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Materials Science and Engineering A 460–461 (2007) 314–323

Study of machinability in boring operation of microalloyedand heat-treated alloy steels

A. Ebrahimi ∗, M.M. MoshksarDepartment of Materials Science and Engineering, School of Engineering, Shiraz University, Shiraz, Iran

Received 15 November 2006; received in revised form 15 January 2007; accepted 16 January 2007

bstract

The machinability of microalloyed steel (38MnVS5) and heat-treated alloy steels (AISI 1045 and AISI 5140) with the same hardness areresented in the paper. In this investigation, the effects of cutting speed, cutting force and feed rate on flank wear land and tool life of TiN coatedarbide inserts in boring operations were studied. The criteria of flank wear (VBB max = 0.3 mm) used, when boring of these steels at different cuttingpeeds, to determine the tool life of inserts. SEM micrographs of flank wear and chips of the steels obtained in boring operations were analyzed.

hip characteristics in machining of these three steels interested. Shear planes and microcracks of the chips in microalloyed steel showed that

he chips of microalloyed steel are regular and discontinuous. The results showed that the machinability of microalloyed steel is better than theeat-treated alloy steels at the identical conditions. 2007 Elsevier B.V. All rights reserved.

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eywords: Machinability; Boring; Microalloyed steel; Flank wear; Chip

. Introduction

Microalloyed medium carbon steels have been widely useds substitutes for quenched and tempered (heat-treated) steelsince they can remarkably reduce production costs of machinedarts by saving heat treatment, logistics cost and lead-time ofroduction process. Machining is a versatile shaping processf major importance for component manufacturing. The mainenefit of microalloyed steels lies in the prospect of impor-ant energy and cost savings in the manufacturing of forgedomponents for automotive applications. In such steels, thetrength levels and other properties achieved after cooling fromot working temperatures are reported to be comparable withhose obtained from conventional quenched and tempered steels1,2].

Microalloying forging steels aim to produce, on controlledir-cooling after forging, a microstructure consisting of pearlitend ferrite, with both microconstituents being strengthened by

he presence of fine precipitates of vanadium carbonitride. Theormation of bainite on cooling is avoided because it is gen-rally regarded as being deleterious to ductility and impact

∗ Corresponding author. Tel.: +98 711 6286531; fax: +98 711 6287294.E-mail address: aliebrahimii@yahoo.com (A. Ebrahimi).

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921-5093/$ – see front matter © 2007 Elsevier B.V. All rights reserved.oi:10.1016/j.msea.2007.01.050

oughness. Microalloyed steels are carbon steels, which achieveheir strength properties in the as-rolled and as-forged conditionithout going through a quenching and tempering treatment.hese steels contain small amounts, generally <0.25 wt%, ofne or more of Cb, V, Ti and Mo which combine with carbonnd/or nitrogen to precipitate and import precipitation hardening3,4]. However, most, if not all, forging go through considerableachining after forging for manufacture of the final part and

he contribution of the machining operation to the total cost ofanufacture is often 50–60%. The machining properties of theicroalloyed steel as compared to heat-treated steel at similar

trength level, is therefore, of great interest in the application ofhese steels.

There are many sets of operation and criteria for machin-bility and the same materials may have a different machiningesponse when a different set of machining operations is used.o deal with this complex situation, the approach adopted in the

iterature has been to address several criteria to assess machin-bility. One or more of the following thus, characterize improvedachinability:

. increased tool life obtainable,

. a higher rate of material removal,

. lower energy/forces, and

. easier chip removability.

A. Ebrahimi, M.M. Moshksar / Materials Science a

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obdmovement with the chip releases particles. The junction betweeneach chip segment can be broken [15–17], but if the particlesadhere to the tool they will eventually alter its geometry shape.The adhesive mechanism depends on the contact force between

Fig. 1. Illustration of forces distribution in boring process.

The enlargement of holes is achieved by using boring oper-tions. The hole diameter is either enlarged with a single insertttached to long boring bar, or with a boring head which hasdiameter equal to the diameter of the hole to be enlarged.

orces distribution along the cutting edge of the boring processs shown in Fig. 1. Excessive static deflections may violate theimensional tolerance of the hole, and vibrations may lead tooor surface, short tool life and chipping of the tool [5–8].

