Correlating surface roughness, tool wear and tool vibration in the milling process of hardened steel...

International Journal of Machine Tools & Manufacture 68 (2013) 1–10

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International Journal of Machine Tools & Manufacture

0890-69

http://d

n Corr

E-m

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

Correlating surface roughness, tool wear and tool vibration in the millingprocess of hardened steel using long slender tools

Marcelo Mendes de Aguiar, Anselmo Eduardo Diniz n, Robson Pederiva

Faculdade de Engenharia Mecanica, CP 6122, Campinas, S ~ao Paulo 13083-860, Brazil

a r t i c l e i n f o

Article history:

Received 5 November 2012

Received in revised form

14 January 2013

Accepted 15 January 2013Available online 31 January 2013

Keywords:

Tool vibration

Tool wear

Surface roughness

High speed milling

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ail address: [email protected] (A.E. D

a b s t r a c t

High speed milling is an operation frequently used in finishing and semi-finishing of dies and molds.

However, when it is necessary to produce molds with deep cavities and/or with small corner radius,

long tools with small diameters are required. This represents a challenge for manufacturing profes-

sionals: how to minimize tool vibration using a tool with such low rigidity and obtain good workpiece

surface quality and long tool lives. This paper attempts to answer this question. Milling experiments on

hardened AISI H13 steel were carried out using integral and indexable insert tools with different tool

overhangs and different diameters. Tool wear, workpiece surface roughness and cutting forces were

measured and these parameters were correlated with the frequency response function (FRF) obtained

with the tools fixed in the machine tool. The main conclusion of this study is that good workpiece

surface roughness allied to long tool lives for long tools with small diameters can be achieved, provided

the tooth passing frequency used in the milling process (and its harmonics) does not produce high FRF

values.

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1. Introduction

Cutting forces directly influence the workpiece precision andsurface quality, the system vibration, cutting power and tool life [1].These forces are influenced by several factors, such as toolgeometry, properties of the workpiece material, cutting conditions,cutting strategy, etc. [2,3].

Using a mathematical model, Dagiloke et al. [4] demonstratedthat a cutting speed of up to 1200 m/min does not influencecutting forces. On the other hand, Schulz and Moriwaki [5] statedthat the force decreases as cutting speed increases. According toFlom and Komanduri [6], cutting forces decrease with increasingcutting speed up to a certain limit. Beyond this point, these forcesgradually increase.

Cutting forces cause deflections in the tool/workpiece/toolfixation/machine system [7], which cause significant geometricalerrors in the machined workpiece [8]. These errors are particu-larly important when a high tool length/diameter ratio (L/D) isused, when the inclination of the machined surface is high andwhen tool wear is significant [9].

Deflections must be controlled mainly in finishing operations,since they impair surface quality and tool life [10]. Kecelj et al.[11] conducted milling experiments using ball nose mills with anL/D ratio of 7 and 10. Their results indicated that tool deflection ishigher when the angle between the machined surface and the

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iniz).

horizontal is small, which is caused by the difficult conditions ofchip formation in the tool’s central region. This occurs particularlywhen low values of depth of cut (ap) are used. These findingsagree with those of Lopez de Lacalle et al. [9,12], who found thehighest errors in surfaces with inclinations of less than 151. Theauthors attributed these results to the slipping effect of smallchips and to cutting distortion when the central portion of thetool is engaged in cutting. However, Oliveira [13] reported adifferent finding, claiming that the tool which cut a surface with a751 angle in relation to the horizontal line (tool in the verticalposition) presented higher deflection than the tool that cut a 451surface, due to the higher radial force of the cut at 751.

A well known model for studying tool deflection is the one thatconsiders the tool fixed in the chuck as an overhanging cylinder[14]. This model does not include dynamic considerations, but itcan make coherent predictions, since in finish operations, inwhich the depths of cut are small, the process is close to stability[12] and tool behavior is quasi-static [15]. According to Xu et al.[16], under stable cutting conditions, static tool deflection is moresignificant than dynamic deflection.

Static tool deflection (d) is calculated by Eq. (1) (consideringthe tool a cylinder), where F is the cutting force perpendicularto the tool axis, E is the Young modulus of the tool’s material, andL3/D4 is the tool’s slenderness coefficient (TSC) [9].

d¼64F

3pE

� �L3

D4

!ð1Þ

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M.M. de Aguiar et al. / International Journal of Machine Tools & Manufacture 68 (2013) 1–102

The machine tool, cutting tool, workpiece and fixation devicesform a complex system of structural elements. During cutting, alarge portion of energy is dissipated through these elements, alsoinducing vibration [17].

