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Technical Brief

# Piezo-Actuated Modulation-Assisted Drilling System With Integrated Force SensingOPEN ACCESS

[+] Author and Article Information
Yang Guo, Seong Eyl Lee

M4 Sciences LLC,
West Lafayette, IN 47906

James B. Mann

M4 Sciences LLC,
Purdue University,
West Lafayette, IN 47907
e-mail: jbmann@m4sciences.com

1Corresponding author.

Manuscript received January 12, 2016; final manuscript received June 5, 2016; published online August 8, 2016. Assoc. Editor: Tony Schmitz.

J. Manuf. Sci. Eng 139(1), 014501 (Aug 08, 2016) (7 pages) Paper No: MANU-16-1033; doi: 10.1115/1.4033929 History: Received January 12, 2016; Revised June 05, 2016

## Abstract

A new electromechanical modulation system designed with piezoelectric stacks for both linear actuation and force sensing functions is described. The system can be adapted for modulation-assisted machining (MAM) drilling processes where a low-frequency (<1000 Hz) sinusoidal oscillation is superimposed directly onto the drilling process, such that the feedrate is modulated. A series of drilling experiments were conducted in Ti6Al4V, 17-4 steel, and Al6061 with the system installed on a CNC machine. The drill displacement, thrust force, and chip morphology were characterized across a range of conventional and MAM drilling conditions. The mechanical response (stiffness) of the system agrees with the design specifications. The system offers new capabilities to control the modulation frequency and amplitude in MAM drilling, while simultaneously measuring the drilling thrust force in real time. The force sensing function enables detection of the intermittent separations between the drill tip and the workpiece surface (occurrence of discrete cutting), providing a method to prescribe and control the modulation conditions necessary for effective MAM drilling. Opportunities for force feedback control and process monitoring in MAM drilling processes are discussed. While the system described emphasizes MAM drilling, the capabilities can be extended to other machining processes.

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## Introduction

Mechanical drilling is one of the most common machining operations and consumes half of the cutting tools used in all the chip-making processes [1]. Enhancing the efficiency and the controllability of mechanical drilling process has implications for both cost reduction and quality improvement across a wide range of manufacturing sectors.

In drilling operations, the cutting action and chip formation occur at the bottom of the hole, and evacuating chips away from the cutting zone is complicated. The difficulty increases as the length (depth) of the drilled hole increases. Without proper evacuation, the drilled chips can bind or jam in the drill flute(s) and inside the drilled hole. This can damage the hole, increase drilling force and torque, and cause drill breakage [2]. The ability to control chip formation and evacuation are primary considerations for improving drilling processes.

Modulation-assisted drilling (MAM drilling) has shown remarkable improvements in drilling performance by altering the chip formation process. In MAM drilling, a low-frequency (<1000 Hz) sinusoidal oscillation is superimposed onto the drill feed motion. This transforms the otherwise continuous cutting into a series of discrete cutting events. Each cutting event creates a discrete chip that can be easily evacuated through the drill flute(s). The intermittent cutting also leads to enhanced effectiveness of cutting fluids and increased cutting tool life [3]. MAM drilling is receiving increasing attention as the solution for complex drilling operations in steel alloys, titanium alloys [4], composite materials [5], and composite metal stacks [6,7]. In addition, drilling with MAM may have benefits in unique applications, such as reducing temperature in drilling of bone [8] or alternative process cooling methods, such as heat-pipe assisted drilling [9].

The development of suitable modulation system is critical for the application of MAM drilling. A range of systems can realize the motion required for MAM drilling, including piezoelectric actuators [10], linear motor drives [4,11], electromagnetic shakers [5], self-excited vibrating structure [12], and bearings with wavy race surfaces [13]. In order for MAM drilling to be effective, the frequency and amplitude of the oscillation need to be correctly prescribed based on machining process variables, i.e., rotational speed and feedrate. A modulation system that can provide the required oscillation and simultaneously measure thrust force will be advantageous for monitoring and controlling the drilling process [14]. The force sensing could offer unique benefits toward controlling and optimizing the modulation conditions in MAM drilling processes. This paper demonstrates the design and experimental testing of a piezo-actuated modulation system with integrated force sensing capabilities.