The tool life of tools is the criterion to explain the machin-bility of these steels. The contact phenomena on the tool flankurface are of interest because understanding them allows thexplanation of tool flank wear and the formation of surfacentegrity of machined surfaces on workpieces. Flank wear isonsidered using an aged Taylor tool life equation (VTn = c)aving a phenomenological nature [9]. Machining workpieceaterials prone to work hardening can generate high tempera-

ures, welding on the tool surface in addition to high resistancef metal removal because of their high shear strength duringachining. These usually lead to a shorter tool life, pronounced

hipping/premature failure of the cutting edge, severe surfacend production of long continuous chips [10]. Since hardnessas a major role on the wear of tools, by using heat treatment,ardness of these steel is equaled.

This paper attempts to study machinability of microalloyed38MnVS5) and heat-treated (AISI 1045 and AISI 5140) steels.n addition to strength and fatigue properties, machinability iscritical property of microalloyed bar steels, since these steelso through considerable machining operation following forging.his work was therefore undertaken to evaluate the machinabil-

ty of microalloyed steel. The results were then compared tohose obtained on normally used alloy steel quenched-temperedo the same hardness. Machinability was evaluated using bor-

ng test with cemented carbide tools. The results show that iningle point boring, the machining of microalloyed steel is supe-ior (at tool life, cutting force, and chip morphology). Studies of

Fs

nd Engineering A 460–461 (2007) 314–323 315

he chip formation process indicate that the microalloyed steeleacts differently in terms of chip thickness and shear flow layer.hus by substitution of microalloyed forging steel for quenched-

empered steel, cost saving, by omitting heat treatment cyclesfter forging and reduced cost of machining of microalloyedteel, is achievable.

. Cutting tool wear

The severe forces taken by the tool’s rake face and flankause cutting tool wear. These loads are accentuated by ther-al phenomena in the high speed cutting process, thus toolaterials must exhibit good hardness and tenacity properties.urthermore, high temperatures can cause physical–chemicalhenomena, which exacerbate tool wear on the face. The typef wear depends on the nature of the tool, machined materiel,utting conditions and the type of cutting [11,12]. It can be dis-inguished between wear in the crater, wear on the flank face,reakage of the cutting edge (chipping), and wear by plasticeformation. Flank wear model is shown in Fig. 2. The presentork involves assessment of flank wear appearing on the cutting

dge of tool.

.1. Wear by abrasive and adhesive

Wrenching of tilted and hard particles on the surface of theachined leads to abrasive wear of the cutting surfaces. The

utting process, continuously releases these particles, typicallyrecipitates or inclusions in the material. The speed of abra-ive wear increases as the abrasive particle quantity and/or sizencrease [13,14]. Consequently, abrasive wear depends on thehemical composition of the machined material and on the pro-esses used to develop the material into inserts. This type ofear is typical of low cutting feeds and speeds.The mechanical wrenching of particles after the development

f microwelding also causes adhesive wear. High local forcesetween the rough surfaces of the chip and tool assist this sol-ering. Forces shear the metallic junctions, and this shearing

ig. 2. Flank wear land characteristics due to ANSI/ASME B94.55M-1985tandard.

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16 A. Ebrahimi, M.M. Moshksar / Materials

he tool and the chip, on the crystal structure of the materialsnd on their flow strains.

.2. A flank wear model for orthogonal metal cutting

It can be found in previous wear models that the hardnessand normal stress σt are two of the main parameters closely

elated to wear and the wear rate is usually proportional to σtnd inversely proportional to H [18–20]. Shaw and Dirke [21]resented an Archard [22] type of equation for adhesive wear:

W = Arc

bZ dL (1)

here dW is the wear volume for the sliding distance dL, Arhe real area of contact, c the height of the postulated plate-likeear particle, b the mean spacing of the asperities and Z is therobability of producing a wear particle per asperity encounter.