Vibration may reach unacceptable levels, particularly whenthere is an inherent lack of rigidity in the system, as in the millingof dies and molds, which frequently requires the use of long toolsto machine deep cavities [18]. Therefore, vibration must beminimized due to its harmful influence on the dimensionalquality and surface texture of the workpiece, on the accelerationof tool wear/damage and on the increased probability of toolbreakage [17,19].

The main types of vibration involved in die and mold millingusing high speed machining (HSM) are forced and self-excitedvibrations [13].

Forced vibrations are those caused by external forces. Theyoccur in all types of machining operations, but are especiallycritical in finishing operations, where shape errors and highvalues of surface roughness are unacceptable. They are even moreharmful when the excitation frequency is close to either thenatural system frequencies or to one of their harmonics, as theymake the cutting unstable [17].

Self-excited vibrations are not caused by external forces, butby forces generated by cutting the material [20]. These vibrationsoccur when the damping capacity of the tool–workpiece–machinesystem is insufficient to absorb the energy transmitted by thecutting [17], generating a self-exciting mechanism duringmachining which causes continuous variations in chip thickness.Initially, one of the structural elements is excited by cutting forcesand a wavy surface generated by the cut produced by one edge isremoved by the next edge, which also leaves a wavy surface dueto structural vibrations [21]. When a phase discrepancy occursbetween the vibration waves left by the cutting edges on thesurface, it produces a regenerative effect that generates evenmore vibration [21]. This phenomenon is known as ‘‘chatter.’’ Thisis the most harmful type of vibration in HSM machining processes[17,22].

Vibration can also be controlled by the use of a more rigid tool andtool fixation. Oliveira [13] studied the influence of two grades of toolmaterial and two types of tool shank (carbide and steel) on tool wear,tool life and workpiece surface roughness life in hardened steelmilling. The carbide shank produced better results than the steelshank. This finding was attributed to the higher rigidity of the carbideshank, which decreased its tendency to vibrate.

Several authors [20,23,24] argue that to prevent chatter andachieve good workpiece surface quality, the frequency of thecutting edge entering the cut during each rotation of the tool(tooth passing frequency) must differ from the natural andharmonic frequencies. The natural frequency is influenced, amongother factors, by the tool overhang (tool length/tool diameterratio), density and Young’s modulus of the tool and tool shankmaterials [20,25].

One of the main goals in finishing operations is to achieve avery low workpiece surface roughness [26]. However, surfaceirregularities, which are always present in all machined parts,depend on several factors. In milling operations, surface qualityimproves at higher cutting speeds. Depth of cut indirectly affectssurface quality, since the cutting force, vibration and cuttingtemperature increase with an increase in the depth of cut. Otherfactors that influence surface roughness are feed, tool nose radius,tool wear, cutting strategy, the tool’s trajectory during cutting,workpiece material, cooling/lubrication system and the dynamicparameters of machining, such as cutting force, tool deflection,vibration and several thermal phenomena [17,27–29].

The use of high cutting speeds and small tool diameters, due tothe small radius of the tool used in cutting dies and molds makes

it mandatory to apply very high tool rotation speeds in high speedmilling finishing operations. Therefore, the feed velocity is high,even when low feed rates are used, which allows for a highnumber of tool passes (with low radial and axial increments)without increasing the cutting time. As a result, good levels ofsurface finish are usual in these processes [9,30–32].

In the milling of inclined flat surfaces with either toroidal orball nose tools, the theoretical surface roughness can be deter-mined in both the transverse and longitudinal directions inrelation to the feed direction. The theoretical roughness perpen-dicular to the feed direction (RthTRANS) is determined by thecombination of tool radius (R), axial increment (ap) and inclina-tion angle of the surface (a), according to Eq. (2) [13,30].

RthTRANS ¼ R�

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffið2RÞ2� ap

sena� �2

4

sð2Þ

In the feed direction, the maximum theoretical roughness(RthLONG) is given by Eq. (3), where fz is the feed per tooth andREF is the effective radius of the tool measured at the point wherethe tool touches the workpiece. REF is related to the surfaceinclination, since the higher the angle of inclination the higherthe effective radius [13].

