TiB2 + Al. A set of samples were prepared with different laser processing parameters and compositions of the precursor in order to study the effect of Al on the coating. Thermal modeling was performed to quantitatively evaluate the maximum temperature and the range of cooling rate for the melting pool. Phase constituents and microstructure were characterized using X-ray diffractometer, optical microscopy, and scanning electron microscopy. Results show that TiB2 dissociated when the Al content reached 30 wt.% or more. The composite coating with the presence of TiB2 shows acicular TiB2 particles embedded in the a-Fe matrix. Coatings produced using precursor of high-Al content reveals a refined cellular structure due to the high-cooling rate induced by short pulse duration. Compared with the steel substrate, microhardness and wear resistance of the coating are improved significantly. ß 2008 Elsevier B.V. All rights reserved.
Keywords: Laser surface engineering TiB2 + Al Wear
1. Introduction Recent years have seen the development of laser surface engineering (LSE) due to its unique features and capabilities in various commercial applications [1–7]. LSE is a versatile coating technique which by employing laser as the heat source modifies a confined area of component. This unique feature leads to localized melting and solidification within a shallow depth and hence makes it possible to modify the surface layer without affecting the bulk of the as-received sample [1,2]. In addition, the high-energy density of laser beam can lead to the melting of partial substrate, resulting in a metallurgical bond between the coating and substrate. Thus, compared with other conventional coating technique, such as thermal spray, electro deposition and hardfacing, LSE provides a coating with higher bonding strength but low-dilution ratio. Although continuous laser has been used extensively in the field of LSE, there is a lack of study about employing pulsed laser to produce coatings. Due to the short pulse width coupled with high-peak power, pulsed laser can not only produce a sufficient melting depth but also cause a cooling rate of the melting pool that is substantially higher than that of continuous laser. Thus it is expected that coating with refined microstructure can be fabricated.
* Corresponding author at: Department of Materials Science and Engineering, The University of Tennessee, 1512 Middle Drive, 326 Dougherty Eng. Building, Knoxville, TN 37996, USA. Tel.: +1 865 974 3609; fax: +1 865 974 4451. E-mail address: [email protected] (N.B. Dahotre). 0169-4332/$ – see front matter ß 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.apsusc.2008.09.010
The most common applications of LSE are to improve resistance to wear and corrosion. For example, using Fe–Cr–Ti– C as precursor, mixed carbide (TiC, M7C3) coatings have been formed with improved wear resistance [8]. Matthews et al. [9] fabricated amorphous coatings on titanium alloy substrate and found that the coating shows high hardness and low-friction coefficients. However, in many cases for components served under severe condition it is the combined effect of wear and corrosion that is damaging and as a result this complexity makes the field of surface engineering much challenging [10]. TiB2 is characterized by a high-melting point, low-specific weight, high hardness, high strength to density ratio, good wear resistance and excellent thermal and chemical stability up to 1700 8C [11,12]. In the past continuous Nd:YAG laser has been used to produce titanium diboride reinforced composite coating on steel [13]. It is shown that such coating is composite in nature (consisting TiB2 particles and a-Fe matrix) and possesses good wear resistance [14]. However, although the reinforcement of TiB2 exhibit high hardness and chemical stability, the a-Fe matrix may become a weak point for the damage to propagate if the coating serves under severely corrosive and high-temperature condition. Thus it is necessary to further modify the coating by judicial selection of the coating materials as well as the processing parameters. Literature shows that Al can be incorporated with iron to produce iron–aluminide alloys with superior high-temperature oxidation and sulfidation resistance, and good wear resistance [15,16]. Besides, the incorporation of Al as alloying element has also been successfully pursued in fabrication of creep-resistant and Al2O3-forming austenitic stainless steel which can be used under
B. Du et al. / Applied Surface Science 255 (2008) 3188–3194
the aggressive oxidizing conditions encountered in energyconversion system [17]. Recently, the fabrication of iron–aluminide coating on less corrosion resistance materials has been investigated. The process of weld overlay was employed to fabricate coating with an aluminum concentration of 15– 17 at.% (8–9 wt.%) for depth of 2 mm and it is found that the aluminum-enriched surface shows an improved oxidation and sulfidation resistance [18]. Corbin et al. produced an iron– aluminide coating on a mild steel substrate using pulsed laser assisted powder deposition (LAPD) [16]. The aim of this paper is to develop a wear, oxidation and corrosion resistant composite coating consisting of TiB2 and iron– aluminide on steel substrate by pulsed Nd:YAG laser using mixtures of Al and TiB2 as precursor. Processing parameters as well as the composition of the precursor were correlated with the phase composition and microstructure of the coating. Microhardness and wear resistance tests were conducted to provide a preliminary evaluation of the coating properties. The corrosion and high-temperature oxidation behaviors of the coating, as well as the scratch test which can be used to measure the adhesion of the coating are being evaluated and will be presented in a separate article.
the precursor and the laser processing parameters used in the current investigation can be found from the list of Table 1. In order to get the cross-sectional view, the coated samples were sectioned perpendicular to the laser track using low-speed diamond saw. Samples for metallurgical observation were processed according to the standard metallographic procedure and etched with nital. Structural characterization was performed on a Philips Norelco Xray Diffractometer (XRD) with Cu Ka radiation (wavelength 1.5418 A˚) operating at 20 kV and 10 mA. The 2u angle ranges from 208 to 1008. A LEO 1525 scanning electron microscope (SEM) was used to study the microstructure of the coatings. Surface roughness measurements of the coated samples were carried out using a Mahr Federal profilometer with a tip scan distance of 5.6 mm. Microhardness measurements were performed on a microhardness tester with a normal load of 300 g applied for 12 s. Evaluation of the wear resistance was carried out using a pin on disk wear tester with a linear speed of 0.0471 m/s and a normal load of 21.7 N. The pin used in the experiment was WC with a diameter of 3 mm. Weight loss was recorded at regular intervals. The wear rate was calculated using the formula of Wear rate ðmg=ðmin cm2 ÞÞ ¼
2. Experimental procedures AISI 1010 steel plates with dimensions of 75 mm 75 mm 6 mm were used as substrate. The steel plates were polished by emery paper followed by washing with acetone to provide a clean surface. Coating material consists of commercial TiB2 powder (99.5% purity, 325 mesh) and Al powder (99.5% purity, 325 mesh) with different weight ratios. The powder mixture was mixed with a proprietary organic binder and then was spray deposited on the steel substrates. The sprayed coupons were dried at 80 8C for about 5 min to remove the moisture. A preplaced precursor about 150 mm thick was obtained for all the coupons. A JK 701 model pulsed Nd:YAG laser with average power of 400 W (pulse duration: 0.5–2 ms; repetition rate: 20 Hz) was used to carry out the laser surface engineering process. Schematic of the laser surface engineering process is shown in Fig. 1. The pulsed laser is operated in the TEM00 mode to offer a laser beam with Gaussion-shaped heat source. The laser beam was delivered by a fiber system then focused using a 120 mm focal length convex lens which provides a spot diameter of 240 mm at focus. During laser processing, the laser beam was defocused on the material to give a beam spot with diameter of approximately 500 mm. Several laser tracks were performed by overlapping each other with an overlapping ratio of 70% to cover a large area. The composition of
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total weight loss ðmgÞ wear time ðminÞ wear track area ðcm2 Þ
(1)
3. Results and discussion 3.1. Thermal modeling The precursor (aluminum and titanium diboride mixed in varying compositions) as well as the substrate material (steel) was modeled for the thermal predictions that were necessary to correlate the observed microstructural features with the processing parameters and material composition. A geometry in the form of a cube representing the precursor (75 mm 75 mm 150 mm) was coupled with another cube representing the substrate (75 mm 75 mm 6 mm). The energy input to the system and the time for which the energy was incident on the material surface were essential for the computational predictions as they affected the temperature profiles and the ensuing cooling rates. The pulsed laser had a beam with a circular cross-section (diameter = 0.05 cm) and as there was a power off period between two successive pulses, the repetition rate and pulse width were taken into account to determine the residence time for which the laser beam interacts with the surface. The residence time was estimated using the following relation [19]: Residence time ðsÞ ¼
repetition rate ðHzÞ pulse width ðsÞ beam diameter ðcmÞ traverse speed ðcm=sÞ (2)
The residence time predicted for the different cases was input to the model along with the power density per unit time given by
Power density ðW=cm2 Þ ¼
Fig. 1. Schematic of experimental setup.
