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Progressive flank wear and machining performance of silver toughened alumina cutting tool inserts
Wear 261 (2006) 885–895
Progressive flank wear and machining performance of silver toughened alumina cutting tool inserts A.K. Dutta a , A.B. Chattopadhyaya b , K.K. Ray c,∗ a
Department of Mechanical Engineering, Bengal Engineering and Science University, Shibpur, Howrah-711103, India b Department of Mechanical Engineering, Indian Institute of Technology Kharagpur, Kharagpur-721302, India c Department of Metallurgical and Materials Engineering, Indian Institute of Technology Kharagpur, Kharagpur-721302, India Received 5 March 2005; received in revised form 25 January 2006; accepted 31 January 2006 Available online 3 March 2006
Abstract Alumina-based composites with different amounts of silver as second phase have been fabricated using conventional powder metallurgical route and their microstructure, density, thermal conductivity, hardness and fracture toughness have been studied. Cutting tool inserts fabricated from these composites were subjected to dry turning operations on 0.45% C steel bars under wide range of cutting parameters. Performance evaluation of the silver toughened alumina inserts was done on the basis of progressive flank wear, and was supplemented by cutting force, surface roughness and chip morphology analyses. Within the investigated range, abrasion and plastic deformation are considered to be the active wear mechanisms for the developed inserts. A comparative examination of the developed composites with those of commercial zirconia toughened alumina (ZTA) inserts indicates that the developed composites have high potential for commercial exploitation. © 2006 Published by Elsevier B.V. Keywords: Wear; Alumina–silver composite; Ceramic tool; Machining
1. Introduction The origin and magnitude of various types of wear are the dominant factors, which govern the machining performance of cutting tools. As a consequence, assessment of the performance of cutting tools are based on their wear behavior [1–8], and the primary assessment is often supplemented by characterization of chip morphology [9,10], cutting force [9,11–13] and roughness of the generated surface [1,3,13,14]. All these factors for assessing the machining performance of a cutting tool are mutually inter-dependent. But the description of machining performance through progressive tool wear is considered to provide the life of a tool unlike the descriptions given by other parameters. For example, 300 m of average flank wear is specified as the maximum permissible limit for single point turning ∗
Corresponding author. Tel.: +91 3222 283278; fax: +91 3222 282280/255303. E-mail addresses: [email protected] (A.K. Dutta), [email protected] (A.B. Chattopadhyaya), [email protected] (K.K. Ray). 0043-1648/$ – see front matter © 2006 Published by Elsevier B.V. doi:10.1016/j.wear.2006.01.038
inserts [15]. Hence studies related to tool wear are essential in the development of any advanced cutting tool material. Alumina-based materials are abundantly used as ceramic cutting tools because of their inherent properties of high hot hardness, abrasion resistance and chemical stability, but the base material suffers from the limitation of low fracture toughness and low thermal shock resistance capability. In order to overcome the limitation of low fracture toughness, advanced alumina ceramics have been developed with the addition of zirconia [5,6,8,9,16], titanium carbide [1,7,17] or silicon carbide [8,18,19]. Toughening of alumina can also be done with the addition of metallic particles [20–31], but the potential of such metal toughened alumina composites has not been carefully explored. Some recent reports [32–35] related to the characteristics of silver toughened alumina indicate that this material can be used for cutting tool applications. Organized studies on tool wear supplemented by characterization of cutting forces, chip morphology and generated surface roughness under varied cutting parameters are considered necessary for assessing the potential of this material for cutting tool applications. These aspects have been examined in this report.
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2. Experimental procedure The alumina–silver composites were prepared from powder mixtures of commercial ␣-alumina (Reynolds, USA, purity 99.8%) and varying amounts of silver oxide (Emark India Ltd., purity 98%). The details of the processing route have been reported earlier [34,35]. In brief, a powder blend of ␣-Al2 O3 (doped with 0.5 wt.% MgO) and Ag2 O was milled for 8 h in an alcohol medium and the blended powder was crushed with the addition of a few drops of polyvinyl alcohol (PVA) solution as binder, and was sieved. The powder mixture was then subjected to uniaxial compaction to produce square shaped green pellets and these were fired at 500 ◦ C for 5 h to burn off the PVA binder. The fired pellets were then presintered at 1100 ◦ C for 2 h followed by final sintering at 1550 ◦ C for 2 h under atmospheric conditions. The heating and cooling rates for the final sintering were kept at 5 ◦ C/min. An identical process route was followed to prepare a series of specimens for three types of composites by altering Ag2 O content as 5, 10 and 15 wt.% and a monolithic alumina doped with MgO. Unlike the composites, the monolithic alumina specimens were sintered at 1600 ◦ C for 2 h. The bulk density, open porosity and the relative density of the different types of pellets were determined using the boiling water method [36] following Archimedes’ principle. At least three samples of each type of material were evaluated for estimation of the average values of porosity and density. Thermal conductivity was measured following a steady state heat flow method [37,38]. In this method, one end of the specimen is attached to a heat sink, while a heater is attached to the other end. The flow of heat (Q) causes two (copper-constantan) thermocouples attached to the specimen to indicate a difference of temperature (T). If the distance between the two thermocouples is L and A is the cross sectional area of the specimen, the thermal conductivity (k) is expressed as k=
QL AT
(1)
The experiments were conducted at room temperature and at a vacuum better than 10−6 Torr. The specimens for metallographic studies were first ground using a 325 grit diamond wheel to make the opposite faces smooth and flat. These were then subjected to standard metallographic polishing by first using different grades of silicon carbide papers and then using different grades (3, 1 and 0.25 m) of diamond pastes. Phase identification in these samples was made by X-ray diffraction analysis using a X-ray diffractometer (PW1729, Philips). The microstructures of the samples were first examined using a light microscope (Versamet-2, Union, Japan) and then using a scanning electron microscope (Jeol, JSM 5800, Japan). Representative photographs of the composites were taken during these examinations. The volume fraction of silver particles in the alumina–silver composites was ascertained by standard point counting technique [39]. The grain size of the alumina matrix was determined by the linear intercept method [40] from SEM photographs of the fractured surfaces. The hardness values of the prepared composites were measured using a Knoop indentor with the help of a microhardness
tester (LECO, DM400, Japan). For each measurement at least 10 readings were taken for estimating the average value. Fracture toughness values of the developed materials were determined by indentation method using the formula suggested by Lankford [41]. The indentation cracks were generated using a Vickers indenter at an indentation load of 98 N. Cutting tool inserts, conforming to ISO specification SNGN 120408, were prepared from the sintered pellets by following the same grinding procedure used for making specimens for microstructural studies and mechanical tests. Nose radius of 0.8 mm was carefully provided on each corner of the inserts using a diamond file of 20–40 m grit size with intermittent examination of the nose through a stereomicroscope (SZ-PT, Olympus, Japan) fitted with a graduated eyepiece. Land width of 0.1 mm at an angle 20◦ was provided on each cutting edge by holding the specimen in a fixture and polishing on 1000-grit size SiC paper. After beveling, the sharp edges were rounded off (to a radius of 20–30 m) by light honing with diamond paste of 3 m size on a lapping machine (Ecomet 3, Buehler, USA). The machining performance of these inserts was examined by plain dry turning of C45 steel (C-0.47%, 200 BHN) in an 11 kW rigid centre lathe (NH22, HMT, India). Turning tests were carried out for a wide range of cutting velocity (150–400 m/min), feed (0.12–0.24 mm/rev.) and depth of cut (1.5 and 2.0 mm). During turning, the tangential (Pz ) and the axial (Px ) components of cutting force were measured using a piezoelectric turning dynamometer (Type: 9257B, Kistler, Switzerland). Tool inserts were withdrawn at regular intervals from the turning operation and the magnitude of the average flank wear at each interruption was measured by an optical microscope. The pattern and the degree of degradation of the turning inserts, used for machining, were examined using a scanning electron microscope. The chips were collected during turning operation under different combinations of cutting parameters for subsequent examination. The roughness of the machined surfaces was measured along the length of the machined workpiece with the help of a Talysurf (Taylor-Hobson, England). 3. Results and discussion The major emphasis in this report is directed to understand the wear behavior of alumina–silver composites during machining. Examination of the nature of the composites and characterization of their properties are necessary supplements. These aspects are presented in the different subsections. 3.1. Nature of the composites Prepared monolithic alumina and different alumina–silver composites are henceforth referred to as Ag-0, Ag-5, Ag-10 and Ag-15 in accordance with the wt.% of Ag2 O in the initial powder mix for convenience of subsequent discussion. The selected weight percentages of Ag2 O powder are expected theoretically to yield nearly 2–6 vol.% of Ag in the fabricated composites. These selections are based on the available information [23,25] that fracture toughness and flexural strength of Al2 O3 –Ag composites do not improve considerably beyond 5–6 vol.% of silver
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Fig. 2. A typical rim and core macrostructure of alumina–silver composite pellet [34].
