Development of Silicon nitride–Aluminum nitride-Yttrium ...



Development of Si3N4-AlN-Y2O3 composite cutting tool for the machining of compacted graphite iron

J.V.C. Souza1, C. Santos,2, C. A. Kelly2, M. V.Ribeiro3, K.Strecker4 , O.M.M. Silva5

1INPE - Av. dos Astronautas,1.758, S. J. Campo s - SP, CEP. 12245-970, Brazil

2FAENQUIL-DEMAR – Polo Urbo Industrial, Gleba AI-6, s/n, Lorena - SP, CEP. 12600-000, Brazil

3FEG-UNESP – Av.Ariberto Ferreira da Cunha, 333, Guaratinguetá - SP, CEP. 12516-410, Brazil

4UFSJ - Praça Frei Orlando 170, São João del Rei - MG, CEP. 36307-352, Brazil

5CTA-IAE/AMR - Pça. Marechal do Ar Eduardo Gomes, 50, S. J. Campo s - SP, CEP. 12228-904, Brazil

vitor@las.inpe.br / candidojvc64@.br

Abstract

In this work, a new composition of the Si3N4 ceramic cutting tools was developed, characterized and applied on dry machining process of compacted graphite iron. Compacted graphite iron (CGI) is athe promising material under study for the upcoming upcoming new generation of high-power diesel engines due to its high strength. It allows increased , including blocks, cylinder-pressures and therefore reduces fuel consumption and a higher power output is possible so far, CGI is not used to a larger extent because of its very difficult machinability when compared to other automotive industry materials. liners, and cylinder heads. Its unique graphite structure yields an increased due to its increased strength compared to grey cast iron (CI), . It allowings an increase in the cylinder-pressures and therefore a better fuel economy and a higher power output are possible. Until now, CGI is not applied in the automotive industry due to its uneconomic machinability, because of a significant decrease in tool life when compared to CI machining. In the literature, there are few detailed informations about machining compacted graphite iron with ceramic cutting tools. Only one paper Most of references give some detailed information when using CBN as the cutting tool, but its known that CBN cutting tool present higher cost when compared with others ceramics cutting tools.. The CGI is not used to a larger extent in large scale production up to now is its very difficult machinability as compared to others automotive industry materials. For this reason, a new ceramics cutting tool with composition of Si3N4 ceramic cutting tool that has unique chemical and mechanical properties has been was developed, applied in the dry machining of CGI and compared with recent literature. The new cutting tool has been ,ment. characterized and applied to dry longitudinal turning of compacted graphite iron. It was also made a comparison of the benefit-cost ratio between the CBN and the Si3N4 ceramic cutting tools. The tool obtained ing methodology to using ed The a powder mixture based on a Si3N4-AlN-Y2O3 powder mixture, aimed to produce aternary system, was developed aiming the obtaining of SiAlON ceramic, a solid solution of AlN and Y2O3 in the Si3N4 structure. Inserts were sintered at 1900 0°C, and characterized by X-Ray Diffraction and Scanning Electron Microscopy. This cutting tool The Si3N4-AlN-Y2O3 ceramic composite cutting tool insert presented a hardness of 21 GPa and a fracture toughness of 5.6 MPam1/2. Phase analysis indicated Furthermore, XRD patterns indicated the presence of α-SiAlON as the predominant crystalline phase and elongated grains were observed in the microstructureSEM analysis. The Ddry machining tests of CGI samples were performed using the newly developed cutting tool at cutting speeds of turningmachining studies have been conducted using the newly developed cutting tool insert on compacted graphite iron (CGI) aat cutting speeds of Vc1=200, Vc2=350 and Vc3=500  m/min, with a feed rate of 0.20  mm/rev and a cut depth of cut of 0.5  mm. In general, crater and flank wear and fracture were observed in the machining tests which appear depending on the mechanical and thermal conditions generated on the wear zones. Maximum length cut was observed at cutting speeds of 200m/min. This work presented innovative results in dry machining process on machining CGI with α-SiAlON. These results are unique, due to few information there was and fill out at information lack in literature.