incident energy ðJÞ pulse width ðsÞ area of incident beam ðcm2 Þ
(3)
The residence time and power density for different processing conditions are presented in Table 1. The heat transfer during laser processing of coating on the steel substrate was modeled using Fourier’s second law of heat transfer
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Table 1 List of sample compositions, processing parameters and roughness Precursor composition (wt.%)
Pulse width (ms)
Repetition rate (Hz)
Scan speed (mm/s)
Pulse energy (J)
Residence time (ms)
Power density (1010 W/m2)
Surface roughness, Ra (mm)
50Al–50TiB2
0.5 1 1.5 2 1.5
20
84.7
4 6 8 10 8
0.059 0.118 0.177 0.236 0.177
4.1 3.1 2.7 2.5 2.7
5.1 3.6 3.4 3.0 10.4 9.6 6.1 3.5
5Al–95TiB2 10Al–90TiB2 20Al–80TiB2 30Al–70TiB2
Table 2 Thermal properties of TiB2 and Al [12,20] Temperature (K)
300 525 725 925
Al
TiB2 Specific feat (J/(kg K))
Thermal conductivity (W/(m K))
Density (103 kg/m3)
Specific heat (J/kg)
Thermal conductivity (W/(m K))
Density (103 kg/m3)
638 972 1059 1110
95 84.8 81 79
4.5
863 1006 1102 1276
225 203 188 185
2.71
Eq. (4) in COMSOL’sTM heat transfer transient mode: " 2 # 2 2 @Tðx; y; z; tÞ kðTÞ @ Tðx; tÞ @ Tðy; tÞ @ Tðz; tÞ ¼ þ þ @t rC p ðTÞ @x2 @y2 @z2
(4)
where k(T) and Cp(T) are the variations in thermal conductivity and specific heat as a function of temperature, r the density of the precursor, T the temperature field, t the time and x, y and z are the spatial directions (Fig. 1). Thermal properties of TiB2 and Al at different temperatures retrieved from references are shown in Table 2. Thermal properties of the relatively compacted precursor were calculated using law of mixture taking the weight percent of each component as the variable. For improved accuracy of calculations, variation of thermal conductivity and specific heat as a function of temperature for the coating and the substrate were considered in the model [12,20]. The latent heat of solidification was accounted for by considering the variation of specific heat as a function of temperature. At time t = 0, the initial temperature of T = T0 = 300 K was applied. The balance between the absorbed laser energy at the surface and the losses due to radiation was given by @Tðx; y; 0; tÞ @Tðx; y; 0; tÞ @Tðx; y; 0; tÞ þ þ kðTÞ @x @y @z ¼ daI es ðTðx; y; 0; tÞ4 T 0 4 Þ; tp ;
d ¼ 0 when t > tp
d ¼ 1 when 0 t (5)
where I is the laser power density predicted above in Eq. (3), k(T) the temperature dependent thermal conductivity of the material (W/(m K)), e the emissivity for thermal radiation 0.7 [20], tp the time at which pulsing is stopped, s the Stefan–Boltzman constant (5.67 108 W/(m2 K4)) and a is the absorptivity of the material. It has been shown that the reflection of a mixture of opaque particles is a function of the relative surface area [21]. Thus, the weight percent composition of the precursor was converted to surface area percent in order to evaluate the absorbtivity of the precursor. Conducting in situ absorptivity measurements in short duration high-energy density dynamic process such as laser–material interaction being extremely difficult, values of absorptivity of individual precursor components (TiB2 and Al) commonly found in literature were used for the calculations [20,21]. The term d takes a
value of 1 when the time, t is less than the pulse on time, tp and it is 0 when the time, t exceeds the pulse time (off-time). Thus the value of d depends on the time, t and ensures that the energy is input to the system only till the pulse is on and cuts off the energy supply after that. The convection at the bottom surface of the sample was given by @Tðx; y; L; tÞ @Tðx; y; L; tÞ @Tðx; y; L; tÞ þ þ kðTÞ @x @y @z ¼ hðTðx; y; L; tÞ T 0 Þ
(6)
where L is the thickness of the sample which was taken as 150 mm for the coating and 6 mm for the substrate and h is the heat transfer coefficient (W/(m2 K)) which was included as a function of temperature [22]. The maximum surface temperatures reached and the corresponding range of cooling rates (slope of the cooling curve at different time instants) were obtained from these computations and are presented in Table 3 for different compositions and laser input energies considered in the study. Although a high value of maximum temperature is found in Table 3, it should be noted that the maximum temperature is the instantaneous temperature at the top surface of the sample. The temperature drops as a function of depth due to the high-cooling rate. This maximum surface temperature is used as an indication of trend in temperature change as a function of processing and coating materials parameters. Besides, it can be seen Table 3 Cooling rates and maximum surface temperature for combinations of composition and input energy Composition (wt.%)
Power density (1010 W/m2)
Cooling rate range (K/s)
Maximum surface temperature (K)
50Al–50TiB2
4.1 3.1 2.7 2.5 2.7
4.2 104 1.1 105 1.9 105 1.5 105 1.2 106 7.6 105 7.1 105 2.4 105
3276 2897 2742 2635 5330 4926 4676 3841
5Al–95TiB2 10Al–90TiB2 20Al–80TiB2 30Al–70TiB2
to to to to to to to to
6.1 107 2.08 107 1.63 107 1.5 107 7.8 107 6.72 107 6.36 107 1.98 107
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Moreover, it can be found that with the increase of TiB2 content in the precursor, the maximum temperature also increases. This is due to the high absorbtivity (0.77) and relatively low-thermal conductivity of TiB2 compared with those of Al (absorbtivity: 0.17) (Table 2) [20,21]. For a given composition of precursor, a reasonable change of the cooling rate was observed for different processing parameters (power density in the present case). 3.2. Coating morphology and microstructure
Fig. 2. Macrograph of coating (5Al–95TiB2, 2.7 1010 W/m2): (a) cross-sectional view and (b) top view.
that very high cooling rate can be reached which is a characteristic of laser processing. Compared with the cooling rate attained using continuous laser (in the order of 103 K/s for sample processed with power density of 0.5 109 to 1.2 109 W/m2 for Fe-based alloy coating [8]), it is apparent that significantly higher cooling rate can be reached using this pulsed laser. During the laser processing, only a confined area is irritated and heated by the laser beam defocused at the surface (500 mm). The rest of cold substrate acts as a massive heat sink and thereby induces the self-quenching effect [10,13], resulting in high-cooling rate of the molten pool. Besides, the short pulse duration (0.5–2 ms, 20 Hz) means very short heating period that also contributes to the high-cooling rate. The cooling rate is also further enhanced by the cover gas (air flown at the rate of 20 l/min) employed in the present case. It is evident that for precursor with the same composition, maximum temperature and cooling rate decreases with the increasing of pulse energy. Although higher pulse energy leads to the increasing of residence time, it also causes a substantial decrease of the power density. Thus, the combination of these two factors results in lower maximum temperature and cooling rate for samples processed with higher pulse energy.
