ORIGINAL ARTICLE
Micromachining of PMMA
—manufacturing of burr-free structures
with single-edge ultra-small micro end mills
Ingo G. Reichenbach1&Martin Bohley1&Fabio J. P. Sousa1,2&Jan C. Aurich1
Received: 5 July 2017 / Accepted: 18 February 2018 / Published online: 10 March 2018 # Springer-Verlag London Ltd., part of Springer Nature 2018
Abstract
Nowadays, the prototypes of microfluidic systems are generally produced via micromilling of thermoplastic polymethyl meth-acrylate (PMMA). The main limitations are the design of micro tools with diameters D≤ 50 μm adapted for each application, and the understanding of the machining process itself. The objective of this research work is to contribute to mastering the process of PMMA micromilling with tool diameters D≤ 50 μm on a 3-axes precision milling machine. For this purpose, the process design must include the complete process chain—from the CAD/CAM data up to the final structure geometry. The main requirements are the manufacture of microfluidic structures with Ra< 60 nm on the groove bottom and a top burr overhang h0< 3μm. Based on the experimental results, milling parameters were established and the influence of the tool geometry on the burr formation was determined. Finally, CAD/CAM machining strategies were recommended.
Keywords Micromachining . Micromilling . Lab-on-chip . Ultra-small micro end mill
1 Introduction
The advances in micromachining encourage the development of new microproducts and the improvement of existing sys-tems [1,2]. In this context, microfluidic lab-on-a-chip (LOC) platforms represent an important toolbox allowing the manip-ulation of small fluidic volumes in life sciences and chemical analytics. Such microfluidic systems offer disposable devices with high throughput, low reagent consumption and reduced analysis time. In addition to these advantages, the application
of polymers as basis material allows mass production with low cost per unit and a wide field of application [3].
The mass production of polymeric LOCs is mostly based on the use of molds. Such molds generally have an inverse geometry of the desired structure, and the development and manufacture of those molds have a primary influence on the production chain. Depending on the microfluidic design, the manufacture of the molds may require the combination of different manufacturing processes such as micromilling, EDM, or laser. This combination allows the optimized pro-duction of molds with dimensions between 20 and 50μm, considering structure sizes and machining time [4].
In a competitive market, the need for innovation and prod-uct variety demands a redprod-uction in the time-to-market. As consequence, the very time-consuming process of mass repli-cation technologies may be the critical point, especially in the prototype phase of the product development. An emerging alternative to reduce cost and time for bringing new products to the market is the direct milling of LOC structures. In fact, several studies have reported the direct milling in polymers using CNC-controlled machine tools as a suitable manufactur-ing process in a prototypmanufactur-ing phase [5]. Kang machined a PMMA microchip with a 200-μm-diameter end mills and ob-served a minimal dimensional error when compared with in-jection molding or hot embossing methods [6].
* Fabio J. P. Sousa fabio.sousa@ect.ufrn.br Ingo G. Reichenbach ingogustav@gmail.com;https://www.mv.uni-kl.de/fbk/home Martin Bohley martin.bohley@mv.uni-kl.de;https://www.mv.uni-kl.de/fbk/home Jan C. Aurich Jan.aurich@mv.uni-kl.de;https://www.mv.uni-kl.de/fbk/home 1
Institute for Manufacturing Technology and Production Systems, University of Kaiserslautern, Gottlieb-Daimler-St,
67663 Kaiserslautern, Germany
2 Escola de Ciências e Tecnologia, Federal University of Rio Grande
do Norte (UFRN), Campus Universitário Lagoa Nova, Natal, RN 59078-970, Brazil
After reporting the low machinability of cyclic olefin co-polymer (COC/Topas) using laser ablation, Bundgaard ma-chined different structures using commercially available end mills and drills. He manufactured channels with 30, 100, and 200μm width and holes with 300 and 500 μm diameters [7]. Jáuregui micromilled PMMA with 350-μm-diameter end mills and concluded that the milling process for a prototype structure is 30 times faster than making one prototype with a lithography process [8]. Despite the very promising contribu-tion of direct end milling to the produccontribu-tion of PMMA proto-types, most studies in this area are limited to the micro end mills’ dimension or the structure complexity [2].
This paper studies the feasibility of the direct manufacture of lab-on-a-chips using FBK-developed precision machine tools and micro end mills. The main goal is to analyze the challenges and the technical response of the direct micro end mill process during producing PMMA prototypes with tool diameters D≤ 50 μm. It presents therefore a comprehensive investigation of the machining process of PMMA structures using a single-edge micro end mill, including the burr forma-tion, the influence of the cutter geometry, machining strate-gies, and the use of CAx-supported machining.