Keywords: Ceramic cutting tools; Flank wear; Dry machining, Finish surface, Compacted graphite iron.

* Corresponding author. Tel.: +55 (12) 3945-6679, 39456989, FAX: +55 (12) 3945-6717

E-mail address: vitor@las.inpe.br / candidojvc64@.br (J.V.C. Souza).

1. Introduction

The machining is a fabrication technique where different materials are removed from a part using a tool with a small hard tip. Machining is a major manufacturing process and plays a key role in the creation of wealth. Machining operations consumes a large amount of money annually worldwide. Over US$ 100 billion is spent annually worldwide on metal part finishing processes such as turning, milling, boring and other cutting operations [1]. In order to quickly fabricate parts, high-speed cutting is desired. These higher speeds, however, lead to a faster degradation of the tool tip, which requires more frequent replacements of the cutting edge [2]. Ceramic cutting tools [1-3] play important roles in the development of advanced manufacture technology, and the increased application of ceramic cutting tools in proportion is inevitable. Types and brands of commonly used ceramic tools were summarized. Compared with traditional high speed steel and hard alloys, advanced ceramic cutting tool has the advantage of high hardness, chemical stability etc[3 and 4].

Silicon nitride (Si3N4) based ceramic tools are widely used as cutting tools because of their high hardness, high thermal conductivity and low thermal expansion coefficient that give better thermal shock resistance than other ceramic materials [5]. Under high-speed machining conditions, the elevated temperatures developed in the cutting zone (tool-chip and tool-workpiece contact), enhance chemical and mechanical wear mechanisms [2-54].

Other important ceramic cutting tool is CBN tool that has had greater wear resistance than others ceramic tools due to their high degree of hardness. They have successfully been used in the high speed machining of many materials [6,12, 725 and 829]. Despite their superior tool life, the cost of CBN cutting tools restricts their more widespread use in industry. Application of CBN tools is highly affected by the high cost. In recent paper N. Camuscu and E. Aslan, (2005) showed that although, the CBN tool exhibited the best cutting performance in terms of flank wear and surface finish, the tools should also be compared from an economic point of view. In milling process of AISI D3 tool steel the volume of material removed with CBN tool was more than twice of that removed with others ceramic tools. However, the price of CBN tool was approximately five times more than that of ceramic tool [9]. This process, only one edge of CBN tool was used for machining as CBN tip was brazed onto only one corner of the insert. However, for other tools, it is generally possible to use the other edges as well. The number of edges, which are suitable for machining, depends on the shape and form of the insert.

α−SiAlON (or α’) is a solid solution of α−Si3N4 of composition MxSi12-(m+n)Al(m+n)OnN16-n, where M represents the metallic cation used for the stabilization of this phase and Si and N are partially substituted by Al and O, respectively, in the crystal structure. This ceramic material exhibits higher creep resistance, hardness and oxidation resistance than β−Si3N4 [5-7]. Its improved high-temperature properties are due mainly to the absence or reduced amount of intergranular secondary phases, because of the incorporation of the metallic cation of the liquid phase into the Si3N4 structure forming α−SiAlON. The metal ions used for the stabilization of the α-Si3N4 structure are also responsible for the densification during the sintering, the improvement of the microstructural characteristics and the mechanical properties of these ceramic materials. Yttrium Y+3 is one of the cations used to stabilize the α’-phase at higher temperatures, generating elongated grains and therefore increased fracture toughness.