A top view and a cross-sectional view of the sample processed using power density of 2.7 1010 W/m2 and precursor of 50Al– 50TiB2 are shown in Fig. 2. A smooth and uniform coating over relatively large area and depth is obtained. Surface roughness values of the coatings are presented in Table 1. The surface roughness decreases with the decreasing of power density. It is well documented that surface roughness of the laser treated coating is influenced by the wetting behavior of coating materials to the substrate, liquid inside the melt pool and overlapping ratio of the laser tracks [16,23]. As can be found from Table 3, higher power density results in higher melt pool temperature, which may lead to severe convective flow in the melting pool (whirls of materials flow caused by convective flow was evidenced by Fig. 3(a)). Thus with the decreasing of power density, less rough surface was obtained due to the less disturbance in the melt pool. Besides, it can also be found that with the increase of Al content in the precursor, surface of the coating becomes smoother. Furthermore, with the increasing of Al content, a larger melt volume can be generated since Al possesses a much lower melting temperature (933 K [20]). This in turn leads to better spreading of the melt liquid that after solidification produces a less rough surface. It should be mentioned that precursor with pure TiB2 was also tried to be deposited on the steel substrate but it was found that part of the coating peeled off thus preventing the formation of a homogeneous and dense coating. This fact also shows that by including Al in the precursor the wetting behavior between TiB2 and iron can be substantially improved. From the cross-sectional view (Fig. 3(a)) it is evident that the sample can be characterized into three parts: composite coating, heat affected zone (HAZ) and substrate. During the pulsed laser
Fig. 3. SEM micrograhs of the composite coating (50Al–50TiB2, 2.7 1010 W/m2): (a) cross-sectional view with low magnification, (b) interface and (c) coating area with high magnification.
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processing the severity of thermal excursions experienced by materials varied from region to region, resulting in these three different areas. Fig. 3(b) shows the interface between the coating and HAZ. It clearly indicates a sound metallurgical joint to the steel substrate was obtained using pulsed laser. Besides, the interface is characterized by a planar growth band. At the interface, the thermal gradient G is very high while the growth rate of crystal is low, resulting in a high value of G/R. According to the general solidification theory, this leads to the planar growth. Across this region, dendrites were found growing into the coating. Fig. 3(c) shows the high-magnification micrograph of the coating area. The microstructure reveals refined cellular structure grains of the size typically in the range of 1–2 mm. This kind of microstructure is associated with a very rapid cooling rate, as is indicated in Table 3. During the pulsed laser processing, laser beam scans the sample with a very high rate, resulting in a very short interaction time. The pulsed laser energy can only generate a melting pool with confined volume while the bulk of the substrate plays the role of heat sink [10,13]. As a result, very fine grains were generated in the coating as is illustrated in Fig. 3(c). The micrographs of coating with retained TiB2 shows distinctly different microstructure (Fig. 4). This coating was processed using pulse energy density of 2.7 1010 W/m2 with 5Al–95TiB2 as precursor. A cross-sectional view shows the sound metallurgical joint of coating to the substrate and a refined HAZ (Fig. 4(a)). Acicular TiB2 particles with a high-aspect ratio which is different from the polygonal TiB2 particles in the precursor were found embedded in the a-Fe matrix (Fig. 4(b)) and also identified by the XRD results (Fig. 6). In addition, from Table 3 it can be found that the temperature of the molten pool is well above the melting point of TiB2 (2790–3225 8C [11,24,25]). Thus it follows that TiB2 is formed by the melting–dissolution–reprecipitation mechanism, i.e. TiB2 melts and dissolves into the molten liquid and during the cooling process TiB2 nucleates and grows into the final reinforcement. This acicular morphology of TiB2 crystal is presumably due to the preferential growth of TiB2 particle along the c axis ([0 0 0 1] direction) during the rapid cooling process of laser surface engineering. Fig. 4(c) reveals different particle morphology along the remelting area caused by the overlapping of laser tracks. It
shows acicular TiB2 particles tend to transform to small blocky morphology. In contrast, the typical acicular and/or blocky shaped TiB2 particles were not observed for all 50Al–50TiB2 samples processed using different laser parameters, indicating the dissociation of this phase. This is in accordance with the XRD results as will be discussed later. 3.3. Phase analysis XRD results are shown in Figs. 5 and 6. The overlay of XRD patterns of precursor (50Al–50TiB2) processed with different parameters is presented in Fig. 5. For the precursor, phase constituents are Al and TiB2, as should be expected. However, after laser processing, phases evolved in the coatings are a-Fe, Fe3Al, Al2O3, B2O3, Ti2O, Ti3AlC and AlB12. TiB2 were not detected within the resolution of XRD. It illustrates that with 50 wt.% Al in the precursor, a substantial amount of TiB2 was dissociated and/or reacted with other element during the laser processing, indicating high-Al content (50 wt.%, 62.5 vol.%) in the precursor influences the presence of TiB2 in the composite coating for different thermal conditions that prevailed under the laser processing parameters employed in the present work. The presence of oxides in the coating indicates that the oxygen was included in the melting pool. Oxygen can come from the environment during the laser processing and react with the alloying element to form oxides. Al, Ti and B are elements that have high affinity with oxygen, thus the following reactions may take place leading to the consumption of TiB2: 2Al þ 3=2O2 ¼ Al2 O3
(7)
2B þ 3=2O2 ¼ B2 O3
(8)
2Ti þ 1=2O2 ¼ Ti2 O
(9)
Besides, the presence of compounds of Ti3AlC and AlB12 indicates the reaction of TiB2 with Al and C. The interaction of laser beam with the material leads to the formation of a melting pool which includes not only the precursor but also a portion of the steel substrate. Thus C elements can present in the composite
Fig. 4. SEM micrographs of the coating (5Al–95TiB2, 2.7 1010 W/m2): (a) a cross-sectional view of the coating, (b) the composite coating morphology, and (c) morphology change of TiB2 in overlap area.
B. Du et al. / Applied Surface Science 255 (2008) 3188–3194
Fig. 5. XRD patterns of the coating processed with precursor of 50Al–50TiB2.
coating. The overall reactions for the formation of Ti3AlC and AlB12 can be concluded as TiB2 þ Al þ C ! Ti3 AlC þ AlB12
(10)
XRD patterns of coatings which were fabricated with different composition of precursor but same processing parameters (power density of 2.7 1010 W/m2) are illustrated in Fig. 6. Peaks of TiB2 can be identified for these samples although the peak intensity of sample processed with precursor containing 30 wt.% Al is very weak. Other compounds such as TiAl, Ti3AlC, AlB12 and Fe3B were also found in the composite coating. The relative proportions of TiB2 in the composite coating were semiquantitatively determined using the following equation [26]: Ii ð%Þ ¼
Ii I1 þ I2
(11)
where Ii is the phase in concern, I1 corresponds to the integrated intensity of the TiB2 (1 0 0) peak and I2 corresponds to the integrated intensity of the a-Fe (2 1 1) peak. These peaks were chosen to avoid overlapping or closely adjacent lines from different
Fig. 6. XRD patterns of the coating with different composition processed with energy density of 2.7 1010 W/m2.
