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Laser-assisted micro-milling of austenitic stainless steel X5CrNi18-10
Journal of Manufacturing Processes 48 (2019) 174–184
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Laser-assisted micro-milling of austenitic stainless steel X5CrNi18-10 a,
a
a
Mohammadali Kadivar *, Bahman Azrhoushang , Ali Zahedi , Claas Müller a b
T
b
Institute of Precision Machining (KSF), Furtwangen University of Applied Sciences, Jakob-Kienzle-Str 17, 78056 Villingen-Schwenningen, Germany Department of Microsystems Engineering — IMTEK, Laboratory for Process Technology, University of Freiburg, Georges-Köhler-Allee 103, 79110 Freiburg, Germany
A R T I C LE I N FO
A B S T R A C T
Keywords: Laser-assisted micro-milling Micro-milling Laser structuring Ultra-short pulsed laser Austenitic stainless steel
This paper presents a novel Laser-Assisted Micro-Milling (LAMM) process of austenitic stainless steel X5CrNi1810. The LAMM process is compared with the conventional micro-milling process. Ultra-short pulsed laser radiation is utilized for the structuring of the workpiece surface prior to the micro-milling process. Different laser structures are produced on the workpiece surface at a constant laser scanning speed with various laser powers and laser line spans. The high performance of the developed process is shown by experimental investigations. The effect of laser structuring on the micro-milling forces and temperature indicated the superior performance of the new LAMM process. Cutting forces and temperature could be reduced by up to 70% and 50%, respectively. The results of conventional micro-milling showed that increasing the cutting speed, at a constant undeformed chip thickness, reduced the micro-milling forces. Increasing the cutting speed from 50 to 250 m/min halved both the trust and normal forces, while it slightly improved the surface roughness. On the other hand, increasing the feed per tooth degraded the surface roughness and increased the cutting forces. Furthermore, in conventional milling the workpiece was subjected to high plastic deformation during the cutting process, while side flow, smeared material, metal debris, and cavities were observed on the workpiece surface.
Introduction High precision parts and miniaturization are required by various high-tech devices. However, the manufacturing processes are faced with new challenges owing to the ever-increasing demands of these components. Non-Mechanical micro-manufacturing processes, like lithography and Electro Discharge Machining (EDM), are mostly expensive and time-consuming. They are also only applicable to some specific sort of materials. Mechanical micro-milling is generally a costeffective process. It offers the capability to produce complex 3D-geometries and micro-structures with tight tolerances in a broad range of materials [1]. It is a commonly used machining process to manufacture micro-tools and molds with micro-features. Nevertheless, further researches are required to optimize the productivity, tool costs and quality standards of the process to increase its application in industry. The difficulties in micro-milling become more noticeable when machining difficult-to-cut materials. Their machining is normally accompanied by considerable tool wear, large cutting forces, and high temperature and undesired burr formation [2–9]. Thepsonthi and Özel [10] experimentally and analytically investigated the tool wear and burr formation in the micro-milling of Ti-6Al-4 V titanium alloy. They showed that the edge radius and feed per tooth are the two most
⁎
important factors regarding cutting forces. Increasing the cutting speed and feed rate increased the cutting temperature. The feed per tooth was the main parameter which affected the surface roughness and burr size, which is also shown by Klocke et al. [11]. They also investigated the influence of the cutting parameters on the micro-milling of steel. They concluded that there is an optimum cutting speed in the case of tool wear. This opposes the idea which assumes the tool life is inversely proportional to the cutting speed. They also showed that excessive burr formation could occur even at an optimal set of parameters. From economical aspects, it is important to reduce the tool wear and burr formation and increase the material removal rate as well as surface quality. One of the effective approaches in this regard is using advanced machining processes (AMPs) such as ultrasonic and laser as an assistant in the process [12]. Li and Wang [13] showed that ultrasonic vibrations could improve the micro-cutting performance of SKD 61 steels (hardness: HRC38) by reducing the tool wear. They used a micro-tool with a diameter of 2 mm and showed that ultrasonic-assisted micro-milling could reduce the burr formation up-to 18% while reducing the tool wear. The surface roughness could be improved noticeably as well. They found that surface quality and the aspect ratio of the generated features could be improved by using this novel method. Melkote et al. [14] developed a hybrid laser-assisted micro-milling processes with a
Corresponding author. E-mail address: [email protected] (M. Kadivar).
https://doi.org/10.1016/j.jmapro.2019.11.002 Received 30 August 2018; Received in revised form 11 September 2019; Accepted 1 November 2019 1526-6125/ © 2019 The Society of Manufacturing Engineers. Published by Elsevier Ltd. All rights reserved.
Journal of Manufacturing Processes 48 (2019) 174–184
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low power Ytterbium-doped continuous wave fiber laser. They chose a hardened A2 tool steel (62 HRc) as the workpiece material and used a TiAlN-coated tungsten carbide 4-flute ball end mills of 250 μm diameter for their experiments. They showed that using laser heating could improve the dimensional accuracy, the rate of tool wear as well as the surface roughness. Ding et al. [15] modeled the LAMM process numerically. They concluded that the workpiece temperature increases in the case of LAMM and the workpiece flow stress drops by about 20–25%. They found a uniform specific cutting energy in the cutting depth of 250 μm. The simulation model also showed that the stable build-up edge can be eliminated at a large ratio of undeformed chip thickness to cutting edge radius or reduced it at a higher ratio during LAMM when the proper continues laser parameters are chosen. They used the LAMM simulation result to empirically model the tool wear of tungsten carbide tools with diameters of 100–300 μm in fine side microcutting of 422SS and the results were compared with the real tool wear measurements. The finite element results were in good agreement with the tool wear measurements. The model showed that the LAMM could reduce BUE with proper heating of the work material prior to micromachining. Kumar and Melkote [16] compared the LAMM with An Ytterbium-doped continuous-wave infra-red fiber laser with Conventional Micro-Milling (CMM) of a hardened A2 tool steel (62HRC). A two-flute, 400 μm diameter, TiAlN coated tungsten carbide micro-milling tool was used for the experiments. The results showed that the cutting forces (up to 60%), tool wear, burr formation, and surface roughness are lower in the case of LAMM, while the material removal rate was about six times larger than the tool manufacturer’s recommendations. Özel and Pfefferkorn [17] used the LAMM for micromilling of AISI 4340 steel with a two-flute carbide endmill with a diameter of 635 μm. They showed that the Pulsed LAMM reduces the cutting forces and rises the temperature in the cutting tool. Apart from a few research such as the one carried out by Kumar et al. [15], where the laser was used mainly to induce thermal cracks on the workpiece surface, the material removal mechanism in the conventional LAMM processes is generally based on a local and time-dependent softening of the workpiece material. Pre-heating of a specific area over the workpiece surface is the main mechanism behind the previously performed LAMM investigations [14–17] (Fig. 1a). The preheating leads to the reduction of the material stiffness and strength, which leads to material softening along the cutting path (as shown in Fig. 1a). Furthermore, only continuous wave or short-pulsed lasers are utilized. This method has several disadvantages. Its main drawback is the induced Heat Affected Zone (HAZ) on the workpiece surface and the difficulty of controlling undesired changes of material properties, which could remain on the finished surface even after the micro-milling process [18,19]. These processes utilize a focused laser beam in front of the milling tool (geometrical constraint), as shown in Fig. 1a, and are only applicable in dry conditions, where using coolant in the process is not
possible [20,21]. The presence of coolant causes an immediate cooling of the laser affected and softened area and makes the process inefficient. Therefore, the tool temperature is generally higher within these classes of LAMM compared to the conventional micro-milling [19]. To reduce the cutting forces and temperature a novel LAMM process, developed by the Institute of Precision Machining (KSF) [20,21], is utilized for the first time for micro-milling of X5CrNi18-10. Furthermore, Controlling the depth of thermal damages with respect to the laser ablation parameters in processing with ultra-short pulsed lasers has not been precisely addressed in the micro-scale till now. The proposed method is a two-stage (structuring and micro-milling) laser-assisted machining process, which can be performed also in wet conditions. In this method, an ultra-short pulsed laser has been used to structure the workpiece surface (stage 1), instead of using a continuous or short-pulsed laser to pre-heat a specific area of the workpiece surface. Afterward, the micro-milling tool was used to remove the structured area to the final dimensions (Fig. 1-b). Utilizing ultra-short pulsed laser induces a neglectable and simultaneously controllable HAZ on the workpiece surface [22] which is the main advantage over the previous LAMM methods. Additionally, the process can be utilized in wet conditions with no restriction on the application of machining fluids. A laser-structured workpiece sample can be successively machined in wet conditions, where the structured zone will be completely removed (stage 2) via chip formation in presence or absence of coolant or lubricant. The workpiece can be structured up to a certain depth according to the laser parameters, and finally be machined with different cutting conditions. Therefore, the machining parameters are completely independent of the laser parameters, and laser-structuring and micromilling processes are completely time-independent. Accordingly, only the depth of laser-induced features is matched with the milling depth of cut. To this end, several laser-generated structures have been induced over the workpiece surface and milling forces and cutting temperatures have been compared in LAMM and CMM process. Experimental procedure Austenitic stainless steel X5CrNi18-10 was chosen as the workpiece material. It has favorable properties such as high corrosion resistance, high strength, and ductility. These properties make it a suitable choice for a wide field of applications like the medicine sector, automotive industry or high precision measurement equipment. However, the machining of this material is challenging. One of the main challenges is burr formation owing to its high toughness. Moreover, it has a high tendency to work hardening and build-up edge formation during machining [23]. The samples were prepared in a block form with dimensions of 30 × 20 × 10 mm. The micro-milling tests were carried out at different cutting speeds (vc) and feed rate (fz). The influence of micro-
Fig. 1. Schematic view of LAMM with the help of (a): softened area (b): structured area. 175
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length to evaluation length, λc/L, of 0.8/4 mm. The average of these 3 measurements was chosen as the value of surface roughness. Fig. 4 shows the procedure and a measurement sample of the surface roughness. Fig. 2a shows the experimental LAMM setup. A high-speed thermography camera (Infratec, ImageIR 8300) was used to measure the process temperature as shown in Fig. 2c. Since the temperature measurement via thermography camera is not possible under wet conditions, the micro-milling tests were performed in dry conditions. However, the utilized laser-structuring approach allows the application of cutting fluids during LAMM. In this study, a two-flute tungsten carbide micro-endmill with a diameter of 500 μm and a helical angle of 30 degrees from Crazy Mill- Mikron Tool was used in a side micro-milling operation (Table 1). Fig. 5a shows the schematic of the side micromilling. Fig. 5 shows the experimental setup for the laser structuring of the workpiece surface consisting of a laser scanner, workpiece, and an air nozzle. The laser scanner focuses the laser beam over the workpiece surface and the air nozzle is responsible for sweeping the ablated debris away. The process parameters are listed in Table 1. Several laser patterns were produced on the workpiece surface which are listed in Table 2.