Despite of advance in cutting tool materials, CGI machining process is very difficult with few reported in literature [[110]5]. The reason for this difficult is not known so far. Few preliminary studies were performed related to CGI machining. However, the results were not conclusive [11][42].. The compacted graphite iron (CGI) is widely discussed as a candidate material for diesel engine motor blocks because of its advantageous combination of mechanical and elastic characteristics [12][1]. However, it is not widely used up today because of the difficult machining process, despite the advances in cutting tool materials [13, 14 and 15]. Few studies on the CGI machining have been reported, however with inconclusive results [16, 17 and 18]. CGI the machiningability of CGI is not a simple taskvery worse (bad). Recent works showed that, depending on the machining parameters, the lifetime of cubic boron nitride (cBN) tools is up to 20 times longer for cast iron (CI) when as compared to CGI [17 and 19][2, 3]. Any Any one report has been found observedon the machining about α-SiAlON cutting tools machining of CGI with a α-SiAlON cutting tool[4], which motivated the present work.. Therefore, we report here on an innovative study on the tool wear on CGI machining. A multimethod approach was used for this purpose, including model experiments to study development, characterization and dry machining application. In all cases, the same experiments were performed twice times in order to identify the differences in the tool wear and to explain the tool lifetimes. However the results data were analyzed, including tool wear rates, surface finish and temperature so that machining parameters which can extend tool life could be identified.

1.2. Properties of Compacted graphite iron

The material properties of cCompacted graphite iron (CGI) offers higherbetter strength and stiffness than gray iron, and improved better castability, machinability, and thermal conductivity than ductile iron, making it an ideal material for components that undergo both mechanical and thermal loading [208]. There is almost always some spheroidal graphite present in CGI, which is referred to as percent nodularity, which is ideally between 0 and 10 %, although up to 20 % is tolerable allowable depending on the application. Compacted graphite iron (CGI) tends to maintain its hold its strength and elongation well properties well up tountil approximately 400 ° °C [139].

Due to its The CGI has a higher tensile strength and higher stiffness, it is difficult to machine. Additionally a , the pearlitic CGI structure is more difficult to machine than ferritic CGI [129]. Therefore, machining under a high torque using stiff machining tools and The properties of CGI necessitate that high torque and stiffness machine tools be used and Dawson [1410] also noteds that 20–30 % higher spindle power is required. Tool life is specially reduced during The greatest losses in tool life occur during turning and cylinder boring [1611]. Dawson et al. [1512] state that during low speed cutting using carbide cutting tooling,, a 50 % reduction in tool life is observed when seen as compared to gray iron machining.

2. Experimental

Commercially available, high -purity α-Si3N4 powder, Y2O3 powder and AlN powders Grade-A powder (H.C. Starck-Germany) were used as starting materialspowders.. A Ppowder batch containing 82.86 wt.% of α-Si3N4, 6.51 wt % Y2O3 and 10.63 wt % AlN was prepared obtained by ball-milling for 4h, in a plastic jar using Si3N4 balls and ethanol as the milling- media. The ball to/ powder weight ratio used was 5:1. After milling the Ppowder mixture was dried at 100°C for 24h and sieved. Green compacts were cold uniaxially pressed in a steel die under under a pressure of 50MPa and further isostatically pressed under a pressure of 300 MPa. The compacts were Pprismatic in shape of approximately green samples (16.4  ×  16.4 × 8  mm3.) were encapsulated in plastic bags and isostatically pressed under pressure of 300MPa for 2 min. The green relative density of the specimen samples was obtained using the relation of the apparent between the measured density (weight/volume) and theoretical density of the powder batch, calculated by the rule of mixtures.

Sintering was carried out in a graphite- resistance furnace (Thermal Technologiesy 1000-4560-FP-20). The specimen compacts were heated up underin vacuum to, at 600 ° °C under a heating rate -10 min. using a heating rate of 5 ° °C/min an isothermal . The lloading time of 10 min in order ow heating rate was employed in order to allow living-out of the the organic binder usedto burn slowly and prevent the damage of the samples. From 600 ° °C to the a final sintering temperature of (1900 0°C-120min), the samples were heated up under a rate of using 15 ° °C/min, injecting as rand maintaining 0,1 MPa nitrogen ate. Aat 1200 0C. At 1900 °C a time of 2h was employed., nitrogen atmosphere (0.1MPa) was used in the sintering and the cooling rate was 15 0C/min.