RthLONG ¼ðf zÞ

2

8nREFð3Þ

However, real roughness values usually differ from theoreticalones [28,29]. Axinte and Dewes [33] observed high values ofsurface roughness generated in a high speed milling operation,which they attributed to tool run-out at high cutting speed, alliedto the vibrational effect of high cutting forces. According toFallbohmer and Scurlock [34], cutting with a tool with a smalllevel of wear may generate lower roughness than cutting with afresh tool. Diniz et al. [35] found similar results in the milling ofH13 steel with a toroidal tool in semi-finishing conditions.A possible explanation for these results is that roughness valuesmay be associated with tool coating defects on the cutting edge,as cited by Oliveira [13], which affect roughness at the beginningof tool life. Depending on the type of wear and its evolution, thesedefects may spread along the entire length of the cutting edge incontact with the workpiece, making it more uniform and therebyimproving the surface roughness value.

The objective of the experiments shown in this work is to findhow to minimize tool vibration using long slender tools andobtain good workpiece surface quality and long tool lives. There-fore, milling experiments using integral carbide and indexablecarbide end mill with high tool overhang are described below.

2. Methods, equipments and materials

Several experiments involving finishing operations using thehigh speed milling technique were performed to determine theinfluence of tool diameter, tool slenderness coefficient (L3/D4,where L is tool overhang and D is tool diameter) and type of tool(integral carbide end mill and indexable carbide tool with a carbidetool shank) on tool wear and workpiece surface roughness.

The machine tool used in the experiments was a 5-axismachining center with 15 kW of power in the main motor, toolrotation between 35 and 25,000 rpm and HSK 63 A system fortool fixation.

The workpieces used in the experiments were made of AISIH13 steel, quenched and tempered to reach 50 HRC of hardness.

Four ball nose end mill tools were used, two integral and twoindexable carbide inserts with carbide tool shanks. The cementedcarbide inserts had a 3 to 4 mm thick PVD coating of TiAlN.The first was an 8 mm tool with insert code KDMB08M0ERGN

M.M. de Aguiar et al. / International Journal of Machine Tools & Manufacture 68 (2013) 1–10 3

grade KC515M, which, according to the manufacturer’s catalog, isemployed in machining steel (class P) and hardened steel (class H)with hardness up to 54 HRc with geometry suitable for finishingoperations, fixed in a tool shank code KDMB08R150A08HN. Thesecond tool was a 12 mm tool with insert code KDMB12M0ERGN,with the same characteristics as the 8 mm diameter inserts, fixed ina tool shank code KDMB12R160A12HNC. Both tools were fixed inthe chuck by cold deformation. The integral carbide end mills had8 and 12 mm diameters, both PVD coated with TiAlN. The two typesof tools (insert and integral) had similar cemented carbide gradesand their tool radial run-out in all the experiments was lower than10 mm.

Due to the tools’ high wear resistance, it was impossible toreach tool wear values that would configure the end of tool life,even after a long cutting time and using a very high cutting speed.Therefore, the experiments ended upon reaching 400 min ofcutting time. Even after such a lengthy cutting time, tool flankwear was less than 0.10 mm. The tool flank wear was measuredusing an optical microscope with 50� magnification.

The cutting speed (vc) and the angle between the milledsurface and the machine tool’s XY plane (a) were kept constantduring all experiments (vc¼500 m/min and a¼751).

The input variables of the experiments were tool diameter (D),tool slenderness coefficient (called TSC in this paper) and type oftool (integral and indexable inserts and ball nose end mills), all ofthem with two levels, which would result in a 23 factorialexperimental design. In preliminary experiments, the conditionwith the indexable insert carbide tool with D¼8 mm andTSC¼45 mm�1 presented very high roughness values from thebeginning of the experiments (Rz¼9.11 and 3.96 mm measuredtransverse and longitudinal to the feed directions, respectively).Therefore, the use of TSC¼45 for the indexable insert tools wasdiscarded, resulting in a final incomplete 23 factorial experimen-tal design. The experimental conditions employed here aredescribed in Table 1. Each experiment was performed twice.

All the tools used in the experiments were extremely sharpand had a very small cutting edge radius, but larger than the chipthicknesses. The integral carbide end mills presented a morepositive rake angle than the indexable carbide insert tools.

