Comparative study of the machinability of austempered and pearlitic ductile irons in drilling...

TECHNICAL PAPER

Comparative study of the machinability of austemperedand pearlitic ductile irons in drilling process

Patrıcia Alves Barbosa • Eder Silva Costa •

Wilson Luiz Guesser • Alisson Rocha Machado

Received: 15 May 2012 / Accepted: 2 April 2013 / Published online: 27 March 2014

� The Brazilian Society of Mechanical Sciences and Engineering 2014

Abstract The aim of this work is to analyze the aus-

tempered ductile iron (ADI) feasibility to substitute the

conventional ductile iron in automotive components man-

ufacture. This analysis was based on machinability criteria

comparison of thrust force, torque and cutting power of two

ADI grades (ISO 800-10 and ISO 1.050-6) with FE 70003

pearlitic ductile iron as-cast. Cutting speed, feed rate and

cutting fluid application technique, besides the work

material, were varied in the drilling tests. The tool used in

the tests was the K20 cemented carbide twist drill coated

with multilayers of TiN/TiAlN with diameter of 10 mm.

The results showed that the ISO 800-10 ADI grade was the

material with highest machinability followed by FE 70003

pearlitic ductile iron as-cast and then by ISO 1.050-6 ADI

grade. It was concluded that the machinability of the

materials studied is thereafter closely related to their

microstructures. Therefore, ADIs obtained by DI partially

austenitized have excellent potential to substitute ductile

irons based on the machinability criteria evaluated.

Keywords Machinability � Drilling process �Austempered ductile iron � Pearlitic ductile iron

Abbreviations

ADI Austempered ductile iron

ANOVA Analyses of variance

DI Ductile cast iron; Pearlitic ductile iron

f Feed rate, mm/rev

FZ Thrust force, N

Lf Feed length, mm

MQF Minimum quantity of fluid

MZ Torque, Nm

p Significance level

Pc Cutting power consumption, W

Vc Cutting speed, m/min

1 Introduction

In the industry, increasing demands for new engineering

materials are driven by the cost as well as the ‘‘strength to

weight ratio’’. In the effort to achieve these requirements

metallurgists either apply new materials with similar

strengths but lower densities, or increase the strength of

traditional material by alloying or by heat treatment. The

option chosen depends on parameters such as the

mechanical and thermal loads or on boundary conditions

such as manufacturing costs, recyclability, public accep-

tance and machinability [12]. In this list are the cast irons,

which almost always offer good machinability and low

manufacturing cost. In this regard, heat treatment of ductile

Technical Editor: Alexandre Abrao.

P. A. Barbosa (&) � E. S. Costa � A. R. Machado

Department of Mechanical Engineering, Federal University of

Uberlandia, Uberlandia, MG 38408-902, Brazil

e-mail: [email protected]

E. S. Costa


A. R. Machado


W. L. Guesser

Tupy S.A., Joinville, SC 89206-900, Brazil


W. L. Guesser

Center of Technological Sciences, Santa Catarina State

University, Joinville, SC 89223-100, Brazil

123

J Braz. Soc. Mech. Sci. Eng. (2015) 37:115–122

DOI 10.1007/s40430-014-0161-z

cast irons using different routes offers a wide range of

properties through matrix microstructure control, making

these cast irons very efficient materials [25].

The history of ductile cast iron (DI), also known as

nodular, has been influenced by a large number of technical

developments that have resulted in materials offering new

business opportunities for the foundry industry [13, 29]. DI

presents good castability, damping capacity, and mechan-

ical properties [high ductility (up to more than 18 %), high

toughness, and high strength (up to 850 MPa)] as well as

acceptable machinability [5]. Because of these good

properties, nodular cast iron has been used in many struc-

tural applications. Camshafts, crankshafts, brake disks,

pump cases, steering parts are major examples of vehicles

parts made of ductile iron [25]. Usually, the nodular can be

machined with the same techniques used for the gray cast

irons due to its similar behavior during machining, despite

the best mechanical properties in relation to the latter [26].

