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河南理工大学万方科技学院毕业设计翻译毕业设计翻译院系名称： 机械设计及其自动化 班 级： 机设10升（一）班 学 号： 0816101012 学生姓名： 汪洋 指导教师： 韩晓明 considering elastic contact forces as external forces on an otherwise free chip. The line n= 5 in Figure 6.5 was deduced for an elastic contact length five times the plastic length .The elastic contact should not be ignored in machining analyses. Slip-line field modelling may also be applied to machining with restricted contact tools(Usui et al., 1964), with chip breaker geometry tools (Dewhurst, 1979), with negative rake tools ( Petryk , 1987), as well as with flank-worn tools (Shi and Ramalingham , 1991), to give an insight into how machining may be changed by non-planar rake face and cutting edge modified tools. Figures 6.6 and 6.7 give examples. Figure 6.6 is concerned with modifications to chip flow caused by non-planar rakefaced tools. As the chip/tool contact length is reduced below its natural value by cutting away the rake face (Figure 6.6(a), the sliding velocity on the remaining rake face is reduced, with the creation of a stagnant zone, and the chip streams into the space created by cutting away the tool. If a chip breaker obstruction, of slope d, is added some distance l B from the cutting edge of a plane tool (Figure 6.6(b), its effect on chip curvature and cutting forces can be estimated . The combination of these effects can give some guidance on the geometrical design of practical chip-breaker geometry tools. The slip-line fields of Figure 6.7 show how, with increasingly negative rake angle, a stagnant zone may develop, eventually (Figure 6.7(c) allowing a split in the flow, with material in the region of the cutting edge passing under the tool rather than up the rake face. The fields in this figure, at first sight, are for tools of an impractically large negative rake angle. However, real tools have a finite edge radius, can be worn or can be manufactured with a negative rake chamfer. The possibility of stagnation that these fields signal , needs to be accomodated by numerical modelling procedures.6.2.4 SummaryIn summary, the slip-line field method gives a powerful insight into the variety of possible chip flows. A lack of uniqueness between machining parameters and the friction stress Between the chip and tool is explained by the freedom of the chip, at any given friction stress level, to take up a range of contact lengths with the tool. Chip equilibrium is maintained for different contact lengths by allowing the level of hydrostatic stress in the field to vary. The velocity fields indicate where there are regions of intense shear, which should be taken into account later in numerical modeling . They also illustrate how velocities might vary in the secondary shear zone, a topic that will be returned to later. They also show a range of variations of normal contact stress on the rake face that is observed in practice. However, a frustrating weakness of the slip-line field approach is that it offers no way, within the limitationsof the rigid perfectly plastic work material model, of removing the non-uniqueness: what does control the chip/tool contact length in a given situation ? Additionally, it can offer no way of taking into account variable flow stress properties of real materials, demonstrated experimentally to have an influence. An alternative modeling ,concentrating on material property variation effects, is introduced in the next section.6.3 Introducing variable flow stress behaviorSlip-line field modelling investigates the variety of chip formation allowed by equilibrium and flow conditions while grossly simplifying