1.1 CAD/CAM and machining strategies
in micromilling
The use of CAM technology and the strategies for tool path generation are an important instrument allowing the optimiza-tion of both, the cutting, and the dynamic behavior of micro end mills. Some researchers investigated the integration of CAD/CAM systems and tool path generation into the micromilling process. Due to size effects, the surface and ge-ometry requirements are strictly dependent on the workpiece, cutting tool, and machine tool combination. Even selecting the machining parameters is challenging in micromilling, since unlike a conventional process, there is limited process knowledge.
Most related publications take the milling process charac-teristic into account and try to reduce its negative effects. The common practice is a manufacturing oriented design of the workpiece trying to avoid sharp corners, and thus preventing posterior damages and structural failures. Abrupt changes of tool direction are also not recommended during the milling process, in order to maintain constant cutting conditions, as well as to stabilize the forces, which in turn is essential for producing thin and delicate features, and for avoiding uncon-trolled tool deflections and vibrations. Therefore, a clear def-inition of the characteristics which are important for the final structure contributes to the selection of the right machining strategy.
Considering the surface and geometry requirements, there are basically two machining regions to focus on: the bottom and side walls of the structures. Using micro end mills, the
bottom of a structure is milled by the axial/face cutter. The surface topology as well as the geometrical depth of the struc-ture is defined by the tool path and the machining parameters. The structure’s side walls are machined with the radial/ peripheral cutter. In this case, the strategy and the machining parameters, which may differ from milling the bottom surface, define the wall quality, the top burr formation, the tool/workpiece deflection behavior, and the final geometrical accuracy [9–11]. A process characteristic is that the down-milled side walls have a better surface finish than the up-milled ones, so that the down-milling tool path is preferable [1].
Typical for the milling process of pouch-shaped structure is the machining from inside out, i.e., a roughing operation followed by finishing [11]. During roughing, the inner mate-rial of the pouch-shaped structure is machined at the highest possible material removal rate Qw(“clearing” of the structure). Finally, the final contours of the structures flanks (outer con-tour) are machined by down-milling (ae< 50% D) with new and thus unworn tools to achieve a smooth flank topography [11].
The performance of LOC can be strongly affected by the presence of burrs, so their characteristics became important criteria of the LOC requirements. An overview about burr formation in micro end milling is presented in the next section, including grouping the mechanisms of burr formation and relevant models.
1.2 Burr formation in micro end milling
The burr formation, like the surface roughness, is of particular importance in micromilling. In the production of prototypical microfluidics, the burrs delimit the sealing capacity of joint microcomponents. If the burrs are too high, a thick gap may result between the joint parts and liquids may leak. The clas-sification, terminology, and common models for describing the burr formation in end milling are presented in Fig. 1. There is a plurality of definitions for burr. Hereafter, they refer to an unwanted amount of material arising by plastic defor-mation on a workpiece edge. This extra material extends be-yond the ideal workpiece edge and may be pendulous or as-sociated to possible waveforms [12].
According to the German standard DIN ISO 13715, periph-eral edges are classified as burred if the measured edge devi-ates more than 50μm from the idealized geometry [13]. This threshold is obviously unsuitable for classifying the edges resulting from micro end mills. The parameters adopted in this work are presented in Fig.1, which were inherited from the German standard.
The adjustment made to this model are based on existing literature about micromilling with D≤ 50 μm [14] and the requirements for the production of microfluidic prototypes [15]. The burr profile and the measurement values in the cross
section are used for the assessment and examination of burrs [16]. Since the preparation of micromachined structures in cross sections for the measurement of burrs is limited due to the feature sizes, usually a quantitative statement by optically detected point clouds is made.
The terms primary burr and secondary burr are typically used for describing burrs. Primary burrs are completely pro-nounced burrs, arising immediately after the separating oper-ation, while secondary burrs are smaller and derive from a primary burr. The transformation of primary to secondary burrs can occur by two principles, depending on the machin-ing process. The primary burr either breaks at its thinnest point (minimum burr thickness bg) or is minimized by deburring operations. Despite the associated reduction in burr size, in neither case, the ideal geometric workpiece edge is reached [17].
In end milling, two relevant engagement conditions prevail (Fig.2): the face-peripheral milling and the slot milling. In face-peripheral milling, six peripheral burrs result (Fig.2a). In slot milling, eight burrs occur in an open groove (Fig.2b). Hashimura et al. divided the types of burr by the location of formation [18]. In this work, the classification in entrance, sideward, and exit burrs was combined with the system of classification by Nakayama and Arai [19]. Table1 provides the system used for burr classification, and Fig.2illustrates the locations of the burrs for face-peripheral and slot milling. The identification of the burrs was based on the following parameters: direction of burr formation (Eb, Ab, and Sb), af-fected edge (F, G), and relative movement between cutting edge and workpiece (GG, GL). By the engagement conditions in end milling process, burrs can be classified as Poisson burr, rollover burr, and tear burr, and all of them can be described by appropriate development mechanisms and models [20]. The Poisson burr is caused by the compression of material at the cutting edge and consists of a prominent bulge displaced aside or reverse to the cutting tool motion. The rollover burr is the remaining material located on the edge of the workpiece that
has been bent by the main cutting force. The tear burr arises when the workpiece material is teared instead of sheared.