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phases. In this calculation the composite coating were assumed to be composed of TiB2 and a-Fe and other phases were neglected. The calculated results show that the relative proportions of TiB2 are 24.1%, 16.02%, 9.16% and 5.5% for precursors containing 5 wt.% Al, 10 wt.% Al, 20 wt.% Al and 30 wt.% Al, respectively. Thus it is evident that the TiB2 content in the coating decreases with the increasing of Al content in the precursor. Taken these results together, it can be found that above the value of 30 wt.% Al concentration in the precursor the TiB2 tends to dissociate and/or react with other elements resulting in the absence of it in the composite coating. In order to retain TiB2 in the coating, low-Al concentration (up to 20 wt.%) should be used. Taken the dilution effect of the steel substrate into consideration, the melting pool can be regarded as a Fe–Al–Ti–B alloy system. Doucakis and Kumar [27] investigated the feasibility of forming refractory diboride particles in a Fe3Al matrix by conventional casting techniques and found that Fe–Al–Ti–B alloy yields a two-phase microstructure of Fe3Al and TiB2. Recently, Krein et al. [28] studied the Fe–Al–Ti–B alloy containing about 30 at.% Al and detected the presence of TiB2 precipitates even though the content of B and Ti is at a very low level (1 at.% B and 0.5 at.% Ti). This result seems in contrast to the current observation. However, temperature during laser processing under the present set of parameters (Table 3) has been nearly equal to or higher than the decomposition/melting temperature of TiB2 (2790–3225 8C) [11,24,25]. In addition, unlike processing techniques adopted in above referenced works [25,26], laser surface engineering being a non-equilibrium synthesis process, it may have caused instantaneous variations of Al activity and other local thermodynamic conditions (thermal gradient, cooling rate, etc.) within the melt to the extent leading to the reaction of TiB2 with Al and C, as well as the consumption of it by oxygen as explained earlier for the certain set of laser and materials processing parameters explored in the present work. 3.4. Microhardness and wear resistance Microhardness profiles of the samples were taken on the crosssectional plane perpendicular to the laser track. Relatively uniform microhardness distributions can be found throughout the coating (Fig. 7). Compared with the steel substrate, microhardness of the coatings is improved significantly. The maximum microhardness exists on the coating processed with precursor of 5Al–95TiB2 which possesses the highest content of TiB2. Besides, the samples with TiB2 retained in the coating generally show a higher hardness value. However, other samples reveal no apparent difference in hardness. As TiB2 possesses the highest hardness (HK0.1 2600 (at 25 8C) [11]) among the reinforcements, it is reasonable that a high-volume fraction of TiB2 in the coating leads to a high microhardness. The wear rate result of the coating is shown in Fig. 8. It is obvious that due to the presence of hard reinforcements, such as TiB2 and iron–aluminide, the laser-coated samples demonstrate considerably better wear resistance than the steel substrate. The hard reinforcements (TiB2, Fe3Al, Al2O3) can play the role of hard barrier to interrupt the ploughing and scratching and therefore improve the wear resistance of the coating. The sample processed using precursor of 20Al–80TiB2 with laser power density of 2.7 1010 W/m2 which possesses optimum combination of surface roughness and hardness among the current set of samples shows the lowest wear rate. It is well established that wear resistance of such composite coating is very complex, depending on the size and volume fraction of reinforcements, as well as the mechanical properties of the matrix [14]. This sample possesses retained TiB2 particles and a relatively soft matrix because of the alloying effect of a proper amount of Al. Thus, this synergetic effect
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4. Conclusions
Fig. 7. Microhardness of the composite coatings.
AISI 1010 steel substrate with preplaced powder mixtures of TiB2 + Al with Al content ranging from 5 to 50 wt.% was processed using pulsed Nd:YAG laser. Thermal modeling results show that the cooling rate of the melting pool ranges from 104 to 107 K/s. With the amount of Al in the precursor lower than 30 wt.%, TiB2 could retain in the coating. However, TiB2 started to dissociate and/ or react with other elements with higher Al content (>30 wt.%) in the precursor resulting in a coating consisted of a-Fe, iron– aluminide, oxides, titanium aluminide and aluminum boride. Microsturcture characterization reveals that while TiB2 retained in the composite coating shows acicular and small blocky morphology, coatings without the presence of TiB2 show a refined cellular structure. Compared with the steel substrate, the microhardness and wear resistance of the coating are improved significantly. It is shown that coating with an optimum combination of hardness (Hv 900) and surface roughness (Ra 6.1) possess minimum wear rate (0.113 mg/(min cm2)). References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18]
Fig. 8. Wear rate of the composite coating.
[19] [20] [21] [22]
leads to the maximum wear resistance of this coating. Although the sample with the most amount of retained TiB2 (5Al–95TiB2, 2.7 1010 W/m2) possesses the highest microhardness value, it does not show the maximum wear resistance due to the poor ductility of the coating and the rough surface which increases the contacting stress between asperities during wear process.
[23] [24] [25] [26] [27] [28]
A. Singh, N. Dahotre, J. Mater. Sci. 39 (2004) 553. I. Goldfarb, M. Bamberger, Scr. Mater. 34 (1996) 1051. A. Agarwal, N. Dahotre, Int. J. Refract. Met. Hard Mater. 17 (1999) 283. T. Miokovic, V. Schulze, O. Vohringer, D. Lohe, Acta Mater. 55 (2007) 589. J. Majumdar, B. Chandra, A. Nath, I. Manna, Mater. Sci. Eng. A 433 (2006) 241. I. Pashby, S. Barnes, B. Bryden, J. Mater. Process. Technol. 139 (2003) 585. T. Yue, T. Li, Surf. Coat. Technol. 202 (2008) 3043. A. Singh, N. Dahotre, Metall. Mater. Trans. A 36 (2004) 797. D. Matthews, V. Ocelik, J. Hosson, Mater. Sci. Eng. A 471 (2007) 155. J. DeHosson, in: N.B. Dahotre, T.S. Sudarshan (Eds.), Intermetallic and Ceramic Coatings, Marcel Dekker Inc., New York, 1999, pp. 307, 311. D. Vallauri, I.C. Atias Adria´n, A. Chrysanthou, J. Eur. Ceram. Soc. 28 (2008) 1697. R. Munro, J. Res. Natl. Inst. Standard Technol. 105 (2000) 705. A. Agarwal, N. Dahotre, Mater. Charact. 42 (1999) 31. A. Agarwal, N. Dahotre, Wear 240 (2000) 144. S. David, J. Horton, C. Mckamey, T. Zacharia, R. Reed, Weld. J. 68 (1989) 372. E. Toyserkani, A. Khajepour, S. Corbin, Laser Cladding, CRC, Florida, 2005, p. 180. Y. Yamamoto, M. Brady, Z. Lu, P. Maziasz, Science 316 (2007) 433. Y. Bhambri, V. Sikka, L. Walker, M. Santella, G. Muralidharan, J.W. Hales, Mater. Sci. Eng. A 394 (2005) 249. A. Samant, S. Harimkar, N. Dahotre, J. Appl. Phys. 102 (2007) 123105. Y. Touloukian, Thermophysical Properties of High Temperature Materials, IFI/ Plenum, New York, 1967, pp. 11, 13. E. Schatz, J. Opt. Soc. Am. A 57 (1967) 941. F. Incropera, D. Dewitt, Fundamentals of Heat and Mass Transfer, John Wiley, New York, 2002, p. 543. H. Li, S. Costil, H. Liao, C. Coddet, J. Therm. Spray Technol. 15 (2006) 191. B. Bhushan, B. Gupta, Handbook of Tribology Materials, Coatings, and Surface Treatments, McCraw Hill, New York, 1991, p. 4.55. J.L. Murry, P.K. Liao, K.E. Spear, Bull. Alloy Phase Diagrams 7 (1986) 550. Y. Greish, P. Brown, J. Biomed. Mater. Res. 67 (2003) 633. T. Doucakis, K. Kumar, Intermetallics 7 (1999) 765. R. Krein, A. Schneider, G. Sauthoff, G. Frommeyer, Intermetallics 15 (2007) 1172.