Usui and Shirankashi [23] rewrote Eq. (1) by introducing theormal stress at the contact surface as follows:

W = σt

H

c

bZ dL (2)

y mathematically combining cZ/b into a coefficient K, a wearquation can be obtained as follow:

W = Kσt dL

H(3)

The wear volume can be converted to the wear land measure-ent VB in Eq. (4) according to Shaw’s analysis [24] with the

ssumption of a zero degree rake angle of the cutting tool:

= 1

2wV 2

B tan θ (4)

here W is the volume worn away on the relief face, VB theeasurement of wear land on the tool, w the width of cut and θ

s the relief angle of the tool (Fig. 3). The apparent area of theool-work contact interface is as follows:

= wVB (5)

It is assumed that the normal stress at the flank face is:

t = Ft = Ft (6)

A wVB

here Ft is the thrust force in cutting operation.It should be noted that VB is a variable, which increases with

he development of flank wear. By substituting Eqs. (4) and (6)

Fig. 3. Schematic of orthogonal cutting with zero rake angle.

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e and Engineering A 460–461 (2007) 314–323

nto Eq. (3) to predict flank wear in cutting can be obtained asollows:

B = K

(2

w2 tan θ

)1/3(FtL

H

)1/3

(7)

r

B = K

(2Vc

w2 tan θ

)1/3(Ftt

H

)1/3

(8)

here K is a coefficient, which can be determined by experi-ent, L the cutting length, H the hardness of the cutting tool,c the cutting speed and t is the cutting time. By combining allarameters except cutting time, t, in the right side of Eq. (8) intocoefficient K1, Eq. (8) was written in a form to demonstrate

he relationship between the flank wear VB and cutting time t ashown in Eq. (9).

B = K1t1/3 (9)

. Experimental procedures

Machinability was evaluated by using boring tests. The bor-ng tests were carried out by using a conventional CNC lathe.utting inserts used in the tests were CNMG120412 (P20) car-ide inserts. The tool holder used was a Hertel MCLNR20.utting conditions selected as follows: depth of cut 1 mm, feed

ate 0.2 and 0.25 mm rev−1 were kept constant and cuttingpeeds in range of 70–200 m min−1 and machining carried inry condition. By using optical video microscope and SEMechnique, the width of flank wear land of the carbide toolsfter boring operation was measured and determined. To under-tand the machinability test results and possible mechanismsehind observed machinability differences, wear of tools andhip samples were examined extensively. Chip morphologyas observed and chip thicknesses were measured. Chips were

xamined microscopically and microhardness readings werebtained across the chip thicknesses.

.1. Trial materials

The trial materials for evaluation machinability were8MnVS5 microalloyed steel and AISI 1045 and AISI 5140eat-treated alloy steels. Since hardness has a major role inool wear, by using heat treatment, hardness of the heat-treatedteels were equaled to the microalloyed steel. By consideringo equilibrium diagram of Fe–%C, the appropriate heat treat-

ent carried out for steel samples of AISI 1045 and AISI 5140.hese samples were quenched tempered steels, and providedy industry after hot rolling. The workpieces were austenitizedt 850 ◦C and then quenched in appropriate oil. The quenchedamples had high-level hardness, thus tempering of these steelsarried out at 600 ◦C for specific time and then cooling in the air.

fter cooling, measured hardness of AISI 1045 and AISI 5140

n BHN criterion was equal to the workpieces of 38 MnVS5245–250 BHN) steel. The chemical compositions of the steelsnvestigated are given in Table 1. Mechanical properties of the

icroalloyed and heat-treated steels are presented in Table 2.

A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323 317

Table 1Chemical composition of the steels used in boring operation

Sample C% Si% Mn% P% S% Cr% Mo% Ni% V% N% Cu%

AISI 1045 0.43 0.3 0.78 0.035 0.035 0.38 0.12 0.06 – – –AISI 5140 0.42 0.29 0.7 0.004 0.01 0.98 0.19 0.07 – – –38MnVS5 0.36 0.56 1.45 0.011 0.041 0.1 0.001 0.08 0.09 0.019 0.19

Table 2Mechanical properties of microalloyed and quenched-tempered steels

Material Y (N/mm2) UTS (N/mm2) r (%) Impact (J/cm2) Hardness BHN

AISI 1045 480 710 18 42 245A3

4

4

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4

o

TR

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1112

1112

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ISI 5140 570 7908MnVS5 585 850

. Results and discussion

.1. Effect of cutting speed and feed rate

Table 3 shows recorded tool life at cutting conditions investi-ated. There was a gradual reduction in tool life with increasingutting speed up to 200 m min−1. These results show that theicroalloyed steel at least is equivalent to, or better than, the

eat-treated steels at the same hardness. Flank wear may be dueo adhesive wear or abrasive wear caused by the hard secondhases in the work material. The presence of Mo in quenched-empered steels is considerable in comparison to microalloyedteel, thus, causes to presence of hard abrasive carbide particlesuch as Mo2C and FeMo2C in the matrix of the heat-treatedlloy steels grades. Table 3 also shows the effect of feed ratehen boring of the steels. An increase in feed rate from 0.2 to.25 mm rev−1 reduced the tool life. Increase in cutting parame-ers indicates that the interaction of process energies, includingemperature, increases generally and can therefore promotearious wear mechanisms, which accelerate tool wear the con-equent reduction in tool life.

