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中原工學(xué)院畢業(yè)設(shè)計(jì)英文翻譯原文
Tool damage
Chapter 3 considered cutting tool minimum property requirements (both mechanical and thermal) to avoid immediate failure. By failure is meant damage so large that the tool has no useful ability to remove work material. Attention is turned, in this chapter, to the mech- anisms and characteristics of lesser damages that accumulate with use, and which eventu- ally cause a tool to be replaced. In reality, there is a continuous spectrum of damage severities, such that there is no sharp boundary between what is to be considered here and what might in practice be described as immediate failure. There is some overlap between this chapter and the previous one.
Chapters 2 and 3 have demonstrated that cutting tools must withstand much higher fric- tion and normal stresses – and usually higher temperatures too – than normal machine tool bearing surfaces. There is, in most cases, no question of avoiding tool damage, but only of asking how rapidly it occurs. The damages of a cutting tool are influenced by the stress and temperature at the tool surface, which in turn depend on the cutting mode – for exam- ple turning, milling or drilling; and the cutting conditions of tool and work material, cutting speed, feed rate, depth of cut and the presence or not of cutting fluid and its type. In Chapter 2, it was described in general that wear is very sensitive to small changes in sliding conditions. In machining, the tool damage mode and the rate of damage are simi- larly very sensitive to changes in the cutting operation and the cutting conditions. While tool damage cannot be avoided, it can often be reduced if its mode and what controls it is understood. Section 4.1 describes the main modes of tool damage.
The economics of machining were introduced in Chapter 1. To minimize machining cost, it is necessary not only to find the most suitable tool and work materials for an oper- ation, but also to have a prediction of tool life. At the end of a tool’s life, the tool must be replaced or reground, to maintain workpiece accuracy, surface roughness or integrity. Section 4.2 considers tool life criteria and life prediction.
4.1 Tool damage and its classification
4.1.1 Types of tool damage
Tool damage can be classified into two groups, wear and fracture, by means of its scale and how it progresses. Wear (as discussed in Chapter 2) is loss of material on an asperity or micro-contact, or smaller scale, down to molecular or atomic removal mechanisms. It usually progresses continuously. Fracture, on the other hand, is damage at a larger scale than wear; and it occurs suddenly. As written above, there is a continuous spectrum of damage scales from micro-wear to gross fracture.
Figure 4.1 shows a typical damage pattern – in this case wear – of a carbide tool, cutting steel at a relatively high speed. Crater wear on the rake face, flank wear on the flank faces and notch wear at the depth of cut (DOC) extremities are the typical wear modes. Wear measures, such as VB, KT are returned to in Section 4.2.
Damage changes, however, with change of materials, cutting mode and cutting condi- tions, as shown in Figure 4.2. Figure 4.2(a) shows crater and flank wear, with negligible notch wear, after turning a medium carbon steel with a carbide tool at high cutting speed. If the process is changed to milling, a large crater wear with a number of cracks becomes the distinctive feature of damage (Figure 4.2(b)). When turning Ni-based super alloys with ceramic tools (Figure 4.2(c)) notch wear at the DOC line is the dominant damage mode while crater and flank wear are almost negligible. Figure 4.2(d) shows the result of turning a carbon steel with a silicon nitride ceramic tool (not to be recommended!). Large crater and flank wear develop in a very short time. In the case of turning b-phase Ti-alloys with a K-grade carbide tool, large amounts of work material are observed adhered to the tool, and part of the cutting edge is damaged by fracture or chipping (Figure 4.2(e)).
4.1.2 Causes of tool damage
Chapter 2.4 outlined the general conditions leading to abrasive, adhesive and chemical wear mechanisms. In the context of cutting tool damage, the importance and occurrence of these mechanisms can be classified by cutting temperature, as shown in Figure 4.3. Three causes of damage are qualitatively identified in the figure: mechanical, thermal and adhesive. Mechanical damage, which includes abrasion, chipping, early fracture and fatigue, is basi- cally independent of temperature. Thermal damage, with plastic deformation, thermal diffu- sion and chemical reaction as its typical forms, increases drastically with increasing temperature. (It should be noted that thermal diffusion and chemical reaction are not the direct cause of damage. Rather, they cause the tool surface to be weakened so that abrasion, mechanical shock or adhesion can then more easily cause material removal.) Damage based on adhesion is observed to have a local maximum in a certain temperature range.
Mechanical damage
Whether mechanical damage is classified as wear or fracture depends on its scale. Figure 4.4 illustrates the different modes, from a scale of less than 0.1 mm to around 100 mm (much greater than 100 mm becomes failure).
Abrasive wear (illustrated schematically in Figure 2.29) is typically caused by sliding hard particles against the cutting tool. The hard particles come from either the work mater- ial’s microstructure, or are broken away from the cutting edge. Abrasive wear reduces the harder is the tool relative to the particles and generally depends on the distance cut (see Section 4.2.2).
Attrition wear occurs on a scale larger than abrasion. Particles or grains of the tool material are mechanically weakened by micro-fracture as a result of sliding interaction with the work, before being removed by wear.
Next in size comes chipping (sometimes called micro-chipping at its small-scale limit). This is caused by mechanical shock loading on a scale that leads to large fluctuations in cutting force, as opposed to the inherent local stress fluctuations that cause attrition.