Fig. 1. X-ray diffraction spectra on alumina–silver pellet at different depths from the top surface (phases: ␣-alumina, no coding; silver, Ag).
content, whereas their hardness monotonically decreases with increase in silver content. Alumina powder was doped with MgO to prevent abnormal grain growth of the matrix [42]. The gray colored Al2 O3 –Ag2 O green compacts turned white on sintering. During sintering Ag2 O is decomposed to metallic Ag leading to the formation of Al2 O3 –Ag system. The phenomena of colour change and the details of the formation of metallic silver have been explained in an earlier report [34]. At temperatures above 1000 ◦ C, Ag starts evaporating from the alumina–silver specimen surfaces. The alumina particles on the surfaces while getting sintered make a barrier to prevent further escape of Ag by evaporation. This phenomenon leads to a silver depleted surface layer and a silver-rich central region in the sintered pellets. Identification of phases at surface layers of different depth by X-ray diffraction analyses illustrated (Fig. 1) the existence of a graded microstructure in the fabricated composites; the amount of silver is observed to increase from the surface towards the central region of a specimen of the alumina–silver composites. The almost silver-free white surface layer is called here as the rim and the silver-rich gray central region is called as the core of the fabricated pellets. A typical rim-core type com-
Fig. 3. A typical fractograph of alumina–silver composite used for grain size measurement.
posite is illustrated in Fig. 2. The rim thicknesses of the different composites were estimated and are shown in Table 1. The apparent porosity and relative density of the fabricated materials are shown in Table 1. The apparent porosity of Ag-10 was found to be the lowest amongst the developed composites. A typical fractograph of an alumina–silver composite, shown in Fig. 3, illustrates the grain size of the alumina matrix. The average grain size of the matrix for all the composites can be designated as 1.4 ± 0.2 m. A comparison of the estimated matrix grain size with that in some earlier reports [23,25] infers the obtained grain size to be finer. Employment of lower sintering temperature, lower initial size of alumina particles and MgO additive are considered to have led to this result.
Table 1 Microstructural characteristic of alumina–silver composites Specimen code
Initial Ag2 O content (wt.%)
Final Ag content (vol.%)
Rim thickness (m)
Apparent porosity (%)
Relative density (%)
Ag-0 Ag-5 Ag-10 Ag-15
0 5.0 10.0 15.0
0.0 1.6 3.38 5.01
0 472 348 295
1.563 1.695 0.855 1.724
98.06 96.19 97.96 94.20
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than that of the conventional Hk values. This observation is in agreement with similar results in soda–lime–silica glass [43] and silicon carbide composites [44]. The fracture toughness values of the developed materials have been determined by indentation technique and the estimated toughness values are given in Table 2. The results in Table 2 indicate that higher amount of silver in the developed composites increases their fracture toughness but decreases their hardness. The increase in fracture toughness (based on H0 values) from Ag-0 to Ag-15 samples was observed to be nearly 160%, but the associated degradation of hardness (H0 ) for similar materials is only about 35%. The higher fracture toughness of silver containing composites has been attributed to the mechanisms of ductile particle stretching and crack deflection.
Fig. 4. Typical distribution of silver particles in an alumina–silver composite.
Microscopic examinations indicated uniform distribution of silver particles in the alumina matrix (Fig. 4). The measured volume fractions of silver in the alumina–silver composites were approximately 10–15% lower than the theoretical estimates. The lower measured values compared to the theoretical estimates of silver in the fabricated composites are attributed to the loss of silver by evaporation during sintering. 3.2. Physical and mechanical properties The estimated values of thermal conductivity of Ag-0, Ag-10 and a commercially available zirconia toughened alumina (ZTA) samples are given in Table 2. These estimates assist to infer that silver addition in alumina improves its thermal conductivity. However, such an inference should not be considered generalized because minor differences in porosity in the materials may alter thermal conductivity in an uncertain way. Hardness values (Hk ) of the developed materials have been determined using a Knoop indentor at different loads and these are found to exhibit indentation size effect [43]. In order to compare the hardness values of the developed materials, the true hardness (H0 ) as suggested by Ray and Dutta [43] and Li et al. [44] were estimated for the developed materials using the following expression [43]: 14229 1/2 1/2 d= (2) P − de Ho where d is the measured diagonal, P the indentation load and de is the amount of relaxation in indentation diagonal. The procedure for estimating H0 has been described elsewhere [34]. The true hardness values for the developed composites are given in Table 2. The results indicate that H0 is generally lower
3.3. Machining tests Ceramic tools undergo crater wear mainly by abrasion, plastic deformation and pull out of the deformed grains [45–50]. Inter-diffusion of chemical species between tool and work piece and the phenomenon of spinnel formation may also contribute to crater formation depending upon the material characteristics of the tool-work pair in intimate contact at the high temperature chip–tool interface [47–49]. Therefore, physico-chemical interaction and friction play vital role in crater wear. Higher conductivity also increases the resistance to crater wear in alumina composite tools [51]. Since silver induces higher thermal conductivity and makes the chip–tool interaction favourable in the alumina–silver ceramic inserts, the white rim of silverfree alumina was removed from the rake surfaces. On the other hand, alumina being more abrasion resistant compared to alumina–silver composite due to the its higher hardness; a thin white band of alumina was deliberately retained at the flanks where abrasion is the dominant wear mechanism. The present machining tests have been done under dry condition to achieve similitude with industrial practice where cutting fluid is generally not used for machining steels by ceramic tools. Amongst the cutting force components, the tangential component (Pz ) is the most significant because it is largest in magnitude and governs the power consumption in cutting. Next important is the axial component (Px ) because it gets considerably influenced by the nature and extent of interaction of the chip–tool interface. The third component i.e. transverse force (Py ) is much less significant because its magnitude is considerably low for wide cutting edge angle. As an example, for the
Table 2 Properties of alumina–silver composites Specimen Code
True hardness, H0 (GPa)
Knoop hardness, Hk (GPa)
IFT (H0 )* (MPa m1/2 )
IFT (Hk )* (MPa m1/2 )
Thermal conductivity, k (kW/mK)
Ag-0 Ag-5 Ag-10 Ag-15 ZTA
16.15 13.3 12.97 10.01 14.8
25.46 19.89 16.49 13.46 18.03
3.38 5.4 7.3 8.8 –
4.4 6.81 8.38 10.36 –
12 – 15.7 – 10.5
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Table 3 Some pertinent details of the machining tests (I) Cutting tools Materials Code
Composition
Make
Ag-0 Ag-5 Ag-10 Ag-15 ZTA
␣-Al2 O3 + 0.5 wt.% MgO ␣-Al2 O3 + 0.5 wt.% MgO + 5 wt.% Ag2 O ␣-Al2 O3 + 0.5 wt.% MgO + 10 wt.% Ag2 O ␣-Al2 O3 + 0.5 wt.% MgO + 15 wt.% Ag2 O ␣-Al2 O3 + ZrO2
Developed Developed Developed Developed CX3, NGK Spark Plug Co. (NTK), Japan
Tool size and shape: SNGN 120408 Tool geometry: −6◦ , −6◦ , 6◦ , 6◦ , 15◦ , 75◦ , 0.8 mm (II) Work specimen Material Size Hardness
C45 steel φ210 mm × 900 mm length 200 BHN
(III) Machine tool
11 kW rigid center lathe, NH22, HMT, India
(IV) Process parameter Cutting velocity (Vc ) Feed (S0 ) Depth of cut (t)
150, 200, 250, 300,350, and 400 m/min 0.12, 0.16, 0.2, and 0.24 mm/rev. 1.5 and 2.0 mm
(V) Environment
Dry
geometry of the tools (as given in Table 3) under investigation, mechanics of turning yields Py ≈ 0.14 Pz . Tool life, which is one of the most significant indices of cutting tool performance, is usually evaluated based on the average width (VB ) of the flank wear at the main cutting edge [2,5]. When VB (as depicted in Fig. 5) reaches a value of 0.3 mm, the life of a tool particularly in turning [15] is said to be exhausted. 3.3.1. On machining chips The type of chips produced during machining operations has significant effect on both the generated surface finish [52,53] and the rate of tool wear [54]. Thus, analyses of chip morphology
Fig. 5. Wear features of cutting tool insert.
