連接片級進模設計【沖孔落料級進?！俊菊f明書+CAD】,沖孔落料級進模,說明書+CAD,連接片級進模設計【沖孔落料級進模】【說明書+CAD】,連接,片級進模,設計,沖孔,落料級進模,說明書,仿單,cad 河南機電高等?？茖W校 學生畢業(yè)設計（論文）中期檢查表 學生姓名 學 號 指導教師 選題情況 課題名稱 連接片級進模設計 難易程度 偏難 適中 √ 偏易 工作量 較大 √ 合理 較小 符合規(guī)范化的要求 任務書 有 √ 無 開題報告 有 √ 無 外文翻譯質量 優(yōu) 良 中 差 學習態(tài)度、出勤情況 好 一般 差 工作進度 快 按計劃進行 慢 中期工作匯報及解答問題情況 優(yōu) 良 中 差 中期成績評定： 所在專業(yè)意見： 負責人： 年 月 日 第7章 沖壓及薄板液壓成形的案例分析： 7.1。案例1：精密反射板的液壓成形 7.1.1。問題陳述 為了空間通信，直徑為12m被稱為平方公里陣列（SKA）的反射鏡/天線陣列被確定了三種方法制造?12米的反射鏡，即： （a） 鋁合金板組裝和一個鋼筋支撐的結構 （b） 復合材料制造形成的反光板 （c） 模具（超高頻- D）內高壓成形的板材 在這些方法中，超高頻- D被認為是最經濟的，并能產生所需的表面光潔度和精度，并且對高壓成形的小尺寸（?50寸）和大到（?12米）反射[Antsos2003]的有限元進行了模擬與檢查。這一分析表明部分液壓成行后的回彈性較大。這導致維護最后反射剖面（圖0.2毫米的均方根（RMS）的耐受性是相當困難的。（7.1）。這是關于確定最佳的模具或模具的幾何形狀，以減少所需內高壓成形?12米的反射回彈得一項調查。 7.1.2。序列形成 目前生產的直徑30英寸較小的反射鏡如圖7.2 第1步：板材由于重力下垂 第2步：板料的頂部和底部之間模具的夾緊 第3步：脹形高端板材模具型腔 7.1.3。目標和方法途徑 這項研究的具體目標是： （a） 預測在反射?12米高壓成形回彈 （b） 確定影響反射器回彈的工具和材料參數 （c） 通過反射的回彈補償估算模具/模具的幾何形狀 下面概述的是任務 任務1：確定使用雙軸片凸起材料AA3003 – O的性能測試 任務2：調查片材各向異性，一個?。?50英寸）內高壓成形反射對回彈和變 薄的影響。 任務3：調查表模摩擦條件下的界面效應在模具板下垂由于重力作用，以及變薄 和內高壓成形反射模具回彈?12米的結構。 任務4：估計?12米反射鏡最佳模具結構，讓沿曲線長度可以很容易地補償回 彈均勻分布 任務5：量化對細化和優(yōu)化模具結構板材厚度的影響 7.1.4。有限元（FE）仿真 使用單元進行了利用商業(yè)有限元代碼PAMSTAMP2000對液壓成形過程模擬圖7.3顯示了在PAMSTAMP有限元模型等軸測視圖2000板殼單元。有限元模擬輸入條件列于表鋁AA3003- O獲得粘性壓力的測試。當剛性空白是仿照彈塑性凸凹模的模板的有限元模擬。由于軸對稱變形和邊界條件只有四分之一的空白部分為模板根據庫侖摩擦定律的接口摩擦系數（μ）= 0.1的假設。對工作表上模夾緊進行了模擬移動，流體（或空氣）壓力是通過使用有限元軟件PAMSTAMP的'Aquadraw'。 7.1.5.內高壓成形材料 各向異性反射板對回彈和變薄分配一個小尺寸（?50寸）有限元進行了兩個為形成?50英寸直徑的反射板的各向異性模擬。初始毛坯直徑= 57.5英寸（?一千四百六十毫米）。系數的各向異性三個方向（0 °，45°和90°）分別輸入到有限元模擬，以評估在形成反射的各向異性對厚度分布和回彈的影響，這些各向異性值選擇上，真正形成條件是可以仿效的。然而，他們不反映??鋁合金3003- O的（圖7.4）的實際值準確的各向異性值鋁合金3003- O的需要從拉伸試驗確定。