The relative density of the sintered samples was determined by the immersion method in distilled water, using Archimedes’ principlemethod in distilled water. Phase analysis was done by X-ray diffraction (XRD), of polished-sections using a Phillips PW-1380/80 dX-ray Diffractometter, Cu-Kα radiation (λ = 1.5418Å), and comparing the diffraction patterns with the JCPDS files. X-ray diffraction analysis was conducted on a cross section of the samples.

For microstructural analysis, the specimensamples were ground, polished, and chemically etched with by a molten 1:1 mixture of NaOH and KOH at 500 °0C for 10 min. Prior to the SEM investigation, the samples were coated with gold.

Crack propagation and fracture surfaces of the specimens were observed using scanning electron microscope (LEO 1450VP) at an accelerating voltage of 20 kV. A gold coating was applied on the surface of the etched samples in order to minimize the surface charging under electron beam.

The hardness was determined by Vicker’s indentations for 30 s under an applied load of 98110 kgf N for 30s. . For statistical reasons, 21 indentations were have been made in each per sample. The fracture toughness was has been determined by the measurement of the crack length created by the Vicker’s indentations. The calculation of the fracture toughness, was calculated done by the relation proposed in Equation (1) [1321], valid for Palmqvist type cracks:

KIC = 0.16(E/H)1/2.F.b-3/2 (1)

Where: KIC = fracture toughness [MPa.m1/2];

E = Young modulus of material [GPa];

HV = Vickers hardness [GPa]; b = crack size [(m] and F = applied load [N].

2.1. Machining tests

All experiments were carried out on a computer numerical controled (CNC) lathe (Romi, Mod. Centur 30D) under dry cutting condition. The sintered specimen ceramic cutting tools were cut and ground to make SNGN120408 cutting tools (12.7 x 12.7 x 4.76  mm, 0.08  mm nose radius and 0.2  mm x 20° chamfer). A tool holder of CSRNR 2525 M 12CEA type tool holder (offset shank with 15°° [75°°] side cutting edge angle, 0° insert normal clearance and 25 mm x 25 mm x 150  mm) was used for the cutting experiments. The cutting performance of the ceramic cutting tool was tested by machining compacted graphite iron.

The cutting tests for machining of compacted graphite iron were performed at cutting speeds of Vc1=200, Vc2=350 and Vc3=500  m/min under with a feed rate of 0.20  mm/rev and a cut depth of cut of 0.5  mm. The CGI material compacted graphite iron presented a tensile strength of 500 MPa, a young`s elastic modulus of 140 GPa, a thermal conductivity of 35 W/m-1-K-1 and a hardness of 225  BHN. Since the workpieces were manufactured by the rolling process, in order to eliminate the problems encountered in the first pass, 1.5–2  mm of the skin was removed with a different separate insert prior to the machining before the experiments with SiAlON cutting tool.. This procedure application reduceds the vibrations since it improves the roundness accuracy of the cutting toolworkpieces. This produce reduced vibrations since the roundness accuracy of the workpiece has been improved.. The dimension of workpiece was material was about 60  mm in internal diameter, 120 mm external diameter 120 and 400  mm in length. Tool he wear of the tools was determined by measuring the wear depth ofon the flank face by using a scanning electron microscopy (LEO-1450 VP) at more than four points of the flank face. and tThe average of them was taken as thea nominal flank wear depth. The For workpiece surface roughness was measured with a profilometer by a meter (Mitutoyo, Surftest 402 series 178). A Flank wear of 0.6 mm (ISO 3685) and variation abrupt variation of Ra and Ry has been used as end tool life criterion.

3. Results

3.1 Specimen preparationCompaction and Sintering

The prossed Compacts presented green specimen a relative density of 60 % of the theoretical density of the powder mixture.

This high compaction-degree improves the densification during the sintering process.

It is well established that the sintering additives react with the SiO2, present in Si3N4 surface of the starting powder, forming a liquid phase during sintering. Subsequently, the α-Si3N4 particles dissolve in this liquid phase, causing a local supersaturation and reprecipitation of β-Si3N4 [14 and 15]. Different sintering additives will lead to the formation of different liquid phases with a composition-specific viscosity at 1900 °C. The lower viscosity of the liquid phase, facilitate the faster the diffusion rate of Si4+ and N3- ions and the higher the transformation rate from α- to β-Si3N4. In the process of transformation of Si3N4 from α- to β-structure, Al3+ and O2- ions can enter the β-Si3N4 structure to replace some Si4+ and N3- ions respectively, and form the α-SiAlON solid solution.