## Background

MAM superimposes low-frequency sinusoidal oscillation (<1000 Hz) onto the tool feed motion, and the conventional feedrate, ho (feed/rev), is modulated as ho + Asin(2πfmt), where fm and A are the oscillation frequency and amplitude, respectively. Under the appropriate oscillation conditions, the mechanics of the chip formation process are altered. The successive cutting paths of the tool cutting edge(s) can occur out-of-phase, resulting in a variation of the undeformed chip thickness (Fig. 1(a)). If the oscillation amplitude is sufficiently high compared to the feedrate, the successive cutting paths will intersect and the cutting is divided into a series of discrete events that ensure discrete chip formation.

The phase difference between the successive cutting paths determines the minimum threshold oscillation amplitude required to achieve discrete cutting. A kinematic model of MAM describing these discrete cutting conditions has already been established for turning or drilling [15,16]. Based on this model, the minimum amplitude threshold occurs at successive cutting paths with 180 deg phase difference (Fig. 1(a)). The frequency and amplitude conditions for discrete cutting corresponding to 180 deg phase difference are given as Display Formula

(1)$fm=nc⋅fr⋅2N+12 N=0, 1, 2,…$
Display Formula
(2)$App≥honc$

where fm is the oscillation frequency, fr is the rotational frequency (rpm/60), nc is the number of cutting edges of the tool, App is the peak-to-peak oscillation amplitude (=2A), and ho is the tool feed per revolution (feedrate). (Note that App will mainly be used and simply be referred to as modulation/oscillation amplitude in the rest of the paper.) To achieve discrete cutting, the oscillation amplitude needs to exceed the conventional undeformed chip thickness (ho/nc).

Other phase difference will require higher oscillation amplitude to achieve discrete cutting. More importantly, discrete cutting is not possible at 0 deg phase difference, since the successive cutting paths are parallel (Fig. 1(b)). Effective MAM drilling can only occur when modulation frequency and amplitude satisfy the discrete cutting conditions.

## Piezo-Actuated Modulation System

Figure 2 shows the schematic of the piezo-actuated modulation system adapted from Ref. [17]. The new system can superimpose an axial oscillation to the drill tool and simultaneously measure the force in the drilling direction (thrust force). Two separate piezostacks (PZT type) of different lengths are aligned axially and preloaded compressively against a moving end by a disk spring of stiffness ks. The preload force (F0) should be sufficient to ensure that both piezostacks are always operating under a compressive stress during drilling. The long piezostack is the linear actuator and generates the axial displacement (u) of the moving end from its preload-balanced position. The displacement is directly proportional to the drive voltage of the piezo-actuator. The stiffness of the piezo-actuator is ka. The short/thin piezostack is used as a sensor to measure the axial force (excluding the preload) exerted on the piezo-actuator (Fa). The force-induced charges on the piezosensor can be converted by a charge amplifier to the voltage signal that is proportional to the applied force. The moving end of the modulation system is supported and guided by a ball spline bearing which allows the linear motion but prevents rotation of the moving end. The drill is rigidly connected to the moving end of the system, and they translate together as a single rigid mass (m).

When the piezo-actuator is not activated (zero drive voltage), a thrust force (Ft) applied on the front of the system will cause a negative displacement of the moving end of the system. This displacement can be related to the thrust force as Display Formula

(3)$Ft=(ks+ka)(−u)$

This shows that the axial stiffness of the system is ks + ka, which is the sum of the stiffness of the spring and the stiffness of the piezo-actuator. In addition, this displacement can also be related to the force exerted on the piezo-actuator (Fa) as Display Formula

(4)$Fa=ka(−u)$

From Eqs. (3) and (4), a relation between Ft and Fa can be obtained as Display Formula

(5)$FtFa=(1+kska)$

Equation (5) is valid only when the piezo-actuator is not activated. This relationship is useful for calibrating the piezoforce sensor.

When the piezo-actuator is activated in MAM drilling, the equation of motion for the oscillating mass (drill and moving end) can be written as Display Formula

(6)$mu¨+bu˙+ksu+Ft−Fa=0$

where $mu¨$ and $bu˙$ are the inertial force and damping force, respectively; ksu is the spring force; and Fa and Ft are the piezo-actuator force and drilling thrust force, respectively. For oscillation at low frequency with small mass and low friction, the inertial and damping terms in Eq. (6) can be neglected compared to the thrust force. Then, Eq. (6) can be simplified to Display Formula

(7)$Ft=Fa−ksu$

Therefore, Ft can be derived from the measurements of Fa and u. However, if the inertial force is significant compared to the thrust force, it cannot be neglected in Eq. (7). Then, a measurement of the inertial force is necessary to accurately obtain the thrust force.