a metals yield behaviour. A complementary approach is to concentrate on the effects of yield stress varying with strain (strain hardening)and in many cases with strain rate and temperature too, while simplifying the modelling of equilibrium and flow. Pioneering work in this area is associated with the name of Oxley. The remainder of this section relies heavily on his work, which is summarized in Mechanics of Machining (Oxley, 1989). Developments may be considered in four phases: firstly experimental and numerical studies of actual chip flows, by the method of visioplasticity; secondly,simplifications allowing analytical relations to be developed between stress variations in the primary shear zone and material flow properties, dependent on strain, strain rate and temperature; thirdly, a consideration of stress conditions in the secondary shear zone; and finally, a synthesis of these, allowing the prediction of chip flow from work material properties.6.3.1 Observations of chip flowsVisioplasticity is the study of experimentally observed plastic flow patterns. In its most complete form, strain rates throughout the flow are deduced from variations of velocity with position, and strains are calculated by integrating strain rates with respect to time along the streamlines of the flow. The temperatures associated with the plastic work are calculated from heat conduction theory. Then, from independent knowledge of the variation of flow stress with the strain, strain rate and temperature, it can be attempted to deduce what the stress variations are throughout the flow and what resultant forces are needed to create the flow. Alternatively, measured values of the forces can be used to deduce how the flow stress varied. Frequently, however, the accuracy of flow measurement is not good enough to support this entire scheme. Nonetheless, useful insights come from only partial success. In the case of plane strain flows, the first step is usually to determine the maximum shear strain rate trajectories of the flow, and from these to construct the slip-line field.Departures of the fields shape from the rules established for perfectly plastic solids(Section 6.2) are commonly observed. Figure 6.8(a) shows an early example of a chip primary shear zone investigated in this way (Palmer and Oxley, 1959). In addition to flow calculations in deriving this field, Palmer and Oxley also applied the force equilibrium constraint, that the slip-lines should intersect the free surface AA at 45. The field is for a mild steel machined at the low cutting speed of 12 mm/min and a feed of 0.17 mm. At the low strain rates and temperatures generated in this case, departures from perfect plasticity are expected to be due only to strain hardening. The strain hardening behaviour of the material was measured in a simple compression test. Two conclusions arise from Figure 6.8 (and from other examples that could have been chosen). First, and most obviously, the entry and exit slip lines OA and OA are of opposing curvature. The field violates equation (6.4). This is a direct effect of work-hardening. Secondly, and less obviously, there is a problem with the constraint placed on the field that the slip-lines should meet the free surface at 45. By revisiting the derivation of equations (6.1) (Appendix 1, Section 1.2.2), and removing the constraint of no strain hardening, it is easy to show that where s1 and s2 are distances along an a and a b slip line respectively. In Figure 6.8(a), as in Figure 6.1, AC is a b line and CA an a line. After estimating the variations of k, k/s1 and k/s2 in the region of AAC, Palmer and Oxley