Development of a modular dynamometer for triaxial cutting force measurement in turning

International Journal of Machine Tools & Manufacture 51 (2011) 34–42

Contents lists available at ScienceDirect

International Journal of Machine Tools & Manufacture

0890-69

doi:10.1

n Corr

E-m

journal homepage: www.elsevier.com/locate/ijmactool

Development of a modular dynamometer for triaxial cutting forcemeasurement in turning

G. Totis n, M. Sortino

DIEGM—Department of Electrical, Management and Mechanical Engineering, University of Udine, Viadelle Scienze, 208, I-33100 Udine, Italy

a r t i c l e i n f o

Article history:

Received 24 November 2009

Received in revised form

28 September 2010

Accepted 1 October 2010Available online 15 October 2010

Keywords:

Turning

Dynamometer

Cutting Forces

Dynamics Identification

55/$ - see front matter & 2010 Elsevier Ltd. A

016/j.ijmachtools.2010.10.001

esponding author. Tel.: +39 0432 558241; fa

ail address: [email protected] (G. Totis).

a b s t r a c t

Accurate and reliable measurement of cutting forces in turning is essential for tool geometry, tool

trajectory and cutting parameters optimization, as well as for tool condition monitoring and machinabilty

testing. In this work, an innovative dynamometer for triaxial cutting force measurement in turning,

specifically designed to be applied at a milling–turning CNC machine tool endowed with an indexable

head, is presented. The device is based on a piezoelectric force ring integrated into a commercial

toolshank, and its modular design allows the easy change of the cutting insert without altering sensor

preload. The prototype device was assembled and experimentally tested by means of static calibration

and dynamic identification, which evidenced good static and dynamic characteristics. Eventually, the

sensor was tested in operating conditions by machining a benchmark workpiece.

& 2010 Elsevier Ltd. All rights reserved.

1. Introduction

Cutting forces are strongly related to cutting process mechanics,and the application of dynamometers for their measurementduring machining is essential for investigating, monitoring andoptimizing the turning process. In the last decades, different typesof dynamometers have been used in industry and researchlaboratories for understanding the principles of chip formation[1], for developing cutting force models [2], as well as for cuttingprocess control [3], tool geometry optimization [4], tool conditionmonitoring—TCM [5–8] and for detection and suppression ofchatter vibrations [9,10].

Generally, dynamometers tend to reduce the machining systemstiffness. Therefore, a trade-off between sensitivity and stiffnesshas to be accepted. Also, the frequency bandwidth of thedynamometer should be as wide as possible in order to get accuratemeasurements of rapidly changing cutting forces, since chipformation, interrupted cutting conditions, chatter and toolbreakage may cause sudden signal variations that have to bereadily detected.

In Table 1, recent research works dealing with cutting forcemeasurement in turning are presented. The table is organized togive the description of the measuring system, the usable frequencybandwidth in operating conditions and the author’s reference.

Comparison between cutting force measuring systems is difficult,due to the lack of information regarding the usable frequency

ll rights reserved.

x: +39 0432 558251.

bandwidth, which is obtained when the sensor is installed in areal machining system, with additional masses – such as thetoolholder or the workpiece – connected to it. The usable frequencybandwidth is significantly narrower with respect to the nominalsensor bandwidth – sensor fixed in a very stiff configuration, withoutadditional masses – specified in the sensor commercial datasheet.

As shown in Table 1, one possible approach for measuring cuttingforces is by applying strain gauges on flexible mechanical parts of themachining system. Alternatively, cutting forces can be measured bypiezoelectric sensors embedded into the machining system.

Strain gauge based devices require special toolholders, endowedwith parallel beam structures for local strain amplification [11–14],or flexible elements such as octagonal rings [15]. Generally, thestrain-gauge dynamometers are affected by low stiffness or limitedfrequency bandwidth.

Scheffer and Heyns [8] mounted the strain gauges directly on acommercial boring bar, without relevant modifications of thetooling system. By doing so, a good frequency bandwidth of3 kHz was achieved, but a sensitivity problem affected the backforce component direction.