There are certain constraints that define the scope of the operation of the modulation system. First, the inertial force resulting from the oscillating mass (m) must be smaller than the preload (F0) of the system, such that the piezostack actuator and sensor are always under compression from the moving end. This gives Display Formula

(8)$2π2fm2Appm

Second, the controlled modulation should be in the nonresonant regime, which typically requires the modulation frequency be set below one-half of the resonant frequency of the system. This gives Display Formula

(9)$fm<14πks+kam$

Equations (8) and (9) must be satisfied to ensure the proper function of the system. They serve as important criteria in the design of the system.

A modulation system based on the concept illustrated in Fig. 2 was constructed for experimental testing. The key specifications of this system are summarized in Table 1. For this system, the preload spring has a stiffness of ks = 1.7 N/μm; the piezo-actuator has a stiffness of ka = 15 N/μm; and a preload force of F0 = 300 N is applied. The oscillation amplitude of the moving end was calibrated with a sinusoidal drive voltage applied to the piezo-actuator across a range of frequencies with no thrust load applied (Ft = 0). In the calibration, the displacement of the moving end was directly measured by a capacitive probe. The results are shown in Fig. 3. The amplitude of oscillation is linearly dependent on the drive voltage amplitude and is negligibly affected by the oscillation frequency. The maximum drive voltage allowable for the piezo-actuator is 150 V, which corresponds to the maximum displacement of u ≈ 60 μm on the moving end of the system. The piezoforce sensor was also calibrated after it was integrated in the modulation system. For the calibration, the system was configured by pressing the moving end against a load washer (Kistler 9021) (Fig. 4(a)). The thrust force (Ft) was measured directly by the load washer, and the resulting force on the piezoforce sensor (Fa) was determined based on Eq. (5). A customized charge amplifier circuit was built to convert the charges on the piezosensor to the voltage signal. The correlation between the exerted force and the generated voltage signal is shown in Fig. 4(b). The linear regression yields the force–voltage relationship for the piezoforce sensor corresponding to 20.1 mV/N.

The oscillating mass of the system is 90 g. Assuming the oscillation amplitude App is at the maximum (60 μm), Eqs. (8) and (9) give the limiting frequency of the system of 1678 Hz and 1085 Hz, respectively.

## Experimental Drilling

The modulation system was installed in a CNC lathe (Miyano BNC-42C) to perform a series of centerline drilling tests with a rotating workpiece (stationary nonrotating drill). These tests included both conventional drilling (Table 2) and MAM drilling (Table 3) of Ti6Al4V, 17-4 stainless steel, and Al6061. All the drilling tests were performed using a 2.6 mm diameter solid carbide twist drill (Mitsubishi MZS0260LB, nc = 2) with application of high-pressure cutting fluid through the drill (Blaser Vascomill 10 at 85 bar). The rotational speed was 3600 rpm (fr = 60 Hz), and the drilled depth was 10 mm for all the tests.

The oscillation of the drill was controlled by a sinusoidal drive voltage applied to the piezo-actuator by an external power amplifier and sinewave generator. A capacitive displacement probe (Capacitec, Ayer, MA) was used to measure the displacement u of the drill/moving end of the modulation system. The piezo-actuator force Fa was measured directly by the internal piezoforce sensor. Both u and Fa signals were collected synchronously via an analog to digital data-acquisition system (National Instrument: NI USB 6229 + LabView) at the sampling rate of 10,000 Hz. The oscillating mass of the system was 90 g. For the modulation conditions listed in Table 3, the inertial force (=2π2fm2Appm) was estimated to a maximum of 3.2 N. This justifies the use of Eq. (7) for the derivation of Ft.

## Results and Discussion

The purpose of the experimental drilling is to test the design of the modulation system and its integrated force sensing function. The measurement results from the drilling tests will demonstrate the modulation and real-time force sensing capabilities of the system.