The blue-marked burrs Eb-G-GG and EB-F-GG on the entrance edge and at Ab-G-GL as well as Ab-F-GL on the exit edge outweigh the lateral deformation, and thus, Poisson burrs arise. These are relatively small and usually not investigated. In slot milling, the groove bottom at Eb-G is mainly influenced by lateral deformation, while Ab-G is mainly formed by tearing the material.
At the yellow burrs Eb-G-GL and Eb-F-GL or at Ab-G-GG and Ab-F-GG, rollover burrs arise. Since the material dis-placement at the tool edge leads to comparatively large burrs, extensive efforts have been made to model such burr forma-tion processes. Mechanisms for this have been worked out for rollover burrs and tear burrs by Ko and Dornfeld [21] and Hashimura et al. [18]. For the green-marked burrs Sb-G-GG and Sb-G-GL, detailed investigations for face milling were made by Chern, whereby five different types of burrs were detected [22]. The burrs Sb-G-GL are identified as rollover of the chip or tearing of workpiece material, depending on tool-workpiece interaction, and are classified as hybrid Poisson burr/tear burr, while the burrs Sb-G-GG are referred to hybrid Poisson burr/rollover burr.
In most cases, the sideward burrs on the flanks (simply defined as top burr) are relevant for the machining of microfluidic structures. The red-marked burrs Sb-F-GG und Sb-F-GL are defined as Poisson burrs in slot milling as well as in face-peripheral milling [20].
The modeling of the sideward burrs on the groove flanks in micro end milling is summarized in Fig.3. Chern et al. exam-ined the burr formation while machining of AL 6061-T6 with a cutter diameter of about D = 100μm [23]. As seen in the figure, sideward burrs arise on the flank. The chip is not formed during the engagement of the micro end mill. The workpiece material is pressed against the tool causing a plas-tically deformed region (Fig.3(a)). Continuing this process, a crack near to the center of the groove is formed (Fig.3(b)).
Fig. 1 a, b Configuration and parameters of the edges in micromilling
The forward motion of the cutting tool also leads to the for-mation of burrs on the lateral edges. These burrs cumulate along the tool path until they become removable by the cutting tool, or until the occurrence of further plastic deformation or crack formations. The position of the crack formation in the workpiece is critical for the emergence of burrs. If the crack formation does not start in the middle, but for example on the up-milling side (Fig.3(f)), the burr formation is more pro-nounced on the down-milling side (Fig.3 (h)) [24]. Piquard et al. [25] examined the burr formation with biocompatible memory-shape alloys—NiTi—with double-edged end mills (D = 800μm) and justify the burr formation with the model proposed by Weinert and Petzoldt [26]. The burr formation is described by the ratio of feed per tooth and emerging burr root thickness bf. If the feed per tooth is constant and bigger than the resulting burr root thickness, no burr is cumulated with continuous processing. Upon reaching a critical cutting capa-bility (e.g., caused by wear), the tool may lose the acapa-bility to machine the workpiece (bf< fz). A small proportion of burrs remain, being latterly displaced to the flank. With further
machining, the tool loses its cutting ability according to DIN 6583 [27], and as result, the burr size increases [26].
General statements about the burr formation in micro end milling are difficult, because the process-machine interaction has a significant impact on the process variables. Also, the properties of the workpiece material, as ductility for example, are crucial for the formation of burrs. In general, the burr formation is minimized with increasing feed per tooth by hav-ing the undeformed chip thickness h much greater than the cutting edge radius rβ(h >> rβ). Furthermore, up-milling min-imizes the burr formation compared to down-milling [28].
2 Machining set-up and preliminary
investigations
Both machine tools for the manufacture and the application of the micro end mills were developed at the Institute for manufacturing technology and production systems (FBK). The micro end mills were manufactured on a desktop 3-axes
Fig. 2 Burr location fora face-peripheral milling andb slot milling
Table 1 System of burr
classification in end milling Direction of formation Cutting edge directly concerned Movement between cutter and workpiece
Parallel to feed direction
Entrance burr Eb Side walls Up-milling GG
Exit burr Ab Peripheral cutter F
Perpendicular to feed direction Bottom Down-milling GL
CNC tool grinding machine with a resolution of 0.1μm and a positioning accuracy of 1μm. The tool blanks were therefore mounted on a horizontally oriented air bearing, with two dia-mond grinding wheels. The final micro end mill geometry was ground according to [29]. Milling was performed using a desktop 3-axes CNC precision milling machine with a resolu-tion of 20.32 nm and a posiresolu-tioning accuracy of 1μm. A 54,000-rpm-range air bearing spindle was vertically installed on the Z-axis. The motions of the work table were executed with X and Y air bearing linear stages with a total travel of 100 mm [29]. Tools with shank diameters of 3 mm and 1/8″ are usable, as well as CAD/CAM generated G-code following the ISO 6983.