milling parameters on the cutting forces and the surface roughness were studied while using conventional micro-milling process under wet conditions (utilizing oil as coolant lubricant). Subsequently, the workpiece was laser-structured using a Yb:YAG picosecond laser (TruMicro5050 from Trumpf) with different laser parameters. Each test was repeated three times and the average values have been reported in the paper. One of the challenges during the structuring is heat dissipation. In using a picosecond laser, pulse energy in the microjoule range is radiated in a very short time period over the workpiece surface. It was already shown in authors previous research [22] that, as the ultra-short pulse length of the laser beam the ratio of the pulse length (10 ps) to the pulse radiation period (1/400 kHz = 2.5 μs) is very small (the relative relaxation time is so large), the imposed heat of a single pulse is dissipated before the next pulse. Therefore, the heat accumulation owing to successive pulsed is negligible. Assuming the extreme case, that the whole laser energy, which is delivered by the removed debris and partially remained inside the workpiece material is absorbed by the workpiece, the overall laser heat input to the workpiece corresponds to a heat source with a few Watts of average power. The assigned pressurized air nozzle within the experiments serves to remove the debris from the laser structuring and compensates the heat accumulation up to this moderate laser power level. A high-precision 5-axis CNC machining center (KERN Pyramid Nano) was used to perform the micro-milling tests (Fig. 2a). The forces were measured by a Kistler dynamometer (type 9256C2) and processed with Kistler 5015 charge amplifiers and recorded by a PC acquisition board and LabView software. The dynamometer has suitable dynamics and sensitivity for micro-machining processes. The presented forces are peak-to-valley forces. A sample of force measurement in the time domain for both tangential and normal forces is presented in Fig. 3. To calculate the micro-milling forces, the forces from each micro-milling test were divided into three parts in the time domain of 6 × 10-3 s including the highest and lowest peaks and valleys. From each time-domain three peak-to-valley forces were measured, and the mean value was reported as the micro-milling force. A confocal microscope (μsurf mobile plus) was integrated into the machine to measure the surface roughness (Fig. 2b). The surface roughness was taken perpendicular to the micro-milling path at three positions (the beginning, the middle, and the end of the cut) with the sampling length of 4.8 mm. In each position several lines perpendicular to the cutting direction, like the black line in Fig. 4b were drawn to achieve a total length of 5 mm shown in Fig. 4c. The surface roughness was finally calculated through these connected lines with cut-off length, λc, of 0.8 mm and sample
Results and discussions The effect of cutting speed on the micro-milling process Fig. 6 shows the effect of cutting speed on the micro-milling forces and surface roughness. Increasing the cutting speed decreased both the cutting forces and the surface roughness. A 70% reduction in both tangential and normal forces could be achieved by increasing the cutting speed from 50 to 250 m/min. High cutting speeds generally lead to high process temperature. Additionally, the temperature gradient in the workpiece in front of the cutting edge in micro-milling is proportional to the cutting speed. This increment is due to the induced material flow by the micro-tool [24]. High cutting temperatures at high cutting speeds ease the material flow and material removal, causing better cutting performance and lower cutting forces. Another reason for the reduction of cutting forces is an increase in the strain-rate, which is also proportional to cutting speed. Eleiche [25] showed that strain-rate has a great effect on the flow stress. The sensitivity of the flow stress to the strain-rate can be categorized by three regions with different deformation mechanisms. When the strain-rate is low, the dominant phenomenon governing the material removal mechanism and flow stress is the dislocation density, which is barely influenced by strain-
Fig. 2. (a): The experimental setup; (b): Integrated confocal microscope, (c): The temperature measurement. 176
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Fig. 3. The cutting forces sample in time domain (vc = 50 m/min, ae = 50 μm, and fz = 4 μm/tooth). Table 1 Process parameters. Parameters
Values
Milling tool Workpiece Cutting speed (vc) Feed per tooth (fz) Radial depth of cut (ae) Axial depth of cut (ap) Coolant Laser power (PL) Laser pattern Laser scanning speed (vL) Pattern distance (L)
2.CMC30.A3Z2.050.1 - Crazy Mill X5CrNi18-10 (1.4301) 50,150, and 150 m/min 1, 3, 4, and 6 μm 50 and 60 μm 500 μm Oil and dry 7.5 and 10 W X 100 mm/s 250 and 500 μm
Merchant’s equation:
∅=
β π α + − 4 2 2
β =tan−1 (
Fn + Ft tanα ) Ft − Fn tanα
(2)
(3)
where β is the friction angle and α is the clearance angle. However, in the case of high-speed cutting, the flow stress is dominated by the effects of strain-rate. Therefore, in high-speed cutting the theory of the strain gradient plasticity is not valid [26]. Thus, increasing the cutting speed causes more average strain-rate ε¯n [27], (following equation), affecting the flow stress and material removal mechanism:
ε¯n =
rate. The flow stress δ, in this regime, can be expressed according to the cutting and thrust forces using the following equation [25]:
(Ft sin∅+ Fn cos∅) sin∅ h d ap
(4)
On the other hand, the yield strength of X5CrNi18-10 stainless steel increases slightly with rising the strain-rate. This increment is due to the strain hardening of the material in terms of a transition from isothermal (low strain-rate) to adiabatic process (high strain-rate) [28]. Moreover, at higher strain-rate the accumulation of the generated heat in the workpiece is more considerable. The plastic deformation within a very short deformation time during micro-milling at high cutting speeds causes higher Stacking Fault Energy (SFE). The higher SFE impedes martensitic transformation. Therefore, the strain hardening is dropped with rising the strain and strain-rate [28]. The aforementioned factors ease the cutting process and lower the cutting forces at higher cutting speeds. Fig. 6-b shows that increasing the cutting speed leads to a slight reduction of surface roughness, despite the uncut chip thickness was kept constant. Reduction of roughness values with increasing the cutting speed could be due to lower built-up edge formation over the
Fig. 4. The surface roughness measurement procedure (a): schematic of the workpiece; (b): surface topography with a confocal microscope (vc = 250 m/ min, ae = 50 μm, and fz = 1 μm/tooth), (c): The surface roughness profile extracted from several parallel lines in b.