Figure 5 shows the measured displacement, thrust force, and the resulting chip morphology for drilling Ti6Al4V at four different conditions: (a) conventional drilling (test C1), (b) in-phase MAM drilling (test M1), (c) out-of-phase MAM drilling with insufficient amplitude (test M2), and (d) out-of-phase MAM drilling with sufficient amplitude (test M5). These tests were conducted with the same drill feedrate (ho = 40 μm/rev). The complete u and Ft data are shown in the left column of Fig. 5. The drilling occurs during t = 0–4 s. Note that the nonzero Ft before and after the MAM drilling (t < 0 s and t > 4 s) in Fig. 5(b)5(d) indicates the inertial and damping forces related to the drill oscillation. Indeed, the inertial and damping forces are negligibly small compared to the drilling thrust force. The magnified variations of u and Ft at the midpoint of the drilling cycle (1.98 s < t < 2.02 s) are shown in the center column of Fig. 5. The corresponding drilling chips are shown in the right column of Fig. 5.

In conventional drilling (Fig. 5(a)), the thrust force is steady at about 209 N. The thrust force causes a negative (compressive) displacement of approximately 12.6 μm, reflecting the finite stiffness of the system. The chips created in conventional drilling are long and continuous. The chips also appear to be folded at a regular interval that is on the same order as the rotational frequency of the drilling process.

For in-phase MAM drilling (Fig. 5(b)), the modulation amplitude is 43 μm, and the thrust force varies periodically from 192 N to 303 N. Since the undeformed chip thickness remains constant during in-phase MAM drilling, the variation of the thrust force is mainly caused by the changes in the relative feed velocity and effective tool rake angle due to modulation of the feedrate. The drilling chips at this condition are continuous and appear analogous to the chips formed in conventional drilling.

Figure 5(c) shows an out-of-phase MAM drilling with the modulation amplitude being 21 μm and the thrust force ranging from 124 N to 301 N. Compared to the in-phase MAM drilling, the thrust force varies in a greater range despite a smaller modulation amplitude. In the out-of-phase MAM drilling, the thrust force is primarily influenced by the periodic changing of the undeformed chip thickness, and secondarily by the changing of feed velocity and effective rake angle. The modulation amplitude in this case is not sufficiently high to achieve discrete cutting. The measured thrust force is always greater than zero, indicating continuous engagement between the drill and the workpiece surface. The cutting is continuous. However, since the undeformed chip thickness varies during each cycle of modulation, thinner (or weaker) sections occur during chip formation. These thin sections can randomly break apart during chip evacuation, so the chips appear discontinuous and variable in size (length).

Figure 5(d) shows an out-of-phase MAM drilling with the modulation amplitude being 36 μm and the thrust force varying from −4 N to 359 N. The modulation amplitude is sufficiently high for discrete cutting to occur. This is evidenced by the measured thrust force which periodically reaches zero during each cycle of modulation, indicating the disengagement between the drill and the workpiece surface. Due to the discrete cutting process, the chips created are not only completely discrete but also of the same size and shape. Such chip morphologies can lead to more reliable chip evacuation and hence improvement of drilling performance.

Additional drilling tests in 17-4 steel and Al6061 alloys show similar characteristics of displacement and force measurements as observed in drilling of titanium alloy. With sufficient modulation amplitude, discrete cutting can be achieved in out-of-phase MAM drilling in these materials. Figure 6 shows the comparison of thrust forces in conventional drilling and corresponding MAM drilling when discrete cutting is achieved for several drilling cases (tests C1–C6 versus tests M5, M8, M11, M14, M17, and M19). The maximum thrust force in MAM drilling is 70% higher for Ti6Al4V, 130% higher for 17-4 steel, and 50% higher for Al606. The mean thrust force in MAM drilling is, however, 10% lower for Ti6Al4V, 30% lower for Al6061 but still 30% higher for 17-4 steel. Among the three different materials, drilling in Ti6Al4V results in the highest thrust force while drilling in Al6061 results in the lowest thrust force.

The measured displacement and force data from the drilling tests provide a practical way to characterize the modulation system. Figure 7(a) shows the relationship between the thrust force and the displacement for conventional drilling tests (data from Table 2). The linear regression indicates a system stiffness of 16.6 N/μm. In MAM drilling, the oscillation amplitude is always reduced under the dynamic thrust load compared to the zero load condition, i.e., App < App0 (see Fig. 5 and Table 3). Again, this is due to the compliance of the modulation system. The variation of the thrust force in MAM drilling results in alternating elastic deformation and recovery of the system, which has the effect of reducing the amplitude of the oscillation delivered by the system. Figure 7(b) shows the relationship between the range of thrust force and the reduction in oscillation amplitude for MAM drilling tests (data from Table 3). The linear regression indicates a system stiffness of 16.0 N/μm. The stiffness of the system deduced from both conventional drilling tests and MAM drilling tests is consistent and in agreement with the design specification (ks + ka = 16.7 N/μm). It should be noted that the reduction in oscillation amplitude in MAM drilling is caused by the static stiffness instead of the dynamic stiffness of the system.

