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中原工學(xué)院畢業(yè)設(shè)計（論文）譯文 畢業(yè)設(shè)計（論文）譯文 題目名稱： 帶提手的桶蓋注塑模具設(shè)計 院系名稱： 機電學(xué)院 班 級： 機自073班 學(xué) 號： 200600314327 學(xué)生姓名： 張軍超 指導(dǎo)教師： 胡 敏 2011 年3月 6.3.2 近似分析中的主剪切帶 雖然在第一剪切帶對靜應(yīng)力變化有一個全面的分析，如圖6.8（b），可能對考慮其形成過程中的可能產(chǎn)生碎片的裂痕是有用的，如果目的只是為了跨越剪切帶來預(yù)測力的傳輸（幅度和方向），它可能不是必要的。例如，如果沿平面OA″，如圖6.9（a），靜應(yīng)力的變化主要以流動應(yīng)力的變化為主，而不是靠旋轉(zhuǎn)中的滑移線場，沿著OA″一個近似的壓力分析，而忽略旋轉(zhuǎn)，可能就足夠了。這是由奧克斯利發(fā)現(xiàn)的方法。 圖6.9（b）結(jié)合圖6.8（b）和6.9（a）方面，顯示了典型的流場邊界，但強調(diào)一個狹窄的圍繞著平面OA″的矩形區(qū)域。在A“處的流體靜應(yīng)力應(yīng)該是有一定的ps的值，然后，通過與公式（2.7）（第2章）推導(dǎo)類比，假設(shè)壓力變化沿著OA″ (長度為s)是由?k/?s1主宰，由此得到的力的方向?穿過OA″是被給予 R的大?。ㄅcD，切削深度）是被計算出從 奧克斯利介紹了如何對涉及方程右邊（6.9a）第二任期材料的加工硬化行為，表現(xiàn)為 和對OA的剪應(yīng)變率，從方程（6.6），以取代方程（6.9a） 術(shù)語Cn可能被認為是一個對PS / KOA″的值的校正，tan(f + l – a)將在任何情況下的應(yīng)變硬化的影響。非唯一性的應(yīng)變硬化情況已經(jīng)被考慮在第6.2節(jié)。在那里，圖6.4給出了一個tan(f + l – a)與F組的范圍變化對于零前角工具的例子。在他的著作中，奧克斯利制約了非硬化關(guān)系允許的范圍，提議 這可以從圖6.4接近允許范圍的上邊界看出。因此，最后， 如圖6.11，低碳鋼σ0 (o) 和 n (?)的變化來源于加工測試，相比壓縮試驗數(shù)據(jù)（一） 在某種程度上，限制了ps/kOA″的變化是有效的，方程（6.9b）和（6.13）可用于研究的應(yīng)變，應(yīng)變率和溫度在主剪切帶的依賴流。史蒂文森和奧克斯利（1969-70，1970-1971）進行了切削實驗對0.13％C鋼其切割速度可達300米/分鐘，并進行了量具力和剪切面角度測試。他們計算n從方程（6.l3），假定C = 5.9。他們計算KOA″從方程（6.9b），并乘以√3到OA上獲得同等流量壓力，他們計算在OA″上的等效應(yīng)變，假設(shè)它是總應(yīng)變的一半，最后得出S0（方程（6.10））。他們還計算了在OA“上的應(yīng)變率和溫度。圖6.11顯示了應(yīng)變速率和溫度的變化，他們導(dǎo)出了s0 and n ，應(yīng)變率和溫度組合成一個單一的功能，被稱為變溫速度，TMOD (K): 有材料科學(xué)的理由（第7章）為什么應(yīng)變速率和溫度可能以這種方式結(jié)合起來。n是一個常數(shù)，取為0.09，而e?—0應(yīng)變率的參考，取為1。 該圖還顯示了數(shù)據(jù)從一個類似的碳素鋼壓縮試驗中的得到的和進一步的數(shù)據(jù)（應(yīng)力強度）第二剪切流的分析報告，這將在6.3.3節(jié)中討論得出的數(shù)據(jù)確定。機加工和壓縮試驗的數(shù)據(jù)是不定量協(xié)議，但有一個質(zhì)的相似性在他們的變化與變溫速度中，支持這一觀點，至少具有切割速度變化的加工力和剪切面角度的有些部分是由于流動與應(yīng)變，應(yīng)變率和溫度應(yīng)力變化。 在剛才所說的一些程序中顯然有一個假設(shè)，在(f + l – a) 中所有的變化是由于在n中的變化；這平行雙面剪切帶模型是足夠的（在實踐中會有所不同應(yīng)變率從切削邊緣到自由面，剪切帶的實際寬度可變）; 及的C實際上是加工過程中的常數(shù)。在以后的工作中，奧克斯利調(diào)查了他的造型靈敏度C的一個變化。A變更到C造成了靜水壓力梯度沿剪切面和從而在切削工具的尖端的正常接觸應(yīng)力，sn,O的變化。添加sn,O來自主剪切面造型約束應(yīng)被視為是相同的，從第二剪切建模（第6.3.3節(jié)）， 他的結(jié)論是同一鋼，他最初給出的值C = 5.9，但在更廣泛的進給，速度和前角切削條件下C可能在3.3和7.1之間變化。有興趣的讀者可以參考法案（1989年）。 6.3.3 第二剪切帶的流動 隨著對部分例外的低速切削試驗像羅斯和奧克斯利（圖6.8），可視塑性研究從來沒有準(zhǔn)確充分的給出信息關(guān)于在第二剪切帶中應(yīng)變率和應(yīng)變分布在具有同等水平的詳細揭示了主剪切帶。當(dāng)然，在高速切削中，內(nèi)部網(wǎng)或其他標(biāo)記必要的流后完全被毀。也沒有任何辦法，相當(dāng)于運用方程在主要區(qū)域（6.13）中推導(dǎo)在第二剪切帶流應(yīng)變的硬化指數(shù)n。所以，即使流動應(yīng)力可推導(dǎo)出材料在那里，一個S0值（方程（6.10））和一個TMOD的估計值的提取可能被認為是不切實際的。然而，圖6.11包含，TMOD的應(yīng)力強度變化，例如塑性流動應(yīng)力變化的信息。使這個數(shù)據(jù)將提交的見解和假設(shè)是值得考慮的。 奧克斯利法案明確提出，在第二剪切帶應(yīng)變硬化將超過1.0的應(yīng)變可以忽略不計。這使得他 從方程(6.10) 和e—= 1中去識別s0 和 s—.。這是在材料的加工建模中的主要問題，返回到7.4章-確定流動應(yīng)力事實上如何在二次切變產(chǎn)生的高應(yīng)變應(yīng)力變化。奧克斯利然后建議S 是與應(yīng)力強度一至或√3tav，在那里tav是在芯片/工具的接觸面上的平均摩擦應(yīng)力（除以接觸面積測量摩擦力獲得）。假如有一個微不足道的彈性接觸的地，區(qū)從加工中的摩擦條件（第2章）考慮這是合理的。奧克斯利認為在他的（羅斯和奧克斯利，1972年）低速觀察的基礎(chǔ)上這事事實，但觀察圖6.5是不支持的。 為了確定TMOD的值，他估計在第二剪切帶中具有代表性的溫度和應(yīng)變率。對于應(yīng)變速率e?—int他認為第二剪切帶的平均寬度dt2，而在這個寬度上芯片的速度從前刀面為0到其體積值Uchip。 因而 他把代表溫度認為是在刀面上的平均溫度，計算其方式類似于方程（2.18），但是考慮到隨溫度變化的熱性能和對那些在二次剪切熱產(chǎn)生的現(xiàn)象并不完全平面而是通過二次分配剪切帶（黑斯廷斯等。