Fig. 6. Chips generated by the inserts while machining steel for different feeds.
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[10,55,56] are important for assessing any machining performance test. Fig. 6 represents a typical set of different morphology of chips produced by the developed inserts as well as by commercial ZTA insert during machining of C45 steel bar under different feeds (S0 ) at cutting velocity (Vc ) of 150 m/min and depth of cut (t) of 2 mm. This figure indicates that all the fabricated alumina–silver inserts produce chips similar to those produced by the commercially available ZTA insert, even at cutting velocity as low as 150 m/min. The morphology of the obtained chips is illustrated for the lowest investigated Vc ; because at low cutting speed the machining process is generally unstable, particularly at the beginning of the cut. This is due to higher friction, irregular brake-in wear of fresh cutting edges, chances of built-up edge formation and fluctuation in the cutting forces. All these factors lead to the generation of chips representative for the worst machining condition. Interestingly, the machining performance of the developed alumina–silver inserts with respect to chip morphology, within the investigated range of cutting velocities, is comparable to that of commercial ZTA inserts. The good machining performance of the developed inserts, in general, can be attributed to the stability of the cutting edges due to their appropriate hardness and improved toughness aided by possible favourable effect of silver on chip–tool interaction. The developed tools produced thin, long and nearly uniform chips at low feeds of 0.12 and 0.16 mm/rev. At higher feeds the chips were observed to be thick and closely curled, which broke into almost half-turn pieces by striking against the tool flank. Only in the case of plain alumina (Ag-0) inserts, the chips gradually became constrained and non-uniform even at lower feed until Vc has been raised to 350 m/min. This phenomenon is attributed to the lack of resistance to micro-chipping and crushing of the fresh edges of this insert. Among the alumina–silver inserts, Ag-10 having silver content of 3.38 vol.% was found to be the most satisfactory with respect to cutting edge stability and chip formation. 3.3.2. On cutting forces Knowledge about the cutting forces in machining assists in designing both the machine tool and the cutting tool, and in achieving optimized cutting conditions. In this investigation the magnitudes of the tangential (Pz ) and the axial (Px ) components of the cutting forces have been examined while turning C-45 steel rod with the developed alumina-based inserts and the commercial ZTA inserts at different cutting conditions (Table 3), keeping the depth of cut as 2.00 mm. Fig. 7 represents typical variations of Pz and Px with cutting velocity at constant feed (S0 ) and depth of cut (t). The magnitudes of Pz and Px are found to decrease with increase in cutting velocity. Cutting velocity affects the cutting force by two opposing mechanisms. On one hand, with increase in cutting velocity the cutting force decreases due to softening of the work material at the shear zones and causes favourable change in the chip–tool interaction. On the other hand, as the cutting velocity increases the tool wear rate also increases, which in turn increases the cutting force. Therefore, as the cutting velocity is increased, the
Fig. 7. Variation of (a) tangential and (b) axial components of force with cutting velocity at constant feed and depth of cut for all inserts.
cutting force increases or decreases depending on the dominant mechanism. Generally, cutting force decreases with increasing cutting velocity until a minimum is reached at a speed characteristic of a given tool-work material combination. Beyond that characteristic speed, the force tends to increase slowly [57]. The magnitudes of Pz and Px are found to increase substantially with the increase in feed for all the inserts. Typical comparative assessment of the cutting force components induced by the Ag-10 and ZTA inserts at the varied cutting velocity and varying feeds are illustrated in Fig. 8. The increase in Pz or Px with increasing S0 is due to higher chip load [11,12]. The tangential forces (Pz ) recorded during machining by different tools indicate marginal differences. Though the size and shape of all the tools were almost identical, possible variation in the geometry of the cutting edges and angles during manufacturing cannot be ruled out. The marginal differences between Pz values generated by different cutting tools are attributed to this uncertainty apart from the variability of tool performance during machining and the possible differences in the chip–tool interactions. An examination on the nature of the variation in Pz and Px , in Fig. 8 reveals that the performances of Ag-10 inserts closely resemble that of ZTA inserts for the investigated range of cutting parameters. The results obtained related to cutting forces in all the experiments also indicate that Ag-10 inserts exhibit the best
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Fig. 8. Variation of (a) tangential and (b) axial force components with cutting velocity at different feeds for Ag-10 and ZTA inserts.
performance amongst the developed tools with respect to the stability of the cutting edges and chip–tool interaction. 3.3.3. On surface roughness Surface finish of a machined job is another important criterion used for the evaluation of machining performance of a cutting tool. Under a given set of process parameters, surface finish depends on the material, geometry of the tool, type of chip produced, work material, and vibration of the machining system [53]. The surface roughness produced in turning operation comprises of two parts: one originates inherently as feed marks depending upon the tool geometry and the magnitude of feed, whereas the other gradually appears due to deterioration of the cutting edges and vibration [53]. Formation of built-up edges (BUE) may also significantly degrade the surface finish. Analysis of surface roughness for assessing machinability performance is normally carried out in finishing conditions. Since the present experiments are carried out at medium cutting conditions, an attempt was made only to examine the possible deterioration of the cutting edges of the tool inserts from surface roughness measurement. An examination of the surface roughness produced by the different tools at identical conditions of cutting and at varied machining times inferred: (a) The r.m.s. values of surface roughness produced by Ag-15 insert was found to be highest. Excessively high surface roughness generated by Ag-15 inserts is due to severe chipping at its nose for its relatively lower hardness. (b) The surface roughness produced by Ag-0 and Ag-5 inserts are more than those produced by ZTA and Ag-10 inserts. This may be attributed to the inferior fracture toughness property of the Ag-0 and Ag-5 inserts. (c) The Ag-10 insert usually produced less surface roughness than that generated by the other tools for all feeds due to
its higher resistance to chipping and wear which can be attributed to favourable combination of its hardness and toughness. It may be mentioned here that the above inferences were made under the circumstances when built-up-edge was not found to occur and the effect of vibration was found to be insignificant. These results, in general, indicate that the stability of the cutting edges for Ag-10 insert is as satisfactory as that of ZTA insert. 3.3.4. On tool wear High performance ceramic tools are so designed that these tools do not fail prematurely by catastrophic brittle fracture or by rapid plastic deformation of the cutting edges. Such tools can be used for reasonably long period and their useful life is governed by gradual and systematic wear [1,2,58]. When wear of a tool reaches a reasonable limit, the tool is considered to have served its purpose and is withdrawn from its further use. The usual pattern of wear of turning tools has been schematically illustrated in Fig. 5. In high speed machining of steel, particularly producing continuous chips, ceramic tools undergo slow crater wear at their rake surface mainly by plastic deformation and pull out of the deformed grains in addition to abrasion. Due to more chemical stability, adhesive and diffusive wear are less significant in ceramic tools unlike that of carbide tools. But spinnel formation with chip material in presence of oxygen may enhance the crater wear. Flank wear is mainly caused by abrasion due to rubbing of the flanks against the work surfaces at relatively lower temperatures. But with the increase in Vc and hence in cutting temperature, the mechanisms that cause crater wear also become operative for the flank wear. The life of turning tools particularly of carbides and ceramic tools is generally assessed on the basis of flank wear; particularly at the principal flank. The flank wear is known to grow systemat-
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Fig. 9. Progressive flank wear with machining time up to 4 min.