A wide frequency bandwidth (2.35 kHz) was also achieved byKim and Kim [14], who developed a triaxial dynamometer based onstrain gauges and piezo film-type accelerometer for measuringstatic and dynamic force components, respectively. However,such device was designed for ultraprecision micromachining(Fmax�15 N) and not for standard applications.

Another deformation based approach was applied by Jin et al.[16], who derived the main cutting force component from thetoolshank deflection measured by an optical fibre. The device wascharacterized by good dynamic properties.

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Table 1State of the art of cutting force measuring systems in turning.

Sensor structure Type Maximum force (N) Maximum bandwidth inoperating conditions (kHz)

Ref.

Strain gauges sticked on a special toolholder close to circular holes enhancing

local deformation

Triaxial 100 Not specified [11]

Strain gauges sticked on parallel beam type toolholder, obtained with new circular

holes design

Triaxial 4000 1.6 (Fx), 0.8 (Fy), 3 (Fz) [12]

Strain gauges attached on toolholder endowed with infra-red data transmission device Monoaxial �1000 Not specified [13]

Combined-type tool dynamometer based on strain gauges and piezo film-type

accelerometer

Triaxial �15 2.35 [14]

Strain gauges mounted on the toolholder (boring bar) Triaxial �1000 3 [8]

Strain gauges attached on flexible octagonal rings embedded into toolholder Triaxial 3500 0.16 (theoretical) [15]

Optical fibre detecting the vibration of the rear part of the toolshank Monoaxial �1000 0.95 [16]

Piezoelectric plate dynamometer Kistler 9121 to be installed on lathe turret Triaxial 3000 (Fx,y),

6000 (Fz)

fno1 [2,7,17,18]

Piezoelectric plate dynamometer Kistler 9257B Triaxial 5000 fno0.7 [18–20]

Piezoelectric plate dynamometer Kistler CompacDyn 9254 Triaxial 500 (Fx,y), 1000 (Fz) Not specified [18,21]

Slim line Kistler 9132B placed under the cutting insert Monoaxial 7000 Not specified [7]

G. Totis, M. Sortino / International Journal of Machine Tools & Manufacture 51 (2011) 34–42 35

A better compromise between stiffness and sensitivity isgenerally achieved by piezoelectric dynamometers [22].

Specifically, plate dynamometers are usually preferred thanksto their high stiffness and good static properties, such as goodlinearity and negligible crosstalk [2,6,7,17–21]. However, suchdevices are affected by the dynamic properties of the machiningsystem, which may cause a poor frequency bandwidth [19].

Plate dynamometers can be easily fixed on the lathe turrets ofCN lathes. On the contrary, modern multi-function CN lathesendowed with rotating turrets or with indexable heads requireadditional massive adaptors for including the plate dynamometerinto the tooling system. Accordingly, the resonance frequency ofthe machining system and hence the usable frequency bandwidthof the sensor are further reduced, as illustrated in Section 2.

Recently, Kirchheim et. al [7] described a turning dynamometercomposed of a monoaxial piezoelectric cell embedded into thetoolshank, just under the cutting insert. In this configuration,sensor dynamics were significantly improved with respect to theconfiguration based on a plate dynamometer. This is in accordanceto the well common principle of sensor location: the closer thesensor to the process, the more sensitive the signal output is withrespect to process variations.

Briefly, no triaxial dynamometers are available that could beintegrated into flexible CN lathes using special adaptors, whichexhibit satisfactory static and dynamic properties. Therefore,the aim of this work was to develop an innovative triaxialdynamometer specifically designed for the afore said application,which may show better static and dynamic characteristics.

First, sensor requirements are illustrated and the design phase isoutlined. The new sensor configuration is compared with state ofthe art configuration including a commercial plate dynamometer.Afterwards, its static properties are determined by performing astatic calibration procedure. Sensor dynamics are then assessed bymeans of pulse testing. Eventually, the sensor is tested in operatingconditions by machining a benchmark workpiece, and measuredcutting force signals are compared with the theoretical trendspredicted by a cutting force model.

