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Experiment and numerical simulation study on laser truing and dressing of bronze-bonded diamond wheel
ARTICLE IN PRESS Optics and Lasers in Engineering 48 (2010) 295–304
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Optics and Lasers in Engineering journal homepage: www.elsevier.com/locate/optlaseng
Experiment and numerical simulation study on laser truing and dressing of bronze-bonded diamond wheel Genyu Chen a, Lifang Mei a,n, Bi Zhang b, Chunrong Yu a, Kangjian Shun a a b
State Key Laboratory of Advanced Design and Manufacturing for Vehicle Body, College of Mechanical and Automobile Engineering, Hunan University, Changsha 410082, China University of Connecticut, Storrs, CT, USA
a r t i c l e in fo
abstract
Article history: Received 8 July 2009 Received in revised form 2 November 2009 Accepted 10 November 2009 Available online 16 December 2009
An acoustic-optic Q-switched YAG laser is newly developed for truing and dressing of bronze-bonded diamond grinding wheels through ablation in the radial direction. The laser ablation merges truing and dressing as a single process and can be easily controlled through an online closed-loop control system. The laser truing and dressing characteristics are investigated by wheel surface topography and a highspeed grinding test. A two-dimensional mathematical model is developed to simulate laser truing and dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by a pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. Both the experimental results and the theoretical analysis indicate that the truing and dressing processes can be simultaneously realized with the laser ablation. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Laser truing and dressing Bronze-bonded diamond grinding wheel Topography Grinding performance Numerical simulation
1. Introduction Diamond superabrasive grinding wheels are widely used in grinding hard brittle non-ferrous metals, horniness alloys and hard brittle non-metals like ceramics, optical glasses and gems, because of their high hardness, thermal resistance and long tool life [1]. Grinding with superabrasive wheels offers a series of advantages such as high grinding efficiency, fine surface quality and finedimensional stability. Due to their superior grinding performances, they are extensively applied to such fields as aerospace, automotive, electronics, medicine and so on [2]. In order to achieve efficient or fine grinding, a specific surface topography on the wheel face should be generated and therefore periodic truing and dressing are required [3]. It is, however, difficult to sharpen these superabrasive wheels with a high degree of accuracy, due to the qualities mentioned above, which could inevitably affect their applications. Therefore, the truing and dressing of superabrasive wheels have been a research topic. In general, the superabrasive grinding wheels can be dressed by means of conventional mechanical dressing methods on the interaction of force, such as diamond rolling methods [4,5] and diamond stylus shaping methods [6,7]. But these truing and dressing methods, especially in the case of diamond wheels, are time-consuming and cause pollution of the working
n
Corresponding author. E-mail addresses: [email protected] (G. Chen), [email protected] (L. Mei). 0143-8166/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.optlaseng.2009.11.006
environment. In addition, mechanical damage of the abrasives is induced by the direct contact between the wheel and the truer and dresser [8]. Truing and dressing tools are easy subjected to wear and their geometry will be altered. It influences the truing and dressing conditions, which in turn affects the wheel wear and the grinding process. In order to solve these problems and achieve a good truing and dressing quality, a series of studies have conducted recently, which include improving and developing new methods, such as stick-aided loose abrasive (SLAD) dressing method [9], Green Carborundum (GC) cup dressing method[10], water jet dressing method [11], and electrical discharge dressing and truing method [12], etc. although these studies provide insight into proper truing and dressing methods and optimum ablation parameters, the rapid wear of truer and dresser and unsatisfactory topography of wheels remain unresolved. Thus, it is necessary to find an appropriate method for truing and dressing the superabrasive wheels. As an alternative to the conventional dressing method, laser truing and dressing is proposed as a new technique with many merits, such as extensive adaptability, high efficiency, non-pollution to environment, non-contact and high geometric accuracies. Laser truing and dressing is a method which uses laser to true and dress grinding wheels. The grinding wheels need to be dressed after being used for a period, even before putting into use. Truing means to remove part bond around the abrasives, so the abrasives can protrude bond. This process engenders enough space for swarf extraction and forms cutting edge. Dressing refers to micromachining on workface of grinding wheels until attaining required
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geometry accuracy. A number of studies have been conducted recently, most of which are focused on dressing of grinding wheels and a few experiments have investigated the truing diamond wheels, such as Jackson et al. [13] made a study of laser dressing of vitrified aluminum oxide grinding wheels. Kang et al. [14,15] conducted a study of tangentially truing metal-bonded diamond wheel by conventional pulsed Nd:YAG laser. Westk¨amper et al. [16] employed the conventional pulsed Nd:YAG laser to dress resinbonded CBN wheel and proposed a method of dressing on wheels in orthogonal direction, while truing in tangential direction. Zuo Dunwen et al. [17] made a study of diamond plate dressed by CW CO2 laser. Moreover, a technique of CO2 laser-assisted truing and dressing for vitrified CBN wheel developed by Zhang and Shin [18]. However, systematic study of truing and dressing on bronzebonded diamond superabrasive wheels by acousto-optic Qswitched pulsed Nd:YAG laser as a single process has not been reported for the moment. In this paper, an acoustic-optic Q-switched Nd:YAG laser is used to ablate a bronze-bonded diamond grinding wheel in the radial direction to realize in-process truing and dressing. The mechanism of selective removal and surface topography of laser truing and dressing wheel are exhaustively analyzed. The laser truing and dressing characteristics of the wheel are also examined and comparison made to a conventionally dressed grinding wheel. The grinding performance of the laser trued and dressed wheel was evaluated in terms of grinding force and workpiece surface roughness. A two-dimensional mathematical model is developed to simulate laser truing and dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by an acousto-optic Q-switched YAG-pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. The laser dressing process is characterized by a small ablation depth and thus very limited heat-affected zone (HAZ) in each pulse. It is considered as an ideal tool for truing and dressing of the superabrasive wheels.
2. Laser truing and dressing experiment
continuous output power of the Nd:Yag laser is 100 W, which is modified by current of krypton discharge source. There is an acousto-optic Q-switched device built in the resonant cavity whose frequency ranges from 0.2 to 50 kHz and therefore a pulse laser with pulse width (t0) from 170 to 560 ns and pulse repetition frequency (f) from 0.2 to 50 kHz can be secured. The output laser with a wavelength of 1.06 mm is irradiated radially on the wheel surface. The focus lens is aberrationless hyperbolical with its focal length 60 mm. A continuous laser beam is sent out from the Nd:YAG laser for wheel surface measurements, and a pulsed laser beam sent from the same YAG laser by acoustic-optic Q-switched driver for grinding wheel ablation. The two laser beams are actually from the same laser source. The Q-switch can automatically shut-off the pulsed laser beam if the grinding wheel surface level is too low, resulting in no material removal from the wheel surface. If the wheel surface level is too high, the Q-switch is turned on to irradiate a strong pulsed laser beam on the grinding wheel surface. Fig. 2 shows a photographic view of the experimental setup, while Table 1 lists laser ablation parameters. Experimental installation of single-pulse ablation on superabrasive wheels is secured by coordination the feed motion and rotation speed of lathe machine. The focus spot diameter is measured to be 0.26 mm and the beam divergent angle is 20 mrad. The output power of continuous laser is measured by power device (model: PC-1000, made in Canada), the surface topography after laser truing and dressing on wheels is measured by light microscope (LM) (model: Leica DMLM, made in England) and scanning electronic microscope (SEM) (model: STEREOSCAN440, made in England), the protruding height of abrasives is measured by surface roughness measuring instrument and Form Talysurf (model: Roughness Talysurf Series 2), the roundness error of wheel is measured by micrometer checker and Taly roundness measuring equipment (model: Taly265). A bronzebonded diamond grinding wheel (denoted as MBD100/120 M100) is chosen as the truing and dressing target. Table 2 provides the physical properties of the bronze bond and diamond abrasives. The dressing characteristics are examined by wheel topography and a grinding test.
2.1. Laser truing and dressing system
2.2. Laser ablation parameters
Fig. 1 shows a schematic drawing of the laser truing and dressing setup for a superabrasive grinding wheel. The maxim
The laser power density (I) directly influences ablation crater depth. For the pulsed laser with a rectangular pulse waveform and
Fig. 1. Schematic of laser truing and dressing setup.