The workpieces were made of PMMA, and pouch-like structures (e.g., storing sections or reservoirs) and LOC chan-nels were selected for being machined.
Three different cutting edge geometries were selected to investigate the effect of the cutting geometry on the machining results. For investigating each cutting edge geometry, three new tools were produced and characterized. These geometries are presented in Fig.4and were namely undrilled tools (λ = 0°) with D = 20 μm, χ´r= 12° and D = 48μm, χ´r= 12°, as well as drilled tools with D = 48μm, χ´r= 1° (Fig.4). The drilled tools are characterized byλ = 30°, which corresponds to the current helix angle when machining PMMA [30].
The minor cutting edge angle was set toχ´r= 12° andχ´r= 1° in case of drilled tools. These parameters were selected based on previous kinematic simulations. While the kinematic roughness on the groove flanks increases exponentially de-pending on the feed per tooth (independent of tilt angleρ), the roughness parameters on the groove bottom are dependent onχ´randρ. The feed per tooth was set as fz= 1, 3, and 6μm to investigate possible contact between the tool’s peripheral
flank face and the workpiece. Additionally, the depth of cut was set to ap= 10 μm. The cutting speed was set to vc= 3.14 m/min, corresponding to a spindle speed of 50,000 rpm for D = 20 μm and vc= 4.71 m/min (spindle speed 30,000 rpm) for D = 48 μm. The reasons for this choice of parameters are given in Figs.5and 6and are related to the interaction between feed per tooth and cutting speed per-formed as preliminary investigations. The feed per tooth fz has a strong influence on the chip formation and on the kine-matic of the tool-workpiece interaction in micromachining. The characterization of the spindle indicates that radial and axial runout is depending on the spindle speed and has to be considered for a stable process.
For this reason, machining tests with different feed rates were carried out at constant cutting speed (vc constant). Accordingly, the influence of different cutting speeds was determined at constant fzvalues.
Figure 5 shows SEM pictures of the groove bottom, the down-milling flank, and both peripheral edges. The cutting speed in this case was set to vc= 1.13 m/min.
As expected, the distance between two adjacent arcuate ridges on the slot bottom (generated by the milling kinematic) corresponded to the level of feed per tooth. Furthermore, the surface becomes rougher with increasing feed per tooth. While the kinematic influence of fz on the groove bottom topography is clear, the groove flank shows no clear correla-tion with the movement of the cutter on the SEM picture. In order to assess the primary burr, the samples were not cleaned after the machining process. With increasing feed per tooth, the burr formation is reduced, while the kinematic roughness increases. With the micro end mill of D = 48μm in diameter, a chip removal could be achieved, despite low cutting speed. Therefore, for further experiments, the spindle speed was set
Fig. 3 Modeling of burr formation during micro end milling
to a value where the radial and axial runout was minimized, namely 30,000 rpm. In micromilling, the positive effects of higher cutting speed have to overcome the negative effects of possible dynamic runout errors. In other words, despite the smaller cutting speed, in terms of machining results, the use of low rotation levels of the spindle-tool-system is better than using higher cutting speeds.
At D = 20μm, an influence of the cutting speed on chip formation could be detected. Figure6presents SEM pictures of milled grooves at three different cutting speeds and each two different feed rates without previous ultrasonic cleaning. Machining was carried out with the same tool without re-clamping at each groove. A classification of chip shape (e.g., swarfs and spiral or discontinuous chips) is not possible with interrupted cutting (milling) [31]. More common is the classi-fication of chip types by Vieregge (e.g., continuous and seg-mented chips) for conventional machining and primarily for ferrous metals [32]. The chip formation of thermoplastics is characterized by the brittle and ductile temperature-dependent behavior of the material. For this reason, the phenomenolog-ical classification of the chip formation of PMMA with D≤ 50μm is carried out according to a brittle and ductile separa-tion behavior.
Figure6shows that brittle behavior occurs at low cutting speed (vc= 0.44 m/min) at a feed per tooth of fz= 0.5μm as well as at fz= 2μm. This behavior is characterized by the low plastic deformation capacity of the material. Without ductile behavior during chip formation, the produced chips do not correlate clearly with the milling kinematics, resulting in poorer surface quality.