ically with time and increases cutting force, cutting temperature and surface roughness. When the average width (VB ) of wear at the principal flank (Fig. 5) reaches a limit, the tool is said to have failed and needs replacement. The limit is usually preset depending upon the functional requirements and objectives. It has become a practice, especially for research and development purposes, that the limiting value of VB is kept as 0.3 mm [15]. All the developed tools as well as the ZTA insert were subjected to initial wear test at high cutting velocity (400 m/min), feed (0.24 mm/rev.) and depth of cut (1.5 mm) for duration of 4 min. While machining steel bar by the different tools, each tool was withdrawn at regular intervals and the nature and extent of their wear and cutting edge conditions were examined under SEM and optical microscope. It was found that the cutting edge conditions and wear rate of the Ag-10 inserts were more or less stable as that of the ZTA insert after their break-in wear stage. The nature of growth of the average flank wear (VB ) measured with the progress of machining up to 4 min at high Vc and S0 is shown in Fig. 9. It is clear from the results depicted in this figure that the ZTA and the Ag-10 inserts attained much less flank wear compared to the other inserts. The Ag-0, Ag-5 and Ag-15 inserts which were found to work quite satisfactorily for the first few seconds of machining with respect to chip formation and cutting forces, were found to exhibit considerable break-in wear within about 1 min as can be seen in Fig. 9. This may be attributed to the fact that initially the cutting edges of these inserts were fresh but with the progress of time, lack of adequate toughness (in Ag-0 and Ag-5 inserts) and hardness (in Ag-15) contributed to rapid break-in wear. However, after the break-in wear stage Ag0, Ag-5 and Ag-15 inserts also showed steady growth in flank wear. From the results and analyses of chip morphology, cutting force, surface roughness and preliminary wear test, Ag-10 inserts emerged out as the best amongst the developed inserts. For confirmation, machining by the ZTA and Ag-10 inserts with fresh cutting edges were repeated under the same Vc , S0 and t, and continued until VB reached the limiting value of 0.3 mm as usual tool life. Both the inserts were again found to work almost in an identical fashion with respect to progressive flank wear (Fig. 10). For both the inserts, roughness of the machined
Fig. 10. Average flank wear for Ag-10 and ZTA inserts while machining C-45 steel up to 12 min.
surface was quite high at the beginning of machining due to break-in wear, which gradually decreased due to smoothening of the cutting tool nose almost in an identical fashion. But after some period of steady machining, the Ag-10 insert seems to have marginally better performance than the ZTA insert with respect to surface roughness which may be attributed to higher fracture toughness and thermal conductivity. The fracture toughness of ZTA with hardness of 18 GPa is known to be 4.0 MPa m1/2 [59]. This is considerably lower than the toughness of 8.38 MPa m1/2 for Ag-10 insert. The thermal conductivity of Ag-10 insert is also higher compared to ZTA as shown in Table 2. However, the tangential and the axial components of the cutting force for the Ag-10 insert were found to be marginally higher than that generated by the ZTA insert (Fig. 11). The physical conditions of the Ag-10 and ZTA inserts after being used for machining for 12 min at high cutting velocity (400 m/min), feed (0.24 mm/rev.) and depth of cut (1.5 mm) are illustrated with a series of photographs in Fig. 12. Both the Ag-10 and ZTA inserts made to almost of similar geometry underwent smooth crater wear associated with grooving wear and chipping of insignificant amount, indicating their stability in machining as shown in Fig. 12(a)–(d). However, the details of the edge chipping procedure have not been examined like the one men-
Fig. 11. Variation of tangential and axial force components with time for Ag-10 and ZTA inserts.
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Fig. 12. Condition of ZTA and Ag-10 inserts after 12 min of machining at Vc = 400 m/min, So = 0.24 mm/rev. and t = 1.5 mm: (a) top view of ZTA insert; (b) top view of Ag-10 insert; (c) 3D view of the cutting edge of ZTA insert; (d) 3D view of the cutting edge of Ag-10 insert; (e) crater surface in ZTA insert; (f) crater surface in Ag-10 insert; (g) flank surface of ZTA insert; (h) flank surface of Ag-10 insert.
tioned by Jawahir et al. [60]. The cutting edge of Ag-10 insert unlike that of the ZTA insert attained some micro-irregularities after machining, possibly due to the mode of manual cutting edge preparation. However, there was little grooving wear at the rake surface and no sign of notching wear at its flank, which indicates that both the tools are sufficiently resistant to chem-
ical wear and macro-chipping. The crater on the rake surface is caused due to abrasion wear of plastically deformed layer at the chip–tool interface. It can be noted from Fig. 12(e) and (f) that compared to the ZTA insert, the Ag-10 insert exhibits lesser degree of plastic deformation. This can be attributed to better chip–tool interaction due to higher thermal conductivity
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of alumina–silver composites and favourable frictional condition. The presence of ridges and groove-marks on the principal flank surface, in Fig. 12(g) and (h), indicates that flank wear has occurred due to abrasion between the tool flank and the work material. Fig. 12(h) depicts relatively deeper wear marks on the flank of the Ag-10 insert. This is attributed to lower hardness of this composite than that of the ZTA. 3.3.5. Some generalization The edge geometry of the laboratory made Al2 O3 –Ag inserts were expectedly of inferior quality in comparison to that of the commercial ZTA inserts. But in spite of this, all the developed composite inserts led to the generation of favourable chip morphology. Ag-10 inserts, despite being relatively less hard, have been found to be comparable with the industrially made ZTA inserts with respect to chip formation, cutting force, surface finish and wear resistance. This fact is well illustrated by the comparative assessment of the results related to progressive flank wear, cutting force components and apparent surface roughness. Addition of silver is thus unambiguously considered to result in enough resistance to chipping and micro-fracturing of the cutting edges of the Ag-10 insert. The improved performance of this alumina–silver composite is attributed to the optimum combination of hardness, fracture toughness and thermal conductivity. Thus, it can be inferred that if alumina–silver turning inserts are manufactured with appropriate silver content and their surfaces and edges are given proper finish as in modern tool manufacturing units, this new type of inserts are expected to provide further improved performance and may even outperform the ZTA type inserts. The enhanced cost due to silver addition is expected to be compensated by the simple and inexpensive route of manufacturing these composites. On an overview, this report demonstrates the development of advanced ceramic cutting tools on the major basis of the study on the progressive flank wear, supplemented by several necessary engineering considerations. 4. Conclusions (i) Addition of silver to alumina leads to potential materials for advanced ceramic cutting tool for machining. (ii) The enhanced toughness and thermal conductivity of the developed composites, due to addition of silver, lead to their improved machining performances. But to achieve the best performance, the amount of silver in the composites needs to be optimized. (iii) The cutting performance of the fabricated Ag-10 insert specifically in terms of progressive flank wear is comparable to that of industrially made ZTA insert. (iv) Abrasion and plastic deformation are the primary operative wear mechanisms in the alumina–silver composites. Acknowledgement The authors gratefully thank M/s NGK Spark Plug Co. (NTK), Japan for supplying the ZTA ceramic inserts used in this investigation.