δ=
vc hd
(1)
where Ft and Fn are the tangential and normal forces, respectively. hd is the undeformed chip thickness, ap indicates the cutting depth. ∅ is the prediction of shear angle which can be determined according to 177
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Fig. 5. (a): Schematic of LAMM process; (b): Produced patterns on the workpiece surface via laser; (c): The setup for the laser structuring of the workpiece. Table 2 Laser patterns. Structures
Laser parameters
Non-structured Pattern 1 Pattern 2 Pattern 3 Pattern 4
off PL = 10 W; L = 500 μm PL = 10 W; L = 250 μm PL = 7.5 W; L = 500 μm PL = 7.5 W; L = 250 μm
various feed per tooth (0.001; 0.003; 0.004 and 0.006 mm) were conducted. It can be stated from Fig. 7 that, the lower feed rate (and lower average chip thickness) causes lower forces, as well as better surface, roughens as it was expected. Higher material removal rates can be achieved at higher feed per tooth rates, resulting in larger milling forces. Li et al. [30] showed that the minimum chip thickness is a dominant factor regarding the surface generation in micro-milling process. The uncut chip thickness is increased by keeping a constant depth of cut while increasing the feed rate (Eq. (5)). In micro-milling process, under definite cutting conditions, the maximum uncut chip thickness can be smaller than the edge radius of the micro-milling tool. Thus, finding the minimum chip thickness hm or critical chip thickness seems essential, under which workpiece material may not be properly removed as chips. This condition causes lower share of chip formation and higher plowing and rubbing in the micro-milling process. With
cutting edge [11]. Weule et al. [29] showed that the built-up edge worsens the surface quality in the micro-cutting process which is already known from macro-cutting. Increased cutting forces cause higher tool wear, larger tool deflection and more considerable built-up-edge [29]. The effect of undeformed chip thickness on micro-milling process The effect of undeformed chip thickness, hd, on the cutting forces and surface roughens was also investigated (Fig. 7). The average chip thickness can be calculated from the following equation:
a hav = fz ⎛ e ⎞ ⎝ d ⎠
(5)
where d is the diameter of the micro-milling tool. Experiments with
Fig. 6. The effect of cutting speed on (a) micro-milling forces, (b) surface roughness. 178
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Fig. 7. The effect of average chip thickness on (a) micro-milling forces, (b) surface roughness.
uncut chip thicknesses below this value, the surface roughness values may fluctuate and in some cases be improved as shown by Li et al. [30]. The minimum chip thickness is determined by the following equation [31]:
β π hm = re ⎛1 − cos( − )⎞ 4 2 ⎝ ⎠
Table 4 ANOVA for Ra.
(6)
where β is the friction angle between the tool and the un-cut workpiece passing under the tool and re is the tool edge radius. According to Eq. (6), the minimum chip thickness can be changed by changing the friction coefficient and tool edge radius. Brinksmeier et al [32] found that the friction coefficient while micro-milling austenitic stainless steel with tungsten carbide tool is about 0.258. The tool edge radius was also measured around 2 μm using a confocal microscope. Accordingly, the minimum chip thickness for this material can be derived from Eq. (6) which is 0.415 μm. As it is shown in Fig. 7-b increasing the chip thickness steadily increases the value of the surface roughness. It implies that the chip thickness is larger than the critical chip thickness.
Factors
DF
Sum of squares (SS)
Mean squares (MS)
F ratio α = 5%
P
Contribution (%)
vc (m/min) fz (μm) Error Total
2 3 12 17
0.003735 0.002841 0.000246
0.001868 0.000947 0.000021
91.01 46.14
0.000 0.000
43.19% 52.28% 4.53% 100%
Analysis of variance (ANOVA) To determine the statistically significant factors and their contribution to the micro-milling forces and surface roughness, analysis of variance (ANOVA) was used. The results are presented in Tables 3 and 4 for the tangential micro-milling force and surface roughness Ra, respectively. The ANOVA analysis was conducted at the level of confidence 95%. The results showed that the feed rate has the largest influence on the micro-milling force with 80% compared to the cutting speed with 16% contribution. However, the most influential factor for the surface roughness is the cutting speed, though the influence of the feed rate with 43% contribution is not neglectable. Both factors, cutting speed and feed rate, had a statistical and physical significance on the cutting force and surface roughness since P value was lower than 0.05.
Fig. 8. different views by SEM.
The cutting mechanism The machined surface was observed under different views (Fig. 8) with SEM microscopy. Fig. 9 shows the machined surface from the top view. The surface was generated with the front end of the cutting tool. The micro-milling feed pattern is clearly observed in Fig. 9-c. The black spots over the surface indicate the cavities, while the regions with white borders indicate the smeared materials. The smeared regions are resulted from a high amount of plastic deformation of the material, during the last chip formation phase in the contact zone. In down milling, the chip thickness approaches zero at the end of the chipping process (the end of engagement arc). In this position, the actual rake angle becomes highly negative because of the cutting-edge radius. When the rake angle is negative, the chips break at a certain position in the cutting path. The remaining material on the cutting path is deformed and smeared by the cutting edge over the surface of the workpiece, rather than being cut as chips. Cavities, metal debris, and feed marks, as well as, smeared material particles could be observed on the micro-milled wall as shown in
Table 3 ANOVA for P-to-V Ft. Factors
DF
Sum of squares (SS)
Mean squares (MS)
F ratio α = 5%
P
Contribution (%)
vc (m/min) fz (μm) Error Total
2 3 12 17
0.25460 1.27230 0.06360
0.127300 0.424100 0.005300
24.02 80.02
0.000 0.000
15.96% 80.04% 4.00% 100.00%
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Fig. 9. SEM Images from the milled surface from top view (view 1in Fig. 6) (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth).
plastification tends to flow through the cutting-edge toward the secondary cutting-edge and finally adheres over the machined surface. There are two mechanisms for the material side flow. The first, the material is squeezed between the flank face of the cutting tool and the finished surface. The second, the material which is encountered by the
Fig. 10. The white line in Fig. 10-c shows the feed mark over the surface of the workpiece, which led to material side flow. The material side flow occurs when the chipping material in the cutting-edge experiences severe stress and thermal loads. These conditions could lead to considerable plastification over the machined surface of the material. This
Fig. 10. SEM images from the milled surface from view 2 (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth). 180
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Fig. 11. SEM images from the milled surface from view 3 (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth).
plastification in the cutting zone flows through the worn trailing edge to the side of the cutting tool [33]. The main reason is the high temperature during the cutting process. Moreover, some material debris was observed over the milled surface (Fig. 10-b). The produced chips during the micro-milling can be welded to the machined surface owing to the high pressure and temperature in the process, which in turn can deteriorate the surface quality of the workpiece. Fig. 9–11 indicates that the material is exposed to high plastic deformation during the micro-milling process. As it is known, during the cutting process the material is extruded to a large extent. The extruded material is pushed to the side and causes burr formation which is marked as plastic deformation in Fig. 11. A large amount of metal debris is also observed after the cutting process in the corner of the machined area, where the corner of the tool formed the cut. Compared to the other areas of the tool, the tool radius (at tool corner) induces much higher cutting temperature and forces, since the corresponding contact length is larger. Higher temperature and forces lead to a larger amount of metal debris in the corner of the cut. Moreover, the chips on the micro-milling tool corner cannot be easily evacuated so they could be welded to the machined surface. The effect of laser structuring on micro-milling process Fig. 12. The effect of laser structuring on the cutting forces in the time domain (vc = 50 m/min, ae = 60 μm, and fz = 6 μm/tooth).
An ultra-short pulsed picosecond laser is utilized to produce specific patterns on the workpiece surface (Table 2). After laser-structuring the structured and non-structured workpieces were micro-milled with cutting speed of 50 m/min, feed per tooth of 6 μm, and depth of cut of 60 μm to generate higher cutting forces and temperature in the process. Since the depth of laser-cut was approximately three times larger than the depth of cut (60 μm), three micro-milling passes were performed to remove the structured area, and the mean values were reported as the results of temperature and cutting forces. The results of the LAMM cutting forces and temperature are compared with the conventional micro-milling and are presented in Figs. 12–14. The tests were performed under dry conditions. The temperature profile presented in Fig. 14 was measured on the machined surface (on the area exactly behind the milling tool). Measuring the cutting contact temperature was not possible since the thermo-camera requires a direct vision over the measured surface area. As Figs. 13 and 14 show, LAMM led to a significant reduction of cutting forces and temperature. Fig.12 presents the cutting forces in a time domain for both CMM and LAMM processes. In CMM process both tangential and normal forces fluctuate around an average value. However, in LAMM process forces do not follow the same trend. This deviation could be due to melted material regions or non-uniform lasergenerated grooves over the surface of the workpiece. Laser-structuring removes material from the surface of the
workpiece and, therefore, the tool-workpiece contact area becomes smaller. Hence, the smaller material volume is removed during the micro-milling process subsequent to the laser structuring, compared to the micro-milling of non-structured surfaces. Additionally, the process becomes more intermittent than conventional micro-milling. Lower contact length and material volume to be cut lead to lower micro-milling forces and temperature. Another reason for the reduction of micromilling forces and temperature lies in the chip morphology. As it is clear in Fig. 15, the LAMM produces shorter, thinner, and segmented chips compared to the CMM, which leads consequently to lower micro-milling forces. Pattern 2 with the laser power of PL = 10 W and the pattern distance of L = 250 μm generated the smallest milling forces and temperature. Moreover, lower laser power leads to larger cutting forces. Increasing the laser power increases the input laser energy density according to the following equation [21]:
EL − input =
PL (dL + wL) vL
(7)
Where, wL is the laser line width, dL indicates the laser line span, and the laser scanning speed is indicated by vL. The amount of material 181
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Fig. 13. The effect of laser structuring on the cutting forces.