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surface topography of the grinding wheel after laser truing and dressing, which in turn affect the wheel grinding performance.
2.3. Combined truing with dressing Due to the thermal–physical and optical differences between the diamond abrasives and the bronze bond, appropriate laser parameters (such as: laser average power Pm, pulse repetition frequency f, pulse duration t0, defocusing amount D) and machine motion parameters (feed motion and rotation speed) were selected to remove the diamond abrasives and the bronze bond to realize the truing and dressing process. In-process measurement with optical triangulation was used with the closed-loop control method for controlling the giant pulse output during the laser ablation process. Because the material removal at each pulse is very small, laser ablation is suitable for precision and simultaneous truing and dressing of grinding wheels.
2.4. Experiment results of laser truing and dressing
Fig. 2. Photograph of experimental setup.
Table 1 Laser ablation parameters. Laser type
Nd:YAG
Laser mode Output power Wavelength Pulse repetition frequency Pulse duration Focused spot diameter
Multi-mode 0–100 W 1.06 mm 0.2–50 kHz 170–560 ns 0.26 mm
Table 2 Physical properties of diamond wheel [19]. Property
Diamond
Bronze bond
Thermal conductivity, l (W/m k) Specific heat, c (J/(kg K)) Thermal diffusivity, a (10 4 m2/s) Vaporization temperature, Tv (K) Absorptivity, A Density, r (kg/m3)
2000 1827 3.114 3550 0.15 3515
41.9 352 0.14 2770 0.38 8620
The diameter of abrasive in bronze-bonded grinding wheel is approximately 106–150 mm. For the sake of obtaining satisfactory surface topography, the protruding proportion is about 20–30% of the abrasive size, therefore the demanding protruding height is 20–45 mm. After truing and dressing with the laser beam under the favorable conditions, the surface of the grinding wheel was observed with a scanning electron microscope (SEM) and a light microscope. Fig. 3 shows the surface topography of diamond wheel before laser truing and dressing. As can be seen from the graph that the wheel surface is flat, and the protuberant abrasive is seldom. The ablation crater is produced when single-pulse laser is irradiated on the wheel. Due to the higher peak power and shorter pulse width of acousto-optic Q-switched pulse laser, each pulse removes material at millimicron individually. When ablation layer was arranged in sequence or overlapped, substantial bronze materials is removed. Finally, the abrasives with sharp-edged are well protruded, which means the achievement of laser truing and dressing. Fig. 4 shows the grinding wheel after the truing and dressing process. It can be observed from the figure that the bronze bond is removed evenly, exposing the diamond abrasives with sharp cutting edges from the wheel surface. An appropriate intensity, which is between the below two intensities, can be used for efficiently removing bronze bond and protruding abrasive. The ablation process does not cause material deterioration and cracking in the wheel surface. Instead, it creates enough space in the wheel surface for chip accommodation. And the diamond abrasive edges are also removed reasonably, which indicates that the diamond wheel is ready for grinding. By setting different intensities, it is possible to adjust the ablation depth reduced on the wheel surface when truing and dressing to a value, which depends on the abrasive size for an optimum abrasive protrusion.
a uniform energy distribution in the circular facula region, the power density is expressed as I¼
E 4E 4pm ¼ ¼ ; S t0 p t0 d2 p t0 f d2
ð1Þ
where pm is the average power of the laser; f is pulse repetition frequency; t0 is pulse duration, d is spot diameter. In the experiment, the affected factor of spot diameter is substituted by defocusing amount D, the spot diameter corresponds to defocusing amount, and the pulse duration corresponds to frequency. Therefore, pulse power density is mainly influenced by three parameters Pm, f and D. The three parameters can affect
Fig. 3. Surface topography of diamond wheel before laser truing and dressing. (a) SEM observation and (b) microscope photo.
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Fig. 4. Surface topography of diamond wheel after laser truing and dressing. (a) SEM observation (I= 45 107 W/cm2, Pm =50 W, f = 1 kHz, D = 0 mm) (I = 35 107 W/cm2, Pm = 40 W, f= 1 kHz, D = 0 mm). (b) Microscope photo (I= 45 107 W/cm2, Pm =50 W, f =1 kHz, D = 0 mm) (I = 35 107 W/cm2, Pm = 40 W, f= 1 kHz, D = 0 mm).
Fig. 5. Profile traces of diamond wheel ablated by laser (about three turns).
Fig. 5 shows the profile traces of diamond wheel ablated by laser at optimum condition (Pm30 W, I= 10.9 107 W/cm2, f2 kHz, D = 0 mm), and it is measured by surface coarseness profiling instrument. Seen from Fig. 5, it verifies that the better surface topography is obtained and the protruding height of abrasives is optimum. Before and after laser truing and dressing, the output of roundness error of the grinding wheel was 120 and 9 mm, respectively. This indicated that based on reasonable laser parameters, laser truing and dressing method can reduce the roundness error of the wheel significantly.
2.5. Comparison with truing and dressing methods The comparison of truing and dressing characteristics in laser and conventional methods is summarized in Table 3. In addition,
laser truing and dressing method is applicable to wheels of various specifications with one laser system by controlling the irradiation parameters such as power density, pulse width and feed rate. Otherwise, several types of truing and dressing tools are necessary corresponding to the wheel to be dressed. One of the negative aspects of the laser truing and dressing method is high initial cost of the laser instrument. However, some technical advantages, such as nonconsumable truer and dresser, short truing and dressing time, applicability to any specified wheels, will recoup the initial investment. Furthermore, lasers can be applied to other conditioning process of wheel surface such as cleaning. For purpose of comparison, the corundum block and continuous-wave (CW) laser dressing are conducted on the wheels, shown in Figs. 6 and 7. For mechanical dressing, the abrasives become chipped or rounded on the interaction of shear and
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Table 3 Comparison of truing and dressing characteristics in laser and conventional mechanical method. Factors
Conventional mechanical method
Laser truing and dressing method
Process time Tool wear Pollution of coolant Mechanical force Thermal damage In-process truing and dressing
Long Large Measurable Considerable Less Difficult
Short None Negligible None Controllable Possible
Fig. 7. Comparison with CW laser dressing (Pm = 40 W, D = 0 mm).
Fig. 6. Comparison with corundum block dressing.
extrusion force. The appearance of slots and insufficient protrusion of abrasives is prone to arise [20]. The deep cracks and undercuts, which cause loosening of the abrasives and reducing the number of sharp cutting edges, are also easily produced in mechanical dressing. Dressing tool is easy subjected to wear and the dressing tool geometry will be altered. It influences the dressing conditions, which in turn affects the wheel wear and the grinding process. For CW laser dressing, some deep tunnels come into being around the grains due to the congregation of heat. While far from the abrasive, the bronze bond remove rates are insufficient. Thereby, both the two dressing methods cannot bring about better surface topography.
Fig. 8. Measurement setup for grinding forces.
is measured by roughness measuring instrument. The laser beam was applied to truing and dressing of high-speed grinding wheel under the conditions of the laser power density of 3.5 108 W/cm2 (Pm = 40 W, f= 1 kHz, D =0). The surface speed of the grinding wheel was set at 50 m/s and feed rate Vw = 4.5 m/min. (Fig. 9)
3.2. Results and discussion 3. Grinding performance of laser trued and dressed grinding wheel In order to examine the effect of laser truing and dressing on superabrasive grinding wheel and the influence of laser parameters on wheel surface topography, the research on grinding performance of laser truing and dressing bronze-bonded grinding wheel is conducted. The grinding force and the workpiece surface roughness are the two important attributes to judge the laser truing and dressing effect. Therefore, this section carries on the analysis to grinding force and surface roughness separately. 3.1. High-speed grinding performance measurement setup A precision surface-grinding machine was used in the experimental investigations. For grinding force measurement, a threephase piezoelectric force transducer together with associated power amplifiers and data acquisition system was used. A schematic drawing of the measurement setup is shown in Fig. 8. The workpiece used for grinding is aluminum ceramics. The workpiece surface roughness obtained after high-speed grinding
When the laser defocusing amount (D =0) and the pulse repetition frequency (f =1 kHz) are fixed, the high-speed grinding performance of grinding wheel varies with the change of the laser power density. It can be seen from Fig. 10 that, the normal grinding force falls first then rises with the increase in laser power density. When the laser power density is increased to approximately 4 108 W/cm2, a turning point appears. The normal force rises again beyond the turning point. This may be interpreted as that at a low power density, the truing and dressing is not fully completed by the laser beam due to its insufficient ablation capability at the low power level. As the laser power density increases to the turning point, its ablation capability becomes sufficiently large for appropriate truing and dressing. At this point, the best truing and dressing results can be obtained. Any further increase in the laser power density beyond this point will result in over-ablation of the abrasive layer on the wheel surface, which directly results in destruction to the abrasives, and thus grinding force rise. Therefore, in truing and dressing the grinding wheel, too low or too high power density in the laser beam does not help establish a good surface topography, resulting in the poor performance of the grinding wheel.