.2. Tool wear and damage

In this investigated, the flank wear mechanisms and damagef the tool examined. Solid carbide inserts used in boring test of

able 3ecorded tool life at different cutting speed in boring of steels

utting speed,(m min−1)

Feed rate, f(mm rev−1)

Tool life (min)

38MnVS5 AISI 1045 AISI 5140

70 0.2 48 43.7 4100 0.2 40 35 30.530 0.2 33.2 26.5 24.660 0.2 25.5 18.6 17.600 0.2 16 10 5

70 0.25 37 30 2800 0.25 28.5 22 2030 0.25 19 16.5 1460 0.25 14.6 11 1000 0.25 6.5 3 3

uttwbfiafmflitfl1wta

16 44 24518.5 55 245

rial materials are shown in Fig. 4 examined by scanning electronicroscope (SEM). Fig. 4 illustrates damaged inserts represent-

ng typically worn tool in boring tests. Mechanisms of flankear land and damage of tool for heat-treated steels are adhe-

ive, abrasive, chipping of cutting edge and plastic deformationFig. 12). For microalloyed steel, flank wear land is affected bybrasive and adhesive. Boring heat-treated AISI 1045 and AISI140 and microalloyed 38MnVS5 steels carried at cutting condi-ion such as: cutting speed 130 m min−1, feed rate 0.2 mm rev−1,epth of cut 1 mm. Typical flake-like oxide debris adhered ono the worn cutting edges of the inserts and generally increasedith increasing cutting conditions as illustrated in Fig. 5. Thisamage occurred in machining of AISI 1045 and AISI 5140teels. Formation and adhesive of the flake-like oxide debris isresult of the metallurgical reaction arising from the interac-

ion between the tool and workpiece material during machining.his is more likely to be influenced by the composition of theork material and temperature generated. The heat-treated alloy

teels with higher chromium content had more flake oxide debrisn the tool face in comparison with microalloyed steel.

.3. Cutting speed and flank wear rate

The flank wear land in during experiment was measured bysing Video Microscope. Fig. 6 shows the comparison betweenhe cutting speed and flank wear rate when machining thesehree grades of steel with PVD TiN coated carbide inserts. Flankear rate is calculated as the average flank wear (mm) dividedy the effective tool life (min). To determine the tool life atve cutting speeds, the flank wear of the tool was measuredfter a constant interval time. The interval time was 120 ± 5 sor any state of machining. During the course of the experi-ents, the tool inserts were checked frequently for wear. Theank wear was measured at any step. The sum of interval times

ndicated the tool life of tool insert. If the flank wear is lesshan the criterion (VBB max = 0.3), repeat these steps until aank wear of 0.3 is achieved for each state. At speed of 70,

00, 130 and 160 m min−1 the flank wear at last interval timeas near to 0.3 mm, thus place in boring holder for attaining

o specific flank wear. Because of low flank wear rate for thebove speeds (slightly for 160), damage of the tool slowly pro-

318 A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323

F r 26 mf

gm(ptwtrpctt

Fc

ooo

4b

ig. 4. Flank wear of cutting tools in boring of (a) 5140 for 24 min, (b) 1045 fo= 0.2 mm rev−1.

ressed, so that in the final step of machining with these tools theacro chipping progressed and then fracture of fatigue occurred

especially in heat-treated steels). Fig. 7 shows typical of thishenomenon in machining of heat-treated alloy steel. In addi-ion, Fig. 6 shows that the microalloyed steel gives best flankear rate when cutting speed increased to 200 m min−1. At cut-

ing speed of 200 m min−1 is seen an abruptly drop in flank wearate. This increase is more for quenched- tempered steels in com-

arison with microalloyed steel. At experiment was seen that forutting condition of 200 m min−1, because of existence the highemperature and high forces (especially for heat-treated steels),he failure (plastic deformation, chipping and then fracture) was

ig. 5. Typical flake-like adhered material in machining of heat-treated steel atutting speed of 70 m min−1 and feed rate of 0.25 mm rev−1.

adcaccowpciroaovstpsw

in and (c) 38MnVS5 for 34 min, cutting speed V = 130 m min−1 and feed rate

ccurred out by progressing of operation. Fig. 8 shows typicalf fracture in cutting tool for heat-treated steel at cutting speedf 200 m min−1.