part. However, in the case of LAMM material debris can be observed more frequently. The white lines in Fig. 15a are determined as the feed mark over the finished surface, which led to material side flow. These feed marks are not observed over the finished surface via LAMM process (Fig. 15a and b). In both cases some cavities can be also seen over the workpiece surface. It could be implied that lower micro-milling forces in addition to lower temperatures in the case of LAMM process caused lower plastic deformation during the cutting process.
removal during laser-structuring can be increased by increasing the laser input energy. Compared to the conventional micro-milling, tangential and normal cutting forces could be reduced by up to 70% and 40%, respectively, via the LAMM. Fig. 14 shows that using higher laser power resulted in a lower cutting temperature during micro-milling. The cutting temperature of LAMM in the case of PL = 7.5 W is almost similar to the conventional micro-milling. The laser-structuring is accompanied by the ablation process. After laser-structuring the ablated material can be deposited and accumulated into the lasered slots. This phenomenon was observed mainly at lower laser powers and caused more cutting temperature in the subsequent micro-milling process. It is also demonstrated that higher laser powers enhance the crack initiation and propagation around the laser structures, owing to the thermal shocks induced by ultra-short pulsed lasers [29]. These cracks can also lead to lower cutting forces and temperature since they reduce material strength and stiffness.
Conclusion A novel and promising laser-assisted micro-milling process utilizing an ultra-short pulsed laser was developed in this study. Different patterns with different laser powers and pattern distances were produced on the surface of the workpiece using a picosecond laser prior to micromilling. The effect of micro-structuring, as well as cutting parameters on the micro-milling of an annealed alloy steel was investigated. The results showed that the cutting speed had a great influence on the cutting forces; however, its influence on the surface roughness was negligible. Increasing the cutting speed form 50 m/min to 250 m/min could reduce the cutting force up to 50%. However, 5 times higher cutting speed just improved the surface quality up to 15%. As it was expected, increasing the feed per tooth in micro-milling increased both the milling forces and surface roughness values. Increasing the material
The effect of laser structuring on material characterization Fig. 16 shows the material characterization of the micro-milled material via LAMM-pattern 2 and CMM. The finished surface of the conventional micro-milling process reveals large material deformation in the form of smeared material over the surface of the micro-milled
Fig. 14. The effect of laser structuring on the cutting temperature. 182
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Fig. 15. Chip morphology (vc = 50 m/min, fz = 6 μm, and ae = 60 μm), (a): CMM; (b): LAMM.
the conventional micro-milling process. This can be explained via smaller contact length and smaller material volume which should be removed by the micro-milling tool. The chip morphology of the LAMM process showed that the chips are smaller and thinner compared to the chips in conventional micro-milling, which also led to lower milling forces. This can be used as an effective and new manufacturing technique to increase the accuracy of the parts, enable larger material removal rates and cause lower tool wear because of lower cutting forces and temperature. Increasing the laser input energy density, increased the material removal rate during the laser-structuring and enhanced the LAMM process by reducing the milling forces and temperature.
removal rate up to 6 times, by increasing the feed per tooth from 1 to 6 μm, increased the micro-milling forces up to 3 times. However, surface roughness did not change significantly. During the micro-milling process, the workpiece was subjected to considerable plastic deformation in form of burr in the edge of workpiece and smeared material. Moreover, several failures such as side flow, smeared material, metal debris, and cavities were observed over the machined surface after micro-milling process. Workpiece structuring using an ultra-short pulsed laser prior to the micro-milling process led to a meaningful reduction of normal and tangential micro-milling forces (70 and 50%, respectively) compared to
Fig. 16. SEM images from the milled surface (vc = 50 m/min, fz = 6 μm, and ae = 60 μm), (a): CMM; (b): LAMM pattern 2, (c): CMM detailed, (d): LAMM pattern 2 detailed. 183
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Furthermore, the smaller distance of the patterns over the surface of the workpiece caused lower milling forces and temperature. The best performance of the laser-structuring was obtained using the laser power of 10 W and the pattern distance of 250 μm. Finding suitable laser parameters is a key factor to have an efficient LAMM process, especially regarding the cutting temperature. Using improper laser parameters may cause an inefficient LAMM process. It was shown that the cutting temperature did not change significantly in LAMM compared to the conventional micro-milling in the case that the laser power was set to 7.5 W. However, increasing the laser power to 10 W could reduce the cutting temperature in LAMM up to 50%. It is worth mentioning that laser-structuring may also help the lubrication and cooling effect during the machining. Hence, future investigations are planned to study the effects of LAMM under wet conditions.
tool wear. J Mater Process Technol 2013;213(4):532–42. [11] Klocke F, Quito F, Arntz K, Souza A. A study of the influence of cutting parameters on micromilling of steel with cubic boron nitride (CBN) tools. In: Maher M, Chiao J, Resnick PJ, editors. SPIE MOEMS-MEMS: micro- and nanofabrication. SPIE; 2009. p. 72040D. [12] Dubey AK, Yadava V. Laser beam machining—a review. Int J Mach Tools Manuf 2008;48(6):609–28. [13] Li K, Wang S. Effect of tool wear in ultrasonic vibration-assisted micro-milling. Proc Inst Mech Eng Part B J Eng Manuf 2013;228(6):847–55. [14] Melkote S, Kumar M, Hashimoto F, Lahoti G. Laser assisted micro-milling of hardto-machine materials. CIRP Ann Manuf Technol 2009;58(1):45–8. [15] Ding H, Shen N, Shin YC. Thermal and mechanical modeling analysis of laser-assisted micro-milling of difficult-to-machine alloys. J Mater Process Technol 2012;212(3):601–13. [16] Kumar M, Melkote SN. Process capability study of laser assisted micro milling of a hard-to-machine material. J Manuf Process 2012;14(1):41–51. [17] Özel T, Pfefferkorn F. Pulsed laser assisted micromilling for Die/Mold manufacturing. ASME 2007 International Manufacturing Science and Engineering Conference 2007:337–42. [18] Chryssolouris G, Anifantis N, Karagiannis S. Laser assisted machining: an overview. J Manuf Sci Eng 1997;119(4B):766. [19] Lei S, Shin YC, Incropera FP. Experimental investigation of thermo-mechanical characteristics in laser-assisted machining of silicon nitride ceramics. J Manuf Sci Eng 2001;123(4):639. [20] Azarhoushang B, Soltani B, Daneshi A. Study of the effects of laser micro structuring on grinding of silicon nitride ceramics. CIRP Ann Manuf Technol 2018;67(1):329–32. [21] Azarhoushang B, Soltani B, Zahedi A. Laser-assisted grinding of silicon nitride by picosecond laser. Int J Adv Manuf Technol 2017;93(5–8):2517–29. [22] Zahedi A, Tawakoli T, Azarhoushang B, Akbari J. Picosecond laser treatment of metal-bonded CBN and diamond superabrasive surfaces. Int J Adv Manuf Technol 2015;76(5–8):1479–91. [23] Biermann D, Steiner M. Analysis of Micro burr formation in austenitic stainless steel X5CrNi18-10. Procedia Cirp 2012;3:97–102. [24] Moriwaki T, Sugimura N, Luan S. Combined stress, Material flow and heat analysis of orthogonal micromachining of copper. CIRP Ann Manuf Technol 1993;42(1):75–8. [25] Eleiche AM, Campbell JD. Strain-rate effects during reverse torsional shear. Exp Mech 1976;16(8):281–90. [26] Wu J, Liu Z. Modeling of flow stress in orthogonal micro-cutting process based on strain gradient plasticity theory. Int J Adv Manuf Technol 2010;46(1-4):143–9. [27] Özel T, Altan T. Determination of workpiece flow stress and friction at the chip–tool contact for high-speed cutting. Int J Mach Tools Manuf 2000;40(1):133–52. [28] Krüger L, Wolf S, Martin U, Martin S, Scheller PR, Jahn A, et al. The influence of martensitic transformation on mechanical properties of cast high alloyed CrMnNisteel under various strain rates and temperatures. J Phys Conf Ser 2010;240:12098. [29] Weule H, Hüntrup V, Tritschler H. Micro-cutting of steel to meet new requirements in miniaturization. CIRP Ann Manuf Technol 2001;50(1):61–4. [30] Li H, Lai X, Li C, Feng J, Ni J. Modelling and experimental analysis of the effects of tool wear, minimum chip thickness and micro tool geometry on the surface roughness in micro-end-milling. J Micromech Microeng 2008;18(2):25006. [31] Son SM, Lim HS, Ahn JH. Effects of the friction coefficient on the minimum cutting thickness in micro cutting. Int J Mach Tools Manuf 2005;45(4–5):529–35. [32] Brinksmeier E, Riemer O, Twardy S. Tribological behavior of micro structured surfaces for micro forming tools. Int J Mach Tools Manuf 2010;50(4):425–30. [33] Kishawy H, Elbestawi M. Effects of process parameters on material side flow during hard turning. Int J Mach Tools Manuf 1999;39(7):1017–30.