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Fig. 11. Relationship between workpiece surface roughness and laser power density.
Fig. 9. Force measurement system.
Fig. 12. Relationship between grinding force and laser pulse repetition frequency.
Fig. 10. Relationship between grinding force and laser power density.
Fig. 11 indicates that at a fixed laser pulse repetition frequency (f=1 kHz) and laser defocusing amount (D =0), workpiece surface roughness increases with the laser power density. This is because at a fixed laser pulse repetition frequency, laser ablation depth is proportional to the laser power density. The higher the power density, the deeper the ablation in an individual pulse, and vice versa. The more ablation depth directly contributes to the amount of abrasive protrusion from the wheel surface, which indirectly increases workpiece surface roughness. When the laser truing and dressing were under the conditions of defocusing amount D = 0 and the power density of I= 2.5 108 W/cm2, the high-speed grinding performance of the diamond wheel was dependent on the laser pulse repetition frequency. It is understood from Fig. 12 that when the power density remains unchanged, the grinding force decrease with the increase in the pulse repetition frequency. Because as the pulse
repetition frequency increases, the number of repetitive ablations also increases, resulting in more truing and dressing on the grinding wheel. The sharpness of the grinding wheel is enhanced due to the effect of repetitive ablations, which in turn reduces grinding force. It can be discovered from Fig. 13 that for a fixed laser power density, the surface roughness of a ground workpiece increases with the laser pulse repetition frequency. The same reason from the above analysis can be applied to this phenomenon. Therefore, with an increase in the pulse frequency, the surface roughness of the grinding wheel increases, this in turn increases the surface roughness of the ground workpiece.
4. Numerical simulation of laser truing and dressing process The laser dressing process is characterized by a small ablation depth and thus very limited heat-affected zone (HAZ) in each pulse. It is considered as an ideal tool for truing and dressing of the bronze-bonded diamond grinding wheels. A two-dimensional mathematical model is developed to simulate laser truing and
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The boundary condition is @T : @n The initial condition is
AI ¼ l
ð3Þ
T9t ¼ 0 ¼ Ta ¼ 300 K;
ð4Þ
where T is temperature at each point in the region; l is the material conductivity; r is material density; cp is material specific heat under constant pressure; A is the grinding wheel surface to the laser absorption coefficient; and I is the incident laser power density. The finite difference method has been popular in temperature field calculations, and it is also convenient in program composition. In the calculation, the array of difference of position variables is in the form of the central differences, for the calculation convenience, the time variable is in the form of the forward array of difference. Therefore, the above heat transfer equation (Eq. (2)) can be transformed into 2-D heat transfer finite difference equations (Eq. (5)). Fig. 13. Relationship between workpiece surface roughness and laser pulse frequency.
Tði; j; kþ 1Þ ¼ Tði; j; kÞ þ þ
Dt a ðDyÞ2
Dt a ðDxÞ2
ðTði þ 1; j; kÞ þ Tði1; j; kÞ2Tði; j; kÞÞ
ðTði; j þ 1; kÞ þ Tði; j1; kÞ2Tði; j; kÞÞ;
x
ð5Þ
bronze bond y
diamond abrasive
Fig. 14. The 2-D geometric model of calculation region within the bronzed-bond diamond wheel.
dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by an acousto-optic Q-switched YAG-pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. 4.1. Establishment of 2-D transient heat transfer As the calculation is in a three-dimensional region, the twodimensional model could be obtained by splitting the grinding wheel along its radial plane. Fig. 14 shows the 2-D geometric model. In order to establish a mathematical model for laser ablation of such a wheel, the following assumptions are made: (1) The bronze-bonded diamond grinding wheel is densely bonded wheel without porosities. (2) Abrasives used in wheel are small balls. (3) The grinding wheel has dense and uniform distribution of abrasives. (4) The laser ablation is considered as a stationary heat source. (5) The pulsed laser heating-up process is assumed to be a two-dimensional heat conduction problem. (6) Surface heat radiation and convection loss from the wheel surface is negligible. The heat transfer equation is expressed as [21,22] ! @T l @2 T @2 T ð2Þ þ ¼ rcp @x2 @y2 @t
2
2
where ((Dt a/(Dx) ) and ((Dt a/(Dy) ) are the grid Fourier factors. The nodal points occupy the whole calculation region, including its periphery. The calculations can be performed based on the Eq. (5). But it should be pointed out that over the periphery of the calculation region, there are four nodal points that are not within the calculation region (is shown in Fig. 15) and must be discussed separately. The representation for Eq. (5), i and j are adopted for numbering the nodal points on x and y axes, and m and n for the maximum values of i and j, respectively. When the nodal point is located at the left boundary of the calculation region, i=1, therefore T(i 1,j,k)=Ta, Ta represents the room temperature; When the nodal point is located at the right boundary of the calculation region, i= m, therefore T(i+ 1,j,k) =Ta; when the nodal point is located at the lower boundary of the calculation region, j= n, therefore T(i, j + 1,k) =Ta; (4) When the nodal point is located at the upper boundary of the calculation region, j= 1, and the laser ablation is directly applied to the surface, which will be discussed in two different cases. The first case is that when the nodal point is within the scope of laser spot, the temperature at the nodal point satisfies Eq. (3). The second case is that when the nodal point is located outside the laser spot, the laser beam does not cause the temperature rise at the nodal point, which is expressed as T(i, j 1,k)= Ta . 4.2. Computation results and discussion In order to analyze material removal and surface topography formation mechanisms in the laser ablation process, MATLAB
Fig. 15. Schematic drawing of nodal point of temperature calculation model.
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software was utilized in the numerical modeling process. When the laser parameters in simulation calculation is fixed that f= 1 kHz, D =0, the relationship between the calculated ablation depth of bronzed-bonded diamond grinding wheel and pulsed laser power density is shown in Fig. 16.
11
Ablation crater depth h(µm)
10 9 8 7
Bronze bond
6
Diamond abrasives
5 4 3 2 1 0 0
10
20
30
40
50
60
70
Pulse laser power density
80
90
100 110
/(107W/cm2)
Fig. 16. Calculated ablation depth in the wheel surface versus laser power density.
Fig. 17. Comparisons between experimental and computational results in laser ablation depth.