2. Sensor development

As evidenced in Section 1, the best triaxial dynamometers forthe outlined application are piezoelectric plate dynamometers,which nevertheless require special adaptors to be embedded intothe tooling system of the flexible CN lathe. In order to evidence thedrawbacks of this configuration, a FE simulation was carried out.

The simulated configuration representing the state of the artincluded a Kistler 9129AA plate dynamometer and a specialHSK63A adapter, see Fig. 1. A commercial toolshank typeSDJCR2525M11 was imagined to be fixed on the plate dynam-ometer using a toolholder made of a light titanium alloy, seeFig. 1(a).

Static stiffness values and main natural frequencies along thedirections X and Z were estimated using the FE model, as illustratedin Fig. 1(b) and (c). The calculations were carried out by COSMOS-Works integrated into SolidWorks CAD environment. The systemwas discretized by a mesh of parabolic tetrahedron elements (10nodes) of maximum sizeE5 mm.

As evidenced by the results reported in Table 2, the frequencybandwidth of the reference dynamometer would be limited to0.5 kHz because of the low natural frequencies of the overallsystem and of the inertial disturbances deriving from the heavymass fixed on the plateform.

In the light of the limitations affecting the referenceconfiguration, new sensor specifications were expressed asfollows:

1.
Triaxial force measurement during standard turning operations(10 NoFco5 kN).
2.
Good static characteristics (relative static error r5%, crosstalkdisturbance r3%, low tool tip receptance �0.1 mm/N along alldirections).
3.
Good dynamic behavior (frequency bandwidth Z1 kHz alongall directions).
4.
Compatibility with different cutting insert geometries (bothpositive and negative, 551 and 801 rhombic inserts).
5.
Compatibility with the tool head of an OKUMA MULTUS B300multi-function machine tool.
6.
Lubricant resistant.
The following criteria were considered during the design andoptimization phase.

�
Static receptances hii measured at cutting insert tip along thethree main directions X, Y and Z should be minimized, in order tolimit undesired tool deflections during machining. � The levers along X, Y and Z directions between insert tip
and sensor centre should be minimized, in order to limitbending and torsion moments acting on the sensor,which may cause undesired sensor rotations or crosstalkeffects [19].

Table 2Comparison between state of the art and new sensor configuration.

Description Symbol Plate dyn. Kistler 9129AA New sensor system

Total mass including adapter M (g) 10269 3642

Mass in front of the sensor m (g) 1068 120

Tool tip static compliance along X hxx (mm/N) 0.047 0.093

Tool tip static compliance along Z hzz (mm/N) 0.052 0.070

First natural frequency along X fnat,x (Hz) 571 1014

First natural frequency along Z fnat,z (Hz) 558 1177

Fig. 1. Reference commercial dynamometer adapted for the application on the tool head of a multi-function CN lathe (a). Evaluation of tool tip static receptance (b) and system

natural frequencies (c) using FE.

G. Totis, M. Sortino / International Journal of Machine Tools & Manufacture 51 (2011) 34–4236

�
Natural frequencies of the whole system including the toolholdershould be maximized and masses in front of the sensor should beminimized in order to improve sensor dynamics.
Different design alternatives were compared by estimating thecorrespondent static receptances and main natural frequenciesusing the FE approach, until the final optimized design illustrated inFig. 2(a) and Table 2 was determined.

FE results showed that the main natural frequencies of theproposed configuration are approximately twice those of thereference configuration, while static receptances are only slightlyincreased. Besides, the inertial mass attached to the sensor isapproximately one order of magnitude smaller than that attachedto the reference dynamometer, supporting the hypothesis that thenew configuration would have a better dynamic behavior than thereference one.

It has to be pointed out that spindle dynamics were not takeninto account by the simulations, since they are unknown and it isvery difficult to include them in the FE model. In general, spindledynamics tend to further ruin the frequency response of the sensor,as explained in [19,20] and shown in Section 4. The degree of thedisturbance is approximately proportional to the inertial mass infront of the sensor. Therefore, the frequency response of thereference dynamometer is expected to be further penalized incomparison with the new configuration.