Acknowledgment The authors would like to gratefully acknowledge the financial support provided for this study by the Forschungsallianz Oberrhein, University of Freiburg, Germany (reference number 33-7533-30). References [1] Hoyle R. Developments in micro and nano engineering and manufacturing. Plast Rubber Compos 2013;37(2–4):50–6. [2] Zhu K, Wong YS, Hong GS. Multi-category micro-milling tool wear monitoring with continuous hidden Markov models. Mech Syst Signal Process 2009;23(2):547–60. [3] Vogler MP, Kapoor SG, DeVor RE. On the modeling and analysis of machining performance in micro-endmilling, part II: cutting force prediction. J Manuf Sci Eng 2004;126(4):695. [4] Vogler MP, DeVor RE, Kapoor SG. On the modeling and analysis of machining performance in micro-endmilling, part I: surface generation. J Manuf Sci Eng 2004;126(4):685. [5] Ucun İ, Aslantas K, Bedir F. An experimental investigation of the effect of coating material on tool wear in micro milling of Inconel 718 super alloy. Wear 2013;300(1–2):8–19. [6] Schueler GM, Engmann J, Marx T, Haberland R, Aurich JC. Burr formation and surface characteristics in micro-end milling of titanium alloys. In: Aurich JC, Dornfeld D, editors. Burrs — analysis, control and removal. Berlin, Heidelberg: Springer Berlin Heidelberg; 2010. p. 129–38. [7] Lu D, Li JF, Rong YM, Sun J, Zhou J, Wang ZQ. A study of exit-burr formation mechanism using the finite element method in micro-cutting of aluminum alloy. KEM 2008;375–376:470–3. [8] Lee K, Stirn B, Dornfeld DA. Burr formation in micro-machining aluminum, 6061T6. In: Inasaki I, editor. Initiatives of precision engineering at the beginning of a millennium. Boston: Kluwer Academic Publishers; 2002. p. 47–51. [9] Inasaki I, editor. Initiatives of precision engineering at the beginning of a millennium. Boston: Kluwer Academic Publishers; 2002. [10] Thepsonthi T, Özel T. Experimental and finite element simulation based investigations on micro-milling Ti-6Al-4V titanium alloy: effects of cBN coating on
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Laser-assisted micro-milling of austenitic stainless steel X5CrNi18-10 a,
a
a
Mohammadali Kadivar *, Bahman Azrhoushang , Ali Zahedi , Claas Müller a b
T
b
Institute of Precision Machining (KSF), Furtwangen University of Applied Sciences, Jakob-Kienzle-Str 17, 78056 Villingen-Schwenningen, Germany Department of Microsystems Engineering — IMTEK, Laboratory for Process Technology, University of Freiburg, Georges-Köhler-Allee 103, 79110 Freiburg, Germany
A R T I C LE I N FO
A B S T R A C T
Keywords: Laser-assisted micro-milling Micro-milling Laser structuring Ultra-short pulsed laser Austenitic stainless steel
This paper presents a novel Laser-Assisted Micro-Milling (LAMM) process of austenitic stainless steel X5CrNi1810. The LAMM process is compared with the conventional micro-milling process. Ultra-short pulsed laser radiation is utilized for the structuring of the workpiece surface prior to the micro-milling process. Different laser structures are produced on the workpiece surface at a constant laser scanning speed with various laser powers and laser line spans. The high performance of the developed process is shown by experimental investigations. The effect of laser structuring on the micro-milling forces and temperature indicated the superior performance of the new LAMM process. Cutting forces and temperature could be reduced by up to 70% and 50%, respectively. The results of conventional micro-milling showed that increasing the cutting speed, at a constant undeformed chip thickness, reduced the micro-milling forces. Increasing the cutting speed from 50 to 250 m/min halved both the trust and normal forces, while it slightly improved the surface roughness. On the other hand, increasing the feed per tooth degraded the surface roughness and increased the cutting forces. Furthermore, in conventional milling the workpiece was subjected to high plastic deformation during the cutting process, while side flow, smeared material, metal debris, and cavities were observed on the workpiece surface.
Introduction High precision parts and miniaturization are required by various high-tech devices. However, the manufacturing processes are faced with new challenges owing to the ever-increasing demands of these components. Non-Mechanical micro-manufacturing processes, like lithography and Electro Discharge Machining (EDM), are mostly expensive and time-consuming. They are also only applicable to some specific sort of materials. Mechanical micro-milling is generally a costeffective process. It offers the capability to produce complex 3D-geometries and micro-structures with tight tolerances in a broad range of materials [1]. It is a commonly used machining process to manufacture micro-tools and molds with micro-features. Nevertheless, further researches are required to optimize the productivity, tool costs and quality standards of the process to increase its application in industry. The difficulties in micro-milling become more noticeable when machining difficult-to-cut materials. Their machining is normally accompanied by considerable tool wear, large cutting forces, and high temperature and undesired burr formation [2–9]. Thepsonthi and Özel [10] experimentally and analytically investigated the tool wear and burr formation in the micro-milling of Ti-6Al-4 V titanium alloy. They showed that the edge radius and feed per tooth are the two most
⁎
important factors regarding cutting forces. Increasing the cutting speed and feed rate increased the cutting temperature. The feed per tooth was the main parameter which affected the surface roughness and burr size, which is also shown by Klocke et al. [11]. They also investigated the influence of the cutting parameters on the micro-milling of steel. They concluded that there is an optimum cutting speed in the case of tool wear. This opposes the idea which assumes the tool life is inversely proportional to the cutting speed. They also showed that excessive burr formation could occur even at an optimal set of parameters. From economical aspects, it is important to reduce the tool wear and burr formation and increase the material removal rate as well as surface quality. One of the effective approaches in this regard is using advanced machining processes (AMPs) such as ultrasonic and laser as an assistant in the process [12]. Li and Wang [13] showed that ultrasonic vibrations could improve the micro-cutting performance of SKD 61 steels (hardness: HRC38) by reducing the tool wear. They used a micro-tool with a diameter of 2 mm and showed that ultrasonic-assisted micro-milling could reduce the burr formation up-to 18% while reducing the tool wear. The surface roughness could be improved noticeably as well. They found that surface quality and the aspect ratio of the generated features could be improved by using this novel method. Melkote et al. [14] developed a hybrid laser-assisted micro-milling processes with a
Corresponding author. E-mail address: [email protected] (M. Kadivar).
https://doi.org/10.1016/j.jmapro.2019.11.002 Received 30 August 2018; Received in revised form 11 September 2019; Accepted 1 November 2019 1526-6125/ © 2019 The Society of Manufacturing Engineers. Published by Elsevier Ltd. All rights reserved.