It can be seen from Fig. 16 that the ablation crater depth increases with the laser power density. Only when the average power density of the laser reaches a value of 30 107 W/cm2 or higher, both the diamond abrasives and the bronze bond can be simultaneously removed, leading to both truing and dressing; otherwise, the material removal only occurs with the bronze bond, leading to dressing only. Meanwhile, under a given laser ablation condition, the removal depth of the bronze bond is larger than that of the diamond abrasives, causing a 10 mm difference between the two. The truing process is realized through material removal in both the diamond abrasives and the bronze bond, while dressing is obtained through the bond removal process. The laser ablation can thus combine both the truing and dressing processes as a single process. In addition, the ablation crater depth is directly related to the laser power density, i.e., the bigger the power density, the deeper the ablation crater. Fig. 17 shows that the simulation curves of the ablation depth coincide well with the experimental results, implying that within the experimental power density range, the mathematical model is verified. Because of a limited laser power density and the requirement for a reduced heat-affected zone in the material removal process, only three relatively low repetition frequencies were selected for the laser in the calculations of the temperature distributions simulation process. At the instant of a pulse finished (namely, t= t0), temperature distributions in the surface of the bronzebonded diamond grinding wheel under ablation by the laser with different parameters are shown in Figs. 18 and 19. Which has manifested the temperature field as well as ablation depth changes with different laser power density and different repetition frequencies. It is observed from Figs. 18 and 19 that the laser ablation generates an extremely small and thin heat-affected zone but a high temperature gradient. Based on the figures, it can be interpreted that the laser ablation has a limited effect on the surface layer of the grinding wheel in terms of material deterioration. The holding ability of the abrasives by the bond material is not compromised by the laser ablation. Due to the fast injection of the pulsed laser energy into the wheel surface forms a great temperature gradient surrounding the ablation spot, which decomposing and removing the bond and abrasive material. Fig. 18 shows the changes of temperature distribution along with power density, when the repetition frequency is 1 kHz. It can be seen from the figure that, the ablation crater depth of abrasive and bond both grows with the increasing of power density, but the removal of bond is greater than it of abrasive. The ablation crater depth is directly related to the laser power density, i.e., the bigger the power density, the deeper the ablation crater. This is because that the absorption of bronze-bond to laser is bigger than it of diamond abrasives, simultaneously the temperature for bronzebond to remove is lower than it for diamond abrasives removing.
Fig. 18. The temperature fields induced by laser ablation under different power density. (a) I = 2.5 108 W/cm2, f= 1 kHz; (b) I = 3.5 108 W/cm2, f= 1 kHz; and (c) I= 4.5 108 W/cm2, f= 1 kHz.
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Fig. 19. The temperature fields induced by laser ablation under different repetition frequencies. (a) I= 3.5 108 W/cm2, f = 0.5 kHz; (b) I= 3.5 108 W/cm2, f= 1 kHz; and (c) I =3.5 108 W/cm2, f= 2 kHz.
Due to these differences, the truing process and dressing process can be simultaneously realized with the laser ablation as a single process. Fig. 19 shows the changes of temperature distribution along with repetition frequency, when the power density is 3.5 108 W/cm2. It can be seen from the figure that for a given power density, the ablation crater depth of bronze-bond increases with the pulse repetition frequency. Simultaneously, the temperature gradient on abrasives surface also increases obviously. This is because as the repetition frequency increases, the laser pulse width increase, and then the single-pulse irradiation time increase. Therefore, the ablation depth in both the bronze bond and diamond abrasives increases.
5. Conclusions In this paper a acousto-optic Q-switched Nd:YAG pulse laser with high power density, short pulse width, and small heataffected zone is independently developed for truing and dressing of a bronze-bonded diamond grinding wheel. This experiment deals with surface topography ablated by single-pulse laser, and examines the truing and dressing characteristic by a grinding test. Finally, develops a finite difference model for simulating the mechanism of laser ablation bronze-bonded diamond wheels. (1) Due to the thermal–physical and optical differences between bond and abrasives, suitable pulse laser parameters are selected to remove bronze bond and diamond abrasives realizing the truing and dressing process. Optical triangulation is incorporated into a measurement system for online closed-loop control of the laser truing and dressing process. (2) Satisfactory surface topography can be obtained under the appropriate conditions. Based on reasonable laser parameters, laser truing and dressing method can reduce the roundness error of the specimen significantly. Compared with the corundum block and CW laser dressing methods, acoustooptic Q-switched pulse laser is more suitable for dressing bronze-bonded diamond superabraisve grinding wheel. (3) In order to examine the effect of laser truing and dressing on superabrasive grinding wheel surface topography, the research on grinding performance of bronze-bonded grinding wheel trued and dressed by pulse laser is conducted. For a given laser pulse repetition frequency and defocusing amount, an optimal power density exists at which grinding force reach its minimum values. Therefore, in truing and dressing the grinding wheel, too low or too high power density in the laser beam does not help establish a good surface topography, resulting in the poor performance of the grinding wheel. (4) A 2-D transient heat transfer finite difference model is developed for simulating laser ablation of bronze-bonded
diamond wheels. The ablation crater depth and temperature field are obtained for single-pulse laser ablation of the bronze bond and diamond abrasives. Results show that, the ablation crater depth increases with the laser power density, when the power density reaches a value of 30 107 W/cm2 or higher, both the diamond abrasives and the bronze bond can be simultaneously removed. The simulation results coincide well with the experimental results. This proved that the truing and dressing are successfully combined as a single process with the laser ablation.
Acknowledgements This project is financially supported by the State Natural Science Foundation of People’s Republic of China (no. 50005005), the State High-Tech Research and Development Plans (863 a state key project) (no. 2007AA042006) and the Development Program of Cheung Kong Scholars and Innovative Team (no. 531105050037). References [1] Webster J, Tricard M. Innovations in abrasive products for precision grinding. Annals of the CIRP 2004;53(2):597–617. [2] Xie X-Z, Chen G-Y, Li L-J. Dressing of resin-bonded superabrasive grinding wheels by means of acousto-optic Q-switched pulsed Nd:YAG laser. Optics & Laser Technology 2004;36:409–19. [3] Klocke F, Konig W. Appropriate conditioning strategies increase the performance capabilities of vitrified-bond CBN grinding wheels. Annals of the CIRP 1995;44(1):305–10. [4] Liu Junhe, Liu Wei. The application of diamond roller in dressing grinding wheel. Bearing 2003;12:43–4. [5] Zhang Hongxia, Chen Zhitong, Chen Wuyi. Dressing performance and dressing mechanism of diamond roller. Journal of Beijing University of Aeronautics and Astronautics 2008;4(7):816–20. [6] Wang Jianjun, Yao Bing. Research on grinding wheel’s dressing accuracy in shape grinding. Diamond & Abrasives Engineering 2002;5(131):48–9. [7] Fu Junjie, Zhang Yongzhong. Grinding wheel dressed by diamond stylus process the shaping gear. Machinery 1994;02:14–5. [8] A. Hosokawa, T. Ueda, T. Yunoki, Laser dressing of metal bonded diamond wheel, Annals of the CIRP 2006. [9] Heiji Yasui, Hiroshi Nakazono, Akira Hosokawa, Makoto Kurusu. A new efficient dressing method of the resin bond CBN wheel. International Journal of the Japan Society for Precision Engineering 1991;25(2):98–109. [10] Zhang Bingpeng, Zhao Wenxiang, Wang Xibin Relationship modeling between granularity of GC cup wheel and dressing efficiency of diamond grinding wheel. Tool Engineering 2006;40(4):31–4. [11] Hirao Masatoshi, Izawa Masaki, Iguchi Nobuaki, et al. Waterjet in-process dressing (1st Report) – dressing property and jet pressure. Journal of Japan Society for Precision Engineering 1995;61(7)1335-1339 1995;61(7). [12] Cai Lanrong, Jia Yan, Hu Dejin. Analysis of temperature field and experimental validation for electrical discharge dressing and truing of electrical diamond grinding wheel. China Mechanical Engineering 2008;19(9):1009–12. [13] Jackson MJ, Robinson NB, Khangar A, Moss R. Laser dressing of vitrified aluminum oxide grinding wheels. British Ceramic Transactions 2003;102(3): 238–245. [14] Kang RK, Yuan JT, Shi XK, Lan J, Ren JX. Study on laser dressing of superabrasive grinding wheels. China Mechanical Engineering 2000;5(11):493–6.
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[15] Kang RK, Yuan JT, Zhang YP, Ren JX. Truing of diamond wheels by laser. Key Engineering Materials 2001;202-203:137–42. ¨ [16] Westkamper E. Grinding assisted by Nd:YAG lasers. The Annals of the CIRP 1995;1(44):317–20. [17] Dunwen Zuo, Yoshihiro Kawano, Toshikazu Yamashita, et al. Effect of CO2 pulsed laser scanning on diamond wheel pieces. Chinese Journal of Mechanical Engineering 1999;35(5):42–5. [18] Zhang C, Shin YC. A novel laser assisted truing and dressing technique for vitrified CBN wheels. International Journal of Machine Tools and Manufacture 2002;42:825–35.