.4. Relation between cutting forces and tool life duringoring

By considering the relation between feed force, thrust forcend resultant force (F2

c = F2t + F2

f ), the main cutting force isetermined. Fig. 9 shows the plot of cutting speed against mainutting force and tool life when boring 38MnVS5, AISI 1045nd AISI 5140 steels with PVD TiN coated inserts. This figurelearly shows a decrease main cutting forces Fc with increasingutting speed. Trent [17] reported that the cutting force decreasesn increasing the cutting speed partly due to the softening oforkpiece material under the high temperature generated andartly due to decreased tool-chip contact are because of thinnerhips produced in single point turning. In orthogonal tests, theres no radial force component measured. The Eq. (7) and (8) areesulted from Eq. (1) to (6). These equations show the relationf flank wear with thrust force and hardness. For each materialt different cutting speed in boring operation, the obtained valuef K1 in Eq. (8) is exerted to Eq. (9). Fig. 10 shows feed forceersus cutting speed. In this figure, is showed that at any cuttingpeed the feed force for microalloyed steel are less than heat-

reated steels. In addition, Fig. 11 shows the relation between K1;redicted thrust force and experimental thrust force with cuttingpeed. The predicted thrust force values have a good correlationith the experimental data. The thrust force in machining of

A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323 319

F0

mtswFb

5

5

btt

Fig. 7. Typical video microscopic picture of failure after progressive machiningon flank face.

Fl

mb

5.2. Chip thickness

Study of the chip formation process provided informa-tion, which can be used to explain the machinability test

ig. 6. Cutting speed vs. flank wear when boring steels at a feed rate of.2 mm rev−1.

icroalloyed steel at different cutting speeds is less than heat-reated alloy steels at identical condition. Increase in cuttingpeed during boring causes higher temperature, increased flankear rate and localized plastic deformation of the cutting edge.ig. 12 shows the plastic deformation of the cutting edge inoring of heat-treated steels.

. Chip formation studies

.1. Chip morphology

In the entire tests carried out, the cutting tools have chipreaker or chip curler in their designs. In machining opera-ion, the heat-treated steels produced continuous chips, whilehe microalloyed steel produced smaller, spiraled chips. The

Ff

ig. 8. Typical SEM of a damaged flank wear land that operate beyond its toolife.

icroalloyed steel thus has better chip shape and should haveetter chip disposability (Fig. 13).

ig. 9. Cutting speed vs. main cutting force and tool life when boring steels ateed rate of 0.25 mm rev−1.

320 A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323

F

rass

5

atsitds

ig. 10. Cutting speed vs. feed force for microalloyed and heat-treated steels.

esults. Fig. 14 shows the chip thicknesses for the heat-treatednd microalloyed steels. In all the cases, the microalloyedteel showed lower chip thickness than the heat-treatedteels.

.3. Chip microhardness

Microhardness in the chip, at the tool edge, in the center andt the outer edge were measured for the microalloyed and heat-reated steels on sample from the boring test. The results arehown in Fig. 15. Although no firm conclusions can be drawn,

t appears that the heat-treated steels may have tempered at chipool interface due to the high temperatures generated. This is evi-enced by lower hardness at the tool edge for the heat-treatedteels. The results also show that while three steels start out

Fig. 12. Plastic deformation of cutting edge for 3 min machining of (a) AISI5140 and (b) AISI 1045 at cutting speed V = 200 m min−1 and feed rate off = 0.25 mm rev−1.

Fig. 11. Relation between K1, thrust force and cutting speed in boring operation of trial materials.