，1980）。在這本書的，方程（2.18）是被修改被一項因子c 如圖6.11 計算(sint, TMOD)的數(shù)據(jù)結(jié)果從這些假設(shè)中得出。他們遵循了預(yù)計從提供了一些支持這些觀點的獨立機機械測試的變化。有一個假設(shè)，因為它需要特別有意思的返回：那就是在芯片/工具界面滑動速度為零。這強烈地影響著雙方的應(yīng)變率的計算，和對溫度c的計算校正的需要。該滑移線場模擬不支持這樣的芯片運動的嚴重下降。如圖6.2，例如，只有在某些情況下和然而僅接近于前端，顯示的滑動速度才降低到零。解決在這些變流動應(yīng)力 和滑移線場上前刀面滑動速度觀點上的沖突，導(dǎo)致對在在高速（溫度影響）加工前刀面的狀況有了深入的了解。 在他的工作，奧克斯利在最接近刀面上確定了兩個二次剪切帶，一個較寬的一個和一個較窄的一個。這窄區(qū)也已經(jīng)被確定被特倫特大學(xué)，特倫特大學(xué)描述它為流區(qū)，當(dāng)它的發(fā)生是由于區(qū)域中扣押之間的芯片和工具（遄達，1991年）發(fā)生。圖6.12（a）表明了在狹窄區(qū)域奧克利斯的測量厚度，為切割速度和進給的范圍，為0.2％C處打開-5 °刀具前角（其他結(jié)果鋼的例子為0.38％C處，鋼和+5 °刀具前角，也可以被證明）。流區(qū)是越薄有越大的切割速度和越低的進給。如果假定的接觸長度L等于芯片厚度t, 發(fā)生在方程式（6.16）是與t（kworkl / Uchip）?一致的。實驗結(jié)果位于在一個平均坡度0.2的線性帶里。流區(qū)位于 圖6.12在進給量（mm）為0.5 (?), 0.25 (+)和0.125 (o)；隨著（a）切割速度流區(qū)厚度的變化。 奧克斯利指出，該流區(qū)的溫度將會降低它的厚度通過因子C（方程（6.16）），并認為其應(yīng)變率會增加稀釋（方程（6.15））。對應(yīng)變速率和溫度這些厚度影響將導(dǎo)致作為一個有厚度的正變溫的速度將是最大的，和剪應(yīng)力最小流量。他建議將采取的厚度，將TMOD的價值最大化。這提供了帶標(biāo)記的價值'理論'在圖6.12（b）那預(yù)測的波段大約50%位于觀察一之上，給予足夠接近有效性的建議。 在第2章（圖2.22（a）項），直接測量出的隨著前刀面溫度摩擦系數(shù)m的變化已提交，為了車削0.45％C鋼。流區(qū)的厚度并沒有被測量在這些測試中。但是，如果實驗關(guān)系如圖6.12（b）的假設(shè)是成立的，圖2.22（a）的數(shù)據(jù)可轉(zhuǎn)化為√3mk (或者sint)在TMOD上的一個依賴。圖6.13顯示了結(jié)果，并且被奧克斯利比較它和0.45％C鋼的使用價值。那兩組數(shù)據(jù)之間的融洽是更好的比在圖6.11的，但并不完美。 注：文章來源Metal_Machining。 6.3.2 Approximate analysis in the primary shear zone Although a complete analysis of hydrostatic stress variations in the primary shear zone, asin Figure 6.8(b), might be useful in considering the possible fracture of chips during theirformation, it might not be necessary if the objective is only to predict the force transmission(the magnitude and direction) across the shear zone. If, for example, along the plane surface OA″ in Figure 6.9(a), variations of hydrostatic stress are dominated by flow stress variations rather than by rotations in the slip-line field, an approximate analysis of stress along OA″, neglecting rotations, might be sufficient. This is the approach developed by Oxley. Figure 6.9(b) combines aspects of Figures 6.8(b) and 6.9(a), showing the boundaries of a typical flow field but emphasizing a narrow rectangular region around the plane OA″.The hydrostatic stress at A″ is supposed to have some value ps. Then, by analogy with the derivation of equation (2.7) (Chapter 2), and after assuming pressure variations along OA″(of length s) are dominated by ?k/?s1, the direction of the resultant force R across OA″ is given by The size of R (with d, the depth of cut) is found from Oxley showed how to relate the second term on the right-hand side of equation (6.9a) to the work-hardening behaviour of the material, expressed as and to the shear strain-rate on OA″, from equation (6.6), in order to replace equation (6.9a) by The term Cn may be thought of as a correction to the value ps/kOA″ that tan(f + l – a) would have in the absence of any strain hardening effects. The non-uniqueness of the nonhardening circumstance has already been considered in section 6.2. There, Figure 6.4 gives a range for