[19] Mei LF, Chen GY, Zhang B, et al. Measurement of YAG laser absorptance by artificial diamond and cubic boron nitride. Optics & Laser Technology 2009;41:770–7. [20] Chen GY, Li LJ, Ma HL, et al. Laser truing and dressing of bronze-bonded diamond grinding wheels by a Q-switched YAG pulse laser. Journal of Mechanical Engineering 2005;41(4):74–179. [21] Yang, SM, Tao, WQ. Heat transfer theory. Beijing: Higher Education Press 2003. [22] Tao, WQ. Numerical heat transfer theory. Xi’an Jiaotong University Publications 1988.
Contents lists available at ScienceDirect
Optics and Lasers in Engineering journal homepage: www.elsevier.com/locate/optlaseng
Experiment and numerical simulation study on laser truing and dressing of bronze-bonded diamond wheel Genyu Chen a, Lifang Mei a,n, Bi Zhang b, Chunrong Yu a, Kangjian Shun a a b
State Key Laboratory of Advanced Design and Manufacturing for Vehicle Body, College of Mechanical and Automobile Engineering, Hunan University, Changsha 410082, China University of Connecticut, Storrs, CT, USA
a r t i c l e in fo
abstract
Article history: Received 8 July 2009 Received in revised form 2 November 2009 Accepted 10 November 2009 Available online 16 December 2009
An acoustic-optic Q-switched YAG laser is newly developed for truing and dressing of bronze-bonded diamond grinding wheels through ablation in the radial direction. The laser ablation merges truing and dressing as a single process and can be easily controlled through an online closed-loop control system. The laser truing and dressing characteristics are investigated by wheel surface topography and a highspeed grinding test. A two-dimensional mathematical model is developed to simulate laser truing and dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by a pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. Both the experimental results and the theoretical analysis indicate that the truing and dressing processes can be simultaneously realized with the laser ablation. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Laser truing and dressing Bronze-bonded diamond grinding wheel Topography Grinding performance Numerical simulation
1. Introduction Diamond superabrasive grinding wheels are widely used in grinding hard brittle non-ferrous metals, horniness alloys and hard brittle non-metals like ceramics, optical glasses and gems, because of their high hardness, thermal resistance and long tool life [1]. Grinding with superabrasive wheels offers a series of advantages such as high grinding efficiency, fine surface quality and finedimensional stability. Due to their superior grinding performances, they are extensively applied to such fields as aerospace, automotive, electronics, medicine and so on [2]. In order to achieve efficient or fine grinding, a specific surface topography on the wheel face should be generated and therefore periodic truing and dressing are required [3]. It is, however, difficult to sharpen these superabrasive wheels with a high degree of accuracy, due to the qualities mentioned above, which could inevitably affect their applications. Therefore, the truing and dressing of superabrasive wheels have been a research topic. In general, the superabrasive grinding wheels can be dressed by means of conventional mechanical dressing methods on the interaction of force, such as diamond rolling methods [4,5] and diamond stylus shaping methods [6,7]. But these truing and dressing methods, especially in the case of diamond wheels, are time-consuming and cause pollution of the working
n
Corresponding author. E-mail addresses: [email protected] (G. Chen), [email protected] (L. Mei). 0143-8166/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.optlaseng.2009.11.006
environment. In addition, mechanical damage of the abrasives is induced by the direct contact between the wheel and the truer and dresser [8]. Truing and dressing tools are easy subjected to wear and their geometry will be altered. It influences the truing and dressing conditions, which in turn affects the wheel wear and the grinding process. In order to solve these problems and achieve a good truing and dressing quality, a series of studies have conducted recently, which include improving and developing new methods, such as stick-aided loose abrasive (SLAD) dressing method [9], Green Carborundum (GC) cup dressing method[10], water jet dressing method [11], and electrical discharge dressing and truing method [12], etc. although these studies provide insight into proper truing and dressing methods and optimum ablation parameters, the rapid wear of truer and dresser and unsatisfactory topography of wheels remain unresolved. Thus, it is necessary to find an appropriate method for truing and dressing the superabrasive wheels. As an alternative to the conventional dressing method, laser truing and dressing is proposed as a new technique with many merits, such as extensive adaptability, high efficiency, non-pollution to environment, non-contact and high geometric accuracies. Laser truing and dressing is a method which uses laser to true and dress grinding wheels. The grinding wheels need to be dressed after being used for a period, even before putting into use. Truing means to remove part bond around the abrasives, so the abrasives can protrude bond. This process engenders enough space for swarf extraction and forms cutting edge. Dressing refers to micromachining on workface of grinding wheels until attaining required
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geometry accuracy. A number of studies have been conducted recently, most of which are focused on dressing of grinding wheels and a few experiments have investigated the truing diamond wheels, such as Jackson et al. [13] made a study of laser dressing of vitrified aluminum oxide grinding wheels. Kang et al. [14,15] conducted a study of tangentially truing metal-bonded diamond wheel by conventional pulsed Nd:YAG laser. Westk¨amper et al. [16] employed the conventional pulsed Nd:YAG laser to dress resinbonded CBN wheel and proposed a method of dressing on wheels in orthogonal direction, while truing in tangential direction. Zuo Dunwen et al. [17] made a study of diamond plate dressed by CW CO2 laser. Moreover, a technique of CO2 laser-assisted truing and dressing for vitrified CBN wheel developed by Zhang and Shin [18]. However, systematic study of truing and dressing on bronzebonded diamond superabrasive wheels by acousto-optic Qswitched pulsed Nd:YAG laser as a single process has not been reported for the moment. In this paper, an acoustic-optic Q-switched Nd:YAG laser is used to ablate a bronze-bonded diamond grinding wheel in the radial direction to realize in-process truing and dressing. The mechanism of selective removal and surface topography of laser truing and dressing wheel are exhaustively analyzed. The laser truing and dressing characteristics of the wheel are also examined and comparison made to a conventionally dressed grinding wheel. The grinding performance of the laser trued and dressed wheel was evaluated in terms of grinding force and workpiece surface roughness. A two-dimensional mathematical model is developed to simulate laser truing and dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by an acousto-optic Q-switched YAG-pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. The laser dressing process is characterized by a small ablation depth and thus very limited heat-affected zone (HAZ) in each pulse. It is considered as an ideal tool for truing and dressing of the superabrasive wheels.
2. Laser truing and dressing experiment
continuous output power of the Nd:Yag laser is 100 W, which is modified by current of krypton discharge source. There is an acousto-optic Q-switched device built in the resonant cavity whose frequency ranges from 0.2 to 50 kHz and therefore a pulse laser with pulse width (t0) from 170 to 560 ns and pulse repetition frequency (f) from 0.2 to 50 kHz can be secured. The output laser with a wavelength of 1.06 mm is irradiated radially on the wheel surface. The focus lens is aberrationless hyperbolical with its focal length 60 mm. A continuous laser beam is sent out from the Nd:YAG laser for wheel surface measurements, and a pulsed laser beam sent from the same YAG laser by acoustic-optic Q-switched driver for grinding wheel ablation. The two laser beams are actually from the same laser source. The Q-switch can automatically shut-off the pulsed laser beam if the grinding wheel surface level is too low, resulting in no material removal from the wheel surface. If the wheel surface level is too high, the Q-switch is turned on to irradiate a strong pulsed laser beam on the grinding wheel surface. Fig. 2 shows a photographic view of the experimental setup, while Table 1 lists laser ablation parameters. Experimental installation of single-pulse ablation on superabrasive wheels is secured by coordination the feed motion and rotation speed of lathe machine. The focus spot diameter is measured to be 0.26 mm and the beam divergent angle is 20 mrad. The output power of continuous laser is measured by power device (model: PC-1000, made in Canada), the surface topography after laser truing and dressing on wheels is measured by light microscope (LM) (model: Leica DMLM, made in England) and scanning electronic microscope (SEM) (model: STEREOSCAN440, made in England), the protruding height of abrasives is measured by surface roughness measuring instrument and Form Talysurf (model: Roughness Talysurf Series 2), the roundness error of wheel is measured by micrometer checker and Taly roundness measuring equipment (model: Taly265). A bronzebonded diamond grinding wheel (denoted as MBD100/120 M100) is chosen as the truing and dressing target. Table 2 provides the physical properties of the bronze bond and diamond abrasives. The dressing characteristics are examined by wheel topography and a grinding test.