At a cutting speed of vc= 1.8 m/min, the result depends on the ratios of cutting speed to feed per tooth and radius of end
mill to feed per tooth, vc/fzand R/fz, respectively. At lower feed per tooth (fz= 0.5μm), there is a ductile separation behavior, resulting in a better surface quality. Reducing the vc/fzratio causes embrittlement of the PMMA. The smaller plastic de-formation capacity at higher feed rates (fz= 2 μm) causes breakage or tearing of the chips during machining, which reduces the surface quality. In addition, the R/fz ratio gets smaller, which increases the probability of contact between the peripheral face flank of the cutter Aαrand the produced surface. On the other hand, by increasing the cutting speed to vc= 3.14 m/min, a prevailing of plastic deformation was ob-served. The chips correlate more clearly with the kinematic of the micromilling process and the surface quality increases.
Cracks could be observed at the end of some grooves. The material in this area is plastically deformed and matches with the burr formation behavior observed by Chern et al. [23]. The material flows from the center of the groove through an emerging sideward crack (Fig.6). Furthermore, there were variations in the burr formation process, depending on the cutting speed. Although the increase of vchas led to ductile behavior and better surface quality, the same tendency was not observed for the burr formation. For instance, machining with D = 20μm and 50,000 rpm (vc= 3.14 m/min) led to ductile separation behavior and an increased probability of burr formation.
According to the results obtained, some limits can be rec-ommended for the cutting parameters. When D = 50μm, a rotation speed of 30,000 rpm with fz≤ 10 μm is preferred, whereas with D = 20 μm, values of 50,000 rpm and fz≤ 8μm should be used. Hereinafter, only the general behavior of the chip formation in slot milling of PMMA with D = 50μm was adopted to describe the development mechanisms
Fig. 4 Geometry of the tools and cutting edge investigated during micro end milling
Fig. 5 Resulting burrs for different levels of feed per tooth
of chip formation and the resulting bottom and flank of the grooves in topographical terms.
To assess behaviors like those mentioned above, a strategy to check the machining process using a trial structure still should be developed. More details on this are given at the next section.
3 CAx-supported machining
of microstructures
A CAD/CAM system is integrated into the process chain of the production of microfluidic structures. With the computer-aided design (CAD) system, the geometrical description of the microfluidic is performed. In a second step, this geometric data is used for the tool path generation for the micro end mill process with the computer-aided manufacturing (CAM) sys-tem when micromilling the microfluidic prototype. When up-dates to the part geometry are being made, the information is transferred to the CAM system automatically. This is useful because several development-construction-manufacturing loops of prototypes are necessary, until a product is ready to be manufactured in series.
To include CAM in CAD systems, there are several possi-bilities according to the desired degree of integration. As a general purpose, CAD systems should be more involved in the process of numeric control (NC) programming. At a low level of integration, CAD and CAM systems are just connect-ed to each other via data interface, which allows the use of the geometry data from the CAD model without retyping. An
integrated CAD/CAM coupling occurs when CAD design functions and NC programming are displayed in a common system [33]. The fully integrated CAM/CAD systems are adopted here because standardized data exchange and inter-faces are not designed for the dimensions in submicron range. The entire process design was carried out in a CAx-system based on 3D CAD models of the structures. For the modeling and the subsequent creation of the respective NC machining programs, CAD/CAM modules of the integrated product de-velopment software Siemens PLM NX8.5 were used.
Figure7shows the steps in the CAD/CAM-based process design for machining microfluidics, including the operation flow of a micromixer. First, the 3D CAD geometry model of the workpiece (Fig.7(a)) has to be defined, and then the raw material to be removed computed. The geometric require-ments define the maximum diameter of the cutting tool. At next, the input data is processed by a CAM program. By selecting appropriate tools and cutting parameters (vf, n, ap, ae), the individual processing steps (roughing and finishing process) and tool paths as well as the machining strategy for each production task are determined (Fig.7(b)). This usually consists of several individual steps and has a significant im-pact on the processing time and quality of the resulting work. Using the associated cutting parameters and tool paths, the CLS data file is generated. The machining operation was also visualized in the Siemens PLM NX environment, and this allows a verification and optimization of the CAM machining strategy before manufacturing. After this, the NC programs were then generated (Fig. 7(c)) [34] using the standardized post processor.
Fig. 6 Phenomenological classification of the chip formation
The tool manufacturing, the tool clamping, and the spindle-tool-system cause runout errors, which affect the flight circle of the cutting edge and the burr formation. Unfortunately, so far, there is no tool measurement system capable of measuring rotating tools of less than 50-μm cutter diameter and that can be integrated into a desktop machine tool. Furthermore, such investigation would be limited to the flight circle, without predicting the behavior of burr formation.