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Progressive flank wear and machining performance of silver toughened alumina cutting tool inserts A.K. Dutta a , A.B. Chattopadhyaya b , K.K. Ray c,∗ a
Department of Mechanical Engineering, Bengal Engineering and Science University, Shibpur, Howrah-711103, India b Department of Mechanical Engineering, Indian Institute of Technology Kharagpur, Kharagpur-721302, India c Department of Metallurgical and Materials Engineering, Indian Institute of Technology Kharagpur, Kharagpur-721302, India Received 5 March 2005; received in revised form 25 January 2006; accepted 31 January 2006 Available online 3 March 2006
Abstract Alumina-based composites with different amounts of silver as second phase have been fabricated using conventional powder metallurgical route and their microstructure, density, thermal conductivity, hardness and fracture toughness have been studied. Cutting tool inserts fabricated from these composites were subjected to dry turning operations on 0.45% C steel bars under wide range of cutting parameters. Performance evaluation of the silver toughened alumina inserts was done on the basis of progressive flank wear, and was supplemented by cutting force, surface roughness and chip morphology analyses. Within the investigated range, abrasion and plastic deformation are considered to be the active wear mechanisms for the developed inserts. A comparative examination of the developed composites with those of commercial zirconia toughened alumina (ZTA) inserts indicates that the developed composites have high potential for commercial exploitation. © 2006 Published by Elsevier B.V. Keywords: Wear; Alumina–silver composite; Ceramic tool; Machining
1. Introduction The origin and magnitude of various types of wear are the dominant factors, which govern the machining performance of cutting tools. As a consequence, assessment of the performance of cutting tools are based on their wear behavior [1–8], and the primary assessment is often supplemented by characterization of chip morphology [9,10], cutting force [9,11–13] and roughness of the generated surface [1,3,13,14]. All these factors for assessing the machining performance of a cutting tool are mutually inter-dependent. But the description of machining performance through progressive tool wear is considered to provide the life of a tool unlike the descriptions given by other parameters. For example, 300 m of average flank wear is specified as the maximum permissible limit for single point turning ∗
Corresponding author. Tel.: +91 3222 283278; fax: +91 3222 282280/255303. E-mail addresses: [email protected] (A.K. Dutta), [email protected] (A.B. Chattopadhyaya), [email protected] (K.K. Ray). 0043-1648/$ – see front matter © 2006 Published by Elsevier B.V. doi:10.1016/j.wear.2006.01.038
inserts [15]. Hence studies related to tool wear are essential in the development of any advanced cutting tool material. Alumina-based materials are abundantly used as ceramic cutting tools because of their inherent properties of high hot hardness, abrasion resistance and chemical stability, but the base material suffers from the limitation of low fracture toughness and low thermal shock resistance capability. In order to overcome the limitation of low fracture toughness, advanced alumina ceramics have been developed with the addition of zirconia [5,6,8,9,16], titanium carbide [1,7,17] or silicon carbide [8,18,19]. Toughening of alumina can also be done with the addition of metallic particles [20–31], but the potential of such metal toughened alumina composites has not been carefully explored. Some recent reports [32–35] related to the characteristics of silver toughened alumina indicate that this material can be used for cutting tool applications. Organized studies on tool wear supplemented by characterization of cutting forces, chip morphology and generated surface roughness under varied cutting parameters are considered necessary for assessing the potential of this material for cutting tool applications. These aspects have been examined in this report.
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2. Experimental procedure The alumina–silver composites were prepared from powder mixtures of commercial ␣-alumina (Reynolds, USA, purity 99.8%) and varying amounts of silver oxide (Emark India Ltd., purity 98%). The details of the processing route have been reported earlier [34,35]. In brief, a powder blend of ␣-Al2 O3 (doped with 0.5 wt.% MgO) and Ag2 O was milled for 8 h in an alcohol medium and the blended powder was crushed with the addition of a few drops of polyvinyl alcohol (PVA) solution as binder, and was sieved. The powder mixture was then subjected to uniaxial compaction to produce square shaped green pellets and these were fired at 500 ◦ C for 5 h to burn off the PVA binder. The fired pellets were then presintered at 1100 ◦ C for 2 h followed by final sintering at 1550 ◦ C for 2 h under atmospheric conditions. The heating and cooling rates for the final sintering were kept at 5 ◦ C/min. An identical process route was followed to prepare a series of specimens for three types of composites by altering Ag2 O content as 5, 10 and 15 wt.% and a monolithic alumina doped with MgO. Unlike the composites, the monolithic alumina specimens were sintered at 1600 ◦ C for 2 h. The bulk density, open porosity and the relative density of the different types of pellets were determined using the boiling water method [36] following Archimedes’ principle. At least three samples of each type of material were evaluated for estimation of the average values of porosity and density. Thermal conductivity was measured following a steady state heat flow method [37,38]. In this method, one end of the specimen is attached to a heat sink, while a heater is attached to the other end. The flow of heat (Q) causes two (copper-constantan) thermocouples attached to the specimen to indicate a difference of temperature (T). If the distance between the two thermocouples is L and A is the cross sectional area of the specimen, the thermal conductivity (k) is expressed as k=
QL AT
(1)
The experiments were conducted at room temperature and at a vacuum better than 10−6 Torr. The specimens for metallographic studies were first ground using a 325 grit diamond wheel to make the opposite faces smooth and flat. These were then subjected to standard metallographic polishing by first using different grades of silicon carbide papers and then using different grades (3, 1 and 0.25 m) of diamond pastes. Phase identification in these samples was made by X-ray diffraction analysis using a X-ray diffractometer (PW1729, Philips). The microstructures of the samples were first examined using a light microscope (Versamet-2, Union, Japan) and then using a scanning electron microscope (Jeol, JSM 5800, Japan). Representative photographs of the composites were taken during these examinations. The volume fraction of silver particles in the alumina–silver composites was ascertained by standard point counting technique [39]. The grain size of the alumina matrix was determined by the linear intercept method [40] from SEM photographs of the fractured surfaces. The hardness values of the prepared composites were measured using a Knoop indentor with the help of a microhardness
tester (LECO, DM400, Japan). For each measurement at least 10 readings were taken for estimating the average value. Fracture toughness values of the developed materials were determined by indentation method using the formula suggested by Lankford [41]. The indentation cracks were generated using a Vickers indenter at an indentation load of 98 N. Cutting tool inserts, conforming to ISO specification SNGN 120408, were prepared from the sintered pellets by following the same grinding procedure used for making specimens for microstructural studies and mechanical tests. Nose radius of 0.8 mm was carefully provided on each corner of the inserts using a diamond file of 20–40 m grit size with intermittent examination of the nose through a stereomicroscope (SZ-PT, Olympus, Japan) fitted with a graduated eyepiece. Land width of 0.1 mm at an angle 20◦ was provided on each cutting edge by holding the specimen in a fixture and polishing on 1000-grit size SiC paper. After beveling, the sharp edges were rounded off (to a radius of 20–30 m) by light honing with diamond paste of 3 m size on a lapping machine (Ecomet 3, Buehler, USA). The machining performance of these inserts was examined by plain dry turning of C45 steel (C-0.47%, 200 BHN) in an 11 kW rigid centre lathe (NH22, HMT, India). Turning tests were carried out for a wide range of cutting velocity (150–400 m/min), feed (0.12–0.24 mm/rev.) and depth of cut (1.5 and 2.0 mm). During turning, the tangential (Pz ) and the axial (Px ) components of cutting force were measured using a piezoelectric turning dynamometer (Type: 9257B, Kistler, Switzerland). Tool inserts were withdrawn at regular intervals from the turning operation and the magnitude of the average flank wear at each interruption was measured by an optical microscope. The pattern and the degree of degradation of the turning inserts, used for machining, were examined using a scanning electron microscope. The chips were collected during turning operation under different combinations of cutting parameters for subsequent examination. The roughness of the machined surfaces was measured along the length of the machined workpiece with the help of a Talysurf (Taylor-Hobson, England). 3. Results and discussion The major emphasis in this report is directed to understand the wear behavior of alumina–silver composites during machining. Examination of the nature of the composites and characterization of their properties are necessary supplements. These aspects are presented in the different subsections. 3.1. Nature of the composites Prepared monolithic alumina and different alumina–silver composites are henceforth referred to as Ag-0, Ag-5, Ag-10 and Ag-15 in accordance with the wt.% of Ag2 O in the initial powder mix for convenience of subsequent discussion. The selected weight percentages of Ag2 O powder are expected theoretically to yield nearly 2–6 vol.% of Ag in the fabricated composites. These selections are based on the available information [23,25] that fracture toughness and flexural strength of Al2 O3 –Ag composites do not improve considerably beyond 5–6 vol.% of silver
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Fig. 2. A typical rim and core macrostructure of alumina–silver composite pellet [34].