Moreover, all the tools used in the tests were subjected tovibrational analysis in order to identify the frequency responsefunction (FRF) of each set, and thus, to determine the naturalfrequency of each system. This procedure was performed on thetools mounted on the spindle using an instrumented impacthammer. The curves of the natural frequency of each set areshown in the graphs in Fig. 1. It can be seen in this figure thestrong influence of both, the tool diameter and the TSC, on thenatural frequencies of the several tool sets tested. For a constanttool diameter, when TSC increased, the natural frequencydecreased. The same occurred for the tool diameter when TSC

was kept constant. The reason for these occurrences was the

Table 1Conditions employed in the experiments.

Experiment vc

[m/min]a[degree]

Rth[lm]

ana

[mm]Tool D

[mm]TSC[mm�1]

1 500 75 0.20 0.10 Integral 8.0 20

2 12.0 20

3 8.0 45

4 12.0 45

5 Indexable

insert

8.0 20

6 12.0 20

a an is the thickness of the material removed perpendicular to the machined

surface.

correlation among these parameters and the rigidity and mass ofthe vibratory system formed by the tool and its fixation in themachine tool. It can be also seen that when the kind of tool(integral or insert) was changed, the variations in the character-istic frequencies of the system were not high. The vertical lines inthe figure, which represent the tooth passing frequencies (TPF)(calculated from the cutting speed, tool diameter and number ofteeth) and their harmonics, indicate the fundamental excitationfrequencies and show how distant they are from the naturalfrequency of the tool and tool fixation system.

The surface roughness produced in the experiments, whichwas measured in the feed direction and perpendicularly to thefeed direction using Rz parameters, was associated with thetheoretical roughness values (Rth—Eqs. (2) and (3)) because thisroughness parameter is sensitive to the presence of high peaksand valleys on the surface [17,36]. The ap was defined usingEq. (2) and fz using Eq. 3. RthTRANS (Eq. (2)) and RthLONG (Eq. (3))were fixed at the same value, 0.2 mm. Using these Equations, andin order to obtain Rth¼0.20 mm in both directions, the values ofap and fz used here were 0.077 and 0.079 mm, respectively, for thetools with D¼8 mm, and 0.095 and 0.096 mm for the tools withD¼12 mm. These values are close to those recommended by thetool supplier, which suggests depth of cut values of around 0.01Dfor finishing operations [37] and also close to the value used byKlocke et al. [38].

The cutting forces (X–Z directions) were measured at thebeginning of the experiments (fresh tool) and after the tool hadbeen cutting for 400 min. These measurements were taken with aKistler 9257B dynamometer connected to a Kistler 5019B signalconditioner and an A/D board to sample the signals entering thecomputer. For a tilt angle of 751, Lopez et al. [39] proposed the useof a sampling frequency of 44 kHz for a tool rotation of10,000 rpm, and 110 kHz for 25,000 rpm. Therefore, in this work,a sampling frequency of 75 kHz was used to acquire signals in allthe experiments, since TPF were 684.7 and 456.7 Hz (or toolrotations of 20,542 and 13,701 rpm) for 8 and 12 mm tools,respectively.

Surface roughness values were measured using a portable rough-ness meter connected to a computer, so as not to have only theroughness values, but also their profiles. They were measured at thebeginning of the experiments, and after 25 and 50 min of cutting.After, roughness measurements were taken at 50-min intervalsduring cutting. At these precise moments, three surface roughnessmeasurements were taken in the directions parallel and perpendi-cular to the feed direction. The values shown in the figures of the nextitem represent the average of three measurements.

3. Results and discussion

Fig. 2 shows the average roughness values over the 400 min ofmachining, obtained from two replicates of the experiments; thedispersion lines represent a standard deviation of 71 (in eachreplicate, roughness was measured three times in each direction).

The values of transverse surface roughness (perpendicular tothe feed direction) were more sensitive to differences in the inputparameters than the longitudinal roughness profiles. This is dueto the greater differences among the roughness curves in each ofthe experiments. Except for the curve of experiment 3, the curvesof longitudinal roughness are very similar to each other. Theresults also show that, over the 400 min evaluated, the transversesurface roughness tended to be higher than in the directionlongitudinal to the feed direction.