Finally, fracture is larger than chipping, and is classified into three types: early stage, unpredictable and final stage. The early stage occurs immediately after beginning a cut if the tool shape or cutting condition is improper; or if there is some kind of defect in the cutting tool or in its edge preparation. Unpredictable fracture can occur at any time if the stress on the cutting edge changes suddenly, for example caused by chattering or an irreg- ularity in the workpiece hardness. Final stage fracture can be observed frequently at the end of a tool’s life in milling: then fatigue due to mechanical or thermal stresses on the cutting edge is the main cause of damage.
Thermal damage – plastic deformation
The plastic deformation type of thermal damage referred to in Figure 4.3 is observed when a cutting tool at high cutting temperature cannot withstand the compressive stress on its cutting edge. It therefore occurs with tools having a high temperature sensitivity of their hardness as their weakest characteristic. Examples are high speed steel tools in general; and high cobalt content cemented carbide tools, or cermet tools, used in severe conditions, particularly at a high feed rate. Deformation of the edge leads to generation of an improper shape and rapid material removal.
Thermal damage – diffusion
Wear as a result of thermal diffusion occurs at high cutting temperatures if cutting tool and work material elements diffuse mutually into each other’s structure. This is well known with cemented carbide tools and has been studied over many years, by Dawihl (1941), Trent (1952), Trigger and Chao (1956), Takeyama and Murata (1963), Gregory (1965), Cook (1973), Uehara (1976), Narutaki and Yamane (1976), Usui et al. (1978) and others. The rates of processes controlled by diffusion are exponentially proportional to the inverse of the absolute temperature q. In the case of wear, different researchers have proposed different pre-exponential factors: Cook (1973) suggested depth wear h should increase with time t (equation 4.1(a)); earlier, Takeyama and Murata (1963) also suggested this and the further possibility of sliding distance s being a more fundamental variable (equation 4.1(b)); later Usui et al. (1978), following the ideas of contact mechanics and wear considered in Chapter 2.4, proposed wear should also increase with normal contact stress sn (equation 4.1(c)). In all these cases, a plot of ln(wear rate) against 1/q gives a straight line, the slope of which is –C2
igure 4.5 shows experimental results for both the crater and flank depth wear rates of a 0.25%C and a 0.46%C steel turned by a P20 grade carbide tool, plotted after the manner of equation (4.1c). Two linear regions are seen: in this case the boundary is at 1/q ≈ 8.5 10–4 K–1 (or q ≈ 1175 K). The slope of the higher temperature data (q > 1175 K) is typi- cal of diffusion processes between steels and cemented carbides (Cook, 1973). The smaller slope at lower temperatures is typical of a temperature dependent mechanical wear process, for example abrasion.
Diffusion can be directly demonstrated at high temperatures in static conditions. Figure 4.6 shows a typical result of a static diffusion test in which a P-grade cemented carbide tool was loaded against a 0.15% carbon steel for 30 min at 1200?C. A metallographic section through the interface between the carbide tool and the steel, etched in 4% Nital (nitric acid and alcohol) shows that the pearlite in the steel has increased from its original level. This means that carbon from the cemented carbide has diffused into the steel. Furthermore, elec- tron probe micro-analysis (EPMA) shows that Co and W from the tool material also diffuse into the steel; and iron from the steel diffuses into the tool material. Many researchers agree that mutual diffusion is the cause of carbide tool diffusion wear, but there is not agreement in detail as to the mechanism that then results in material removal.
Naerheim and Trent (1977) have proposed that the wear rates of both WC-Co (K-grade) and WC-(Ti,Ta,W)C-Co (P-grade) cemented carbides are controlled by the rate of diffusion of tungsten (and Ti and Ta) and carbon atoms together into the work material, as indicated in Figure 4.7. This view is based on transmission electron microscope (TEM) observations on crater wear that show no structural changes in the tool’s carbide grains within a distance of 0.01 mm of the tool–chip interface. The slower wear of P-grade than K-grade materials is explained by slower diffusion in the former than the latter case. Naerheim and Trent state that, in their cutting tests, pulled-out carbide grains were not observed adhering to the underside of chips. This was not Uehara’s (1976) experience. He collected chips after turn- ing a 0.47% C steel with a K-grade or a P-grade tool, dissolved the chips in acid to extract adhered carbides and finally passed the solution through a 0.1m filter, to classify the carbide sizes. With K-grade tools, he only observed carbides less than 0.1 mm in size, in accord with Trent. However, with P-grade tools he observed carbides greater than 0.1 mm in size. This suggests a different wear mechanism for K- and P-type materials.
Other examples of diffusion wear are the severe wear of diamond cutting tools, silicon nitride ceramic tools and SiC whisker reinforced alumina ceramic tools when machining steel. Carbon, silicon and nitrogen all diffuse easily in iron at elevated temperatures; and silicon nitride and silicon carbide dissolve readily in hot iron.
Thermal diffusion wear of carbide tools can be decreased if a layer acting as a barrier to diffusion is deposited on the tool. There are two types of layer in practice: one is as provided by coated tools; the other is a protective oxide layer deposited on the wear surfaces during cutting special deoxidized steels (for example Ca-deoxidized steels), commonly known as a ‘belag’ layer.
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