Fig. 1. X-ray diffraction spectra on alumina–silver pellet at different depths from the top surface (phases: ␣-alumina, no coding; silver, Ag).
content, whereas their hardness monotonically decreases with increase in silver content. Alumina powder was doped with MgO to prevent abnormal grain growth of the matrix [42]. The gray colored Al2 O3 –Ag2 O green compacts turned white on sintering. During sintering Ag2 O is decomposed to metallic Ag leading to the formation of Al2 O3 –Ag system. The phenomena of colour change and the details of the formation of metallic silver have been explained in an earlier report [34]. At temperatures above 1000 ◦ C, Ag starts evaporating from the alumina–silver specimen surfaces. The alumina particles on the surfaces while getting sintered make a barrier to prevent further escape of Ag by evaporation. This phenomenon leads to a silver depleted surface layer and a silver-rich central region in the sintered pellets. Identification of phases at surface layers of different depth by X-ray diffraction analyses illustrated (Fig. 1) the existence of a graded microstructure in the fabricated composites; the amount of silver is observed to increase from the surface towards the central region of a specimen of the alumina–silver composites. The almost silver-free white surface layer is called here as the rim and the silver-rich gray central region is called as the core of the fabricated pellets. A typical rim-core type com-
Fig. 3. A typical fractograph of alumina–silver composite used for grain size measurement.
posite is illustrated in Fig. 2. The rim thicknesses of the different composites were estimated and are shown in Table 1. The apparent porosity and relative density of the fabricated materials are shown in Table 1. The apparent porosity of Ag-10 was found to be the lowest amongst the developed composites. A typical fractograph of an alumina–silver composite, shown in Fig. 3, illustrates the grain size of the alumina matrix. The average grain size of the matrix for all the composites can be designated as 1.4 ± 0.2 m. A comparison of the estimated matrix grain size with that in some earlier reports [23,25] infers the obtained grain size to be finer. Employment of lower sintering temperature, lower initial size of alumina particles and MgO additive are considered to have led to this result.
Table 1 Microstructural characteristic of alumina–silver composites Specimen code
Initial Ag2 O content (wt.%)
Final Ag content (vol.%)
Rim thickness (m)
Apparent porosity (%)
Relative density (%)
Ag-0 Ag-5 Ag-10 Ag-15
0 5.0 10.0 15.0
0.0 1.6 3.38 5.01
0 472 348 295
1.563 1.695 0.855 1.724
98.06 96.19 97.96 94.20
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than that of the conventional Hk values. This observation is in agreement with similar results in soda–lime–silica glass [43] and silicon carbide composites [44]. The fracture toughness values of the developed materials have been determined by indentation technique and the estimated toughness values are given in Table 2. The results in Table 2 indicate that higher amount of silver in the developed composites increases their fracture toughness but decreases their hardness. The increase in fracture toughness (based on H0 values) from Ag-0 to Ag-15 samples was observed to be nearly 160%, but the associated degradation of hardness (H0 ) for similar materials is only about 35%. The higher fracture toughness of silver containing composites has been attributed to the mechanisms of ductile particle stretching and crack deflection.
Fig. 4. Typical distribution of silver particles in an alumina–silver composite.
Microscopic examinations indicated uniform distribution of silver particles in the alumina matrix (Fig. 4). The measured volume fractions of silver in the alumina–silver composites were approximately 10–15% lower than the theoretical estimates. The lower measured values compared to the theoretical estimates of silver in the fabricated composites are attributed to the loss of silver by evaporation during sintering. 3.2. Physical and mechanical properties The estimated values of thermal conductivity of Ag-0, Ag-10 and a commercially available zirconia toughened alumina (ZTA) samples are given in Table 2. These estimates assist to infer that silver addition in alumina improves its thermal conductivity. However, such an inference should not be considered generalized because minor differences in porosity in the materials may alter thermal conductivity in an uncertain way. Hardness values (Hk ) of the developed materials have been determined using a Knoop indentor at different loads and these are found to exhibit indentation size effect [43]. In order to compare the hardness values of the developed materials, the true hardness (H0 ) as suggested by Ray and Dutta [43] and Li et al. [44] were estimated for the developed materials using the following expression [43]: 14229 1/2 1/2 d= (2) P − de Ho where d is the measured diagonal, P the indentation load and de is the amount of relaxation in indentation diagonal. The procedure for estimating H0 has been described elsewhere [34]. The true hardness values for the developed composites are given in Table 2. The results indicate that H0 is generally lower
3.3. Machining tests Ceramic tools undergo crater wear mainly by abrasion, plastic deformation and pull out of the deformed grains [45–50]. Inter-diffusion of chemical species between tool and work piece and the phenomenon of spinnel formation may also contribute to crater formation depending upon the material characteristics of the tool-work pair in intimate contact at the high temperature chip–tool interface [47–49]. Therefore, physico-chemical interaction and friction play vital role in crater wear. Higher conductivity also increases the resistance to crater wear in alumina composite tools [51]. Since silver induces higher thermal conductivity and makes the chip–tool interaction favourable in the alumina–silver ceramic inserts, the white rim of silverfree alumina was removed from the rake surfaces. On the other hand, alumina being more abrasion resistant compared to alumina–silver composite due to the its higher hardness; a thin white band of alumina was deliberately retained at the flanks where abrasion is the dominant wear mechanism. The present machining tests have been done under dry condition to achieve similitude with industrial practice where cutting fluid is generally not used for machining steels by ceramic tools. Amongst the cutting force components, the tangential component (Pz ) is the most significant because it is largest in magnitude and governs the power consumption in cutting. Next important is the axial component (Px ) because it gets considerably influenced by the nature and extent of interaction of the chip–tool interface. The third component i.e. transverse force (Py ) is much less significant because its magnitude is considerably low for wide cutting edge angle. As an example, for the
Table 2 Properties of alumina–silver composites Specimen Code
True hardness, H0 (GPa)
Knoop hardness, Hk (GPa)
IFT (H0 )* (MPa m1/2 )
IFT (Hk )* (MPa m1/2 )
Thermal conductivity, k (kW/mK)
Ag-0 Ag-5 Ag-10 Ag-15 ZTA
16.15 13.3 12.97 10.01 14.8
25.46 19.89 16.49 13.46 18.03
3.38 5.4 7.3 8.8 –
4.4 6.81 8.38 10.36 –
12 – 15.7 – 10.5
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Table 3 Some pertinent details of the machining tests (I) Cutting tools Materials Code
Composition
Make
Ag-0 Ag-5 Ag-10 Ag-15 ZTA
␣-Al2 O3 + 0.5 wt.% MgO ␣-Al2 O3 + 0.5 wt.% MgO + 5 wt.% Ag2 O ␣-Al2 O3 + 0.5 wt.% MgO + 10 wt.% Ag2 O ␣-Al2 O3 + 0.5 wt.% MgO + 15 wt.% Ag2 O ␣-Al2 O3 + ZrO2
Developed Developed Developed Developed CX3, NGK Spark Plug Co. (NTK), Japan
Tool size and shape: SNGN 120408 Tool geometry: −6◦ , −6◦ , 6◦ , 6◦ , 15◦ , 75◦ , 0.8 mm (II) Work specimen Material Size Hardness
C45 steel φ210 mm × 900 mm length 200 BHN
(III) Machine tool
11 kW rigid center lathe, NH22, HMT, India
(IV) Process parameter Cutting velocity (Vc ) Feed (S0 ) Depth of cut (t)
150, 200, 250, 300,350, and 400 m/min 0.12, 0.16, 0.2, and 0.24 mm/rev. 1.5 and 2.0 mm
(V) Environment
Dry
geometry of the tools (as given in Table 3) under investigation, mechanics of turning yields Py ≈ 0.14 Pz . Tool life, which is one of the most significant indices of cutting tool performance, is usually evaluated based on the average width (VB ) of the flank wear at the main cutting edge [2,5]. When VB (as depicted in Fig. 5) reaches a value of 0.3 mm, the life of a tool particularly in turning [15] is said to be exhausted. 3.3.1. On machining chips The type of chips produced during machining operations has significant effect on both the generated surface finish [52,53] and the rate of tool wear [54]. Thus, analyses of chip morphology
Fig. 5. Wear features of cutting tool insert.