A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323 321

Fa

ahets

Fs

Ff

ig. 13. Morphology of the chips obtained at cutting speed V = 100 m min−1

nd feed rate of f = 0.2 mm rev−1.

t the same hardness level in workpiece, the heat-treated steels

arden more than the microalloyed steel and the hardness gradi-nt extends further into the chip. Both these observations suggesthat the heat-treated steels harden more than the microalloyedteel.

5

a

Fig. 15. Microhardness of the chips at various p

ig. 16. SEM micrograph of chips obtained in boring of (a) AISI 1045, (b) AISI= 0.2 mm rev−1.

ig. 14. Relation between cutting speed and chip thickness when boring ofteels.

.4. Shear flow layer and microcracks of chips

Fig. 16 shows chip samples for AISI 1045 and AISI 5140nd 38MnVS5 steels. The heavily deformed shear flow layer

ositions that obtained in boring of steels.

5140, and (c) 38MnVS5 at cutting speed V = 100 m min−1 and feed rate of

322 A. Ebrahimi, M.M. Moshksar / Materials Science and Engineering A 460–461 (2007) 314–323

(Cont

ifeuttslccd

6

asTd

Fig. 16.

s greater in the heat-treated steels than in microalloyed steelor boring test. It is seen that, the spaces between flow lay-rs in heat-treated steels are less than microalloyed steel. Thepper surface of the chips produced from the heat-treated andhe microalloyed steels in the boring test. It is seen that whilehese steels have jagged edges, there is the appearance ofmooth, small cracks in the heat-treated steels while the microal-

oyed steel shows massive and large cracking occurring in thehip. Because the massive cracks in microalloyed steel, thehips of this steel are discontinues and the shear planes easilyistinguish.

cp

d

inued ).

. Conclusions

This work set out to investigate and compare the machin-bility of microalloyed steel (38MnVS5) and heat-treatedteels (AISI 1045 and AISI 5140) during boring tests.he relation between cutting speed and work material toetermine tool life and flank wear studied. Study of the

hips obtained of machining was also the focus of thisaper.

From this investigate; the following conclusions can berawn:

nce a

1

2

3

4

5

A

8o

R

[[

[

[[

[

[[

[[

[20] B.M. Kramer, J. Vac. Sci. Technol. A 4 (6) (1986) 2870.

A. Ebrahimi, M.M. Moshksar / Materials Scie

. Increasing cutting speed and feed rate cause to decrease intool life. The reduction of tool life for heat-treated steelsis more than microalloyed steel, because the presence ofthe hard abrasive carbide particles (in matrix of heat-treatedsteels) causes to decrease the tool life.

. Microalloyed steel has better flank wear rate in com-parison with heat-treated steels. Flank wear criterion(VBB max = 0.3 mm) attained in higher tool life for microal-loyed steel, thus flank wear land of microalloyed steel isregular.

. Flank wear mechanisms and damage of tools in boring ofAISI 1045 and AISI 5140 steels are complicate. By study ofthe SEM pictures, the mechanisms such as abrasive, adhe-sive, chipping, deformation plastic were the main factors inworn tools for boring of heat-treated steels whereas, abrasiveand adhesive were the wear mechanisms for microlalloyedsteel. The plastic deformation in heat-treated steels showsthe existence of high temperature and stress in contact zone.In all the cases by progressive operation beyond the tool lifefailure and fracture of tool inserts occurred.

. The measurement of forces and chip thicknesses provide evi-dence of a basic difference in the chip formation behavior.The microalloyed steel has lower chip thicknesses leadingto smaller contact lengths, which in turn result in lowercutting forces. It is seen, that the ferrite-pearlite structureconcentrates the deformation into a thin localized region. Theheat-treated martensite, on the other hand, has a large shearflow zone. This results in a lower chip shear angle, thickerchips and higher forces.

. The ratio of the thrust force to feed force is higher forthe heat-treated (AISI 1045 and AISI 5140) steels than formicroalloyed (38MnVS5) steel. This is also an indication ofhigher forces in the secondary shear zone and agrees withobserved thicker chips in the heat-treated steels. Chip micro-

scopic examination indicates also that the ferrite-pearlitemicrostructure causes easier initiation of microfractures andthus limits the flow resistance. This leads to easier chip for-mation and machining.

[[[[

nd Engineering A 460–461 (2007) 314–323 323

cknowledgement

The financial support of this project through grant number3-EN-1689-C271 by Vice-Chancellors in charge of researchf Shiraz University is appreciated.

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