Austempered ductile iron (ADI) is a heat-treated ductile

iron, which was developed quite recently. Due to its

attractive properties such as good ductility, substantial

toughness, good wear resistance and high fatigue strength

with low cost [4, 18–20], ADI has rapidly emerged as an

important engineering material. Nowadays, ADI is con-

stantly used in automotive, military, earth moving and

agricultural equipment, and railway industries. ADI is used

in many structural applications, in situations involving

impact, fatigue and wear, such as in spring hanger of

trucks, transport components in coal mining and gear [3,

11, 14]. In the automotive industry, ADI is widely used for

the manufacturing of chassis and suspension system parts,

as well as engine components, such as crankshafts, valves

and connecting rods [27].

Regarding to contact fatigue strength, Dommarco and

Sanvande [9] investigated the resistance of ADI and

found great influence of the austempering temperature.

Higher hardness and contact fatigue strength were

observed in the ADI austempered at a lower temperature

of 240 �C (compared to samples austempered at up to

300 �C). Brunetti et al. [3], however, point out that the

specimen preparation mode (whether gridding or polish-

ing) has an influence on the results when using ball-on-

flat test system. Higher strength was found for polished

specimens.

As to abrasive wear resistance of ductile cast iron, Velez

et al. [28] have studied the performance of several cast

irons with different matrix microstructures (ferrite, pearlite,

bainite and martensite) in a single pass pendulum device,

and found that the resistances of the materials depend on

their mechanical properties and test conditions. Pearlite

matrix showed best performance among materials with

hardness up to 300 HV; bainite was the best material

among those with hardness up to 450 HV and materials

with hardness above 500 HV; the martensite did not per-

form well due to its lack of toughness. Sahin et al. [20] also

studied the abrasive wear behavior of ductile iron and the

mechanisms involved in pin-on-disk tests. Several dual

matrix structures were investigated (austempered and

quenched). The abrasive wear resistance was higher for the

austempered specimens than for the quenched ones. In all

the austempered specimens, the abrasive wear increased

with austempering time.

As compared to conventional ductile iron and steels with

similar strengths, ADI has shown several economic

advantages, namely, in automotive industry [16]. However,

it is very difficult to machine due to its relatively high

hardness, which derives from its particular microstructure

obtained after heat treatment [7, 10]. Another fact that can

account for the poor machinability of ADI is the presence

of retained austenite in its microstructure. The content of

this retained austenite increases with reduced austempering

time [4]. Retained austenite can be transformed into mar-

tensite during machining [23]. Because of this poor

machinability, Brandenberg (2001, apud [2]) suggests to

machine ADI components in three different routes: (1)

machining before heat treatment, (2) rough machining prior

to heat treatment and finish machining thereafter, and (3)

machining after heat treatment. The choice among these

various ways depends upon the dimensional requirements,

costs and manufacturing time. An alternative to the tradi-

tional austempering heat treatment would be to perform

partial austenitizing so that the final microstructure would

consist of primary ferrite and ausferrite. Such dual micro-

structure will then allow an affordable machinability of

ADI components [14]. This novel heat treatment of ADI

together with many other possible options of controlling

the ADI microstructure has stimulated research on the

machinability of ADI.

Among all the conventional machining processes

involved in the manufacturing industry, about 25 % of

machining time is spent on drilling operations [22]. This

shows that drilling is one of the most important

machining processes in the manufacturing industry.

Accordingly, researches on the machinability of ADI,

particularly in the drilling process [17, 25], are thus of

utmost importance for the automotive industry in the

context where ADI is proposed as an alternative material

to nodular cast iron.

The graphite particles embedded in the microstructure of

cast irons act already as a solid lubricant during machining.

Thus, in cast iron machining, cutting fluids are not usually

recommended, due to their many environmental hazards.

However, mineral oil emulsions are still recommended for

drilling. They can be useful for cooling and helping chip

removal from the cutting area. Since the thermal conduc-

tivity of ADI is lower than that of many steels, cutting fluid

116 J Braz. Soc. Mech. Sci. Eng. (2015) 37:115–122

123

application during drilling would be required to reduce the

temperature ahead of cutting area [2].