the variation of tan(f + l – a) with f, for the example of a zero rake angle tool. In his work, Oxley constrained the range of allowable non-hardening relations, to propose that This can be seen in Figure 6.4 to be close to the upper boundary of the allowable range.Then, finally, Fig. 6.11 Variations of σ0 (o) and n (?) for a low carbon steel, derived from machining tests, compared with compression test data (—) To the extent that constraining the variations of ps/kOA″ is valid, equations (6.9b) and (6.13) may be used to investigate the strain, strain-rate and temperature dependence of flow in the primary shear zone. Stevenson and Oxley (1969–70, 1970–71) carried out turning tests on a 0.13%C steel at cutting speeds up to around 300 m/min, measuring tool forces and shear plane angles. They calculated n from equation (6.l3), assuming C = 5.9. They calculated kOA″ from equation (6.9b), and multiplied it by √3 to obtain the equivalent flow stress on OA″; they calculated the equivalent strain on OA″, assuming it to be half the total strain; and finally derived s0 (equation (6.10)). They also calculated the strain rate and temperature on OA″. Figure 6.11 shows the variations with strain rate and temperature they derived for s0 and n. Strain rate and temperature are combined into a single function, known as the velocity modified temperature, TMOD (K): There are materials science reasons (Chapter 7) why strain rate and temperature might be combined in this way. n is a material property constant that was taken to be 0.09, and e?—0 is a reference strain rate that was taken to be 1. The figure also shows data derived from compression tests on a similar carbon steel and further data (sint) determined from the analysis of secondary shear flow, which will be discussed in Section 6.3.3. The data for machining and compression tests are not in quantitative agreement, but there is a qualitative similarity in their variations with velocity modified temperature that supports the view that at least some part of the variation of machining forces and shear plane angles with cutting speed is due to the variation of flow stress with strain, strain rate and temperature. There are clearly a number of assumptions in the procedures just described: that all the variation in (f + l – a) is due to variation in n; that the parallel-sided shear zone model is adequate (strain rates in practice will vary from the cutting edge to the free surface, as the actual shear zone width varies); and that C really is a constant of the machining process. In later work, Oxley investigated the sensitivity of his modelling to variations of C. A change to C causes a change to the hydrostatic stress gradient along the primary shear plane and hence to the normal contact stress on the tool at the cutting edge, sn,O. Adding the constraint that sn,O derived from the primary shear plane modelling should be the same as that from secondary shear modelling (Section 6.3.3), he concluded – for the same steel for which he had initially given the value C = 5.9, but over a wider range of feed, speed and rake angle cutting conditions – that C might vary between 3.3 and 7.1. The interested reader is referred to Oxley (1989). 