The mean roughness produced by using 12 mm diameter endmills showed no significant variations, regardless of the TSC valueused in the experiments. The lengths in balance (Lt) were 78.25

050

100150200250300350400450500

Hertz

INTEGRAL D=8,00 TSC=20

050

100150200250300350400450500

m/s

2 /New

ton

m/s

2 /New

ton

m/s

2 /New

ton

m/s

2 /New

ton

m/s

2 /New

ton

Hertz


XYTPF and harmonics

050

100150200250300350400450500

Hertz


XYTPF and harmonics

050

100150200250300350400450500

Hertz


XYTPF and harmonics

050

100150200250300350400450500

Hertz

INSERT D=8,00 TSC=20XYTPF and harmonics

050

100150200250300350400450500

Hertz

INSERT D=12,00 TSC=20

XYTPF and harmonics

XYTPF and harmonics

Fig. 1. FRF curves of tool/tool-shank/machine system in each experimental condition.


and 101.40 mm for TSC¼20 and 45 mm�1, respectively. This factdemonstrates the possibility of using end mills of this diameterwith lengths in balance for machining deep areas withoutimpairing the surface quality either at the beginning of tool lifeor after 400 min of milling.

The tool diameter does not affect the results when TSC¼20 isused, since low values of surface roughness were obtained withboth tool diameters, particularly in the direction longitudinal tothe feed, with values less than 1.00 mm (Rz).

Most of the tested conditions resulted in low roughness values.An analysis of Fig. 2 indicates that minor variations in roughness,as well as slight standard deviations occurred during the 400 minof the experiments. Therefore, it can be stated that the low toolwear (the tool wear behavior will be analyzed later in this paper)did not affect surface roughness. Moreover, an analysis of theevolution of the roughness curves in experiments 1 and 4 reveals

that the values were lower in both directions when the tool hadalready been in operation for 400 min. In experiment 1, thesurface roughness transverse and longitudinal to the feed direc-tion began with 1.49 and 0.89 mm Rz, respectively, and after400 min showed values of 1.45 and 0.69 mm Rz. In experiment 4,the surface roughness started at 2.14 and 0.73 mm Rz and endedafter 400 min of cutting with values of 1.96 and 0.63 mm Rz. Thesedecreases, albeit slight, demonstrate that, especially under theseconditions, the tools could be used for much longer periods andstill maintain the quality of the machined surfaces.

In contrast to a majority of the experiments, in experiment 3,the use of the integral end mill with 8 mm diameter and TSC¼45resulted in high roughness values with the fresh tool, showing anaverage Rz of 3.14 mm transverse to the feed direction and of1.94 mm in the longitudinal direction, obtained at the beginningof the experiment. During machining, the surface roughness

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

8.00

9.00

10.00

0 25 50 100 150 200 250 300 350 400

Rz

Tran

sver

sal [

µm]

Machining time [min]

INTEGRAL D=8,00 TSC=20INTEGRAL D=12,00 TSC=20INTEGRAL D=8,00 TSC=45INTEGRAL D=12,00 TSC=45INSERT D=8,00 TSC=20INSERT D12,00 TSC=20

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

8.00

9.00

10.00

0 25 50 100 150 200 250 300 350 400

Rz

Long

itudi

nal [

µm]

Machining time [min]

INTEGRAL D=8,00 TSC=20INTEGRAL D=12,00 TSC=20INTEGRAL D=8,00 TSC=45INTEGRAL D=12,00 TSC=45INSERT D=8,00 TSC=20INSERT D12,00 TSC=20

Fig. 2. Roughness, Rz, values obtained in the experiments.

7488

71 6748

62

0

20

40

60

80

100

120

D=8,00 D=12,00 D=8,00 D=12,00 D=8,00 D=12,00

TSC=20 TSC=45 TSC=20

INTEGRAL INDEXABLE INSERT

Flan

k w

ear

[µm

]

Fig. 3. Flank wear presented by ball nose end mills after 400 min of cutting.


values increased further and visible vibration marks appeared onthe machined surface. The low surface quality generated bymachining in this experiment is evident when observing thecurve of increasing surface roughness and also the high standarddeviations, which indicate significant variations in surface rough-ness. Thus, the conditions used in experiment 3 would not complywith the quality requirements for surface machining of dies andmolds, and several analyses were conducted to investigate thecauses of this substantial increase in roughness in this experi-ment, which will be discussed later.

Tool wear may strongly affect the quality of machined sur-faces. Thus, the wear (VBB max) of the tools used in the experimentwas measured at two cutting edges of each of the tools used intwo replicates after 400 min of cutting, and the results are shownin Fig. 3. These values represent the average wear of the fouredges (two edges of each tool, in two replicates of each experi-ment). The dispersion lines represent a standard deviation of 71.

Based on Fig. 3, it can be stated that all the tools showed lowerflank wear, including the tool used in experiment 3, than thatobtained in other experiments.