Fig. 6. Chips generated by the inserts while machining steel for different feeds.
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[10,55,56] are important for assessing any machining performance test. Fig. 6 represents a typical set of different morphology of chips produced by the developed inserts as well as by commercial ZTA insert during machining of C45 steel bar under different feeds (S0 ) at cutting velocity (Vc ) of 150 m/min and depth of cut (t) of 2 mm. This figure indicates that all the fabricated alumina–silver inserts produce chips similar to those produced by the commercially available ZTA insert, even at cutting velocity as low as 150 m/min. The morphology of the obtained chips is illustrated for the lowest investigated Vc ; because at low cutting speed the machining process is generally unstable, particularly at the beginning of the cut. This is due to higher friction, irregular brake-in wear of fresh cutting edges, chances of built-up edge formation and fluctuation in the cutting forces. All these factors lead to the generation of chips representative for the worst machining condition. Interestingly, the machining performance of the developed alumina–silver inserts with respect to chip morphology, within the investigated range of cutting velocities, is comparable to that of commercial ZTA inserts. The good machining performance of the developed inserts, in general, can be attributed to the stability of the cutting edges due to their appropriate hardness and improved toughness aided by possible favourable effect of silver on chip–tool interaction. The developed tools produced thin, long and nearly uniform chips at low feeds of 0.12 and 0.16 mm/rev. At higher feeds the chips were observed to be thick and closely curled, which broke into almost half-turn pieces by striking against the tool flank. Only in the case of plain alumina (Ag-0) inserts, the chips gradually became constrained and non-uniform even at lower feed until Vc has been raised to 350 m/min. This phenomenon is attributed to the lack of resistance to micro-chipping and crushing of the fresh edges of this insert. Among the alumina–silver inserts, Ag-10 having silver content of 3.38 vol.% was found to be the most satisfactory with respect to cutting edge stability and chip formation. 3.3.2. On cutting forces Knowledge about the cutting forces in machining assists in designing both the machine tool and the cutting tool, and in achieving optimized cutting conditions. In this investigation the magnitudes of the tangential (Pz ) and the axial (Px ) components of the cutting forces have been examined while turning C-45 steel rod with the developed alumina-based inserts and the commercial ZTA inserts at different cutting conditions (Table 3), keeping the depth of cut as 2.00 mm. Fig. 7 represents typical variations of Pz and Px with cutting velocity at constant feed (S0 ) and depth of cut (t). The magnitudes of Pz and Px are found to decrease with increase in cutting velocity. Cutting velocity affects the cutting force by two opposing mechanisms. On one hand, with increase in cutting velocity the cutting force decreases due to softening of the work material at the shear zones and causes favourable change in the chip–tool interaction. On the other hand, as the cutting velocity increases the tool wear rate also increases, which in turn increases the cutting force. Therefore, as the cutting velocity is increased, the
Fig. 7. Variation of (a) tangential and (b) axial components of force with cutting velocity at constant feed and depth of cut for all inserts.
cutting force increases or decreases depending on the dominant mechanism. Generally, cutting force decreases with increasing cutting velocity until a minimum is reached at a speed characteristic of a given tool-work material combination. Beyond that characteristic speed, the force tends to increase slowly [57]. The magnitudes of Pz and Px are found to increase substantially with the increase in feed for all the inserts. Typical comparative assessment of the cutting force components induced by the Ag-10 and ZTA inserts at the varied cutting velocity and varying feeds are illustrated in Fig. 8. The increase in Pz or Px with increasing S0 is due to higher chip load [11,12]. The tangential forces (Pz ) recorded during machining by different tools indicate marginal differences. Though the size and shape of all the tools were almost identical, possible variation in the geometry of the cutting edges and angles during manufacturing cannot be ruled out. The marginal differences between Pz values generated by different cutting tools are attributed to this uncertainty apart from the variability of tool performance during machining and the possible differences in the chip–tool interactions. An examination on the nature of the variation in Pz and Px , in Fig. 8 reveals that the performances of Ag-10 inserts closely resemble that of ZTA inserts for the investigated range of cutting parameters. The results obtained related to cutting forces in all the experiments also indicate that Ag-10 inserts exhibit the best
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Fig. 8. Variation of (a) tangential and (b) axial force components with cutting velocity at different feeds for Ag-10 and ZTA inserts.
performance amongst the developed tools with respect to the stability of the cutting edges and chip–tool interaction. 3.3.3. On surface roughness Surface finish of a machined job is another important criterion used for the evaluation of machining performance of a cutting tool. Under a given set of process parameters, surface finish depends on the material, geometry of the tool, type of chip produced, work material, and vibration of the machining system [53]. The surface roughness produced in turning operation comprises of two parts: one originates inherently as feed marks depending upon the tool geometry and the magnitude of feed, whereas the other gradually appears due to deterioration of the cutting edges and vibration [53]. Formation of built-up edges (BUE) may also significantly degrade the surface finish. Analysis of surface roughness for assessing machinability performance is normally carried out in finishing conditions. Since the present experiments are carried out at medium cutting conditions, an attempt was made only to examine the possible deterioration of the cutting edges of the tool inserts from surface roughness measurement. An examination of the surface roughness produced by the different tools at identical conditions of cutting and at varied machining times inferred: (a) The r.m.s. values of surface roughness produced by Ag-15 insert was found to be highest. Excessively high surface roughness generated by Ag-15 inserts is due to severe chipping at its nose for its relatively lower hardness. (b) The surface roughness produced by Ag-0 and Ag-5 inserts are more than those produced by ZTA and Ag-10 inserts. This may be attributed to the inferior fracture toughness property of the Ag-0 and Ag-5 inserts. (c) The Ag-10 insert usually produced less surface roughness than that generated by the other tools for all feeds due to
its higher resistance to chipping and wear which can be attributed to favourable combination of its hardness and toughness. It may be mentioned here that the above inferences were made under the circumstances when built-up-edge was not found to occur and the effect of vibration was found to be insignificant. These results, in general, indicate that the stability of the cutting edges for Ag-10 insert is as satisfactory as that of ZTA insert. 3.3.4. On tool wear High performance ceramic tools are so designed that these tools do not fail prematurely by catastrophic brittle fracture or by rapid plastic deformation of the cutting edges. Such tools can be used for reasonably long period and their useful life is governed by gradual and systematic wear [1,2,58]. When wear of a tool reaches a reasonable limit, the tool is considered to have served its purpose and is withdrawn from its further use. The usual pattern of wear of turning tools has been schematically illustrated in Fig. 5. In high speed machining of steel, particularly producing continuous chips, ceramic tools undergo slow crater wear at their rake surface mainly by plastic deformation and pull out of the deformed grains in addition to abrasion. Due to more chemical stability, adhesive and diffusive wear are less significant in ceramic tools unlike that of carbide tools. But spinnel formation with chip material in presence of oxygen may enhance the crater wear. Flank wear is mainly caused by abrasion due to rubbing of the flanks against the work surfaces at relatively lower temperatures. But with the increase in Vc and hence in cutting temperature, the mechanisms that cause crater wear also become operative for the flank wear. The life of turning tools particularly of carbides and ceramic tools is generally assessed on the basis of flank wear; particularly at the principal flank. The flank wear is known to grow systemat-
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Fig. 9. Progressive flank wear with machining time up to 4 min.