Therefore, the determination of the dynamic runout er-ror was performed directly by the machining of some test structures. Here, a short distance (500μm) was machined onto a separate, non-face-milled sample of the same mate-rial with the same spindle speed and feed rate, and then measured with optical microscopy. The measured groove width wkcorresponded to the expected process result, tak-ing errors like material behavior, cutter diameter, and runout into account. Accordingly, a tool with D = 48 ± 2 μm and D = 20 ± 1 μm can generate groove widths be-tween 46 to 54μm and 14 to 26 μm, respectively, when machining test structures. For microfluidic channels to be machined by slot milling (wk= Deff), the tolerances were set to Deff= 48 ± 2μm and Deff= 20 ± 1. For structures that are wider than De f f, the tool path with D = wk is recalculated (Fig.7(c’)) and the cutter diameter D is com-pensated in the CAD/CAM by the determined groove width wk (Fig. 7 (b’)). Thus, the dimensional variations of the structures are reduced to the accuracy of the machine and tool wear. This procedure allows the precise produc-tion of prototype microfluidics on a desktop milling
machine. However, this approach is based on the trial and error principle, which is discussed in the following section.
3.1 Work preparation and post-processing
Before micromachining, the PMMA samples were face-milled to guarantee a constant depth of cut and thus con-stant structure depths. After individually clamping the plastic samples, the whole area of 75.5 × 25.5 mm2 was face-milled by a double-edged end mill with D = 3 mm by the company Kemmer Präzision. The water-cooled ma-chining was carried out at vf= 60 mm/min and a spindle speed of 30,000 rpm. To achieve the best possible rough-ness parameters, the face milling process took place in unidirectional feed direction without traveling around the workpiece only by up-milling and without traveling around the workpiece. This means that the tool was not lifted (Z-axis is not moving) to minimize positioning errors, but only moved in the X-Y plane. The milling was performed parallel to the 75-mm long side of the sample with ap= 7μm and ae= 30% D. This procedure mitigates the nega-tive influence of an alternating direction, which causes a regular change between up- and down-milling, as well as the lack of Z-axis motion [35]. The roughness measured along the feed direction was Ra= 0.04 ± 0.01μm and Rz= 0.36 ± 0.06 μm, and the roughness perpendicular to the feed direction was Ra= 0.23 ± 0.06 μm and Rz= 3.38 ± 0.88 μm (each roughness measurement was Gauss-filtered with a cutoff of 0.008 mm).
Fig. 7 Work steps of the CAD/ CAM system
Chips and burrs arising during machining have to be re-moved from the workpiece before using it as a LOC (post-processing). In addition, the presence of chips and burrs hin-ders the measurement of the groove by a confocal microscope and the assessment of the groove flanks by SEM. For cleaning the machined microstructures, no chemical solution (usually with acetone) could be used, because the surface would be-come chemically attacked. Instead, ultrasonic cleaning with water was performed.
The cavitation caused by the ultrasonic cleaning can contrib-ute to the minimization of burrs/burr removal. Here, the time in contact with the cleaning medium and the aspect ratio of the ribs (hs:ws) are the main variables for this process. Figure8 summarizes the characteristic behavior of PMMA 7N in the ultrasonic cleaning with water at 35 kHz. While burrs and chips are not removed after a short exposure time tr, depending on the aspect ratio of the ribs and the duration of the ultrasonic cleaning, delamination may occur. Figure8 also presents the resulting aspect ratio of the ribs as a function of cleaning time. A structure with ws= 50 μm and hs= 50 μm could be micromilled and cleaned with an exposure time of ultrasonic cleaning of tr= 5 min. A structure with ws= 50μm and h = 120μm would be destroyed after the same duration time (tr= 5 min), because of the excessive oscillation of the structure due to the bigger aspect ratio of the ribs (hs:ws= 2.4 instead of 1).
After the ultrasonic bath, the samples were examined by the confocal microscope (type μsurf explorer, NanoFocus AG). Since the possible resolution increases with increasing magnification, a lens of type 260 XS (× 60/0.9 NA) was used for the measurements. To achieve the numerical aperture (NA) of 0.9 and to get enough light reflected from the measurement object into the lens, the working distance of the lens was 0.4 mm. The lateral resolution was about 0.5μm, and the vertical resolution was 2 nm in a measurement field of 260 × 260μm2. To analyze the obtained points, the software MountainsMap, from Digital Surf, was used. The roughness parameters were measured parallel to the feed direction.