Contents lists available at ScienceDirect
Applied Surface Science journal homepage: www.elsevier.com/locate/apsusc
Pulsed laser synthesis of ceramic–metal composite coating on steel Baoshuai Du a,b, Anoop N. Samant a, Sameer R. Paital a, Narendra B. Dahotre a,* a b
Department of Materials Science and Engineering, The University of Tennessee, Knoxville, TN 37996, USA School of Materials Science and Engineering, Shandong University, Jinan 250061, China
A R T I C L E I N F O
A B S T R A C T
Article history: Received 9 July 2008 Received in revised form 9 August 2008 Accepted 4 September 2008 Available online 12 September 2008
A pulsed Nd:YAG laser was employed to modify the surface properties of AISI 1010 steel with precursor of TiB2 + Al. A set of samples were prepared with different laser processing parameters and compositions of the precursor in order to study the effect of Al on the coating. Thermal modeling was performed to quantitatively evaluate the maximum temperature and the range of cooling rate for the melting pool. Phase constituents and microstructure were characterized using X-ray diffractometer, optical microscopy, and scanning electron microscopy. Results show that TiB2 dissociated when the Al content reached 30 wt.% or more. The composite coating with the presence of TiB2 shows acicular TiB2 particles embedded in the a-Fe matrix. Coatings produced using precursor of high-Al content reveals a refined cellular structure due to the high-cooling rate induced by short pulse duration. Compared with the steel substrate, microhardness and wear resistance of the coating are improved significantly. ß 2008 Elsevier B.V. All rights reserved.
Keywords: Laser surface engineering TiB2 + Al Wear
1. Introduction Recent years have seen the development of laser surface engineering (LSE) due to its unique features and capabilities in various commercial applications [1–7]. LSE is a versatile coating technique which by employing laser as the heat source modifies a confined area of component. This unique feature leads to localized melting and solidification within a shallow depth and hence makes it possible to modify the surface layer without affecting the bulk of the as-received sample [1,2]. In addition, the high-energy density of laser beam can lead to the melting of partial substrate, resulting in a metallurgical bond between the coating and substrate. Thus, compared with other conventional coating technique, such as thermal spray, electro deposition and hardfacing, LSE provides a coating with higher bonding strength but low-dilution ratio. Although continuous laser has been used extensively in the field of LSE, there is a lack of study about employing pulsed laser to produce coatings. Due to the short pulse width coupled with high-peak power, pulsed laser can not only produce a sufficient melting depth but also cause a cooling rate of the melting pool that is substantially higher than that of continuous laser. Thus it is expected that coating with refined microstructure can be fabricated.
* Corresponding author at: Department of Materials Science and Engineering, The University of Tennessee, 1512 Middle Drive, 326 Dougherty Eng. Building, Knoxville, TN 37996, USA. Tel.: +1 865 974 3609; fax: +1 865 974 4451. E-mail address: [email protected] (N.B. Dahotre). 0169-4332/$ – see front matter ß 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.apsusc.2008.09.010
The most common applications of LSE are to improve resistance to wear and corrosion. For example, using Fe–Cr–Ti– C as precursor, mixed carbide (TiC, M7C3) coatings have been formed with improved wear resistance [8]. Matthews et al. [9] fabricated amorphous coatings on titanium alloy substrate and found that the coating shows high hardness and low-friction coefficients. However, in many cases for components served under severe condition it is the combined effect of wear and corrosion that is damaging and as a result this complexity makes the field of surface engineering much challenging [10]. TiB2 is characterized by a high-melting point, low-specific weight, high hardness, high strength to density ratio, good wear resistance and excellent thermal and chemical stability up to 1700 8C [11,12]. In the past continuous Nd:YAG laser has been used to produce titanium diboride reinforced composite coating on steel [13]. It is shown that such coating is composite in nature (consisting TiB2 particles and a-Fe matrix) and possesses good wear resistance [14]. However, although the reinforcement of TiB2 exhibit high hardness and chemical stability, the a-Fe matrix may become a weak point for the damage to propagate if the coating serves under severely corrosive and high-temperature condition. Thus it is necessary to further modify the coating by judicial selection of the coating materials as well as the processing parameters. Literature shows that Al can be incorporated with iron to produce iron–aluminide alloys with superior high-temperature oxidation and sulfidation resistance, and good wear resistance [15,16]. Besides, the incorporation of Al as alloying element has also been successfully pursued in fabrication of creep-resistant and Al2O3-forming austenitic stainless steel which can be used under
B. Du et al. / Applied Surface Science 255 (2008) 3188–3194
the aggressive oxidizing conditions encountered in energyconversion system [17]. Recently, the fabrication of iron–aluminide coating on less corrosion resistance materials has been investigated. The process of weld overlay was employed to fabricate coating with an aluminum concentration of 15– 17 at.% (8–9 wt.%) for depth of 2 mm and it is found that the aluminum-enriched surface shows an improved oxidation and sulfidation resistance [18]. Corbin et al. produced an iron– aluminide coating on a mild steel substrate using pulsed laser assisted powder deposition (LAPD) [16]. The aim of this paper is to develop a wear, oxidation and corrosion resistant composite coating consisting of TiB2 and iron– aluminide on steel substrate by pulsed Nd:YAG laser using mixtures of Al and TiB2 as precursor. Processing parameters as well as the composition of the precursor were correlated with the phase composition and microstructure of the coating. Microhardness and wear resistance tests were conducted to provide a preliminary evaluation of the coating properties. The corrosion and high-temperature oxidation behaviors of the coating, as well as the scratch test which can be used to measure the adhesion of the coating are being evaluated and will be presented in a separate article.