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low power Ytterbium-doped continuous wave fiber laser. They chose a hardened A2 tool steel (62 HRc) as the workpiece material and used a TiAlN-coated tungsten carbide 4-flute ball end mills of 250 μm diameter for their experiments. They showed that using laser heating could improve the dimensional accuracy, the rate of tool wear as well as the surface roughness. Ding et al. [15] modeled the LAMM process numerically. They concluded that the workpiece temperature increases in the case of LAMM and the workpiece flow stress drops by about 20–25%. They found a uniform specific cutting energy in the cutting depth of 250 μm. The simulation model also showed that the stable build-up edge can be eliminated at a large ratio of undeformed chip thickness to cutting edge radius or reduced it at a higher ratio during LAMM when the proper continues laser parameters are chosen. They used the LAMM simulation result to empirically model the tool wear of tungsten carbide tools with diameters of 100–300 μm in fine side microcutting of 422SS and the results were compared with the real tool wear measurements. The finite element results were in good agreement with the tool wear measurements. The model showed that the LAMM could reduce BUE with proper heating of the work material prior to micromachining. Kumar and Melkote [16] compared the LAMM with An Ytterbium-doped continuous-wave infra-red fiber laser with Conventional Micro-Milling (CMM) of a hardened A2 tool steel (62HRC). A two-flute, 400 μm diameter, TiAlN coated tungsten carbide micro-milling tool was used for the experiments. The results showed that the cutting forces (up to 60%), tool wear, burr formation, and surface roughness are lower in the case of LAMM, while the material removal rate was about six times larger than the tool manufacturer’s recommendations. Özel and Pfefferkorn [17] used the LAMM for micromilling of AISI 4340 steel with a two-flute carbide endmill with a diameter of 635 μm. They showed that the Pulsed LAMM reduces the cutting forces and rises the temperature in the cutting tool. Apart from a few research such as the one carried out by Kumar et al. [15], where the laser was used mainly to induce thermal cracks on the workpiece surface, the material removal mechanism in the conventional LAMM processes is generally based on a local and time-dependent softening of the workpiece material. Pre-heating of a specific area over the workpiece surface is the main mechanism behind the previously performed LAMM investigations [14–17] (Fig. 1a). The preheating leads to the reduction of the material stiffness and strength, which leads to material softening along the cutting path (as shown in Fig. 1a). Furthermore, only continuous wave or short-pulsed lasers are utilized. This method has several disadvantages. Its main drawback is the induced Heat Affected Zone (HAZ) on the workpiece surface and the difficulty of controlling undesired changes of material properties, which could remain on the finished surface even after the micro-milling process [18,19]. These processes utilize a focused laser beam in front of the milling tool (geometrical constraint), as shown in Fig. 1a, and are only applicable in dry conditions, where using coolant in the process is not
possible [20,21]. The presence of coolant causes an immediate cooling of the laser affected and softened area and makes the process inefficient. Therefore, the tool temperature is generally higher within these classes of LAMM compared to the conventional micro-milling [19]. To reduce the cutting forces and temperature a novel LAMM process, developed by the Institute of Precision Machining (KSF) [20,21], is utilized for the first time for micro-milling of X5CrNi18-10. Furthermore, Controlling the depth of thermal damages with respect to the laser ablation parameters in processing with ultra-short pulsed lasers has not been precisely addressed in the micro-scale till now. The proposed method is a two-stage (structuring and micro-milling) laser-assisted machining process, which can be performed also in wet conditions. In this method, an ultra-short pulsed laser has been used to structure the workpiece surface (stage 1), instead of using a continuous or short-pulsed laser to pre-heat a specific area of the workpiece surface. Afterward, the micro-milling tool was used to remove the structured area to the final dimensions (Fig. 1-b). Utilizing ultra-short pulsed laser induces a neglectable and simultaneously controllable HAZ on the workpiece surface [22] which is the main advantage over the previous LAMM methods. Additionally, the process can be utilized in wet conditions with no restriction on the application of machining fluids. A laser-structured workpiece sample can be successively machined in wet conditions, where the structured zone will be completely removed (stage 2) via chip formation in presence or absence of coolant or lubricant. The workpiece can be structured up to a certain depth according to the laser parameters, and finally be machined with different cutting conditions. Therefore, the machining parameters are completely independent of the laser parameters, and laser-structuring and micromilling processes are completely time-independent. Accordingly, only the depth of laser-induced features is matched with the milling depth of cut. To this end, several laser-generated structures have been induced over the workpiece surface and milling forces and cutting temperatures have been compared in LAMM and CMM process. Experimental procedure Austenitic stainless steel X5CrNi18-10 was chosen as the workpiece material. It has favorable properties such as high corrosion resistance, high strength, and ductility. These properties make it a suitable choice for a wide field of applications like the medicine sector, automotive industry or high precision measurement equipment. However, the machining of this material is challenging. One of the main challenges is burr formation owing to its high toughness. Moreover, it has a high tendency to work hardening and build-up edge formation during machining [23]. The samples were prepared in a block form with dimensions of 30 × 20 × 10 mm. The micro-milling tests were carried out at different cutting speeds (vc) and feed rate (fz). The influence of micro-
Fig. 1. Schematic view of LAMM with the help of (a): softened area (b): structured area. 175
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length to evaluation length, λc/L, of 0.8/4 mm. The average of these 3 measurements was chosen as the value of surface roughness. Fig. 4 shows the procedure and a measurement sample of the surface roughness. Fig. 2a shows the experimental LAMM setup. A high-speed thermography camera (Infratec, ImageIR 8300) was used to measure the process temperature as shown in Fig. 2c. Since the temperature measurement via thermography camera is not possible under wet conditions, the micro-milling tests were performed in dry conditions. However, the utilized laser-structuring approach allows the application of cutting fluids during LAMM. In this study, a two-flute tungsten carbide micro-endmill with a diameter of 500 μm and a helical angle of 30 degrees from Crazy Mill- Mikron Tool was used in a side micro-milling operation (Table 1). Fig. 5a shows the schematic of the side micromilling. Fig. 5 shows the experimental setup for the laser structuring of the workpiece surface consisting of a laser scanner, workpiece, and an air nozzle. The laser scanner focuses the laser beam over the workpiece surface and the air nozzle is responsible for sweeping the ablated debris away. The process parameters are listed in Table 1. Several laser patterns were produced on the workpiece surface which are listed in Table 2.
milling parameters on the cutting forces and the surface roughness were studied while using conventional micro-milling process under wet conditions (utilizing oil as coolant lubricant). Subsequently, the workpiece was laser-structured using a Yb:YAG picosecond laser (TruMicro5050 from Trumpf) with different laser parameters. Each test was repeated three times and the average values have been reported in the paper. One of the challenges during the structuring is heat dissipation. In using a picosecond laser, pulse energy in the microjoule range is radiated in a very short time period over the workpiece surface. It was already shown in authors previous research [22] that, as the ultra-short pulse length of the laser beam the ratio of the pulse length (10 ps) to the pulse radiation period (1/400 kHz = 2.5 μs) is very small (the relative relaxation time is so large), the imposed heat of a single pulse is dissipated before the next pulse. Therefore, the heat accumulation owing to successive pulsed is negligible. Assuming the extreme case, that the whole laser energy, which is delivered by the removed debris and partially remained inside the workpiece material is absorbed by the workpiece, the overall laser heat input to the workpiece corresponds to a heat source with a few Watts of average power. The assigned pressurized air nozzle within the experiments serves to remove the debris from the laser structuring and compensates the heat accumulation up to this moderate laser power level. A high-precision 5-axis CNC machining center (KERN Pyramid Nano) was used to perform the micro-milling tests (Fig. 2a). The forces were measured by a Kistler dynamometer (type 9256C2) and processed with Kistler 5015 charge amplifiers and recorded by a PC acquisition board and LabView software. The dynamometer has suitable dynamics and sensitivity for micro-machining processes. The presented forces are peak-to-valley forces. A sample of force measurement in the time domain for both tangential and normal forces is presented in Fig. 3. To calculate the micro-milling forces, the forces from each micro-milling test were divided into three parts in the time domain of 6 × 10-3 s including the highest and lowest peaks and valleys. From each time-domain three peak-to-valley forces were measured, and the mean value was reported as the micro-milling force. A confocal microscope (μsurf mobile plus) was integrated into the machine to measure the surface roughness (Fig. 2b). The surface roughness was taken perpendicular to the micro-milling path at three positions (the beginning, the middle, and the end of the cut) with the sampling length of 4.8 mm. In each position several lines perpendicular to the cutting direction, like the black line in Fig. 4b were drawn to achieve a total length of 5 mm shown in Fig. 4c. The surface roughness was finally calculated through these connected lines with cut-off length, λc, of 0.8 mm and sample