The shape of the edge wear can influence surface roughnesssince this shape is transferred to the machined surface duringthe cutting process. This influence is greater perpendicularly to

the feed direction since, in the longitudinal direction, the direc-tion of measurement of the roughness profile and the tool’srotation attenuate this effect. Fig. 4 shows the edge radius ofthe tools used in the experiments, enabling the identification ofthe shape of the worn edge after 400 min of use. Here, each tool isrepresented by the edge showing the highest wear in eachreplicate.

The images indicate that only the edge of the tool used inreplicate 1 of experiment 1 was slightly altered from its originalprofile. This may have influenced the formation of the surfaceprofile, especially in the direction transverse to the feed. More-over, the two tools used in this experiment showed delaminationof the coating, exposing the substrate. This occurred at one of thetwo edges of each tool. The tool used in replicate 1 of experiment4 showed a slight change from its original shape and minorsuccessive chipping across the rake surface, which could have anegative effect on roughness. However, this was not confirmed bythe roughness curve in Fig. 2.

The other tools exhibited essentially uniform flank wear,including the tools used in experiment 3, particularly whencompared with the wear obtained in other experiments. There-fore, the wear shape analysis also does not explain the significantincrease in roughness occurred in experiment 3 during the400 min of milling shown in Fig. 2.

Another analysis to explain this high surface rough-ness obtained in experiment 3 was to verify the cutting forces,since they may also influence the quality of machined surfaces.The three orthogonal components (X–Z) of the cutting forces weremeasured in all the evaluated conditions. Toh [2] states that thecomponent transverse to the feed direction (Fy in this work) ismore sensitive to the detection of regenerative vibration, due tothe reduced damping ratio, other than the other two axes. There-fore, the average peak values of the Fy component were consideredin this analysis, as indicated in Fig. 5

According to this figure, in each case, the Fy values were higherafter 400 min than at the beginning of the experiments. Only inexperiment 5, after the tool had already been cutting for 400 min,was the value very similar to that obtained when machining withthe fresh tool. It is not clear whether the Fy value in experiment3 differed from the values recorded in other experiments.A comparison of the raw signal of Fy from all the experimentsrevealed that the behavior of the curve obtained in experiment3 was different, as depicted in Fig. 6, which indicates a typicalsample of Fy for the other experiments (Fig. 6a) and a typicalsample of Fy for experiment 3 (Fig. 6b).

Kin

dTS

CD

iam

eter

Exp

erim

ents

Replica 1 Replica 2In

tegr

al20

8 112 2

458 3

12 4

Inse

rt20

8 512 6

100 µm 100 µm

100 µm 100 µm

100 µm100 µm

100 µm 100 µm

100 µm 100 µm

100 µm 100 µm

Fig. 4. Microscopic images of wear on the ball nose end mills used in the experiments.

71.7 75.1 74.1 70.3 78.1 68.584.5 85.8

98.0 88.4 79.7 88.2

0

20

40

60

80

100

120

140

Experiment 1 Experiment 2 Experiment 3 Experiment 4 Experiment 5 Experiment 6

D=8,00 D=12,00 D=8,00 D=12,00 D=8,00 D=12,00



Fy [N

]

New

400 minutes

Fig. 5. Fy in the experiments.


During stable cutting, the signal is periodic with two peaks ofdifferent amplitudes at each rotation of the tool, which demon-strate radial run-out between the two edges of the tool. In otherwords, Fig. 6a shows that, due to the radial run-out of the tool,one cutting edge cuts more material than the other. However,a comparison of the forces obtained in different tool rotationsshows stability. Due to the small chip cross section area, a smalltool radial run-out caused by imperfections either of the tool orof the tool fixation and also caused by the tool deflectiongenerated by cutting forces, made the actual chip cross section

and, consequently, the cutting force to vary in each tool revolu-tion, as it is seen in Fig. 6a. However, this tool run-out, as can beseen in Figs. 2 and 3, was neither able to damage surfaceroughness, nor stimulate flank wear. On the other hand, inexperiment 3, the signal of the Y component of the cutting forces(Fig. 6b) shows different amplitudes not only in a single toolrotation but also when different rotations are compared. Thevibration that caused this cutting force behavior also caused thehighest roughness values obtained in experiment 3, which areillustrated in Fig. 2.