The amplitude reduction in MAM drilling may pose some complications in setting the drive voltage of the piezo-actuator to the appropriate level needed to achieve discrete cutting. The effective oscillation amplitude delivered by the system is dependent on the thrust force during the MAM drilling process, which cannot be precisely predicted ahead of time. Therefore, it is desirable to design the modulation system with sufficient stiffness for targeted drilling applications in order to minimize the amplitude reduction. The piezostack linear actuator is the most compliant component in the system, so modifications to the piezostack actuator will have the most influence on the stiffness. Typically, piezostacks with larger cross-sectional area and shorter length will increase the stiffness; however, these features also influence the characteristic displacement and force capacity of a particular piezostack actuator. Since infinite stiffness is not realistic in the mechanical system, the reduction of amplitude that occurs due to system compliance cannot be completely eliminated. The amplitude reduction has also been noted in other types of modulation systems [18].

The force measuring capability of the system provides a direct method for determining when discrete cutting occurs in MAM drilling. Specifically, when the lower bound of the thrust force reaches zero, discrete cutting is ensured irrespective of the workpiece material. Figure 8 shows the variation of the lower bound of the thrust force with the modulation amplitude for MAM drilling in four different cases (tests M3–M5, M9–M11, M15–M17, and M12–M14). In each case, the lower bound of the thrust force decreases with increasing modulation amplitude. Eventually, the lower bound of the thrust force reaches zero when the modulation amplitude is increased above a certain threshold. This is reflected by data points a–d in each drilling case. The chips corresponding to these points are shown in Fig. 8. These chips of nearly the same discrete morphology in each of the materials confirm the occurrence of discrete cutting.

The process of adjusting the modulation amplitude to achieve discrete cutting can be automated by implementing a real-time force feedback control on the modulation amplitude. This should be readily achievable on the current modulation system with the force sensing capability by using the voltage signal from the piezoforce sensor to control the drive voltage of the piezo-actuator. The force feedback control will be able to address the complication arising from amplitude reduction in MAM drilling due to system compliance and ensure the effective MAM condition to be implemented. This will simplify the application of MAM drilling technology.

## Implications

The piezo-actuated electromechanical modulation system is a practical configuration for industrial implementation of MAM drilling. The system provides a method of directly controlling the modulation conditions in conjunction with a real-time force sensing function. The new modulation system offers unique opportunities for real-time force feedback control and process monitoring in MAM drilling processes. More importantly, the design will enable extensive investigations of drilling with MAM (and simultaneous measurement of thrust force) that will be necessary for the development of a related process control algorithm.

The force sensing capability of the system also enables opportunities for drilling process monitoring and optimization. The related signal processing from the force sensor can be used for a drilling performance optimization routine, assessing tool wear, or detecting drill breakage. In addition, the force-based control can be extended to the machine tool controller, enabling rotational speed and feedrate parameters of the drilling process to be adaptively controlled [19] and optimized based on process thrust force.

The system described is focused on applications in drilling with modulation; however, the configuration can also be useful in a range of conventional machining processes (e.g., drilling, turning, boring, and reaming), where the direct measurement of thrust force could be beneficial for process optimization or monitoring tool breakage, tool wear, or machining process stability.

## Conclusions

This paper has shown a fundamental design approach and configuration with related methodology for implementing a piezomechanical system in MAM. The experimental testing of the system based on a series of drilling tests shows that the mechanical response of the system is consistent with the design. Importantly, the force sensing function of the system is capable of measuring the drilling thrust force and enables the detection of the intermittent separations between the drill tip and the workpiece surface (or the occurrence of discrete cutting). This provides a direct method for prescribing and controlling the oscillation amplitude required to achieve effective MAM drilling. A future study in drilling with the piezo-actuated modulation system will investigate the effects of machining process parameters, including materials, drill size, drilling speed, and feedrate.

## Acknowledgements

This work was supported by the National Science Foundation Grant No. 0822879-IIP (M4 Sciences).