Knowledge of drilling forces is of major significance,

because they can affect the cutting power, the ability to

achieve tight tolerances, as well as the cutting temperature

and tool wear [8]. During metal cutting, a drill is basically

subjected to torsion and compressive stresses due, respec-

tively, to its rotation and thrust motions. Thereby,

machinability study of ADI based on torque and thrust

force (Fz) developed will be particularly relevant for dril-

ling operations.

In this sense, the aim of the present work is to analyze

the ADI feasibility to substitute the conventional ductile

iron in automotive components manufacture. This analysis

was based on machinability criteria comparison of thrust

force, torque (MZ), and cutting power (Pc) of two ADI

grades (ISO 1.050-6 and ISO 800-10) with FE 70003

pearlitic ductile iron as-cast. The first ADI is composed of

an ausferrite matrix, while the second ADI has a dual

ferrite—ausferrite matrix. The cutting conditions (cutting

speed, and feed rate), cutting fluid application technique,

and work material were varied in drilling tests.

2 Materials and methods

2.1 Workpieces and cutting condition

Workpieces of nodular cast irons, whose chemical com-

position is shown in Table 1, were used under three dif-

ferent matrix types:

• DI (FE 70003): pearlitic ductile iron as-cast;

• ADI I (ISO 1.050-6 ADI): pearlitic ductile iron fully

austenitized at 900 �C for 2 h and austempered at 360 �C

for 2 h;

• ADI II (ISO 800-10 ADI): pearlitic ductile iron par-

tially austenitized into critical zone at 780 �C for 5 h and

austempered at 360 �C for 2 h.

The austempering temperature was kept constant at

360 �C to maintain similar microstructural features of

ausferrite in ADIs. Mo and Cu were introduced in the

chemical composition of nodular cast irons to provide

hardenability (Table 1).

Bars of 104-mm diameter were produced by continuous

casting by Tupy S.A. Disk samples with thickness of

32 mm were cut out of these bars. Thereafter, these disk

samples were faced on a lathe to maintain a standard

thickness of 30 mm. This facing helped avoiding the

slippage of drills at the beginning of the drilling operation.

Table 2 shows the matrix type, graphite characteriza-

tion, and hardness of the work materials used in the current

investigation studied. Figure 1 shows the typical micro-

structures of them.

TiN/TiAlN multilayers coated K20 grade Sandvik Cor-

omant CoroDrill� Delta C-840 cemented carbide twist

drills with diameter of 10 mm, helix and point angles of

30.6 � and 143 �, respectively, were used for all the

experiments. In each drilling test, a new tool was used.

2.2 Experimental procedure

The drilling tests were carried out in a Discovery 760—

Romi CNC machine center. This machine tool has 15 kVA

of the apparent power, and 10,000 rpm of maximum

spindle rotation speed. The trials followed two full factorial

planning 24, considering two quantitative variables [cutting

speed (Vc) and feed rate (f)], and two qualitative variables

(work material and cutting fluid application technique).

The first full factorial planning and the analyses of variance

(ANOVA) by Statistica 6.0 software were used to compare

the machinability of the two ADI grades and the significant

variables effects on the test responses evaluated with a

confidence interval of 95 and 5 % of significance level.

Then, results from each ADI grade were compared to those

from as-cast pearlitic ductile iron. Table 3 shows the input

variables at two levels.

Table 1 Pearlitic ductile

iron—FE 70003—chemical

composition (%wt)

C Si Mn P S Cr Cu Mo Mg

3.600 2.500 0.250 0.050 0.015 0.060 0.460 0.23 0.040–0.050

Table 2 Matrix type, graphite

characterization and Brinell

hardness of work materials

Material Matrix type Graphite Brinell

hardness

Shape Nodularity

(%)

Nodule

(p/mm2)

Grain

size

HB 5/750

DI FE 70003 Pearlite with *2 % of ferrite VI 90 182 5–6 269

ADI I ISO 1.050-6 Ausferrite VI 90 195 5–6 288

ADI II ISO 800-10 Ausferrite with ferrite and

residual spheroidized perlite

VI 90 218 5–6 203

J Braz. Soc. Mech. Sci. Eng. (2015) 37:115–122 117

123

Through holes were machined with total feed length (Lf)

of 35 mm to allow the main cutting edges of the drill to

exit completely out of the button face of the disks (5 mm).