the precursor and the laser processing parameters used in the current investigation can be found from the list of Table 1. In order to get the cross-sectional view, the coated samples were sectioned perpendicular to the laser track using low-speed diamond saw. Samples for metallurgical observation were processed according to the standard metallographic procedure and etched with nital. Structural characterization was performed on a Philips Norelco Xray Diffractometer (XRD) with Cu Ka radiation (wavelength 1.5418 A˚) operating at 20 kV and 10 mA. The 2u angle ranges from 208 to 1008. A LEO 1525 scanning electron microscope (SEM) was used to study the microstructure of the coatings. Surface roughness measurements of the coated samples were carried out using a Mahr Federal profilometer with a tip scan distance of 5.6 mm. Microhardness measurements were performed on a microhardness tester with a normal load of 300 g applied for 12 s. Evaluation of the wear resistance was carried out using a pin on disk wear tester with a linear speed of 0.0471 m/s and a normal load of 21.7 N. The pin used in the experiment was WC with a diameter of 3 mm. Weight loss was recorded at regular intervals. The wear rate was calculated using the formula of Wear rate ðmg=ðmin cm2 ÞÞ ¼
2. Experimental procedures AISI 1010 steel plates with dimensions of 75 mm 75 mm 6 mm were used as substrate. The steel plates were polished by emery paper followed by washing with acetone to provide a clean surface. Coating material consists of commercial TiB2 powder (99.5% purity, 325 mesh) and Al powder (99.5% purity, 325 mesh) with different weight ratios. The powder mixture was mixed with a proprietary organic binder and then was spray deposited on the steel substrates. The sprayed coupons were dried at 80 8C for about 5 min to remove the moisture. A preplaced precursor about 150 mm thick was obtained for all the coupons. A JK 701 model pulsed Nd:YAG laser with average power of 400 W (pulse duration: 0.5–2 ms; repetition rate: 20 Hz) was used to carry out the laser surface engineering process. Schematic of the laser surface engineering process is shown in Fig. 1. The pulsed laser is operated in the TEM00 mode to offer a laser beam with Gaussion-shaped heat source. The laser beam was delivered by a fiber system then focused using a 120 mm focal length convex lens which provides a spot diameter of 240 mm at focus. During laser processing, the laser beam was defocused on the material to give a beam spot with diameter of approximately 500 mm. Several laser tracks were performed by overlapping each other with an overlapping ratio of 70% to cover a large area. The composition of
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total weight loss ðmgÞ wear time ðminÞ wear track area ðcm2 Þ
(1)
3. Results and discussion 3.1. Thermal modeling The precursor (aluminum and titanium diboride mixed in varying compositions) as well as the substrate material (steel) was modeled for the thermal predictions that were necessary to correlate the observed microstructural features with the processing parameters and material composition. A geometry in the form of a cube representing the precursor (75 mm 75 mm 150 mm) was coupled with another cube representing the substrate (75 mm 75 mm 6 mm). The energy input to the system and the time for which the energy was incident on the material surface were essential for the computational predictions as they affected the temperature profiles and the ensuing cooling rates. The pulsed laser had a beam with a circular cross-section (diameter = 0.05 cm) and as there was a power off period between two successive pulses, the repetition rate and pulse width were taken into account to determine the residence time for which the laser beam interacts with the surface. The residence time was estimated using the following relation [19]: Residence time ðsÞ ¼
repetition rate ðHzÞ pulse width ðsÞ beam diameter ðcmÞ traverse speed ðcm=sÞ (2)
The residence time predicted for the different cases was input to the model along with the power density per unit time given by
Power density ðW=cm2 Þ ¼
Fig. 1. Schematic of experimental setup.
incident energy ðJÞ pulse width ðsÞ area of incident beam ðcm2 Þ
(3)
The residence time and power density for different processing conditions are presented in Table 1. The heat transfer during laser processing of coating on the steel substrate was modeled using Fourier’s second law of heat transfer
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Table 1 List of sample compositions, processing parameters and roughness Precursor composition (wt.%)
Pulse width (ms)
Repetition rate (Hz)
Scan speed (mm/s)
Pulse energy (J)
Residence time (ms)
Power density (1010 W/m2)
Surface roughness, Ra (mm)
50Al–50TiB2
0.5 1 1.5 2 1.5
20
84.7
4 6 8 10 8
0.059 0.118 0.177 0.236 0.177
4.1 3.1 2.7 2.5 2.7
5.1 3.6 3.4 3.0 10.4 9.6 6.1 3.5
5Al–95TiB2 10Al–90TiB2 20Al–80TiB2 30Al–70TiB2
Table 2 Thermal properties of TiB2 and Al [12,20] Temperature (K)
300 525 725 925
Al
TiB2 Specific feat (J/(kg K))
Thermal conductivity (W/(m K))
Density (103 kg/m3)
Specific heat (J/kg)
Thermal conductivity (W/(m K))
Density (103 kg/m3)
638 972 1059 1110
95 84.8 81 79
4.5
863 1006 1102 1276
225 203 188 185
2.71
Eq. (4) in COMSOL’sTM heat transfer transient mode: " 2 # 2 2 @Tðx; y; z; tÞ kðTÞ @ Tðx; tÞ @ Tðy; tÞ @ Tðz; tÞ ¼ þ þ @t rC p ðTÞ @x2 @y2 @z2
(4)
where k(T) and Cp(T) are the variations in thermal conductivity and specific heat as a function of temperature, r the density of the precursor, T the temperature field, t the time and x, y and z are the spatial directions (Fig. 1). Thermal properties of TiB2 and Al at different temperatures retrieved from references are shown in Table 2. Thermal properties of the relatively compacted precursor were calculated using law of mixture taking the weight percent of each component as the variable. For improved accuracy of calculations, variation of thermal conductivity and specific heat as a function of temperature for the coating and the substrate were considered in the model [12,20]. The latent heat of solidification was accounted for by considering the variation of specific heat as a function of temperature. At time t = 0, the initial temperature of T = T0 = 300 K was applied. The balance between the absorbed laser energy at the surface and the losses due to radiation was given by @Tðx; y; 0; tÞ @Tðx; y; 0; tÞ @Tðx; y; 0; tÞ þ þ kðTÞ @x @y @z ¼ daI es ðTðx; y; 0; tÞ4 T 0 4 Þ; tp ;
d ¼ 0 when t > tp
d ¼ 1 when 0 t (5)
where I is the laser power density predicted above in Eq. (3), k(T) the temperature dependent thermal conductivity of the material (W/(m K)), e the emissivity for thermal radiation 0.7 [20], tp the time at which pulsing is stopped, s the Stefan–Boltzman constant (5.67 108 W/(m2 K4)) and a is the absorptivity of the material. It has been shown that the reflection of a mixture of opaque particles is a function of the relative surface area [21]. Thus, the weight percent composition of the precursor was converted to surface area percent in order to evaluate the absorbtivity of the precursor. Conducting in situ absorptivity measurements in short duration high-energy density dynamic process such as laser–material interaction being extremely difficult, values of absorptivity of individual precursor components (TiB2 and Al) commonly found in literature were used for the calculations [20,21]. The term d takes a
value of 1 when the time, t is less than the pulse on time, tp and it is 0 when the time, t exceeds the pulse time (off-time). Thus the value of d depends on the time, t and ensures that the energy is input to the system only till the pulse is on and cuts off the energy supply after that. The convection at the bottom surface of the sample was given by @Tðx; y; L; tÞ @Tðx; y; L; tÞ @Tðx; y; L; tÞ þ þ kðTÞ @x @y @z ¼ hðTðx; y; L; tÞ T 0 Þ
(6)
where L is the thickness of the sample which was taken as 150 mm for the coating and 6 mm for the substrate and h is the heat transfer coefficient (W/(m2 K)) which was included as a function of temperature [22]. The maximum surface temperatures reached and the corresponding range of cooling rates (slope of the cooling curve at different time instants) were obtained from these computations and are presented in Table 3 for different compositions and laser input energies considered in the study. Although a high value of maximum temperature is found in Table 3, it should be noted that the maximum temperature is the instantaneous temperature at the top surface of the sample. The temperature drops as a function of depth due to the high-cooling rate. This maximum surface temperature is used as an indication of trend in temperature change as a function of processing and coating materials parameters. Besides, it can be seen Table 3 Cooling rates and maximum surface temperature for combinations of composition and input energy Composition (wt.%)
Power density (1010 W/m2)
Cooling rate range (K/s)
Maximum surface temperature (K)
50Al–50TiB2
4.1 3.1 2.7 2.5 2.7
4.2 104 1.1 105 1.9 105 1.5 105 1.2 106 7.6 105 7.1 105 2.4 105
3276 2897 2742 2635 5330 4926 4676 3841
5Al–95TiB2 10Al–90TiB2 20Al–80TiB2 30Al–70TiB2
to to to to to to to to
6.1 107 2.08 107 1.63 107 1.5 107 7.8 107 6.72 107 6.36 107 1.98 107
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Moreover, it can be found that with the increase of TiB2 content in the precursor, the maximum temperature also increases. This is due to the high absorbtivity (0.77) and relatively low-thermal conductivity of TiB2 compared with those of Al (absorbtivity: 0.17) (Table 2) [20,21]. For a given composition of precursor, a reasonable change of the cooling rate was observed for different processing parameters (power density in the present case). 3.2. Coating morphology and microstructure
Fig. 2. Macrograph of coating (5Al–95TiB2, 2.7 1010 W/m2): (a) cross-sectional view and (b) top view.