Regarding the new prototype, it is composed of a modifiedtoolshank of 25�25 mm2 section made of Ck45 steel, a cutting

force sensor and a modular cartridge. Specifically, a triaxial forcesensor Kistler 9251A was selected because of its sensitivity andrange of application, as well as its high stiffness, see Table 3. Themodular cartridge comprises an intermediate element made ofTi6Al4V alloy, a toolshank tip obtained from a commercial tool-shank and the cutting insert. The toolshank tip can be easilyremoved from the intermediate element without disassemblingthe device, thus preserving sensor’s preload. Different commercialtoolshank types can be applied, hence different insert geometriesare compatible with the device. Four unified toolshank types wereprepared to be fixed on the intermediate element, as follows:

(a)
SCLCR2525M12 for positive 801 rhombic inserts typeCCMT1204;
(b)
SDJCR2525M11 for positive 551 rhombic inserts typeDCMT11T3 (Fig. 2(a));
(c)
DCLNR2525M12-2 for negative 801 rhombic inserts typeCNMG1204;
(d)
DDJNR2525M15-2 for negative 551 rhombic inserts typeDNMG1504.
The proposed compact design allowed to minimize the leversbetween the insert tip and the centre of the sensor top surface, inaccordance with the indications derived in [20] for avoidingcrosstalk effects. Also, the levers were almost equal for all theconfigurations.

Fig. 2. (a) Exploded view of the new sensor configuration; (b) final prototype mounted on the CNC machine tool; (c) static and (d) dynamic analysis carried out by means of

a FE software.

Table 3Nominal and measured static characteristics.

Sensor technicaldata

Sensitivity Fx, Fy pC/N �8

Fz �4

Range Fx, Fy kN 72.5

Fz 75

Linearity Fx, Fy, Fz %FSO r71

Crosstalk Fz-Fx, Fy % r71

Fx2Fy r73

Fx, Fy-Fz r73

Stiffness kxx, kyy N/mm �1000

kzz �2600

Sensor installed inthe toolholder

Admissible

range

Fx kN 70.7

Fy 70.5

Fz 71.4

Static relative

errors

max9ex,rel9 % r3.05

max9ey,rel9 r5.32

max9ez,rel9 r5.91

Crosstalk Fz-Fx, Fy % r71.81

Fx2Fy r72.94

Fx, Fy-Fz r71.43


The whole device was fixed on an HSK63A toolholder, whichwas clamped on the spindle of an OKUMA MULTUS B300W toolhead, see Fig. 2(b).

The sensor was connected to a charge amplifier Kistler type5073A311 using lubricant resistant cables. A measuring board NIDAQ-6024E was employed for signal acquisition. Eventually, thesystem was assembled and then preloaded up to 25 kN, asrecommended by Kistler.

3. Static calibration

During the experimental validation, a Sandvik DCMX11T304WFGC4215 insert (nose radius re¼0.4 mm) was fixed on the dedicatedtoolshank (SDJCR2525M11), which was further clamped on theintermediate element. The experimental trials consisted of staticcalibration, pulse tests and cutting tests. Sensor signals weresampled at a frequency of 20 kHz. The working environment fordata analysis was MathWorks MATLAB.

Static calibration was performed using calibrated weights ofknown masses which were applied to dynamometer tip, onedirection at a time, according to the standard procedure adoptedfor instance by Jin et al. in [16]. The quasi-static forces were

Fig. 3. Experimental setup for pulse tests at spindle without (a) and with (b) the complete tooling system including the new sensor.


compared with the known input forces in order to evaluate thestatic characteristics of the dynamometer. By doing so, admissibleforce ranges, as well as static relative errors and crosstalk dis-turbances were evaluated. Results of this phase are summarized inTable 3.