Fig. 1 shows the XRD pattern of athe sintered cutting tool samples. The α-SiAlON phase (42-0251) is the predominant crystalline phase besides some while residual β- Si3N4 (33-1160) and (Y2Si3O3N4) (30-1460). This result indicates an initial reaction between Y2O3 and Si3N4 to forming Y2Si3O3N4, which slowly disappears with increasing sintering time by its dissolution in the liquid phase [22, 23 and 24]. Furthermore, dissolution of Al3+ and O2 ions in the phases were detected. Some α-Si3N4 structure lead to the formation of α-SiAlON.

dissolves and precipitates as β-Si3N4 when its solubility limit in the liquid is exceeded [16]. The X-ray data indicated an initial reaction between Y2O3 and Si3N4 to form Y2Si3O3N4 which slowly disappears with increased heating time by dissolution in the liquid distributed through the sample [16]. Solution-reprecipitation for Si3N4 with Y2O3 and AlN is controlled by interfacial reaction, as explained in previous works [17 and 18]. The interfacial reaction increases with increasing the solid-liquid interfacial surface. An increase in the interfacial surface was expected in the sample due to the homogeneous distribution of sintering aids within the green bodies, at consequently, the phase transformation was enhanced. The dissolution of Y2Si3O3N4 in the liquid phase was also enhanced by increasing the solid-liquid interfacial surface, promoted the phase transformation and the dissolution of crystalline phases.

Figure  1

3.3 Mechanical properties

The relative densityies of all specimen prepared exceeded the samples after sintering presented values higher than 97 % of the theoretical density., for all composition, reducing the effect of the densification on the mechanical properties.

The hardness and fracture toughness of the pressureless sintering samples is given along of the text. In literature, there is clearly a pattern to the mechanical properties as a function of transformation of α ↔ β phase change. Thus, the change of α and β contents resulting from the forward and reverse phase transformations clearly affects the mechanical properties in a reliable fashion. As the forward α → β transformation proceeds during temperature of 1500 -1800°C, thereby, decreasing the hardness of the ceramic. Since both the α- and β- grains have elongated morphologies, it would be expected that the toughness would not change significantly, possibly decreasing as a result of the agglomeration of the grain boundary phase. In fact, the α→ β transformation is accompanied by an increase in toughness, supporting the observation that the transformed β- has a greater aspect ratio than the parent α- phase. The greater aspect ratio enhances the toughening mechanisms of the phase, thereby, increasing the fracture toughness of the material.

As It can also be seen in Fig. 2, increasing grain size leads to decrease of the hardness of the that the hardness shows a general trend to microstructure of material, agreement to Shin [25]. Fracture toughness also decreased due to the smaller aspect ratio of the grains, despite their increased grain size. In the general, α-SiAlON presents typically higher hardness in comparison to β-Si3N4, because of the reduced amount of intergranular secondary phases. The sintered samples in this work presented a fracture toughness of the reduced amount of intergranular, secondary phases. The sintered samples prepared in this work presented a This is primarily a result of grain size, which will tend to slight decrease the hardness of the material. Therefore grain size will tend to increase toughness in microstructures with elongated morphologies. However, the aspect ratio of the β- grains may decrease with heated at a rate, amount and type additives, decreasing the toughness, thereby, counteracting the effect of grain growth.

In the general form, α-SiAlON present typically higher hardness values, superior than β-Si3N4. Sintered samples studies in this work presented fracture toughness and hardness of KIC=5.64 ± 0.12 MPa.m1/2 and HV=21.1 GPa ± 0.15 GPa,, respectively. In comparison with The results presented in this work, indicates that these ceramics present high toughness, compared with different commercial silicon nitride cutting toolstools [3], with a which presented hardness of near to 16 to-19GPa and a fracture toughness of 5 to-6MPam1/2 the α-SiAlON materials exhibit equal toughness fracture but increased hardness[2].