Results and discussions The effect of cutting speed on the micro-milling process Fig. 6 shows the effect of cutting speed on the micro-milling forces and surface roughness. Increasing the cutting speed decreased both the cutting forces and the surface roughness. A 70% reduction in both tangential and normal forces could be achieved by increasing the cutting speed from 50 to 250 m/min. High cutting speeds generally lead to high process temperature. Additionally, the temperature gradient in the workpiece in front of the cutting edge in micro-milling is proportional to the cutting speed. This increment is due to the induced material flow by the micro-tool [24]. High cutting temperatures at high cutting speeds ease the material flow and material removal, causing better cutting performance and lower cutting forces. Another reason for the reduction of cutting forces is an increase in the strain-rate, which is also proportional to cutting speed. Eleiche [25] showed that strain-rate has a great effect on the flow stress. The sensitivity of the flow stress to the strain-rate can be categorized by three regions with different deformation mechanisms. When the strain-rate is low, the dominant phenomenon governing the material removal mechanism and flow stress is the dislocation density, which is barely influenced by strain-
Fig. 2. (a): The experimental setup; (b): Integrated confocal microscope, (c): The temperature measurement. 176
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Fig. 3. The cutting forces sample in time domain (vc = 50 m/min, ae = 50 μm, and fz = 4 μm/tooth). Table 1 Process parameters. Parameters
Values
Milling tool Workpiece Cutting speed (vc) Feed per tooth (fz) Radial depth of cut (ae) Axial depth of cut (ap) Coolant Laser power (PL) Laser pattern Laser scanning speed (vL) Pattern distance (L)
2.CMC30.A3Z2.050.1 - Crazy Mill X5CrNi18-10 (1.4301) 50,150, and 150 m/min 1, 3, 4, and 6 μm 50 and 60 μm 500 μm Oil and dry 7.5 and 10 W X 100 mm/s 250 and 500 μm
Merchant’s equation:
∅=
β π α + − 4 2 2
β =tan−1 (
Fn + Ft tanα ) Ft − Fn tanα
(2)
(3)
where β is the friction angle and α is the clearance angle. However, in the case of high-speed cutting, the flow stress is dominated by the effects of strain-rate. Therefore, in high-speed cutting the theory of the strain gradient plasticity is not valid [26]. Thus, increasing the cutting speed causes more average strain-rate ε¯n [27], (following equation), affecting the flow stress and material removal mechanism:
ε¯n =
rate. The flow stress δ, in this regime, can be expressed according to the cutting and thrust forces using the following equation [25]:
(Ft sin∅+ Fn cos∅) sin∅ h d ap
(4)
On the other hand, the yield strength of X5CrNi18-10 stainless steel increases slightly with rising the strain-rate. This increment is due to the strain hardening of the material in terms of a transition from isothermal (low strain-rate) to adiabatic process (high strain-rate) [28]. Moreover, at higher strain-rate the accumulation of the generated heat in the workpiece is more considerable. The plastic deformation within a very short deformation time during micro-milling at high cutting speeds causes higher Stacking Fault Energy (SFE). The higher SFE impedes martensitic transformation. Therefore, the strain hardening is dropped with rising the strain and strain-rate [28]. The aforementioned factors ease the cutting process and lower the cutting forces at higher cutting speeds. Fig. 6-b shows that increasing the cutting speed leads to a slight reduction of surface roughness, despite the uncut chip thickness was kept constant. Reduction of roughness values with increasing the cutting speed could be due to lower built-up edge formation over the
Fig. 4. The surface roughness measurement procedure (a): schematic of the workpiece; (b): surface topography with a confocal microscope (vc = 250 m/ min, ae = 50 μm, and fz = 1 μm/tooth), (c): The surface roughness profile extracted from several parallel lines in b.
δ=
vc hd
(1)
where Ft and Fn are the tangential and normal forces, respectively. hd is the undeformed chip thickness, ap indicates the cutting depth. ∅ is the prediction of shear angle which can be determined according to 177
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Fig. 5. (a): Schematic of LAMM process; (b): Produced patterns on the workpiece surface via laser; (c): The setup for the laser structuring of the workpiece. Table 2 Laser patterns. Structures
Laser parameters
Non-structured Pattern 1 Pattern 2 Pattern 3 Pattern 4
off PL = 10 W; L = 500 μm PL = 10 W; L = 250 μm PL = 7.5 W; L = 500 μm PL = 7.5 W; L = 250 μm
various feed per tooth (0.001; 0.003; 0.004 and 0.006 mm) were conducted. It can be stated from Fig. 7 that, the lower feed rate (and lower average chip thickness) causes lower forces, as well as better surface, roughens as it was expected. Higher material removal rates can be achieved at higher feed per tooth rates, resulting in larger milling forces. Li et al. [30] showed that the minimum chip thickness is a dominant factor regarding the surface generation in micro-milling process. The uncut chip thickness is increased by keeping a constant depth of cut while increasing the feed rate (Eq. (5)). In micro-milling process, under definite cutting conditions, the maximum uncut chip thickness can be smaller than the edge radius of the micro-milling tool. Thus, finding the minimum chip thickness hm or critical chip thickness seems essential, under which workpiece material may not be properly removed as chips. This condition causes lower share of chip formation and higher plowing and rubbing in the micro-milling process. With
cutting edge [11]. Weule et al. [29] showed that the built-up edge worsens the surface quality in the micro-cutting process which is already known from macro-cutting. Increased cutting forces cause higher tool wear, larger tool deflection and more considerable built-up-edge [29]. The effect of undeformed chip thickness on micro-milling process The effect of undeformed chip thickness, hd, on the cutting forces and surface roughens was also investigated (Fig. 7). The average chip thickness can be calculated from the following equation:
a hav = fz ⎛ e ⎞ ⎝ d ⎠
(5)
where d is the diameter of the micro-milling tool. Experiments with
Fig. 6. The effect of cutting speed on (a) micro-milling forces, (b) surface roughness. 178
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Fig. 7. The effect of average chip thickness on (a) micro-milling forces, (b) surface roughness.
uncut chip thicknesses below this value, the surface roughness values may fluctuate and in some cases be improved as shown by Li et al. [30]. The minimum chip thickness is determined by the following equation [31]:
β π hm = re ⎛1 − cos( − )⎞ 4 2 ⎝ ⎠
Table 4 ANOVA for Ra.
(6)
where β is the friction angle between the tool and the un-cut workpiece passing under the tool and re is the tool edge radius. According to Eq. (6), the minimum chip thickness can be changed by changing the friction coefficient and tool edge radius. Brinksmeier et al [32] found that the friction coefficient while micro-milling austenitic stainless steel with tungsten carbide tool is about 0.258. The tool edge radius was also measured around 2 μm using a confocal microscope. Accordingly, the minimum chip thickness for this material can be derived from Eq. (6) which is 0.415 μm. As it is shown in Fig. 7-b increasing the chip thickness steadily increases the value of the surface roughness. It implies that the chip thickness is larger than the critical chip thickness.
Factors
DF
Sum of squares (SS)
Mean squares (MS)
F ratio α = 5%
P
Contribution (%)
vc (m/min) fz (μm) Error Total
2 3 12 17
0.003735 0.002841 0.000246
0.001868 0.000947 0.000021
91.01 46.14
0.000 0.000
43.19% 52.28% 4.53% 100%
Analysis of variance (ANOVA) To determine the statistically significant factors and their contribution to the micro-milling forces and surface roughness, analysis of variance (ANOVA) was used. The results are presented in Tables 3 and 4 for the tangential micro-milling force and surface roughness Ra, respectively. The ANOVA analysis was conducted at the level of confidence 95%. The results showed that the feed rate has the largest influence on the micro-milling force with 80% compared to the cutting speed with 16% contribution. However, the most influential factor for the surface roughness is the cutting speed, though the influence of the feed rate with 43% contribution is not neglectable. Both factors, cutting speed and feed rate, had a statistical and physical significance on the cutting force and surface roughness since P value was lower than 0.05.
Fig. 8. different views by SEM.
The cutting mechanism The machined surface was observed under different views (Fig. 8) with SEM microscopy. Fig. 9 shows the machined surface from the top view. The surface was generated with the front end of the cutting tool. The micro-milling feed pattern is clearly observed in Fig. 9-c. The black spots over the surface indicate the cavities, while the regions with white borders indicate the smeared materials. The smeared regions are resulted from a high amount of plastic deformation of the material, during the last chip formation phase in the contact zone. In down milling, the chip thickness approaches zero at the end of the chipping process (the end of engagement arc). In this position, the actual rake angle becomes highly negative because of the cutting-edge radius. When the rake angle is negative, the chips break at a certain position in the cutting path. The remaining material on the cutting path is deformed and smeared by the cutting edge over the surface of the workpiece, rather than being cut as chips. Cavities, metal debris, and feed marks, as well as, smeared material particles could be observed on the micro-milled wall as shown in
Table 3 ANOVA for P-to-V Ft. Factors
DF
Sum of squares (SS)
Mean squares (MS)
F ratio α = 5%
P
Contribution (%)
vc (m/min) fz (μm) Error Total
2 3 12 17
0.25460 1.27230 0.06360
0.127300 0.424100 0.005300
24.02 80.02
0.000 0.000
15.96% 80.04% 4.00% 100.00%
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Fig. 9. SEM Images from the milled surface from top view (view 1in Fig. 6) (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth).