Thrust force (Fz) and torque (MZ) were online monitored

during drilling tests with a Kistler type 9123C rotating

dynamometer using a Kistler type 5223 signals conditioner

and Labview� 6.5 signal processing software by National

Instrument Brazil Ltd, in an acquisition rate of 2 kHz. The

cutting power was indirectly determined through the cur-

rent and voltage of the electric signal acquired from the

spindle motor of the machine center using NW-SCD-50-R

Hall-effect transducers and Labview� 6.5 in an acquisition

rate of 2 kHz. The cutting power (Pc) was calculated by the

difference between power values recorded when running a

drilling test and when running free the machine center.

Figures 2 and 3 show schematically the setups used to

monitor, respectively, the thrust force and torque as well as

the cutting power.

3 Results and discussion

3.1 Thrust force results

Thrust force Fz results per material for different cutting

conditions are plotted in Fig. 4. From this figure, it can be

seen that ADI II—ISO 800-10 grade was the material that

presented the lowest values of Fz, and therefore the best

machinability. However, ADI I and DI have showed great

similarity considering the thrust force, regardless the lubri-

cooling system and cutting condition used.

ANOVA results of thrust force presented in Table 4

show that there is a significant difference between the ADI

grades (p \ 0.05). In Table 5 it can also be seen that there

Fig. 1 Microstructures of materials tested: a DI—pearlitic ductile iron (FE 70003) with pearlitic matrix; b ADI I—ISO 1.050-6 ADI grade with

ausferrite matrix; c ADI II—ISO 800-10 ADI grade with ausferrite and primary ferrite

Table 3 Matrix type, graphite characterization and Brinell hardness

of work materials

Variable Level (-1) Level (?1)

Vc (m/min) 110 140

f (mm/rev) 0.15 0.25

Material ADI II/DI ADI I/DI

Cutting fluid Flood (982 l/h) MQF (50 ml/h)

Fig. 2 Schematic setup for

thrust force and torque signals

acquisition and measurement


123

is a significant difference between the DI and ADI II.

However, no significant difference between the DI and

ADI I was observed in Table 6. The feed rate (f) was the

variable that has always significant influence on thrust

force—Fz. Both cutting speed (Vc) and cutting fluid

application technique did not significantly influence on the

response.

Workpiece material hardness increases drilling resis-

tance, i.e., the higher the shear strength, the higher the cutting

forces [15]. Thus, when drilling ADI I (288 HB) instead of

ADI II (203 HB), an average increase in thrust force of

Fig. 3 Schematic setup for

power consumption signal

measurement

Fig. 4 Thrust force—Fz— a flood cooling, b MQF

Table 4 ANOVA of full factorial planning 24 for Fz to comparison

of ADI I with ADI II

Effect SE p

Mean 1,842.688 16.23858 0.000000

Vc (m/min) 50.625 32.47716 0.179787

f (mm/rev.) 594.625 32.47716 0.000009

Cutting fluid application

technique

-17.875 32.47716 0.605739

Material 834.875 32.47716 0.000002

Vc 9 f -56.125 32.47716 0.144538

Vc 9 cutting fluid application

technique

-76.625 32.47716 0.064803

Table 5 ANOVA of full factorial planning 24 for FZ to comparison

of DI with ADI II

Effect SE p

Mean 1,829.563 17.10658 0.000000

Vc (m/min) 31.625 34.21316 0.397721

f (mm/rev.) 631.875 34.21316 0.000009


technique

66.875 34.21316 0.108024

Material 808.625 34.21316 0.000003


123

835 N should be expected (Table 4). Likewise, when

machining DI (269 HB) instead of ADI II (203 HB), an

average rise of Fz about 809 N should be expected (Table 5).