ically with time and increases cutting force, cutting temperature and surface roughness. When the average width (VB ) of wear at the principal flank (Fig. 5) reaches a limit, the tool is said to have failed and needs replacement. The limit is usually preset depending upon the functional requirements and objectives. It has become a practice, especially for research and development purposes, that the limiting value of VB is kept as 0.3 mm [15]. All the developed tools as well as the ZTA insert were subjected to initial wear test at high cutting velocity (400 m/min), feed (0.24 mm/rev.) and depth of cut (1.5 mm) for duration of 4 min. While machining steel bar by the different tools, each tool was withdrawn at regular intervals and the nature and extent of their wear and cutting edge conditions were examined under SEM and optical microscope. It was found that the cutting edge conditions and wear rate of the Ag-10 inserts were more or less stable as that of the ZTA insert after their break-in wear stage. The nature of growth of the average flank wear (VB ) measured with the progress of machining up to 4 min at high Vc and S0 is shown in Fig. 9. It is clear from the results depicted in this figure that the ZTA and the Ag-10 inserts attained much less flank wear compared to the other inserts. The Ag-0, Ag-5 and Ag-15 inserts which were found to work quite satisfactorily for the first few seconds of machining with respect to chip formation and cutting forces, were found to exhibit considerable break-in wear within about 1 min as can be seen in Fig. 9. This may be attributed to the fact that initially the cutting edges of these inserts were fresh but with the progress of time, lack of adequate toughness (in Ag-0 and Ag-5 inserts) and hardness (in Ag-15) contributed to rapid break-in wear. However, after the break-in wear stage Ag0, Ag-5 and Ag-15 inserts also showed steady growth in flank wear. From the results and analyses of chip morphology, cutting force, surface roughness and preliminary wear test, Ag-10 inserts emerged out as the best amongst the developed inserts. For confirmation, machining by the ZTA and Ag-10 inserts with fresh cutting edges were repeated under the same Vc , S0 and t, and continued until VB reached the limiting value of 0.3 mm as usual tool life. Both the inserts were again found to work almost in an identical fashion with respect to progressive flank wear (Fig. 10). For both the inserts, roughness of the machined
Fig. 10. Average flank wear for Ag-10 and ZTA inserts while machining C-45 steel up to 12 min.
surface was quite high at the beginning of machining due to break-in wear, which gradually decreased due to smoothening of the cutting tool nose almost in an identical fashion. But after some period of steady machining, the Ag-10 insert seems to have marginally better performance than the ZTA insert with respect to surface roughness which may be attributed to higher fracture toughness and thermal conductivity. The fracture toughness of ZTA with hardness of 18 GPa is known to be 4.0 MPa m1/2 [59]. This is considerably lower than the toughness of 8.38 MPa m1/2 for Ag-10 insert. The thermal conductivity of Ag-10 insert is also higher compared to ZTA as shown in Table 2. However, the tangential and the axial components of the cutting force for the Ag-10 insert were found to be marginally higher than that generated by the ZTA insert (Fig. 11). The physical conditions of the Ag-10 and ZTA inserts after being used for machining for 12 min at high cutting velocity (400 m/min), feed (0.24 mm/rev.) and depth of cut (1.5 mm) are illustrated with a series of photographs in Fig. 12. Both the Ag-10 and ZTA inserts made to almost of similar geometry underwent smooth crater wear associated with grooving wear and chipping of insignificant amount, indicating their stability in machining as shown in Fig. 12(a)–(d). However, the details of the edge chipping procedure have not been examined like the one men-
Fig. 11. Variation of tangential and axial force components with time for Ag-10 and ZTA inserts.
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Fig. 12. Condition of ZTA and Ag-10 inserts after 12 min of machining at Vc = 400 m/min, So = 0.24 mm/rev. and t = 1.5 mm: (a) top view of ZTA insert; (b) top view of Ag-10 insert; (c) 3D view of the cutting edge of ZTA insert; (d) 3D view of the cutting edge of Ag-10 insert; (e) crater surface in ZTA insert; (f) crater surface in Ag-10 insert; (g) flank surface of ZTA insert; (h) flank surface of Ag-10 insert.
tioned by Jawahir et al. [60]. The cutting edge of Ag-10 insert unlike that of the ZTA insert attained some micro-irregularities after machining, possibly due to the mode of manual cutting edge preparation. However, there was little grooving wear at the rake surface and no sign of notching wear at its flank, which indicates that both the tools are sufficiently resistant to chem-
ical wear and macro-chipping. The crater on the rake surface is caused due to abrasion wear of plastically deformed layer at the chip–tool interface. It can be noted from Fig. 12(e) and (f) that compared to the ZTA insert, the Ag-10 insert exhibits lesser degree of plastic deformation. This can be attributed to better chip–tool interaction due to higher thermal conductivity
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of alumina–silver composites and favourable frictional condition. The presence of ridges and groove-marks on the principal flank surface, in Fig. 12(g) and (h), indicates that flank wear has occurred due to abrasion between the tool flank and the work material. Fig. 12(h) depicts relatively deeper wear marks on the flank of the Ag-10 insert. This is attributed to lower hardness of this composite than that of the ZTA. 3.3.5. Some generalization The edge geometry of the laboratory made Al2 O3 –Ag inserts were expectedly of inferior quality in comparison to that of the commercial ZTA inserts. But in spite of this, all the developed composite inserts led to the generation of favourable chip morphology. Ag-10 inserts, despite being relatively less hard, have been found to be comparable with the industrially made ZTA inserts with respect to chip formation, cutting force, surface finish and wear resistance. This fact is well illustrated by the comparative assessment of the results related to progressive flank wear, cutting force components and apparent surface roughness. Addition of silver is thus unambiguously considered to result in enough resistance to chipping and micro-fracturing of the cutting edges of the Ag-10 insert. The improved performance of this alumina–silver composite is attributed to the optimum combination of hardness, fracture toughness and thermal conductivity. Thus, it can be inferred that if alumina–silver turning inserts are manufactured with appropriate silver content and their surfaces and edges are given proper finish as in modern tool manufacturing units, this new type of inserts are expected to provide further improved performance and may even outperform the ZTA type inserts. The enhanced cost due to silver addition is expected to be compensated by the simple and inexpensive route of manufacturing these composites. On an overview, this report demonstrates the development of advanced ceramic cutting tools on the major basis of the study on the progressive flank wear, supplemented by several necessary engineering considerations. 4. Conclusions (i) Addition of silver to alumina leads to potential materials for advanced ceramic cutting tool for machining. (ii) The enhanced toughness and thermal conductivity of the developed composites, due to addition of silver, lead to their improved machining performances. But to achieve the best performance, the amount of silver in the composites needs to be optimized. (iii) The cutting performance of the fabricated Ag-10 insert specifically in terms of progressive flank wear is comparable to that of industrially made ZTA insert. (iv) Abrasion and plastic deformation are the primary operative wear mechanisms in the alumina–silver composites. Acknowledgement The authors gratefully thank M/s NGK Spark Plug Co. (NTK), Japan for supplying the ZTA ceramic inserts used in this investigation.
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