Figure  2

3.4. Effects of cutting speed on tool life

As the tool life criteria, a value of 0.6 mm of average flank wear land (VBB) for sintered cutting tools was used. During the cutting experiments, the cutting processes wasere paused after in each pass to measure for the average flank wear was measured. As the tool life criteria on average flank wear of 0.6 mm was used.

The cCutting speed effects (Vc) present different effects onto the tool life, tool wear, temperature and surface quality. While a higher speed decreases the cutting time, thus increasing the production rate, it also leads to increased tool wear and arise under an increased cutting speed [1 and 9]. Basically, in decreasing the cutting time and consequently, increasing the production rate, cutting speed is the most important factor. On the other hand, increase in cutting speed makes the tool wear fast and reduces the tool life. By the increase in tool wear, contact area between the tool and the workpiece increases and in consequence, higher cutting forces arise [16].

Different modes of tool failure modes were including rake face wear, flank wear and breakage (fracture) were observed in this work, including rake face wear, flank wear and fracture of the tool. Under the highest cutting speed of 500 m/mim, tool wear on the rake face predominated and, therefore, tool life is determined by the wear deepening until edge fracture results, as illustrated in Fig 6.study. Furthermore, it can be seen that the tool wear behavior changes under different cutting speeds. In the conventional cutting speed range from 200 to 350 m/min the tool wear an the rake face occurred in the form of pits, called crater, at some distance from cutting edge, while the wear on the rake face driving machining under 500 m/min was adjacent to the cutting edge. Higher cutting speeds may cause diffusion, micro-adhesion and plastic deformation due to the increased temperatures on the rake face, this aggravating abrasive tool wear. The α-SiAlON tool had a better performance under the cutting speeds of 200 and 350 m/min as under 500 m/min on turning CGI, due to the lower temperatures generated under these conditions. In Fig. 3 the effect of the cutting speed on the tool life, established as 0.6 mm flank wear, is shown.

Fig. 3

The flank wear increased with increasing cutting length made all three cutting speeds studied. This behavior was more obvious under the highest cutting speed of 500 m/min. Under cutting speeds of 200 and 350 m/min the wear behavior in regard to the cutting length was similar, with the former performing slightly better than the latter. This may be seen was clearly comparing Fig. 4 and 5, which show the flank wear of the tools at maximum cutting length. The flank wear rate curves, shown in Fig. 3, for the cutting speeds 200 and 350 m/min are almost parallel and very close to each other, indicating that both speeds cause similar flank wear characteristics. The negative impact a higher speed, such as 500 m/min, on tool wear is obvious from Fig. 6. As the cutting speed increases, flank wear increases also. This is more pronounced under a cutting speed of 500 m/min, as there is a sudden jump in the flank wear rate, shown in Fig. 3, while the rates under the speeds of 200 and 350 m/min remain almost constant between 0 and 1100 m. Furthermore, no flank wear up to 500 m under the lower cutting speeds has been observed, which is attributed to an only limited temperature rise due to friction, thus not affecting the transverse rupture strength. However, the α-SiAlON tool is not suitable for the high speed machining of CGI under a cutting speed of 500 m/min, leading to a catastrophic failure as can be seen in Fig. 6. The maximum length of the CGI machined under a cutting speed of 200 m/min has been 2192 m, under 350 m/min 1603 m and under 500 m/min 1568 m.

These tool wear patterns in turning compacted graphite iron suggested that the tool wear mechanisms were diffusion, attrition, and wear by possibly chemical interaction.