concluded, from the application of equation (6.5), that the hydrostatic pressure at A could not equal the shear yield stress of the work hardened material at A, as it should according to the further constraint imposedby the free surface boundary condition there. Palmer and Oxley resolved the contradiction by suggesting that plastic flow was not steady at the free surface. The smoothed free surface in Figure 6.8(a) is, in reality, corrugated and therefore the slip-lines should not be constrained to intesect the smoothed profile at 45. The result of a later study (Roth and Oxley, 1972), still at low cutting speed to exclude the effects of strain rate and temperature on flow stress now also including an estimate of the secondary shear zone shape is shown in Figure 6.8(b). At A, the entry boundary OA is still made to intersect the free surface at 45: there, continuity of flow ensures that the free surface slope is known (velocity discontinuities cannot exist in a hardening material discontinuities that would occur in a non-hardening material are broadened into narrow zones). However, a free surface constraint has not been placed on the exit boundary direction at A; and no attempt has been made to detail the field within the near-surface region AAC. Roth and Oxley applied equations (6.5) to the calculation of hydrostatic stress along all the field boundaries, assuming that at A its value was that of the shear yield stress there. These are shown in the figure. Along the entry boundary OA, hydrostatic stress variations are dominated by the effect of work hardening. Integration of the hydrostatic and shear stresses with respect to distance along OA gives the force acting across it. Inclusion or work hardening gives a value of 1.77 k N (in line with experiment), while omitting it gives 3.19 k N, in a grossly different direction. Over the exit boundaries BD and DA, where strain hardening has reduced the rate of change of shear flow stress across the slip lines, the variations approach those expected of a non-hardening material. They depend on the direction changes along the lines. The exit region OBDA is visually similar in this example to the non-hardening slip-line field proposed by Dewhurst (Figure 6.2(c). The whole field is this, with the primary shear plane replaced by a work hardening zone of finite width. In a parallel series of experiments, Stevenson and Oxley (196970, 197071) extended the direct observations of chip flows to higher cutting speeds, but with a changed focus, to assess how large might be the strain rate and temperature variations in the primary shear zone. Figure 6.9(a) is a sketch of the streamlines that they observed when machining a 0.13%C free-machining steel at a cutting speed of 105 m/min and a feed of 0.26 mm. Figure 6.10 shows, for a range of cutting speeds, the derived variations of maximum shear strain rate along a central streamline, such as a a in Figure 6.9(a). The peak of maximum shear strain rate is observed to occur close to the line OA that would be described as the shear plane in a shear plane model of the machining process. The peak maximum shear strain rate was measured to vary in proportion to the notional primary shear plane velocity (from equation (2.3) and inversely as the length s of the shear plane (assumed to be f / sin f):In this case, the best-fit constant of proportionality C is 5.9. In many practical machining operations, peak shear strain rates are of the order of 104/s. It is interesting to consider the value of C = 5.9 in the light of the length-to-width ratio of the primary shear zone, equal to 2, derived in Chapter 2 from Figure 2.10 and