2.1. Laser truing and dressing system
2.2. Laser ablation parameters
Fig. 1 shows a schematic drawing of the laser truing and dressing setup for a superabrasive grinding wheel. The maxim
The laser power density (I) directly influences ablation crater depth. For the pulsed laser with a rectangular pulse waveform and
Fig. 1. Schematic of laser truing and dressing setup.
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surface topography of the grinding wheel after laser truing and dressing, which in turn affect the wheel grinding performance.
2.3. Combined truing with dressing Due to the thermal–physical and optical differences between the diamond abrasives and the bronze bond, appropriate laser parameters (such as: laser average power Pm, pulse repetition frequency f, pulse duration t0, defocusing amount D) and machine motion parameters (feed motion and rotation speed) were selected to remove the diamond abrasives and the bronze bond to realize the truing and dressing process. In-process measurement with optical triangulation was used with the closed-loop control method for controlling the giant pulse output during the laser ablation process. Because the material removal at each pulse is very small, laser ablation is suitable for precision and simultaneous truing and dressing of grinding wheels.
2.4. Experiment results of laser truing and dressing
Fig. 2. Photograph of experimental setup.
Table 1 Laser ablation parameters. Laser type
Nd:YAG
Laser mode Output power Wavelength Pulse repetition frequency Pulse duration Focused spot diameter
Multi-mode 0–100 W 1.06 mm 0.2–50 kHz 170–560 ns 0.26 mm
Table 2 Physical properties of diamond wheel [19]. Property
Diamond
Bronze bond
Thermal conductivity, l (W/m k) Specific heat, c (J/(kg K)) Thermal diffusivity, a (10 4 m2/s) Vaporization temperature, Tv (K) Absorptivity, A Density, r (kg/m3)
2000 1827 3.114 3550 0.15 3515
41.9 352 0.14 2770 0.38 8620
The diameter of abrasive in bronze-bonded grinding wheel is approximately 106–150 mm. For the sake of obtaining satisfactory surface topography, the protruding proportion is about 20–30% of the abrasive size, therefore the demanding protruding height is 20–45 mm. After truing and dressing with the laser beam under the favorable conditions, the surface of the grinding wheel was observed with a scanning electron microscope (SEM) and a light microscope. Fig. 3 shows the surface topography of diamond wheel before laser truing and dressing. As can be seen from the graph that the wheel surface is flat, and the protuberant abrasive is seldom. The ablation crater is produced when single-pulse laser is irradiated on the wheel. Due to the higher peak power and shorter pulse width of acousto-optic Q-switched pulse laser, each pulse removes material at millimicron individually. When ablation layer was arranged in sequence or overlapped, substantial bronze materials is removed. Finally, the abrasives with sharp-edged are well protruded, which means the achievement of laser truing and dressing. Fig. 4 shows the grinding wheel after the truing and dressing process. It can be observed from the figure that the bronze bond is removed evenly, exposing the diamond abrasives with sharp cutting edges from the wheel surface. An appropriate intensity, which is between the below two intensities, can be used for efficiently removing bronze bond and protruding abrasive. The ablation process does not cause material deterioration and cracking in the wheel surface. Instead, it creates enough space in the wheel surface for chip accommodation. And the diamond abrasive edges are also removed reasonably, which indicates that the diamond wheel is ready for grinding. By setting different intensities, it is possible to adjust the ablation depth reduced on the wheel surface when truing and dressing to a value, which depends on the abrasive size for an optimum abrasive protrusion.
a uniform energy distribution in the circular facula region, the power density is expressed as I¼
E 4E 4pm ¼ ¼ ; S t0 p t0 d2 p t0 f d2
ð1Þ
where pm is the average power of the laser; f is pulse repetition frequency; t0 is pulse duration, d is spot diameter. In the experiment, the affected factor of spot diameter is substituted by defocusing amount D, the spot diameter corresponds to defocusing amount, and the pulse duration corresponds to frequency. Therefore, pulse power density is mainly influenced by three parameters Pm, f and D. The three parameters can affect
Fig. 3. Surface topography of diamond wheel before laser truing and dressing. (a) SEM observation and (b) microscope photo.
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Fig. 4. Surface topography of diamond wheel after laser truing and dressing. (a) SEM observation (I= 45 107 W/cm2, Pm =50 W, f = 1 kHz, D = 0 mm) (I = 35 107 W/cm2, Pm = 40 W, f= 1 kHz, D = 0 mm). (b) Microscope photo (I= 45 107 W/cm2, Pm =50 W, f =1 kHz, D = 0 mm) (I = 35 107 W/cm2, Pm = 40 W, f= 1 kHz, D = 0 mm).
Fig. 5. Profile traces of diamond wheel ablated by laser (about three turns).
Fig. 5 shows the profile traces of diamond wheel ablated by laser at optimum condition (Pm30 W, I= 10.9 107 W/cm2, f2 kHz, D = 0 mm), and it is measured by surface coarseness profiling instrument. Seen from Fig. 5, it verifies that the better surface topography is obtained and the protruding height of abrasives is optimum. Before and after laser truing and dressing, the output of roundness error of the grinding wheel was 120 and 9 mm, respectively. This indicated that based on reasonable laser parameters, laser truing and dressing method can reduce the roundness error of the wheel significantly.
2.5. Comparison with truing and dressing methods The comparison of truing and dressing characteristics in laser and conventional methods is summarized in Table 3. In addition,
laser truing and dressing method is applicable to wheels of various specifications with one laser system by controlling the irradiation parameters such as power density, pulse width and feed rate. Otherwise, several types of truing and dressing tools are necessary corresponding to the wheel to be dressed. One of the negative aspects of the laser truing and dressing method is high initial cost of the laser instrument. However, some technical advantages, such as nonconsumable truer and dresser, short truing and dressing time, applicability to any specified wheels, will recoup the initial investment. Furthermore, lasers can be applied to other conditioning process of wheel surface such as cleaning. For purpose of comparison, the corundum block and continuous-wave (CW) laser dressing are conducted on the wheels, shown in Figs. 6 and 7. For mechanical dressing, the abrasives become chipped or rounded on the interaction of shear and
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Table 3 Comparison of truing and dressing characteristics in laser and conventional mechanical method. Factors
Conventional mechanical method
Laser truing and dressing method
Process time Tool wear Pollution of coolant Mechanical force Thermal damage In-process truing and dressing
Long Large Measurable Considerable Less Difficult
Short None Negligible None Controllable Possible
Fig. 7. Comparison with CW laser dressing (Pm = 40 W, D = 0 mm).
Fig. 6. Comparison with corundum block dressing.
extrusion force. The appearance of slots and insufficient protrusion of abrasives is prone to arise [20]. The deep cracks and undercuts, which cause loosening of the abrasives and reducing the number of sharp cutting edges, are also easily produced in mechanical dressing. Dressing tool is easy subjected to wear and the dressing tool geometry will be altered. It influences the dressing conditions, which in turn affects the wheel wear and the grinding process. For CW laser dressing, some deep tunnels come into being around the grains due to the congregation of heat. While far from the abrasive, the bronze bond remove rates are insufficient. Thereby, both the two dressing methods cannot bring about better surface topography.