The feed rate influence on thrust force was clear and within

expectation. The higher the feed rate, the higher the thrust

force. Increasing feed rate (f) increases the areas of the primary

and secondary shear planes. Higher forces and power are thus

required to form the chip. The feed velocity is also increased in

a direct proportion, thereby also contributing to the increase of

thrust force. Accordingly, the increase of feed rate from 0.15

to 0.25 mm/rev should raise the thrust force by about 595 N

when comparing the two ADI grades (Table 4), by about

632 N when comparing DI and ADI II (Table 5), and by about

614 N when comparing DI and ADI I.

3.2 Torque results

MZ results per material for different cutting conditions are

plotted in Fig. 5 and its ANOVA analysis is presented in

Table 7.

The ANOVA results of the thrust force—Fz are shown

in Tables 4, 5, 6.

When comparing, respectively, the two ADI grades as

well as DI with ADI II, the ANOVA of torque (MZ)

showed the same trend as that of thrust force—Fz. How-

ever, the torque appeared more sensitive than the thrust

force when comparing DI (269 HB) with ADI I (288 HB)

as shown in Fig. 5, suggesting a significant difference

between the materials (Table 7). Accordingly, the corre-

sponding increase in workpiece material hardness should

lead to a torque rise of about 1.46 Nm. The feed rate also

has a significant influence on torque, promoting an average

increase of 2.44 Nm when its level increases from 0.15 to

0.25 mm/rev. Similarly as in Fz results, the cutting speed

and cutting fluid application technique did not significantly

influence on MZ.

Table 6 ANOVA of full factorial planning 24 for FZ to comparison

of DI with ADI I

Effect SE

(Rev B)

p

Mean 2,247.000 20.34223 0.000000

Vc (m/min) 64.750 40.68446 0.172369

f (mm/rev.) 614.000 40.68446 0.000023


technique

32.500 40.68446 0.460633

Material -26.250 40.68446 0.547226

Fig. 5 Torque—MZ— a flood cooling, b MQF

Table 7 ANOVA of full factorial planning 24 for MZ to comparison

of DI with ADI I

Effect SE p

Mean 7.03813 0.239682 0.000001

Vc (m/min) 0.50375 0.479363 0.341431

f (mm/rev.) 2.43875 0.479363 0.003810


technique

-0.20125 0.479363 0.692041

Material 1.45875 0.479363 0.028649

Fig. 6 Cutting Power—Pc— a flood cooling, b MQF


123

3.3 Cutting power results

The cutting power results for different cutting conditions

are shown in Fig. 6.

The higher machinability of the ADI II material can also

be clearly seen from the cutting power results shown in

Fig. 6. ADI I presented the lower machinability. The

machinability criterion of cutting power—Pc revealed that

the most severe drilling conditions occurred when using the

higher cutting speed and feed rate (140 m/min and

0.25 mm/rev, respectively) with the flood cooling.

The ANOVA results of the cutting power, with a con-

fidence interval of 95 and 5 % of significance level are

shown in Tables 8, 9, 10.

It is clear that there is a significant difference between

the two ADI materials (Table 8). By order of importance,

the parameters that significantly affect the cutting power

were the workpiece material, feed rate and cutting speed.

The cutting power should increase by 882 W when drilling

ADI I (288 HB) instead of ADI II (203 HB). It should

increase by 662 W when the feed rate increases from 0.15

to 0.25 mm/rev or by 554 W when the cutting speed

increases from 110 to 140 m/min.

The work material impact on the cutting power is

directly related to its effect on the torque. As the workpiece

material hardness increases, the cutting power should be

raised when planning to drill a hole. The feed rate has a

direct relationship with thrust force, as seen in Tables 4, 5,

6. If the effect of thrust force on the cutting power is

established, the feed rate could have similar effects on the

cutting power [15]. The impact of cutting speed on the

cutting power should be expected somehow taking into

account the relationship linking these two quantities, i.e.,

since Pc in cutting processes is calculated by the product of

cutting force and the cutting speed [24].