In order to have a better visualization of the differencesbetween stable and unstable conditions, Fig. 7 was built. It showsFy peak values at the beginning and end of the experiments, usingpolar coordinates during 30 tool rotations. As the instability of theprocess increases, the difference between the shape of its polarcoordinate graphic and a perfect circle also increases. Because thisgraphic shows the force against tool rotation and not againstcutting time like in Fig. 6, it makes easier for the reader tounderstand the force variation along the rotations. The differencebetween two successive points depicts the tool’s radial run-out.

The curves obtained in most of the experiments are symme-trical to a circle passing through the average values of the peaks.Moreover, the peak forces obtained with the fresh tool showlower values than the same tool after 400 min of cutting. Inexperiment 5, this difference is practically nonexistent.

Tool wear not was responsible for increasing the roughnessvalues in the experiments since the flank wear values (Fig. 3)were very low and the tool nose shapes (Fig. 4) were close to the

30

50

70

90

110

130

150

0 1000 2000 3000 4000 5000

Fy [N

]

30

50

70

90

110

130

150

0 1000 2000 3000 4000 5000

Fy [N

]

1 rotation

20 rotations

1 rotation

20 rotations

Fig. 6. Acquisition signals: (a) Stable condition (experiment 1), (b) Unstable condition (experiment 3).


original ones after 400 min of cutting. However, even low wearvalues led to higher tool run-out values, as can be seen in Fig. 7.This figure shows that the difference between cutting forces ateach rotation of the tool (Fy at a given point minus Fy at the nextpoint) were higher when tool had already cut during 400 min.

In experiment 3, instability of the cutting process is clear, sincethere is a significant dispersion of the peak forces in bothmoments, at the beginning (fresh tool) and mainly at the end ofthe experiment. The points form a polygon very distant from acircle, and this distance is even greater in the curve obtained fromthe tool after 400 min of cutting. This indicates that instabilityoccurred in this process from the beginning of the experiments(fresh tool) and was enhanced by the low tool flank wear after400 min of cutting.

All the auxiliary data for this analysis leads to the conclusionthat only the vibration of the cutting process in experiment3 affected the roughness results, since the other analyzed factors– Fy and tool wear – were at levels similar to those obtained inother experiments. Polli [20] stated that high amplitude vibra-tions occur when the harmonics of the TPF approach the system’snatural frequency. This fact may also explain this result since thehighest peak in the FRF curve of the tool used in experiment 3 was1369 Hz, which is the second harmonic of TPF (see Fig. 1).

Fig. 8 shows the FRF values of each tool obtained at a frequencyequal to twice the TPF (FRF in the second harmonic of the tool’snatural frequency). These values were obtained from Fig. 1 and arerelated to the value shown on the curves where the second verticalgreen line (twice the TPF) crosses the FRF curve.

The energy values were low in most of the experiments. Again,the exception was experiment 3, in which a FRF of 456.2 m/s2/Nwas obtained at the frequency of 1370 Hz (very close to the secondharmonic—1369.4 Hz), in the Y direction. Moreover, experiment4 showed a higher FRF than experiment 3 in the X direction, but the

total vibration energy of experiment 4 in the second harmonic ofTPF was much lower than that obtained in experiment 3. This iswhat probably caused the instability in the cutting process,resulting in the high roughness of the machined surface.

These results thus demonstrate that the main reason for thehigher roughness values in experiment 3 was the tool instability,as evidenced in Figs. 6–8. However, the toughness of the tool usedin this process sufficed to prevent its early damage or catastrophicfailure, enabling it to cut for at least 400 min.

It is interesting to compare the results of experiment 3 withexperiment 1. Both were performed with the same tool andcutting speed (same tool revolution) and same TPF. Therefore,both had the same excitation frequency. However, as the toolused in experiment 1 had a shorter length (L) and, consequently,lower values of FRF at the second harmonic of TPF, it was able towithstand the excitation caused by the cutting forces withoutinstability in the process.

As mentioned in the ‘‘Methods, equipments and materials’’section, a complete experiment was not performed in the condi-tion using the indexable insert tool with D¼8 mm andTSC¼45 mm�1, because it resulted in higher roughness valuesthan those obtained in experiment 3, even using a fresh tool.However, Fig. 9a shows the FRF curve of the tool, and Fig. 9bcompares the energy values in the second harmonic of the toolused in this condition with those of the integral ball nose end millof D¼8 mm and TSC¼45 mm�1 (experiment 3).