Fig. 8. Measurement setup for grinding forces.
is measured by roughness measuring instrument. The laser beam was applied to truing and dressing of high-speed grinding wheel under the conditions of the laser power density of 3.5 108 W/cm2 (Pm = 40 W, f= 1 kHz, D =0). The surface speed of the grinding wheel was set at 50 m/s and feed rate Vw = 4.5 m/min. (Fig. 9)
3.2. Results and discussion 3. Grinding performance of laser trued and dressed grinding wheel In order to examine the effect of laser truing and dressing on superabrasive grinding wheel and the influence of laser parameters on wheel surface topography, the research on grinding performance of laser truing and dressing bronze-bonded grinding wheel is conducted. The grinding force and the workpiece surface roughness are the two important attributes to judge the laser truing and dressing effect. Therefore, this section carries on the analysis to grinding force and surface roughness separately. 3.1. High-speed grinding performance measurement setup A precision surface-grinding machine was used in the experimental investigations. For grinding force measurement, a threephase piezoelectric force transducer together with associated power amplifiers and data acquisition system was used. A schematic drawing of the measurement setup is shown in Fig. 8. The workpiece used for grinding is aluminum ceramics. The workpiece surface roughness obtained after high-speed grinding
When the laser defocusing amount (D =0) and the pulse repetition frequency (f =1 kHz) are fixed, the high-speed grinding performance of grinding wheel varies with the change of the laser power density. It can be seen from Fig. 10 that, the normal grinding force falls first then rises with the increase in laser power density. When the laser power density is increased to approximately 4 108 W/cm2, a turning point appears. The normal force rises again beyond the turning point. This may be interpreted as that at a low power density, the truing and dressing is not fully completed by the laser beam due to its insufficient ablation capability at the low power level. As the laser power density increases to the turning point, its ablation capability becomes sufficiently large for appropriate truing and dressing. At this point, the best truing and dressing results can be obtained. Any further increase in the laser power density beyond this point will result in over-ablation of the abrasive layer on the wheel surface, which directly results in destruction to the abrasives, and thus grinding force rise. Therefore, in truing and dressing the grinding wheel, too low or too high power density in the laser beam does not help establish a good surface topography, resulting in the poor performance of the grinding wheel.
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Fig. 11. Relationship between workpiece surface roughness and laser power density.
Fig. 9. Force measurement system.
Fig. 12. Relationship between grinding force and laser pulse repetition frequency.
Fig. 10. Relationship between grinding force and laser power density.
Fig. 11 indicates that at a fixed laser pulse repetition frequency (f=1 kHz) and laser defocusing amount (D =0), workpiece surface roughness increases with the laser power density. This is because at a fixed laser pulse repetition frequency, laser ablation depth is proportional to the laser power density. The higher the power density, the deeper the ablation in an individual pulse, and vice versa. The more ablation depth directly contributes to the amount of abrasive protrusion from the wheel surface, which indirectly increases workpiece surface roughness. When the laser truing and dressing were under the conditions of defocusing amount D = 0 and the power density of I= 2.5 108 W/cm2, the high-speed grinding performance of the diamond wheel was dependent on the laser pulse repetition frequency. It is understood from Fig. 12 that when the power density remains unchanged, the grinding force decrease with the increase in the pulse repetition frequency. Because as the pulse
repetition frequency increases, the number of repetitive ablations also increases, resulting in more truing and dressing on the grinding wheel. The sharpness of the grinding wheel is enhanced due to the effect of repetitive ablations, which in turn reduces grinding force. It can be discovered from Fig. 13 that for a fixed laser power density, the surface roughness of a ground workpiece increases with the laser pulse repetition frequency. The same reason from the above analysis can be applied to this phenomenon. Therefore, with an increase in the pulse frequency, the surface roughness of the grinding wheel increases, this in turn increases the surface roughness of the ground workpiece.
4. Numerical simulation of laser truing and dressing process The laser dressing process is characterized by a small ablation depth and thus very limited heat-affected zone (HAZ) in each pulse. It is considered as an ideal tool for truing and dressing of the bronze-bonded diamond grinding wheels. A two-dimensional mathematical model is developed to simulate laser truing and
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The boundary condition is @T : @n The initial condition is
AI ¼ l
ð3Þ
T9t ¼ 0 ¼ Ta ¼ 300 K;
ð4Þ
where T is temperature at each point in the region; l is the material conductivity; r is material density; cp is material specific heat under constant pressure; A is the grinding wheel surface to the laser absorption coefficient; and I is the incident laser power density. The finite difference method has been popular in temperature field calculations, and it is also convenient in program composition. In the calculation, the array of difference of position variables is in the form of the central differences, for the calculation convenience, the time variable is in the form of the forward array of difference. Therefore, the above heat transfer equation (Eq. (2)) can be transformed into 2-D heat transfer finite difference equations (Eq. (5)). Fig. 13. Relationship between workpiece surface roughness and laser pulse frequency.
Tði; j; kþ 1Þ ¼ Tði; j; kÞ þ þ
Dt a ðDyÞ2
Dt a ðDxÞ2
ðTði þ 1; j; kÞ þ Tði1; j; kÞ2Tði; j; kÞÞ
ðTði; j þ 1; kÞ þ Tði; j1; kÞ2Tði; j; kÞÞ;
x
ð5Þ
bronze bond y
diamond abrasive
Fig. 14. The 2-D geometric model of calculation region within the bronzed-bond diamond wheel.
dressing of bronze-bonded diamond grinding wheel. Based on the model, the ablation depth produced on bond and diamond abrasives by an acousto-optic Q-switched YAG-pulsed laser under different parameters as well as the temperature field on grinding wheel surface produced in the laser ablation process are numerically simulated. 4.1. Establishment of 2-D transient heat transfer As the calculation is in a three-dimensional region, the twodimensional model could be obtained by splitting the grinding wheel along its radial plane. Fig. 14 shows the 2-D geometric model. In order to establish a mathematical model for laser ablation of such a wheel, the following assumptions are made: (1) The bronze-bonded diamond grinding wheel is densely bonded wheel without porosities. (2) Abrasives used in wheel are small balls. (3) The grinding wheel has dense and uniform distribution of abrasives. (4) The laser ablation is considered as a stationary heat source. (5) The pulsed laser heating-up process is assumed to be a two-dimensional heat conduction problem. (6) Surface heat radiation and convection loss from the wheel surface is negligible. The heat transfer equation is expressed as [21,22] ! @T l @2 T @2 T ð2Þ þ ¼ rcp @x2 @y2 @t
2
2
where ((Dt a/(Dx) ) and ((Dt a/(Dy) ) are the grid Fourier factors. The nodal points occupy the whole calculation region, including its periphery. The calculations can be performed based on the Eq. (5). But it should be pointed out that over the periphery of the calculation region, there are four nodal points that are not within the calculation region (is shown in Fig. 15) and must be discussed separately. The representation for Eq. (5), i and j are adopted for numbering the nodal points on x and y axes, and m and n for the maximum values of i and j, respectively. When the nodal point is located at the left boundary of the calculation region, i=1, therefore T(i 1,j,k)=Ta, Ta represents the room temperature; When the nodal point is located at the right boundary of the calculation region, i= m, therefore T(i+ 1,j,k) =Ta; when the nodal point is located at the lower boundary of the calculation region, j= n, therefore T(i, j + 1,k) =Ta; (4) When the nodal point is located at the upper boundary of the calculation region, j= 1, and the laser ablation is directly applied to the surface, which will be discussed in two different cases. The first case is that when the nodal point is within the scope of laser spot, the temperature at the nodal point satisfies Eq. (3). The second case is that when the nodal point is located outside the laser spot, the laser beam does not cause the temperature rise at the nodal point, which is expressed as T(i, j 1,k)= Ta . 4.2. Computation results and discussion In order to analyze material removal and surface topography formation mechanisms in the laser ablation process, MATLAB
Fig. 15. Schematic drawing of nodal point of temperature calculation model.
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software was utilized in the numerical modeling process. When the laser parameters in simulation calculation is fixed that f= 1 kHz, D =0, the relationship between the calculated ablation depth of bronzed-bonded diamond grinding wheel and pulsed laser power density is shown in Fig. 16.
11
Ablation crater depth h(µm)
10 9 8 7
Bronze bond
6
Diamond abrasives
5 4 3 2 1 0 0
10
20
30
40
50
60
70
Pulse laser power density
80
90
100 110
/(107W/cm2)
Fig. 16. Calculated ablation depth in the wheel surface versus laser power density.
Fig. 17. Comparisons between experimental and computational results in laser ablation depth.