that very high cooling rate can be reached which is a characteristic of laser processing. Compared with the cooling rate attained using continuous laser (in the order of 103 K/s for sample processed with power density of 0.5 109 to 1.2 109 W/m2 for Fe-based alloy coating [8]), it is apparent that significantly higher cooling rate can be reached using this pulsed laser. During the laser processing, only a confined area is irritated and heated by the laser beam defocused at the surface (500 mm). The rest of cold substrate acts as a massive heat sink and thereby induces the self-quenching effect [10,13], resulting in high-cooling rate of the molten pool. Besides, the short pulse duration (0.5–2 ms, 20 Hz) means very short heating period that also contributes to the high-cooling rate. The cooling rate is also further enhanced by the cover gas (air flown at the rate of 20 l/min) employed in the present case. It is evident that for precursor with the same composition, maximum temperature and cooling rate decreases with the increasing of pulse energy. Although higher pulse energy leads to the increasing of residence time, it also causes a substantial decrease of the power density. Thus, the combination of these two factors results in lower maximum temperature and cooling rate for samples processed with higher pulse energy.
A top view and a cross-sectional view of the sample processed using power density of 2.7 1010 W/m2 and precursor of 50Al– 50TiB2 are shown in Fig. 2. A smooth and uniform coating over relatively large area and depth is obtained. Surface roughness values of the coatings are presented in Table 1. The surface roughness decreases with the decreasing of power density. It is well documented that surface roughness of the laser treated coating is influenced by the wetting behavior of coating materials to the substrate, liquid inside the melt pool and overlapping ratio of the laser tracks [16,23]. As can be found from Table 3, higher power density results in higher melt pool temperature, which may lead to severe convective flow in the melting pool (whirls of materials flow caused by convective flow was evidenced by Fig. 3(a)). Thus with the decreasing of power density, less rough surface was obtained due to the less disturbance in the melt pool. Besides, it can also be found that with the increase of Al content in the precursor, surface of the coating becomes smoother. Furthermore, with the increasing of Al content, a larger melt volume can be generated since Al possesses a much lower melting temperature (933 K [20]). This in turn leads to better spreading of the melt liquid that after solidification produces a less rough surface. It should be mentioned that precursor with pure TiB2 was also tried to be deposited on the steel substrate but it was found that part of the coating peeled off thus preventing the formation of a homogeneous and dense coating. This fact also shows that by including Al in the precursor the wetting behavior between TiB2 and iron can be substantially improved. From the cross-sectional view (Fig. 3(a)) it is evident that the sample can be characterized into three parts: composite coating, heat affected zone (HAZ) and substrate. During the pulsed laser
Fig. 3. SEM micrograhs of the composite coating (50Al–50TiB2, 2.7 1010 W/m2): (a) cross-sectional view with low magnification, (b) interface and (c) coating area with high magnification.
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processing the severity of thermal excursions experienced by materials varied from region to region, resulting in these three different areas. Fig. 3(b) shows the interface between the coating and HAZ. It clearly indicates a sound metallurgical joint to the steel substrate was obtained using pulsed laser. Besides, the interface is characterized by a planar growth band. At the interface, the thermal gradient G is very high while the growth rate of crystal is low, resulting in a high value of G/R. According to the general solidification theory, this leads to the planar growth. Across this region, dendrites were found growing into the coating. Fig. 3(c) shows the high-magnification micrograph of the coating area. The microstructure reveals refined cellular structure grains of the size typically in the range of 1–2 mm. This kind of microstructure is associated with a very rapid cooling rate, as is indicated in Table 3. During the pulsed laser processing, laser beam scans the sample with a very high rate, resulting in a very short interaction time. The pulsed laser energy can only generate a melting pool with confined volume while the bulk of the substrate plays the role of heat sink [10,13]. As a result, very fine grains were generated in the coating as is illustrated in Fig. 3(c). The micrographs of coating with retained TiB2 shows distinctly different microstructure (Fig. 4). This coating was processed using pulse energy density of 2.7 1010 W/m2 with 5Al–95TiB2 as precursor. A cross-sectional view shows the sound metallurgical joint of coating to the substrate and a refined HAZ (Fig. 4(a)). Acicular TiB2 particles with a high-aspect ratio which is different from the polygonal TiB2 particles in the precursor were found embedded in the a-Fe matrix (Fig. 4(b)) and also identified by the XRD results (Fig. 6). In addition, from Table 3 it can be found that the temperature of the molten pool is well above the melting point of TiB2 (2790–3225 8C [11,24,25]). Thus it follows that TiB2 is formed by the melting–dissolution–reprecipitation mechanism, i.e. TiB2 melts and dissolves into the molten liquid and during the cooling process TiB2 nucleates and grows into the final reinforcement. This acicular morphology of TiB2 crystal is presumably due to the preferential growth of TiB2 particle along the c axis ([0 0 0 1] direction) during the rapid cooling process of laser surface engineering. Fig. 4(c) reveals different particle morphology along the remelting area caused by the overlapping of laser tracks. It
shows acicular TiB2 particles tend to transform to small blocky morphology. In contrast, the typical acicular and/or blocky shaped TiB2 particles were not observed for all 50Al–50TiB2 samples processed using different laser parameters, indicating the dissociation of this phase. This is in accordance with the XRD results as will be discussed later. 3.3. Phase analysis XRD results are shown in Figs. 5 and 6. The overlay of XRD patterns of precursor (50Al–50TiB2) processed with different parameters is presented in Fig. 5. For the precursor, phase constituents are Al and TiB2, as should be expected. However, after laser processing, phases evolved in the coatings are a-Fe, Fe3Al, Al2O3, B2O3, Ti2O, Ti3AlC and AlB12. TiB2 were not detected within the resolution of XRD. It illustrates that with 50 wt.% Al in the precursor, a substantial amount of TiB2 was dissociated and/or reacted with other element during the laser processing, indicating high-Al content (50 wt.%, 62.5 vol.%) in the precursor influences the presence of TiB2 in the composite coating for different thermal conditions that prevailed under the laser processing parameters employed in the present work. The presence of oxides in the coating indicates that the oxygen was included in the melting pool. Oxygen can come from the environment during the laser processing and react with the alloying element to form oxides. Al, Ti and B are elements that have high affinity with oxygen, thus the following reactions may take place leading to the consumption of TiB2: 2Al þ 3=2O2 ¼ Al2 O3
(7)
2B þ 3=2O2 ¼ B2 O3
(8)
2Ti þ 1=2O2 ¼ Ti2 O
(9)
Besides, the presence of compounds of Ti3AlC and AlB12 indicates the reaction of TiB2 with Al and C. The interaction of laser beam with the material leads to the formation of a melting pool which includes not only the precursor but also a portion of the steel substrate. Thus C elements can present in the composite
Fig. 4. SEM micrographs of the coating (5Al–95TiB2, 2.7 1010 W/m2): (a) a cross-sectional view of the coating, (b) the composite coating morphology, and (c) morphology change of TiB2 in overlap area.