In this case, the FRF value on the Y axis at the second harmonicof the TPF is lower than that obtained in experiment 3, but ismuch higher than those obtained in the other experiments (seeFig. 8). However, what might explain the higher roughness valuesthan in experiment 3 is the high FRF in the X direction (thehighest amongst all the experiments). Therefore, in terms of totaltool vibration in the second harmonic of the TPF, this preliminary

EXPERIMENT 1

Integral

D=8,00

TSC=20

Fresh tool400 minutes170

85

0Fy

[N]

EXPERIMENT 2

Integral

D=12,00

TSC=20


85

0

Fy [N

]

EXPERIMENT 3

Integral

D=8,00

TSC=45


85

0

Fy [N

]

EXPERIMENT 4

Integral

D=12,00

TSC=45


85

0Fy

[N]

EXPERIMENT 5

Insert

D=8,00

TSC=20


85

0

Fy [N

]

EXPERIMENT 6

Insert

D=12,00

TSC=20


85

0

Fy [N

]

Fig. 7. Polar coordinates of Fy peaks.


experiment produced the highest value, causing considerable toolinstability, and hence, high surface roughness, precluding the useof this set of conditions.

Again, it is evident that high amplitude vibrations occur whenthe harmonics of TPF are close to the system’s natural frequency,

as stated by Polli [20]. In experiment 3, the roughness valuesobtained in cutting with the fresh tool were Rz¼3.14 and 1.94 mmmeasured, respectively, in the directions transverse and long-itudinal to the feed, and in the preliminary experiment, thesevalues were Rz¼9.11 and 3.96 mm.

Experiment 1 Experiment 2 Experiment 3 Experiment 4 Experiment 5 Experiment 6

D=8,00 D=12,00 D=8,00 D=12,00 D=8,00 D=12,00


INTEGRAL INDEXABLE INSERTEnergyX 44.6 20.2 125.3 203.6 32.2 15.9EnergyY 34.1 18.1 456.2 160.1 2.4 15.3

0.050.0

100.0150.0200.0250.0300.0350.0400.0450.0500.0

m/s

2 /New

ton

Fig. 8. Energy values in the second harmonic of the natural frequency of the tool/

tool-shank/machine system.

050

100150200250300350400450500

m/s

2 /N

ewto

n

Hertz

INSERT D=8,00 TSC=45XYTPF and harmonics

Experiment 3 Preliminary Experiment

D=8,00 D=8,00

TSC=45 TSC=45


Energy X 125.3 291.6

Energy Y 456.2 305.4

0

50

100

150

200

250

300

350

400

450

500

m/s

2 /N

ewto

n

Fig. 9. FRF of the tool/tool-shank/machine system of the preliminary experiment

(a), and energy values in the second harmonic of experiment 3 and preliminary

experiment (b).


4. Conclusions

Based on the results of this work, several conclusions can bedrawn from the milling of AISI H13 steel with a hardness of 50HRC (in conditions similar to those used here):

–
Wear was not a problem for the tools since, even after a longcutting time (400 min), flank wear was very slight and the toolnose shape was not unduly damaged. Surface roughness didnot increase significantly with cutting time in most of theconditions tested in this work.
–
Albeit slight, wear was responsible for the increase in tool run-out as the cutting time proceeded. However, the higher toolrun-out did not increase the surface roughness in most of theexperiments.
–
The experiment using the integral carbide tool with diameterD¼8 mm and tool slenderness coefficient TSC¼45 was theexception among the experiments. When this set of conditionswas used, the second harmonic of the tooth passing frequency(TPF) was a frequency with a high FRF value, which led to theoccurrence of tool instability. This instability augmented theworkpiece surface roughness when compared with thatobtained in the other experiments.
-
In spite of being minimal, tool wear contributed to the increaseof tool instability in this experiment, which caused the work-piece surface roughness to increase over cutting time, unlikewhat occurred in the other experiments.
–
Despite the instability, the tool used in this experiment wasnot damaged. Therefore, the tool is sufficiently tough to with-stand the high vibrations occurring in these conditions.
–
Using TSC¼20, regardless of the tool diameter, and D¼12 mm,regardless of the TSC, it is possible to perform finishingoperations and achieve high workpiece surface quality andlong tool life with both integral tools and indexable carbideinsert tools.Good workpiece surface roughness allied to long tool life oflong slender tools can be achieved provided the tooth passingfrequency used in the milling process (and its harmonics) doesnot produce high FRF values, thus preventing the occurrence ofinstability during milling.
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Correlating surface roughness, tool wear and tool vibration in the milling process of hardened steel...

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