A slight rotation of the tool tip—sensor system was observedwhen the force applied along Y direction exceeded 500 N, causing adisturbance in the cutting force signals. This was likely due to thetorque caused by the pushing force and by insufficient friction atthe sensor—intermediate element interface. Greater forces wereachieved along the other two directions without any yielding of thedevice. Accordingly, the sensor can be applied only for semi-finishing and finishing operations. However, maximum relativeerrors are smaller than 6% and crosstalk disturbances are smallerthan 3% along all directions, in good accordance with othermeasurements performed by the authors [20].

4. Dynamic identification

The identification of machining system and sensor dynamicswas accomplished by means of pulse tests [23], as illustrated inFig. 3. The instrumented tooling system was excited using animpact hammer Dytran type 5800B4, (sensitivity 2.41 mV/N)connected to a Kistler amplifier type 5134B. In addition, twoeddy current probes Micro-Epsilon type ES1 (measuring range1 mm, sensitivity �15 mV/mm) connected to eddyNCDT 3010-Mcontrollers were employed for measuring the tooling systemvibrations along the main directions of the force sensorreference frame.

The dynamics of the spindle without the tooling system and thedynamics of the overall system including the instrumented toolingsystem were evaluated as follows. Impulsive forces were applied atthe chosen excitation point, one direction at a time. For eachdirection of excitation i, the hammer signal Fin,i and the vibrationsignals Uj (j¼x,y,z) were acquired. Afterwards, modal analysis wasperformed according to Ljung [24], in order to derive the dynamicreceptances

HjiðjoÞ ¼UjðjoÞ

Fin,iðjoÞð1Þ

representing the flexibility of the system along the jth directionwhen it is pushed along the ith direction.

Spindle dynamic receptances along the X and Z radial directionswere quite similar. For the sake of simplicity, an average experi-mental transfer function in the radial direction was calculated,which was nevertheless noisy because of the high spindle stiffness.Therefore, a mathematical model interpolating the average transferfunction was determined, see Fig. 4(a) and (b), evidencing someimportant resonances ranging from 0.5 up to 2 kHz.

Similarly, direct receptances Hii(jo) of the overall systemmeasured at the tool tip were calculated, see Fig. 4(c) and (d).

Coherence analysis was carried out in order to evaluateinput-output correlation at each pulsation [25], evidencing thatvibration measurements were reliable only in the range150–2500 Hz, because of a bad signal to noise ratio affectinglow and high frequency ranges. For this reason, static receptancevalues

hji ¼HjiðjoÞjo ¼ 0 ð2Þ

were estimated from the dynamic receptance curves in a neighbor-hood of o¼200(2p) rad/s. By doing so, the experimental recep-tance matrix was obtained and it was compared with thereceptance matrix computed by FE, as follows:

H¼ ½hji� ¼

0:135 �0:032 �0:037

�0:032 0:044 �0:026

�0:037 �0:026 0:079

264

375ðEXPÞ

�

0:093 �0:024 �0:023

�0:024 0:049 �0:025

�0:023 �0:025 0:070

264

375ðFEMÞ

ð3Þ

where the measuring units are (mm/N). In general, there is a goodagreement between experimental and FE values, although thelatter are generally underestimated because of the simplificationsinvolved in the FE model. In particular, FE values do not take intoaccount the translational and flexural receptance of the spindle.A better correspondence was assessed in Matlab by combining theflexibility of the tooling systems with the experimental flexibility ofthe spindle, thus confirming the validity of the FE model.

Regarding the dynamic behavior, main resonances of the overallspindle-tooling system are located in the range 0.5–2 kHz. Accord-ing to FEM calculations, the first natural frequencies of the toolingsystem fixed at an infinitely stiff spindle should be about 1 kHz.Therefore, the resonance peaks located at 0.5 kHz derived from

Fig. 4. Amplitude (a) and phase (b) of the spindle without the tooling system in the radial direction. Amplitude (c) and phase (d) of tool tip dynamic receptance. Amplitude (e)

and phase (f) of dynamometer ETFEs.


spindle dynamics coupled with the mass of the tooling system,considered as a rigid body, while the high frequency modes are theresult of the complex dynamic interaction between the spindle andthe tooling system vibration modes. On the other side, the effect ofspindle dynamics on the Y direction (axial direction of the spindle-tooling system) is moderate, because of the extremely high spindlestiffness in that direction.