It can be seen from Fig. 16 that the ablation crater depth increases with the laser power density. Only when the average power density of the laser reaches a value of 30 107 W/cm2 or higher, both the diamond abrasives and the bronze bond can be simultaneously removed, leading to both truing and dressing; otherwise, the material removal only occurs with the bronze bond, leading to dressing only. Meanwhile, under a given laser ablation condition, the removal depth of the bronze bond is larger than that of the diamond abrasives, causing a 10 mm difference between the two. The truing process is realized through material removal in both the diamond abrasives and the bronze bond, while dressing is obtained through the bond removal process. The laser ablation can thus combine both the truing and dressing processes as a single process. In addition, the ablation crater depth is directly related to the laser power density, i.e., the bigger the power density, the deeper the ablation crater. Fig. 17 shows that the simulation curves of the ablation depth coincide well with the experimental results, implying that within the experimental power density range, the mathematical model is verified. Because of a limited laser power density and the requirement for a reduced heat-affected zone in the material removal process, only three relatively low repetition frequencies were selected for the laser in the calculations of the temperature distributions simulation process. At the instant of a pulse finished (namely, t= t0), temperature distributions in the surface of the bronzebonded diamond grinding wheel under ablation by the laser with different parameters are shown in Figs. 18 and 19. Which has manifested the temperature field as well as ablation depth changes with different laser power density and different repetition frequencies. It is observed from Figs. 18 and 19 that the laser ablation generates an extremely small and thin heat-affected zone but a high temperature gradient. Based on the figures, it can be interpreted that the laser ablation has a limited effect on the surface layer of the grinding wheel in terms of material deterioration. The holding ability of the abrasives by the bond material is not compromised by the laser ablation. Due to the fast injection of the pulsed laser energy into the wheel surface forms a great temperature gradient surrounding the ablation spot, which decomposing and removing the bond and abrasive material. Fig. 18 shows the changes of temperature distribution along with power density, when the repetition frequency is 1 kHz. It can be seen from the figure that, the ablation crater depth of abrasive and bond both grows with the increasing of power density, but the removal of bond is greater than it of abrasive. The ablation crater depth is directly related to the laser power density, i.e., the bigger the power density, the deeper the ablation crater. This is because that the absorption of bronze-bond to laser is bigger than it of diamond abrasives, simultaneously the temperature for bronzebond to remove is lower than it for diamond abrasives removing.
Fig. 18. The temperature fields induced by laser ablation under different power density. (a) I = 2.5 108 W/cm2, f= 1 kHz; (b) I = 3.5 108 W/cm2, f= 1 kHz; and (c) I= 4.5 108 W/cm2, f= 1 kHz.
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Fig. 19. The temperature fields induced by laser ablation under different repetition frequencies. (a) I= 3.5 108 W/cm2, f = 0.5 kHz; (b) I= 3.5 108 W/cm2, f= 1 kHz; and (c) I =3.5 108 W/cm2, f= 2 kHz.
Due to these differences, the truing process and dressing process can be simultaneously realized with the laser ablation as a single process. Fig. 19 shows the changes of temperature distribution along with repetition frequency, when the power density is 3.5 108 W/cm2. It can be seen from the figure that for a given power density, the ablation crater depth of bronze-bond increases with the pulse repetition frequency. Simultaneously, the temperature gradient on abrasives surface also increases obviously. This is because as the repetition frequency increases, the laser pulse width increase, and then the single-pulse irradiation time increase. Therefore, the ablation depth in both the bronze bond and diamond abrasives increases.
5. Conclusions In this paper a acousto-optic Q-switched Nd:YAG pulse laser with high power density, short pulse width, and small heataffected zone is independently developed for truing and dressing of a bronze-bonded diamond grinding wheel. This experiment deals with surface topography ablated by single-pulse laser, and examines the truing and dressing characteristic by a grinding test. Finally, develops a finite difference model for simulating the mechanism of laser ablation bronze-bonded diamond wheels. (1) Due to the thermal–physical and optical differences between bond and abrasives, suitable pulse laser parameters are selected to remove bronze bond and diamond abrasives realizing the truing and dressing process. Optical triangulation is incorporated into a measurement system for online closed-loop control of the laser truing and dressing process. (2) Satisfactory surface topography can be obtained under the appropriate conditions. Based on reasonable laser parameters, laser truing and dressing method can reduce the roundness error of the specimen significantly. Compared with the corundum block and CW laser dressing methods, acoustooptic Q-switched pulse laser is more suitable for dressing bronze-bonded diamond superabraisve grinding wheel. (3) In order to examine the effect of laser truing and dressing on superabrasive grinding wheel surface topography, the research on grinding performance of bronze-bonded grinding wheel trued and dressed by pulse laser is conducted. For a given laser pulse repetition frequency and defocusing amount, an optimal power density exists at which grinding force reach its minimum values. Therefore, in truing and dressing the grinding wheel, too low or too high power density in the laser beam does not help establish a good surface topography, resulting in the poor performance of the grinding wheel. (4) A 2-D transient heat transfer finite difference model is developed for simulating laser ablation of bronze-bonded
diamond wheels. The ablation crater depth and temperature field are obtained for single-pulse laser ablation of the bronze bond and diamond abrasives. Results show that, the ablation crater depth increases with the laser power density, when the power density reaches a value of 30 107 W/cm2 or higher, both the diamond abrasives and the bronze bond can be simultaneously removed. The simulation results coincide well with the experimental results. This proved that the truing and dressing are successfully combined as a single process with the laser ablation.
Acknowledgements This project is financially supported by the State Natural Science Foundation of People’s Republic of China (no. 50005005), the State High-Tech Research and Development Plans (863 a state key project) (no. 2007AA042006) and the Development Program of Cheung Kong Scholars and Innovative Team (no. 531105050037). References [1] Webster J, Tricard M. Innovations in abrasive products for precision grinding. Annals of the CIRP 2004;53(2):597–617. [2] Xie X-Z, Chen G-Y, Li L-J. Dressing of resin-bonded superabrasive grinding wheels by means of acousto-optic Q-switched pulsed Nd:YAG laser. Optics & Laser Technology 2004;36:409–19. [3] Klocke F, Konig W. Appropriate conditioning strategies increase the performance capabilities of vitrified-bond CBN grinding wheels. Annals of the CIRP 1995;44(1):305–10. [4] Liu Junhe, Liu Wei. The application of diamond roller in dressing grinding wheel. Bearing 2003;12:43–4. [5] Zhang Hongxia, Chen Zhitong, Chen Wuyi. Dressing performance and dressing mechanism of diamond roller. Journal of Beijing University of Aeronautics and Astronautics 2008;4(7):816–20. [6] Wang Jianjun, Yao Bing. Research on grinding wheel’s dressing accuracy in shape grinding. Diamond & Abrasives Engineering 2002;5(131):48–9. [7] Fu Junjie, Zhang Yongzhong. Grinding wheel dressed by diamond stylus process the shaping gear. Machinery 1994;02:14–5. [8] A. Hosokawa, T. Ueda, T. Yunoki, Laser dressing of metal bonded diamond wheel, Annals of the CIRP 2006. [9] Heiji Yasui, Hiroshi Nakazono, Akira Hosokawa, Makoto Kurusu. A new efficient dressing method of the resin bond CBN wheel. International Journal of the Japan Society for Precision Engineering 1991;25(2):98–109. [10] Zhang Bingpeng, Zhao Wenxiang, Wang Xibin Relationship modeling between granularity of GC cup wheel and dressing efficiency of diamond grinding wheel. Tool Engineering 2006;40(4):31–4. [11] Hirao Masatoshi, Izawa Masaki, Iguchi Nobuaki, et al. Waterjet in-process dressing (1st Report) – dressing property and jet pressure. Journal of Japan Society for Precision Engineering 1995;61(7)1335-1339 1995;61(7). [12] Cai Lanrong, Jia Yan, Hu Dejin. Analysis of temperature field and experimental validation for electrical discharge dressing and truing of electrical diamond grinding wheel. China Mechanical Engineering 2008;19(9):1009–12. [13] Jackson MJ, Robinson NB, Khangar A, Moss R. Laser dressing of vitrified aluminum oxide grinding wheels. British Ceramic Transactions 2003;102(3): 238–245. [14] Kang RK, Yuan JT, Shi XK, Lan J, Ren JX. Study on laser dressing of superabrasive grinding wheels. China Mechanical Engineering 2000;5(11):493–6.
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