Due to the dimensions (D≤ 50 μm, RSm << 0.04 mm) and since the measurements were not carried out with a stylus profilometer [36], the arithmetic average roughness Raand the average roughness Rzper roughness profile were calculat-ed with an 8-μm Gaussian filter. For the study of the rough-ness in the groove bottom, an area was defined in the middle of the groove, which extended at least about 70% of the groove width, independent of the cutter diameter. The size of this area and the deduced number of roughness profiles therefore depended on the cutter diameter. When D = 48μm was used, the area (width by length) was about 46 × 90μm2 and 90 roughness profiles were extracted. When machining with a tool D = 20μm, the area was 15 × 90 μm2and 30 roughness profiles were collected. Those roughness profiles were used to ensure a representative roughness characteriza-tion of the groove over the entire width.
By cleaning the samples, it was possible to characterize the secondary burrs at the flanks (Fig.2, marks F-GG and Sb-F-GL) hereafter simply called GG and GL. The burr height was calculated as the average of about 509 profile sections, each one with 174 points collected over the entire measure-ment field (260 × 260μm2). This average value was used as the evaluation parameter for the burr formation.
The groove depth is measured perpendicular to the feed direction and corresponds to the distance between the face-milled surface and the middle of the groove measurement to take account of the roundness of the groove bottom, which is a function of the spindle’s tilt angle. The groove width is also measured on the groove bottom, due to burr formation and possible erosion of the workpiece’s outer edges (Fig.1, with a < 0μm).
4 Influence of the cutter geometry
The influence of the cutter geometry on the resulting work was investigated for D = 20 μm, λ = 0°, χ´r= 12°; D = 48μm, λ = 0°, χ´r= 12°; and D = 48μm, λ = 30°, χ´r= 1°. Figure9shows the groove width, the groove depth, the rough-ness parameters Raand Rz, and the burr height, depending on the feed per tooth. When compared to the nominal dimen-sions, the results of the groove width and depth present the deviations of the flight circle of the microtool’s cutting edge and the depth of cut. The existing differences in groove width were classified as non-critical for the tool application. This was based on the results from machining test structures and the compensation of the deviation (Deff≠ D) by the CAM environment. However, when compared to the nominal di-mension, the deviations of groove width indicate a potential of improvement regarding the face milling of the sample.
The Ra and Rzvalues of samples machined by χ´r= 12° tools increased continuously as a function of fz. When D = 20 μm, Ra and Rzvalues are larger than with D = 48 μm, especially for high feed rates. This is due to the ratio of vc/fz and R/fz at D = 20μm. These results are in agreement with those presented in the section of preliminary investigation. Reducing the ratio vc/fzleads to embrittlement of PMMA dur-ing machindur-ing [37], whereby the roughness parameters in-crease. In addition, the ratio R/fzgets smaller, which increases the probability of contact between the peripheral face flank of the cutter Aαrand the workpiece, resulting in a poorer ma-chining quality.
When compared to the cutter geometry {D = 48μm, χr´ = 12°, λ = 0}, the cutter geometry {D = 48 μm, χr´ = 1°, λ = 30°} is expected to favor the chip formation and the chip removal by its helix angle. Besides, it is also expected to improve the roughness parameters Raand Rzat the groove bottom by the angle χr´. This cutter represents therefore an adapted tool geometry for machining PMMA. As predicted by
the kinematic roughness, the Ra and Rzvalues of samples machined byλ = 30° with χr´ = 1° tools were significantly better than those of samples machined byλ = 0° with χr´ = 12° tools. The helix angle also increased the passive force.
Regarding burr formation, the dependence of the burr height on the feed rate was much stronger when D = 48μm, χr´ = 1°,λ = 30° than with D = 48 μm, χr´ = 12°,λ = 0° tools. Atλ = 30° and fz= 1μm, the height of burrs was in range of
Fig. 9 Influence of the cutter geometry on the surface quality Fig. 8 Effect of ultrasonic baths
on micromilled structures of PMMA
h0= 3μm and thus did not meet the requirements adopted. At higher feed rates, the heights of burrs were below h0= 3μm and also, the standard deviation was smaller, indicating a more stable cutting behavior at higher feed rates. This confirms the behavior of the Poisson burr described by Weinert and Petzoldt [26]. While at smaller feed rates, the burr root thick-ness was greater than the feed per tooth, a total removal could be reached at higher feed rates.
It must be recalled that the behavior presented here refers to unworn tools. By tool wear, the burr formation would be more clearly pronounced atλ = 30°. The devel-oped tools with λ = 30° require the increase of the feed per tooth to meet the requirements regarding the burrs. Smaller spindle speeds and consequently lower vffor an increase of the feed per tooth were not set due to the rotor dynamic behavior at 30,000 rpm, which was the best level for the micromilling machine used. Finally, smaller
spindle speeds would have negative effects on the mate-rial removal rate Qw, resulting in reduced efficiency of the micromilling process. Besides, machining with λ = 30° at 30,000 rpm requires a feed rate of vf> 120 mm/min to minimize the burr height.