Eventually, by comparing the hammer signal Fin,i with thedynamometer signals Fdyn,j, the Empirical Transfer FunctionEstimates – ETFEs – of the dynamometer were obtained asfollows:

WjiðjoÞ ¼Fdyn,jðjoÞFin,iðjoÞ

ð4Þ

Amplitude and phase of direct ETFEs Wii(jo) are given in Fig. 4(e)and (f), respectively, while cross ETFEs Wji(jo) were negligiblewithin the frequency range of interest. As expected, the resonancefrequencies of dynamometer ETFEs Wii(jo) coincide with theresonance frequencies detected in the dynamic receptances Hii(jo).However, the resonance peaks below 1 kHz are significantlyattenuated in Wii(jo) with respect to Hii(jo). A relevant deviationof the ETFEs amplitude from 73 dB bandwidth is observed onlyabove 1 kHz, even if 73 dB crossing frequencies strictly are 656 Hzalong X, 544 Hz along Z and 984 Hz along Y. By means of properfiltering, the usable frequency bandwidth can be easily extended to1 kHz for all directions [26]. On the contrary, dynamics compensa-tion above 500 Hz is not possible for the massive state of the artconfiguration examined in Section 2, due to its low-pass filterbehavior.

A qualitative explanation of sensor dynamics and ETFEs char-acteristics can be found in [19]. Basically, sensor measurements aredisturbed by the forces of inertia exchanged with the mass in frontof it. However, this effect was reduced by minimizing the mass in

front of the sensor. Moreover, a kind of self-compensation oflow-frequency resonances is carried out by means of almostoverlapped antiresonances deriving from sensor positioning closeto the cutting insert.

By doing so, it is possible to attenuate the low-frequencydynamic disturbances – always present and usually characterizedby resonance peaks below 1 kHz – deriving from the mechanicalsystem supporting the tooling system.

5. Application of the dynamometer in a complex turningoperation

In this section, an example of sensor application is illustrated.The aim is to show dynamometer potential for in-process mea-surement and cutting force model development.

For this purpose, cutting forces were measured during acomplex machining operation on a benchmark workpiece, andthe measured signals were compared with the analytical trendsestimated by a mechanistic cutting force model.

In detail, workpiece material was Ck45 carbon steel, while thepositive 551 rhombic insert of the previous phases was applied, seeFig. 5. The tool path consisted of several passes of profile turningwith different feeds and depths of cut, as illustrated in Fig. 6. Thefeed and the depth of cut were varied in order to compare theexperimental and estimated cutting forces on a larger variety ofcutting conditions. A constant cutting speed vc¼350 m/min wasapplied for all the passes, and no lubricant was used.

In order to estimate the theoretical cutting forces, the classicalshearing and ploughing cutting force model was adopted [27].According to the model, the resultant cutting force is the sum ofthe infinitesimal cutting forces acting at each infinitesimal part ofthe engaged cutting edge, as illustrated in Fig. 7. Specifically, the


infinitesimal tangential and normal forces dFc and dFn can beexpressed as a linear combination of the local uncut chip area dA

and of the infinitesimal cutting edge length dl, by means of theshearing (kcs, kns ) and ploughing (kcp, knp) coefficients, respectively,as follows:

dFc ¼ kcsdAþkcpdl

dFn ¼ knsdAþknpdl

(ð5Þ

The uncut chip area dA can be approximated by

dA¼ hðlÞdlffi f sinwðlÞdl ð6Þ

where f is the instantaneous feed per revolution and w is the localcutting edge angle. After projection along the tangential, axial andradial directions and integration along the engaged cutting edgeone can obtain

8>>>>>><>>>>>>:

Fc ¼ kcsfapþkcpL

Ff ¼ knsf

Z L

0sin2wðlÞdlþknp

Z L

0sinwðlÞdl

Fp ¼ knsf

Z L

0sinwðlÞcoswðlÞdlþknp

Z L

0coswðlÞdl

ð7Þ

where Fc, Ff and Fp are the main cutting force, the feed force and theback force components, respectively.