## Nomenclature

• App =

peak-to-peak modulation amplitude during drilling (μm)

• $App0$ =

• fm =

modulation frequency (Hz)

• fr =

workpiece rotational frequency (Hz)

• Fa =

piezo-actuator force (N)

• Ft =

drilling thrust force (N)

• F0 =

• ho =

feed per revolution (μm/rev)

• ka =

stiffness of piezo-actuator (N/μm)

• ks =

• nc =

number of cutting edges

• u =

axial displacement of moving end of modulation system (μm)

• Vpp =

peak-to-peak amplitude of drive voltage (V)

## References

SME, 2010, Fundamental Manufacturing Processes Study Guide, DV09PUB3, Society of Manufacturing Engineers, Dearborn, MI.
Shaw, M. C. , 1984, Metal Cutting Principles, Clarendon Press, Oxford, UK.
Guo, Y. , Mann, J. B. , Yeung, H. , and Chandrasekar, S. , 2012, “ Enhancing Tool Life in High-Speed Machining of Compacted Graphite Iron (CGI) Using Controlled Modulation,” Tribol. Lett., 47(1), pp. 103–111.
Okamura, K. , Sasahara, H. , Segawa, T. , and Tsutsumi, M. , 2006, “ Low-Frequency Vibration Drilling of Titanium Alloy,” JSME Int. J. Ser. C., 49(1), pp. 76–82.
Sadek, A. , Attia, M. H. , Meshreki, M. , and Shi, B. , 2013, “ Characterization and Optimization of Vibration-Assisted Drilling of Fibre Reinforced Epoxy Laminates,” CIRP Ann. Manuf. Technol., 62(1), pp. 91–94.
Pecat, O. , and Brinksmeier, E. , 2014, “ Tool Wear Analyses in Low Frequency Vibration Assisted Drilling of CFRP/Ti6Al4V Stack Material,” Procedia CIRP, 14, pp. 142–147.
Park, K.-H. , Beal, A. , Kim, D.-W. , Kwon, P. , and Lantrip, J. , 2013, “ A Comparative Study of Carbide Tools in Drilling of CFRP and CFRP-Ti Stacks,” ASME J. Manuf. Sci. Eng., 136(1), p. 014501.
Sui, J. , Sugita, N. , and Mitsuishi, M. , 2015, “ Thermal Modeling of Temperature Rise for Bone Drilling With Experimental Validation,” ASME J. Manuf. Sci. Eng., 137(6), p. 061008.
Zhu, L. , Jen, T.-C. , Liu, Y.-B. , Zhao, J.-W. , Liu, W.-L. , and Yen, Y.-H. , 2014, “ Cutting Tool Life Analysis in Heat-Pipe Assisted Drilling Operations,” ASME J. Manuf. Sci. Eng., 137(1), p. 011008.
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Mann, J. B. , Saldana, C. J. , Guo, Y. , Yeung, H. , Compton, W. D. , and Chandrasekar, S. , 2012, “ Effects of Controlled Modulation on Surface Textures in Deep-Hole Drilling,” SAE Int. J. Mater. Manuf., 6(1), pp. 24–32.
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## References