For the high-speed machining (Vc3), tool wear on the rake face predominates and therefore tool life is determined with the wear deepening until edge fracture results. Tool wear on the rake face when in the high-speed turning of compacted graphite iron using α-SiAlON ceramic tool is shown in Fig. 6. From this figure, it can be seen that the tool wear on rake face is different from crater wear in Vc1 and Vc2. In the conventional cutting speed range (Vc1 and Vc2), the tool wear on the rake face occurred in the form of a pit called the crater, which was formed at some distance from the cutting edge, while the tool wear on rake face during high-speed machining was adjacent to the cutting edge. Experimental results showed that increasing the cutting speed even further led to the increase of the wear area. This mode of tool wear is mainly due to too high cutting temperature on the rake face. Extreme high cutting speed leads to very high temperature (838–886 °C) occurring at the vicinity of the main cutting edge where the maximum depth of tool wear on rake face occurs. The hardness of inter-granular phase of the tool materials decreases at such high cutting temperatures, which aggravates the abrasive wear on tool rake face. The high cutting temperature also resulted in possibly diffusion, micro-adhesion, micro-plastic deformation, etc. The tool-chip contact length is shorter in high-speed machining than in the conventional cutting speed, which causes the cutting force to be concentrated adjacent to the main cutting edge. The softer cutting edge due to high temperature under the concentrated cutting force near the cutting edge leads to micro-deformation and micro-deflection. This is one of the main factors, in particular, for α-SiAlON tool wear. Additionally, the turning process may cause mechanical impact and thermal differences on α-SiAlON tool. Mechanical and thermal differences are non-negligible factors to in forming this kind of tool wear morphology. Thus, the combined effect of cutting force and cutting temperature are the main factor to lead to this type of tool wear on the rake face at high cutting speed (Vc3). The matching of mechanical, physical and chemical properties between the tool and the workpiece materials at high temperature is a very important factor in the high-speed machining process. It was clearly observed in this study that both the wear zone for α-SiAlON tool was great to Vc3. It was shown that the α-SiAlON tool had a better performance in Vc1 and Vc2 on turning compacted graphite iron through its high wear-resistance and high-temperature capability.

As indicated above 0.6 mm of flank wear was regarded as the tool life criterion for α-SiAlON tool. In the experiments in which the cutting speed varies, as the cutting speed changes the tool life changes. In Fig. 3, the effects of cutting speeds onto the tool life are shown. The performance of α-SiAlON cutting tool in turning on compacted graphite iron was observed experimentally. In the experiments, of ceramic cutting tool were taken into consideration throughout their tool life at certain cutting speed. During the measurement of VB, some crater wear was also observed, but the amount of crater wear was negligible until the flank face reaches to 0.4 mm. The experiments were continued to observe the affects of crater wear and it was recognized that the crater wear started to be effective after VB reaches to 0.50 mm and after that it was the crater wear that causes the tool fracture (breakage), as seen Fig. 6.

Figure 3

The Fig. 3 shows the flank wear progression with increasing cutting length for all three different cutting speeds. This was more obvious at the cutting speed of 500 m/min than at lower cutting speeds, see in Fig. 3. The Vc1 and Vc2 exhibited similar wear behaviors to part cutting length with the former performing slightly better than the latter. These can be seen more clearly in Fig. 4 and 5, which shows the average flank wear rates obtained by corresponding maximum cutting length (m). It is clearly seen in Fig. 3 that flank wear rate curves for Vc1 and Vc2 are almost exactly parallel to and very near each other indicating that both have similar flank wear characteristics. Therefore Vc1 behaved slightly better than Vc2 and at all other speeds. In facts, when Fig. 4 and 5 are studied, it can be claimed seen that Vc1 micrographic present to be slightly better than Vc2 micrographic. The negative impact of high cutting speed on tool wear is obvious in Fig. 6. As the cutting speed increases, all conditions undergo more flank wear. This is more pronounced for Vc3 than the other two, as there is a sudden jump in flank wear rate curve to Vc3. At lower speeds, more steady increases in flank wear rates are observed with increasing cutting lengths. Flank wear rates for Vc1 and Vc2 remain almost constant between 0 and 1100 m, see in Fig. 3. This pattern of wear produces wear lands on the flank face of the tool on account of the abrasive action of the machined surface. However fact different can been observed, no flank wear of 0 at 500m to Vc1 and Vc2. This event has been attributed to the small friction effect of chip which leads to the results of the local little temperature and pressure on the edge of tools, and hence it creates the low thermal impact. This low thermal impact permits the combinative effect of cutting tool material and results in not flank wear of tool. Low thermal impact cause the stability of yield strength of tool and lead to not mechanisms wear. This behavior was confirmed by the replicated test. In Fig. 4 and 5, it is noted that the wear land is of uniform width. The widest of the wear land in compacted graphite iron machining, is in the vicinity of the tool nose. This is due to the same facts as described above for tool wear on the rake face. In this study, flank wear was observed throughout the whole of the machining tests. Conform Fig. 6, the tool material present low transverse rupture strength and fracture toughness is prone to fracture. The transverse rupture strength for α-SiAlON tool is lower than traditional high-speed steel and carbide cutting tools. Thus, α-SiAlON tool is not suitable for the high-speed machining (up 500m/min) of compacted graphite iron. It can be seen from Fig. 6 that both the main and the minor cutting edges were broken down completely at the last pass cutting. The catastrophic failure of the α-SiAlON tool (see Fig. 6) was observed at the last of the cutting process, while only a normal wear pattern occurred in before last pass, when machining the compacted graphite iron with an α-SiAlON tool (see Fig.4 and 5).