Fig. 6. Tool path for the b

Fig. 5. Experimental setup for cutting tests.

Eventually, analytical cutting force trends were calculated byapplying the model coefficients listed in Table 4, which wereestimated in a preliminary phase and are similar to values reportedin other publications and technical reports for Ck45.

In general, there is a good agreement between the experimentaland predicted cutting force trends, as shown in Fig. 8. The scatterdiagrams of the predicted force components against the measuredforce components together with square correlation values aregiven in Fig. 9. The feed force Ff and cutting force Fc componentsare affected by a small systematic error, while the back force Fp isconsiderably underestimated, as evidenced by the slope of thelinear interpolating curve of Fig. 9(b). This may suggest that thecurrent model assumptions are not adequate for representing allthe force components at the same time with good accuracy.However, the experimental and analytical instantaneous trendsare well correlated, demonstrating the sensor readiness for detect-ing sudden cutting force variations during a real machiningprocess. For instance, an abrupt increase in both measured and

enchmark workpiece.

Fig. 7. Cutting force components modeling.

Table 4Cutting force coefficients of Ck45.

Coefficient kcs kcp kNs kNp

Value 1809 N/mm2 28.47 N/mm 723.1 N/mm2 73.51 N/mm

Fig. 8. Comparison between measured and predicted cutting force trends when machining the benchmark workpiece.

Fig. 9. Correlation between measured and predicted cutting forces.


predicted cutting force trends is present at the end of each pass.This was due to critical cutting conditions characterized by smalluncut chip thickness and long engaged cutting edge length, whichwere caused by the selection of tool trajectories parallel to theworkpiece profile. In these conditions, long chip was also observed,evidencing that instantaneous cutting conditions were outside thechip breaking region. The discrepancies in the force peaks observedat the end of each pass are partly due to the instantaneous cuttingconditions – which lie outside the calibration ranges – and partly tonumerical errors occurred during the simulation of the cuttingprocess. Accordingly, the confidence intervals illustrated in Fig. 9were estimated by excluding the outliers corresponding to thesecritical conditions.

In the light of this comparison, the sensor can be successfullyused for characterizing complex turning operations, for cuttingforce model calibration and improvement, as well as for othermonitoring purposes. Also, it has to be noticed that the staticthresholds identified during static calibration were slightlyexceeded during the cutting tests without altering system config-uration and preload, confirming the applicability of the sensor in awide range of cutting conditions.

6. Conclusions

In this work, an innovative dynamometer for triaxial cuttingforce measurement in turning, and specifically designed to beapplied on modern CNC lathes endowed with indexable tool headsor with rotating turrets, was developed.

The prototype of the dynamometer was statically calibrated andits dynamic response was determined by means of pulse tests.The device evidenced good static behavior when applied for semi-finishing and finishing operations, and its frequency bandwidths inthe three main directions were 656 Hz along X, 544 Hz along Z and984 Hz along Y. For the sake of comparison, the frequency responseof a commercial device, adapted for this application and estimatedby FE modeling under ideal clamping conditions, was less than500 Hz. Moreover, the usable frequency bandwidth of the newsystem can be easily extended to 1000 Hz for all directions bymeans of simple signal filtering, whereas the massive structure ofthe commercial device eliminates all the signal information above500 Hz, due to its low-pass filter behavior.

To test the applicability of the device in operating conditions,a complex shape benchmark workpiece was machined and


measured cutting forces were compared with theoretical valuesobtained from mathematical modeling. This test demonstrated thecapability of the new sensor to follow rapid signal variations inactual production conditions.

It would be of further interest to improve the geometry of thedevice in order to increase the force measuring range and toinvestigate methodologies for further widening its frequencybandwidth.

Acknowledgements

This work was supported by KEYMEC, Centre for Innovation,Research and Training in the field of Mechanics, located in thePonte Rosso Industrial Area of Pordenone, Italy, which is gratefullyacknowledged.

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Development of a modular dynamometer for triaxial cutting force measurement in turning