SME, 2010, Fundamental Manufacturing Processes Study Guide, DV09PUB3, Society of Manufacturing Engineers, Dearborn, MI.
Shaw, M. C. , 1984, Metal Cutting Principles, Clarendon Press, Oxford, UK.
Guo, Y. , Mann, J. B. , Yeung, H. , and Chandrasekar, S. , 2012, “ Enhancing Tool Life in High-Speed Machining of Compacted Graphite Iron (CGI) Using Controlled Modulation,” Tribol. Lett., 47(1), pp. 103–111.
Okamura, K. , Sasahara, H. , Segawa, T. , and Tsutsumi, M. , 2006, “ Low-Frequency Vibration Drilling of Titanium Alloy,” JSME Int. J. Ser. C., 49(1), pp. 76–82.
Sadek, A. , Attia, M. H. , Meshreki, M. , and Shi, B. , 2013, “ Characterization and Optimization of Vibration-Assisted Drilling of Fibre Reinforced Epoxy Laminates,” CIRP Ann. Manuf. Technol., 62(1), pp. 91–94.
Pecat, O. , and Brinksmeier, E. , 2014, “ Tool Wear Analyses in Low Frequency Vibration Assisted Drilling of CFRP/Ti6Al4V Stack Material,” Procedia CIRP, 14, pp. 142–147.
Park, K.-H. , Beal, A. , Kim, D.-W. , Kwon, P. , and Lantrip, J. , 2013, “ A Comparative Study of Carbide Tools in Drilling of CFRP and CFRP-Ti Stacks,” ASME J. Manuf. Sci. Eng., 136(1), p. 014501.
Sui, J. , Sugita, N. , and Mitsuishi, M. , 2015, “ Thermal Modeling of Temperature Rise for Bone Drilling With Experimental Validation,” ASME J. Manuf. Sci. Eng., 137(6), p. 061008.
Zhu, L. , Jen, T.-C. , Liu, Y.-B. , Zhao, J.-W. , Liu, W.-L. , and Yen, Y.-H. , 2014, “ Cutting Tool Life Analysis in Heat-Pipe Assisted Drilling Operations,” ASME J. Manuf. Sci. Eng., 137(1), p. 011008.
Toews, H. G. , Compton, W. D. , and Chandrasekar, S. , 1998, “ A Study of the Influence of Superimposed Low-Frequency Modulation on the Drilling Process,” Precis. Eng., 22(1), pp. 1–9.
Chhabra, P. N. , Ackroyd, B. , Compton, W. D. , and Chandrasekar, S. , 2002, “ Low-Frequency Modulation-Assisted Drilling Using Linear Drives,” Proc. Inst. Mech. Eng., Part B, 216(3), pp. 321–330.
Guibert, N. , Paris, H. , and Rech, J. , 2008, “ A Numerical Simulator to Predict the Dynamical Behavior of the Self-Vibratory Drilling Head,” Int. J. Mach. Tools Manuf., 48(6), pp. 644–655.
Jallageas, J. , K’nevez, J.-Y. , Chérif, M. , and Cahuc, O. , 2012, “ Modeling and Optimization of Vibration-Assisted Drilling on Positive Feed Drilling Unit,” Int. J. Adv. Manuf. Technol., 67(5), pp. 1205–1216.
Mann, J. B. , Lee, S. , and Guo, Y. , 2015, “ Tool Holder Actuator and Real-Time Force Monitoring for Machining Process Control,” Provisional U.S. Patent Application, June 17.
Mann, J. B. , Guo, Y. , Saldana, C. , Compton, W. D. , and Chandrasekar, S. , 2011, “ Enhancing Material Removal Processes Using Modulation-Assisted Machining,” Tribol. Int., 44(10), pp. 1225–1235.
Mann, J. B. , Saldana, C. J. , Guo, Y. , Yeung, H. , Compton, W. D. , and Chandrasekar, S. , 2012, “ Effects of Controlled Modulation on Surface Textures in Deep-Hole Drilling,” SAE Int. J. Mater. Manuf., 6(1), pp. 24–32.
Mann, J. B. , Chandrasekar, S. , and Compton, W. D. , 2009, “ Tool Holder Assembly and Method for Modulation-Assisted Machining,” U.S. Patent No. 7,587,965.
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## Figures

Fig. 1

(a) Out-of-phase and (b) in-phase sinusoidal cutting paths in MAM

Fig. 2

Schematic of piezo-actuated modulation system showing the configuration of piezo-actuator, force sensor, preload spring, and moving end. The spring exerts compressive preload.

Fig. 3

Calibration of piezoelectric actuator of the modulation system: drive voltage–displacement relation for the actuator

Fig. 4

Calibration of piezoelectric force sensor of the modulation system: (a) schematic of calibration configuration and (b) force–voltage relation for the sensor and charge amplifier system

Fig. 5

Measured displacement, thrust force, and the resulting chip morphology in drilling Ti6Al4V: (a) conventional drilling (test C1), (b) in-phase MAM drilling (test M1), (c) out-of-phase MAM drilling with insufficient amplitude (test M2), and (d) out-of-phase MAM drilling with sufficient amplitude where discrete cutting occurs (test M5)

Fig. 6

Comparison of thrust force between conventional drilling and MAM drilling at modulation conditions where discrete cutting occurs

Fig. 7

(a) Relationship between the thrust force and the system displacement in conventional drilling and (b) relationship between the range of thrust force and the reduction in oscillation amplitude in MAM drilling. Both reflect system stiffness of approximately 16 N/μm.

Fig. 8

Variation of the lower bound of thrust force (Ft_min) with the oscillation amplitude (App) in MAM drilling and discrete chips corresponding to data points a–d

## Tables

Table 1 Key specifications of the system
Table 2 Conventional drilling tests (fr = 60 Hz and nc = 2)
Table 3 MAM drilling tests (fr = 60 Hz and nc = 2)

## Discussions

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