plastification tends to flow through the cutting-edge toward the secondary cutting-edge and finally adheres over the machined surface. There are two mechanisms for the material side flow. The first, the material is squeezed between the flank face of the cutting tool and the finished surface. The second, the material which is encountered by the
Fig. 10. The white line in Fig. 10-c shows the feed mark over the surface of the workpiece, which led to material side flow. The material side flow occurs when the chipping material in the cutting-edge experiences severe stress and thermal loads. These conditions could lead to considerable plastification over the machined surface of the material. This
Fig. 10. SEM images from the milled surface from view 2 (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth). 180
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Fig. 11. SEM images from the milled surface from view 3 (vc = 250 m/min, ae = 50 μm, and fz = 4 μm/tooth).
plastification in the cutting zone flows through the worn trailing edge to the side of the cutting tool [33]. The main reason is the high temperature during the cutting process. Moreover, some material debris was observed over the milled surface (Fig. 10-b). The produced chips during the micro-milling can be welded to the machined surface owing to the high pressure and temperature in the process, which in turn can deteriorate the surface quality of the workpiece. Fig. 9–11 indicates that the material is exposed to high plastic deformation during the micro-milling process. As it is known, during the cutting process the material is extruded to a large extent. The extruded material is pushed to the side and causes burr formation which is marked as plastic deformation in Fig. 11. A large amount of metal debris is also observed after the cutting process in the corner of the machined area, where the corner of the tool formed the cut. Compared to the other areas of the tool, the tool radius (at tool corner) induces much higher cutting temperature and forces, since the corresponding contact length is larger. Higher temperature and forces lead to a larger amount of metal debris in the corner of the cut. Moreover, the chips on the micro-milling tool corner cannot be easily evacuated so they could be welded to the machined surface. The effect of laser structuring on micro-milling process Fig. 12. The effect of laser structuring on the cutting forces in the time domain (vc = 50 m/min, ae = 60 μm, and fz = 6 μm/tooth).
An ultra-short pulsed picosecond laser is utilized to produce specific patterns on the workpiece surface (Table 2). After laser-structuring the structured and non-structured workpieces were micro-milled with cutting speed of 50 m/min, feed per tooth of 6 μm, and depth of cut of 60 μm to generate higher cutting forces and temperature in the process. Since the depth of laser-cut was approximately three times larger than the depth of cut (60 μm), three micro-milling passes were performed to remove the structured area, and the mean values were reported as the results of temperature and cutting forces. The results of the LAMM cutting forces and temperature are compared with the conventional micro-milling and are presented in Figs. 12–14. The tests were performed under dry conditions. The temperature profile presented in Fig. 14 was measured on the machined surface (on the area exactly behind the milling tool). Measuring the cutting contact temperature was not possible since the thermo-camera requires a direct vision over the measured surface area. As Figs. 13 and 14 show, LAMM led to a significant reduction of cutting forces and temperature. Fig.12 presents the cutting forces in a time domain for both CMM and LAMM processes. In CMM process both tangential and normal forces fluctuate around an average value. However, in LAMM process forces do not follow the same trend. This deviation could be due to melted material regions or non-uniform lasergenerated grooves over the surface of the workpiece. Laser-structuring removes material from the surface of the
workpiece and, therefore, the tool-workpiece contact area becomes smaller. Hence, the smaller material volume is removed during the micro-milling process subsequent to the laser structuring, compared to the micro-milling of non-structured surfaces. Additionally, the process becomes more intermittent than conventional micro-milling. Lower contact length and material volume to be cut lead to lower micro-milling forces and temperature. Another reason for the reduction of micromilling forces and temperature lies in the chip morphology. As it is clear in Fig. 15, the LAMM produces shorter, thinner, and segmented chips compared to the CMM, which leads consequently to lower micro-milling forces. Pattern 2 with the laser power of PL = 10 W and the pattern distance of L = 250 μm generated the smallest milling forces and temperature. Moreover, lower laser power leads to larger cutting forces. Increasing the laser power increases the input laser energy density according to the following equation [21]:
EL − input =
PL (dL + wL) vL
(7)
Where, wL is the laser line width, dL indicates the laser line span, and the laser scanning speed is indicated by vL. The amount of material 181
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Fig. 13. The effect of laser structuring on the cutting forces.
part. However, in the case of LAMM material debris can be observed more frequently. The white lines in Fig. 15a are determined as the feed mark over the finished surface, which led to material side flow. These feed marks are not observed over the finished surface via LAMM process (Fig. 15a and b). In both cases some cavities can be also seen over the workpiece surface. It could be implied that lower micro-milling forces in addition to lower temperatures in the case of LAMM process caused lower plastic deformation during the cutting process.
removal during laser-structuring can be increased by increasing the laser input energy. Compared to the conventional micro-milling, tangential and normal cutting forces could be reduced by up to 70% and 40%, respectively, via the LAMM. Fig. 14 shows that using higher laser power resulted in a lower cutting temperature during micro-milling. The cutting temperature of LAMM in the case of PL = 7.5 W is almost similar to the conventional micro-milling. The laser-structuring is accompanied by the ablation process. After laser-structuring the ablated material can be deposited and accumulated into the lasered slots. This phenomenon was observed mainly at lower laser powers and caused more cutting temperature in the subsequent micro-milling process. It is also demonstrated that higher laser powers enhance the crack initiation and propagation around the laser structures, owing to the thermal shocks induced by ultra-short pulsed lasers [29]. These cracks can also lead to lower cutting forces and temperature since they reduce material strength and stiffness.
Conclusion A novel and promising laser-assisted micro-milling process utilizing an ultra-short pulsed laser was developed in this study. Different patterns with different laser powers and pattern distances were produced on the surface of the workpiece using a picosecond laser prior to micromilling. The effect of micro-structuring, as well as cutting parameters on the micro-milling of an annealed alloy steel was investigated. The results showed that the cutting speed had a great influence on the cutting forces; however, its influence on the surface roughness was negligible. Increasing the cutting speed form 50 m/min to 250 m/min could reduce the cutting force up to 50%. However, 5 times higher cutting speed just improved the surface quality up to 15%. As it was expected, increasing the feed per tooth in micro-milling increased both the milling forces and surface roughness values. Increasing the material
The effect of laser structuring on material characterization Fig. 16 shows the material characterization of the micro-milled material via LAMM-pattern 2 and CMM. The finished surface of the conventional micro-milling process reveals large material deformation in the form of smeared material over the surface of the micro-milled
Fig. 14. The effect of laser structuring on the cutting temperature. 182
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Fig. 15. Chip morphology (vc = 50 m/min, fz = 6 μm, and ae = 60 μm), (a): CMM; (b): LAMM.
the conventional micro-milling process. This can be explained via smaller contact length and smaller material volume which should be removed by the micro-milling tool. The chip morphology of the LAMM process showed that the chips are smaller and thinner compared to the chips in conventional micro-milling, which also led to lower milling forces. This can be used as an effective and new manufacturing technique to increase the accuracy of the parts, enable larger material removal rates and cause lower tool wear because of lower cutting forces and temperature. Increasing the laser input energy density, increased the material removal rate during the laser-structuring and enhanced the LAMM process by reducing the milling forces and temperature.
removal rate up to 6 times, by increasing the feed per tooth from 1 to 6 μm, increased the micro-milling forces up to 3 times. However, surface roughness did not change significantly. During the micro-milling process, the workpiece was subjected to considerable plastic deformation in form of burr in the edge of workpiece and smeared material. Moreover, several failures such as side flow, smeared material, metal debris, and cavities were observed over the machined surface after micro-milling process. Workpiece structuring using an ultra-short pulsed laser prior to the micro-milling process led to a meaningful reduction of normal and tangential micro-milling forces (70 and 50%, respectively) compared to
Fig. 16. SEM images from the milled surface (vc = 50 m/min, fz = 6 μm, and ae = 60 μm), (a): CMM; (b): LAMM pattern 2, (c): CMM detailed, (d): LAMM pattern 2 detailed. 183
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Furthermore, the smaller distance of the patterns over the surface of the workpiece caused lower milling forces and temperature. The best performance of the laser-structuring was obtained using the laser power of 10 W and the pattern distance of 250 μm. Finding suitable laser parameters is a key factor to have an efficient LAMM process, especially regarding the cutting temperature. Using improper laser parameters may cause an inefficient LAMM process. It was shown that the cutting temperature did not change significantly in LAMM compared to the conventional micro-milling in the case that the laser power was set to 7.5 W. However, increasing the laser power to 10 W could reduce the cutting temperature in LAMM up to 50%. It is worth mentioning that laser-structuring may also help the lubrication and cooling effect during the machining. Hence, future investigations are planned to study the effects of LAMM under wet conditions.
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