B. Du et al. / Applied Surface Science 255 (2008) 3188–3194
Fig. 5. XRD patterns of the coating processed with precursor of 50Al–50TiB2.
coating. The overall reactions for the formation of Ti3AlC and AlB12 can be concluded as TiB2 þ Al þ C ! Ti3 AlC þ AlB12
(10)
XRD patterns of coatings which were fabricated with different composition of precursor but same processing parameters (power density of 2.7 1010 W/m2) are illustrated in Fig. 6. Peaks of TiB2 can be identified for these samples although the peak intensity of sample processed with precursor containing 30 wt.% Al is very weak. Other compounds such as TiAl, Ti3AlC, AlB12 and Fe3B were also found in the composite coating. The relative proportions of TiB2 in the composite coating were semiquantitatively determined using the following equation [26]: Ii ð%Þ ¼
Ii I1 þ I2
(11)
where Ii is the phase in concern, I1 corresponds to the integrated intensity of the TiB2 (1 0 0) peak and I2 corresponds to the integrated intensity of the a-Fe (2 1 1) peak. These peaks were chosen to avoid overlapping or closely adjacent lines from different
Fig. 6. XRD patterns of the coating with different composition processed with energy density of 2.7 1010 W/m2.
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phases. In this calculation the composite coating were assumed to be composed of TiB2 and a-Fe and other phases were neglected. The calculated results show that the relative proportions of TiB2 are 24.1%, 16.02%, 9.16% and 5.5% for precursors containing 5 wt.% Al, 10 wt.% Al, 20 wt.% Al and 30 wt.% Al, respectively. Thus it is evident that the TiB2 content in the coating decreases with the increasing of Al content in the precursor. Taken these results together, it can be found that above the value of 30 wt.% Al concentration in the precursor the TiB2 tends to dissociate and/or react with other elements resulting in the absence of it in the composite coating. In order to retain TiB2 in the coating, low-Al concentration (up to 20 wt.%) should be used. Taken the dilution effect of the steel substrate into consideration, the melting pool can be regarded as a Fe–Al–Ti–B alloy system. Doucakis and Kumar [27] investigated the feasibility of forming refractory diboride particles in a Fe3Al matrix by conventional casting techniques and found that Fe–Al–Ti–B alloy yields a two-phase microstructure of Fe3Al and TiB2. Recently, Krein et al. [28] studied the Fe–Al–Ti–B alloy containing about 30 at.% Al and detected the presence of TiB2 precipitates even though the content of B and Ti is at a very low level (1 at.% B and 0.5 at.% Ti). This result seems in contrast to the current observation. However, temperature during laser processing under the present set of parameters (Table 3) has been nearly equal to or higher than the decomposition/melting temperature of TiB2 (2790–3225 8C) [11,24,25]. In addition, unlike processing techniques adopted in above referenced works [25,26], laser surface engineering being a non-equilibrium synthesis process, it may have caused instantaneous variations of Al activity and other local thermodynamic conditions (thermal gradient, cooling rate, etc.) within the melt to the extent leading to the reaction of TiB2 with Al and C, as well as the consumption of it by oxygen as explained earlier for the certain set of laser and materials processing parameters explored in the present work. 3.4. Microhardness and wear resistance Microhardness profiles of the samples were taken on the crosssectional plane perpendicular to the laser track. Relatively uniform microhardness distributions can be found throughout the coating (Fig. 7). Compared with the steel substrate, microhardness of the coatings is improved significantly. The maximum microhardness exists on the coating processed with precursor of 5Al–95TiB2 which possesses the highest content of TiB2. Besides, the samples with TiB2 retained in the coating generally show a higher hardness value. However, other samples reveal no apparent difference in hardness. As TiB2 possesses the highest hardness (HK0.1 2600 (at 25 8C) [11]) among the reinforcements, it is reasonable that a high-volume fraction of TiB2 in the coating leads to a high microhardness. The wear rate result of the coating is shown in Fig. 8. It is obvious that due to the presence of hard reinforcements, such as TiB2 and iron–aluminide, the laser-coated samples demonstrate considerably better wear resistance than the steel substrate. The hard reinforcements (TiB2, Fe3Al, Al2O3) can play the role of hard barrier to interrupt the ploughing and scratching and therefore improve the wear resistance of the coating. The sample processed using precursor of 20Al–80TiB2 with laser power density of 2.7 1010 W/m2 which possesses optimum combination of surface roughness and hardness among the current set of samples shows the lowest wear rate. It is well established that wear resistance of such composite coating is very complex, depending on the size and volume fraction of reinforcements, as well as the mechanical properties of the matrix [14]. This sample possesses retained TiB2 particles and a relatively soft matrix because of the alloying effect of a proper amount of Al. Thus, this synergetic effect
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4. Conclusions
Fig. 7. Microhardness of the composite coatings.
AISI 1010 steel substrate with preplaced powder mixtures of TiB2 + Al with Al content ranging from 5 to 50 wt.% was processed using pulsed Nd:YAG laser. Thermal modeling results show that the cooling rate of the melting pool ranges from 104 to 107 K/s. With the amount of Al in the precursor lower than 30 wt.%, TiB2 could retain in the coating. However, TiB2 started to dissociate and/ or react with other elements with higher Al content (>30 wt.%) in the precursor resulting in a coating consisted of a-Fe, iron– aluminide, oxides, titanium aluminide and aluminum boride. Microsturcture characterization reveals that while TiB2 retained in the composite coating shows acicular and small blocky morphology, coatings without the presence of TiB2 show a refined cellular structure. Compared with the steel substrate, the microhardness and wear resistance of the coating are improved significantly. It is shown that coating with an optimum combination of hardness (Hv 900) and surface roughness (Ra 6.1) possess minimum wear rate (0.113 mg/(min cm2)). References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18]
Fig. 8. Wear rate of the composite coating.
[19] [20] [21] [22]
leads to the maximum wear resistance of this coating. Although the sample with the most amount of retained TiB2 (5Al–95TiB2, 2.7 1010 W/m2) possesses the highest microhardness value, it does not show the maximum wear resistance due to the poor ductility of the coating and the rough surface which increases the contacting stress between asperities during wear process.
[23] [24] [25] [26] [27] [28]
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