6.3.3 Flow in the secondary shear zone With the partial exception of slow speed cutting tests like those of Roth and Oxley (Figure 6.8), visioplasticity studies have never been accurate enough to give information on strain rate and strain distributions in the secondary shear zone on a par with the level of detail revealed in the primary shear zone. Certainly at high cutting speeds, grids or other internal markers necessary for following the flow are completely destroyed. Nor is there any way, equivalent to applying equation (6.13) in the primary zone, of deducing the strain hardening exponent n for flow in the secondary shear zone. So, even if a flow stress could be deduced for material there, the extraction of a s0 value (equation (6.10)) and the estimation of a TMOD value for it might be thought to be impractical.Yet Figure 6.11 contains, in the variation of sint with TMOD, such plastic flow stress information. The insights and assumptions that enabled this data to be presented are worth considering. Oxley explicitly suggested that in the secondary shear zone strain-hardening would be negligible above a strain of 1.0. This allowed him, from equation (6.10) with e—= 1, to identify s0 with s—. It is a major issue in materials’ modelling for machining – and is returned to in Chapter 7.4 – to determine how in fact flow stress does vary with strain at the high strains generated in secondary shear. Oxley then suggested that s— is the same as sint, or √3tav, where tav is the average friction stress over the chip/tool contact area (obtained by dividing the friction force by the measured contact area). This is reasonable, from considerations of the friction conditions in machining (Chapter 2), provided there is a negligible elastic contact region. Oxley argued that this was the case, on the basis of his (Roth and Oxley, 1972) low speed observations, but the observations of Figure 6.5 do not support that. To determine a TMOD value, he estimated representative temperatures and strain rates inthe secondary shear zone. For the strain rate e?—int he supposed the secondary shear zone tohave an average width dt2, and that the chip velocity varied from zero at the rake face toits bulk value Uchip across this width. Then He took the representative temperature to be the average at the rake face, calculated ina manner similar to equation (2.18), but allowing for the variation of work thermal properties with temperature and for the fact that heat generated in secondary shear is not entirely planar but is distributed through the secondary shear zone (Hastings et al., 1980). In the notation of this book, equation (2.18) is modified by a factor c The calculated (sint, TMOD) data in Figure 6.11 result from these assumptions. That they follow the variations expected from independent mechanical testing gives