equation (2.7). The average shear strain rate may be roughly half the peak rate. It is also the total shear strain divided by the time for material to pass through the primary zone. This time is the width of the zone divided by the work velocity normal to the plane, namely U work sin f. An easy manipulation equates the length-to-width ratio to C / 2, or about 3 in this case. A consistent view emerges of a primary shear region in which the strain rates do in fact peak along a plane OA but which in its totality may not be as narrow compared with its length as is commonly believed. Temperature rises in the primary zone have already been considered in Chapter 2.Stevenson and Oxley used the same approach described there to obtain the total temperature rise from the measured cutting forces resolved on to the shear plane. In the notation of this book, combining equations (2.4a), (2.5c) and (2.14), and remembering that only a fraction (1 b) of generated heat flows into the chip However, as will be seen in the next section, there is a particular interest in the temperature rise in the plane OA where the strain rate is largest. Stevenson and Oxley took the temperature along OA to beWhere h can range from 0 to 1. Usually, they took it to equal 1, but this is not consistent with OA being upstream of the exit boundary of the primary zone. They commented that lower values (0.7 to 0.95) might be better (Oxley, 1989). 考虑弹性接触力为外力在另一个自由芯片上的作用。线n= 5在图6.5推导出对于弹性接触长度五倍塑料的长度。弹性接触分析在加工过程中不应该被忽略。 滑移线场模型也可用于有限制的联系工具加工(1964年高庆宇等)，与芯片断路器几何工具(杜赫斯特,1979年),与负前角工具(Petryk,1987), 以及与侧翼陈腐的工具(史和Ramalingham,1991), 给予了怎样的加工可以由非平面倾斜变化的洞察力脸和尖端修改工具。图6.6和6.7给出了例子。图6.6是关于修改芯片流量非平面倾斜工具。由于该芯片/工具的接触长度，下面是随它的自然价值的减少而远离前刀面(图6.6(a)，对剩余的前刀面滑动速度降低，与创建一个停滞区,该芯片流到这个空间创造的被切掉的工具。如果断屑阻塞、边坡失稳d,是在一个平面刀具的切削刃前面加了一些距离IB(图6.6(b)，其对芯片的曲率和切力的影响可以预计。这些效果组合可以提供一些实用的断屑槽几何工具的几何设计指导。滑移线场显示,随着越来越多的6.7%负前角,可以形成一个停滞区，最终(图6.7(c),允许分裂流与材料对该区域在车底的切削刃的工具而不是上升的前刀面。在这个图的领域,乍一看,是为了负前角工具不切实际地增大角度。然而,真正的工具有一个有限的刃口半径,可以穿也可制造一个负斜面槽。停滞的可能性,这些领域的信号是需要通过数值模拟的程序来归纳的。6.2.4总结总之,滑移线场给出了一个可以观察各种各样的芯片流动的可能性的方法。一个缺乏独特性的摩擦应力加工参数的芯片和工具可以解释为自由的芯片,在任何给定的摩擦应力水平下,采取了一系列的接触长度的工具。芯片是保持在外静压力的时候允许不同接触长度的平衡。该速度场表明那里有强烈的剪切区域,这应考虑到后面的数值模拟。他们还说明了如何在速度可能会有所不同的二次剪切区,将回到后面的课题。他们还表现出对前刀面的接触应力的变化时正常的，在实践中可以观察到。然而,一个令人沮丧的弱点，滑移线场的方法没有提供在完全的刚性材料模型的塑性中消除非唯一性，局限性：在某种情况下，什么是芯片/工具接触长度的控制？此外,它可以提供没有考虑应力变流量特性采取的方法,通过试验可以证明它的影响。另一种模型,对材料特性的变化的影响，在下一个部分集中介绍。图6.6 与切割滑移线场模型（a）零接触限制倾斜和（b）断屑几何工具。杜赫斯特(1979）图6.7 芯片流量使用工具，从（a）至（c）越来越倾斜。Petryk(1987)6.3介绍可变的流动应力的行为滑移线场模拟考察了各种芯片的形成和流动的平衡条件允许时可以简化一个金属的屈服行为。一个互补的方法是集中于产量的应变（应变硬化）和应变率和温度太多的情况下应变力的影响，同时简化了平衡流动模型。开拓这方面的工作是奥克斯利。本节的其余部分在很大程度上依赖于他的工作，这是在总结了加工机械(奥克斯利,1989年9月初版)。也可以被认为是发展的四个阶段:首先实验和数值研究了实际芯片流动,由visioplasticity方法;其次,简化，使分析应力之间关系的变化发展的主要剪切带和材料的流动性能,应变、应变率和温度依赖型;第三,考虑应力条件在二次剪切带;最后,综合这些,使芯片流动特性可以预测。6.3.1芯片流动的观测Visioplasticity是实验观察到的塑性流动模式的研究。在其最完整的形式，整个流程应变率推导出速度与位置的变化，通过整合和应变方面的应变率沿水流流线的时间计算。与塑料相关的工作温度从热传导理论计算。然后，从流动与应变，应变率和温度应力变化独立的知识，可以推断出什么导致应力变化是整个流程以及由此产生的变化是需要创造的流动。另外,测量值的大小可以用来推断如何流动应力变化。通常情况下，然而，流量测量精度不够好来支持这整个的研究。不过，有益的见解来自只有部分成功。在平面应变流动的情况下,第一步通常是确定最大剪应变的流量轨迹,并从这些构建滑移线场。为理想塑性固体建立了从该领域的规则出发的模型（6.2节）是经常观察到的。图6.8（a）所示的一个芯片主要剪这种方式(帕尔默和奥克斯利,1959年)研究区的早期范例。除了在这一领域的流动计算，帕尔默和奥克斯利还采用了力平衡约束，即滑线应在45 相交自由表面局部。该部分于低切削速度为12毫米/分钟和塑料加工的0.17毫米温和钢。在低应变速率这种情况下产生的温度，完美塑性出发，预计仅是有应变硬化产生的。该材料的应变硬化行为是衡量一个简单的压缩试验。 从图6.8得到两个结论(以及其他的例子,本来是可以选择的)。首先，最明显的是，进入和退出滑移线和OA自动工作的反向是曲率。违反本场方程（6.4）。这是与工作直接影响的硬化。 其次,不太明显, 有一个与约束问题在线场上放置该滑线应符合在45自由表面。通过方程的推导(6.1)(见附录1、剖面1.2.2 ),以及取消没有应变硬化约束,很容易证明 其中S1和S2是沿着一个与AB的距离分别的滑移线场。在图6.8（a），如图6.1，交流是AB线和CA的一条线。估计在以后的AAC，帕尔默和奥克斯利研究下，的K / s1和的K /S2的变化得出结论，从应用方程（6.5）,这在一个静水压力可能不等于剪切屈服应力在A的工作硬化材料
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