Just as expected, the cutting-power ANOVA results when

comparing, respectively, the two ADIs (Table 8) as well as

DI and ADI II (Table 9) are consistent. When drilling DI

instead of ADI II, the cutting power should increase to about

544 W. Increasing the feed rate from 0.15 to 0.25 mm/rev

should raise the cutting power of about 574 W. Likewise,

increasing the cutting speed from 110 to 125 m/min should

increase the cutting power to about 476 W. Thus, these

trends revealed that, as compared to the cutting speed and

workpiece material, the feed rate has a bigger impact on the

cutting power. The cutting fluid application technique

remains without significant influence on the responses.

Similarly, when comparing DI with ADI I, the torque

ANOVA results (Table 7) together with the cutting power

ANOVA results shown in Table 10 show that when drilling

ADI I (288 HB) instead of DI (269 HB), the cutting power

should increase to about 399 W. Likewise, these results

show that when increasing the feed rate from 0.15 to

0.25 mm/rev, the cutting power should increase to about

649 W. They also show that when increasing the cutting

speed from 110 to 125 m/min, the cutting power should

increase to about 650 W.

It emerged that the thrust force, torque and cutting

power—Pc—results (Figs. 4, 5, 6) showed similar trends as

expected, the cutting power being somehow proportional to

the tangential cutting force (torque) and the latter

depending on the workpiece material shear strength [1, 6].

As the cutting fluid application technique did not show

significant influence on the machinability parameters

studied (force torque and power consumption), this could

suggest that minimum quantity of fluid (MQF) is an

interesting choice because of its much more environmen-

tally friend system [21]. However, in terms of production,

before any move towards this direction, the MQF system

has to prove good enough in terms of tool life and surface

quality as well.

Table 8 ANOVA of full factorial planning 24 for Pc to comparison

of ADI I with ADI II

Effect SE p

Mean 1,879.563 55.3207 0.000000

Vc (m/min) 554.125 110.6414 0.004076

f (mm/rev.) 662.375 110.6414 0.001865


technique

-70.875 110.6414 0.549994

Material 882.375 110.6414 0.000500


of DI with ADI II

Effect SE p

Mean 1,710.250 16.26672 0.000000

Vc (m/min) 475.750 32.53344 0.000027

f (mm/rev.) 573.500 32.53344 0.000011


technique

-2.000 32.53344 0.953362

Material 543.750 32.53344 0.000014


of DI with ADI I

Effect SE p

Mean 2,151.438 59.6271 0.000000

Vc (m/min) 605.625 119.2542 0.003840

f (mm/rev.) 649.125 119.2542 0.002841


technique

-62.125 119.2542 0.624645

Material 338.625 119.2542 0.036267


123

4 Conclusions

The thrust force, torque and cutting power increased in the

following order of the workpiece materials tested: ADI II,

DI and ADI I. This sequence respects the hardness

increasing order of these materials (203, 269, and 288 HB).

ADI II which presents a dual matrix of primary ferrite and

ausferrite outperforms the two other workpiece materials

with regard to machinability in drilling process. Thus,

partially austenitized ADIs are contribution that should

guide and strengthen the plans of substituting, at least in

the automotive industry, pearlitic ductile iron to ADI.

ANOVA analyses allowed identifying the material

machinability differences, showing also that feed rate is the

most impacting cutting parameter on the output variables

investigated.

In spite of cutting fluid application is required during

ADI drilling to reduce the temperature ahead of the cutting

area, statistically it did not affect thrust force, torque and

cutting power results significantly. The different techniques

evaluated showed similar results. MQF system is likely to

be the choice if only environmental approach is the goal.

However, in terms of productivity other machinability

criteria, such as tool life and surface roughness should also

be considered before any decision.

Acknowledgments The authors are grateful to Capes, CNPq, Fap-

emig, and IFM for their financial supports. They are also indebted to

Tupy S.A., Sandvik do Brazil S.A., and LEPU/UFU for providing,

respectively, the work materials, cutting tools and infrastructure

necessary for handling this research.

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