Tool breakage is very dangerous due to the influence of area force in high speed in the turning process. The crater moves closer to the cutting edge as the cutting speed is increased, eventually damaging the cutting edge at a cutting speed of 500 m/min. Security measures to prevent the launch of a broken tool should be taken during the design of high-speed machine tools.

Taking the flank wear value of 0.6 mm as the tool life criterion, the maximum metal length machined with Vc1 is 2192 m, with Vc2 is 1603 m and with Vc3 is 1568 m.

Figure 4

Figure 5

Figure 6

These results show that the high hardness, α-SiAlON phase predominance and machining condition promoted important performance when machining compacted graphite iron.

Surface roughness results are shown in Fig. 7 and 8. Surface roughness measurements were taken at every each pass. Three readings were taken at three different points on the circumference, which were 120° apart. The values plotted on the graphs are the average of these three readings.

Figure 7

Figure 8

In Fig. 7 and 8 show the variation of the average observed that the surface roughness with increasing cutting speed. Surface roughness increased with increasing cutting length for all cutting speeds. Under the highest cutting speed of 500 m/min a linear increase has been observed, while under a speed of 200 or 350 m/min roughness tends to stabilize. The surface roughness is related to the flank wear, but other factors such as inhomogeneity of the cast material, random distribution of graphite nodules it may have some affect on the surface quality, too. Despite the stable appearance of the surface roughness curves, the last points in most of the curves exhibit slightly higher roughness values than the first points, except under a cutting speed of 500 m/min. Therefore, it may be stated that wear tools generally produce an inferior surface quality. It is believed that if the flank wear reached high values, e.g. > 0.5 mm, it would have a move pronounced effect on surface roughness, while this effect is less important for a flank wear similar than 0,3 mm.

were always continuously increase with increasing cutting length. Continuous increase in surface roughness with increasing cutting length was observed of manner linear in Vc3. However, in some other cases the surface quality obtained with Vc1 and Vc2 tend stable with increasing cutting length which is related to progression of flank wear. Therefore surface quality deteriorated up to a certain value of cutting length is improved. It is therefore difficult to draw general conclusions from these results about the influence of flank wear on surface roughness. As a matter of fact surface roughness is quite a complicated subject. Many other factors beside flank wear may have some effect on the quality of machined surfaces, for example inhomogeneity of cast materials, random distribution of graphite nodules, etc. However, despite the stable appearance of surface roughness curves, the last points in most of the curves have slight greater roughness values than the first points except for Vc3. Therefore, it would not be wrong to state that worn tools generally produce worse surface quality. It is believed that if flank wear reached higher values, e.g., >0.5 mm, it would have a more apparent negative effect on surface roughness. On the other hand, as long as flank wear remains below an acceptable limit, i.e., ................
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