some support to these insights. There is one assumption to which it is particularly interesting to return: that is, that the sliding velocity at the chip/tool interface is zero. This strongly influences both the calculated strain rate and the need for the correction, c, to the temperature calculation. The slip-line field modelling does not support such a severe reduction of chip movement. Figure 6.2, for example, shows sliding velocities reduced to zero only in some circumstances and then only near to the cutting edge. Resolving the conflict between these variable flow stress and slip-line field views of rake face sliding velocities leads to insight into conditions at the rake face during high speed (temperature affected) machining. In his work, Oxley identified two zones of secondary shear, a broader one and a narrower one within it, closest to the rake face. This narrower zone has also been identified by Trent who describes it as the flow-zone and, when it occurs, as a zone in which seizure occurs between the chip and tool (Trent, 1991). Figure 6.12(a) shows Oxley’s measurements of the narrower zone’s thickness, for a range of cutting speeds and feeds, for the example of a 0.2%C steel turned with a –5° rake angle tool (other results, for a 0.38%C steel and a +5° rake tool, could also have been shown). The flow-zone is thinner the larger the cutting speed and the lower the feed. In Figure 6.12(b), the observations are replotted against t(kwork/(Uwork f))?. This is the same as (kworkl/Uchip)?, which occurs in equation (6.16), if it is assumed that the contact length l is equal to the chip thickness t. The experimental results lie within a linear band of mean slope 0.2. The flow-zone lies Fig. 6.12 Variation of flow-zone thickness with (a) cutting speed, at feeds (mm) of 0.5 (?), 0.25 (+) and 0.125 (o); and (b) replotted to compare with theory (see text) Oxley pointed out that the temperature of the flow zone would reduce the thicker it was, through the factor c (equation (6.16)); and that its strain rate would increase the thinner it was (equation (6.15)). These influences of thickness on strain rate and temperature would result in there being a thickness for which the velocity modified temperature would be a maximum, and the shear flow stress a minimum (provided TMOD was above about 620 K for the example in Figure 6.11). He proposed that the thickness would take the value that would maximize TMOD. This gives the band of values labelled ‘Theory’ in Figure 6.12(b). The predicted band lies about 50% above the observed one, sufficiently close to give validity to the proposal. In Chapter 2 (Figure 2.22(a)), direct measurements of the variation of friction factor mwith rake face temperature were presented, for turning a 0.45%C steel. Flow-zone thicknesswas not measured in those tests. However, if the experimental relationship shown in Figure 6.12(b) is assumed to hold, the data of Figure 2.22(a) can be converted to a dependenceof √3mk (or sint) on TMOD. 注：文章來源Metal_Machining。 11