In addition to the machining of linear channels, LOCs include user-specific structures, such as meanders and mags, which require synchronized movement of the axes. Although a feed rate of vf> 120 mm/min was set in the NC code, in practice this value was not reached due to dynamic limits while machining complex structures by interpolation of the axes and short distances to accelerate (< 0.5 mm). The structure was effectively machined with fz< 4 μm, which has a negative effect on the work results. Since tools without a helix angle (λ = 0°) meet the requirements, re-garding both burr formation and roughness parameters, they have been preferred.
Fig. 10 a, b Groove structure of a microfluidic droplet obtained by direct micromilling
5 Machining strategies and results
This section explains the benefits and feasibility of the devel-oped concept based on microfluidic structures. Two examples illustrate that small radii and structural dimensions can be produced by direct milling.
The structures described in text and as CAD models here were not machined by a conventional processing sequence. In order to avoid a tool change during the production of the structure, first the contour of the structure has been machined. Thus, initially the flanks of the structure were machined by slot milling with sharp and unworn microtools to reduce burr formation. At the same time, the effective tool diameter started in unworn condition and decreases during processing.
Whereas the effective cutter diameter has a relevant impact on the structure contour and hence on the resulting dimensions during the milling process of internal regions, a minor influ-ence was observed by the oscillations of the ratio Deff/ae. Here, the engagement width aewas set at 30% of the cutter diameter. In addition, the use of worn tools still meets the requirements of roughness on the groove bottom. By this ma-chining strategy, a good work result could be achieved, re-garding burr formation on the upper structure edges, as well as uniform roughness values at both flanks and bottom.
The groove structure presented in Fig.10consists of a neck used in microfluidic droplets. It was machined by a single tool with D = 20 μm. The amount of removed material was 0.00427 mm3. The insert at the bottom left of the figure shows a machined groove structure with a“flow focus” arrangement for the separation of individual fluid drops.
Figure11presents a machined structure used for mixing several fluids. The structure with an aspect ratio AR 1:1 was completely machined by slot milling with D = 48μm and a-p= 25μm. In some areas with circular interpolation, motion marks can be observed on the flank, whereas no motion marks occurred in linear motion. The structure meets the quality requirements on flank and groove bottom. The amount of material removed was about 0.0425 mm3and thus, far below the evaluated benchmark for the tool life volume at fz= 2μm, D = 48 μm, χr´ = 12°, λ = 0° of volumes greater than 0.45 mm3.
6 Conclusions and outlook
By defining the mechanisms of burr formation according to the burr location, a holistic view of optimization strategies to avoid burrs in micro end milling could be provided. The effects of ultrasonic bath and the aspect ratio of the rib as function of the ultrasonic cleaning exposure time described herein are guidelines to set cleaning parameters. A phe-nomenological classification of the chip formation was al-so presented in detail.
For the micromilling of PMMA with tools D = 48μm, the spindle used already provides sufficient cutting speed at 7500 rpm. To achieve low roughness values and low burr formation, the effects of axial and radial runout are more im-portant than of using higher cutting speeds. For this reason, for D = 48μm, the spindle speed was not set to 50,000 rpm, but to 30,000 rpm with minimal axial (< 1μm) and radial (< 3 μm) runout. For D = 20μm, the cutting speed was crucial, and 50,000 rpm was preferred.
At constant spindle speed, an increasing feed per tooth fz leads to the slope of the kinematic roughness with a trend towards reduction of burr formation. This is in agreement with the Weinert and Petzoldt model for Poisson burr formation [26].
The results show that although the tool D = 48μm; χ´r= 1°; λ = 30° leads to better surface roughness on the bottom surface, the resulting burr formation is more pronounced. In principle, this effect can be compensated by using a higher feed per tooth. However, in most cases, this higher feed per tooth cannot be reached when milling a microfluidic structure. Since the process often leads to a periodic acceleration and stopping in shorter distances and consequently the pro-grammed feed rate is not achieved, the positive properties of χ´r= 1°;λ = 30° are ultimately not exploited. The structures described here were not machined by the conventional pro-cessing sequence from the inside out, but from the outside in, leading to better side walls, with less burrs, and better bottom roughness.
Finally, based on the available knowledge, microfluidic r e l e v a n t s t r u c t u r e s w e r e s e l e c t e d a n d m a c h i n e d . Accordingly, the entire chain—from the CAD data file up to the microfluidic structures for prototypical applications—was presented and discussed.
The chosen micro end mills and milling strategies represent a competitive process at the prototyping level by reducing the time-to-market.
After successful machining of microfluidic structures, the next step is to validate the findings and procedure for another machines and materials.
Funding information This research was funded by the German Research Foundation (DFG) within the Collaborative Research Center 926 “Microscale Morphology of Component Surfaces.”
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