希爾的標準是1948年的產量用來代表在有限元模擬板材各向異性 圖7.4：沿軋制方向（0°），對角方向（45°）和橫向方向（90°）用有限元模擬塑性的變比 板材各向異性的影響中形成的細化反射分布： 即使形成了各向異性材料特性被用來描述了有限元模擬片材（圖7.5）。間伐比例并沒有明顯改變沿圓周的一部分，對材料性能各向同性和各向異性的有限元模擬說明有限元模擬預計在高壓成形的一部分最大減薄了14％最大減薄觀測到是該地區(qū)靠近頂端的圓頂，但不是在圓頂的頂點。由于壓應力沿對社會形成的板材何處接觸到這些地點的模具表面徑向表行事不變，細化超過150毫米的中心曲線長度。 材料各向異性的回彈 ——觀察更多的回彈有限元各向同性和各向異性模型相比，（圖7.6） ——對于各向異性的情況下，不同的回彈沿圓周的一部分。這種變化是由于回彈在沿軋制屈服應力和橫向的滾動方向的各向異性常數屈服準則中引入了變異 圖7.5：有限元模擬預測各向異性材料的稀疏分布插圖：圖為?50寸長）的 反射曲線變形示意圖 圖7.6：各向異性板材有限元模擬形成?50英寸的反射回彈 7.1.6：?12米反射工藝參數和刀具幾何參數對回彈的影響 高壓成形?12米的反射鏡有限元模擬進行了使用代碼PAMSTAMP2G到研究了工藝參數，即以下效果： （a） 由于模具其自身重量導致下垂 （b）片材和模具界面摩擦條件 （c）模具的幾何形狀（法蘭角） 液壓成形零件的回彈： 圖7.7：為假定模具有限元模擬示意圖 表7.2：刀具幾何形狀在模擬中使用 ?12米天線模具結構設計在圖7.7和表7.2顯示了以目前的模具成型制造中使用更小的反射器為基礎內高壓成形有限元天線進行的模擬三個階段。在第一階段，該表將死腔由于其本身重量的重力下垂進行模擬（重力模擬）。在第二階段，由被認為是（夾緊模具板控股模擬）。第三階段，進行回彈模擬后其次進行內高壓成形過程有限元模擬，表7.3顯示了有限元模擬矩陣，研究在?12米反射回彈（a）界面條件及（b）模具的幾何形狀的影響。三種不同的接口之間的表和法蘭在模具的幾何形狀不同的角度摩擦條件和模具幾何形狀考慮對摩擦條件影響。 表7.3：有限元模擬矩陣來確定影響界面摩擦及模具幾何尺寸對回彈（法蘭角），并在?12m變薄形成反射 研究法蘭回彈影響的角度： 在與該模具的法蘭角的增加形成了部分回彈。在?12米反射器，Z軸方向的最小為10毫米，觀察回彈位移為30度角，而在Z法蘭?16毫米的法蘭60度角位移觀測方向為最大回彈。 回彈法蘭角在30 °形成的部分是非線性分布。這是很難修改模具，以補償非線性回彈。因此，這是傾向于選擇與模具結構法蘭角45度回彈的變化曲線長度，并且可以很容易地補償上模具幾何形狀。 隨著75 °法蘭角，徑向回彈比其他法蘭角較?。?0°，45°和60°）。因此，在模具上的選擇75 °法蘭角，被認為是形成和薄鋼板厚度0.25英寸?12米的最佳反射 圖7.8：比較Z位移在換用不同的角度法蘭液壓成形零件回彈的模具結構（初始板材厚度=0.25英寸） 圖7.9：在換用不同的角度法蘭液壓成形模具結構部分（初始板厚=0.25英寸） 比較回彈的徑向位移。 在最后的反射回彈中工具之間摩擦片的影響: 圖7.10顯示了與法蘭不同的摩擦角為60度的條件反射形成的細化比較，細化在不改變摩擦條件的顯著變化。 在液壓成形，片自由凸起（統(tǒng)一拉伸）到上模腔，并逐漸從外圍向中心接觸到上模，表是夾在邊緣以避免在高壓成形過程中的任何物質運動進入體腔。 因為，不存在負債表和上模之間的相對運動。因此，效果對形成中的一部分變薄摩擦是微不足道的。 圖7.10：預測不同回彈的?12米法蘭角60°（初始板材厚度=0.25英寸）反射界面摩擦條件比較。 7.1.7。摘要和結論 在這項研究中，有限元（FE）板材液壓成形與模具（超高頻- D）的?50英寸和?12米的反射進行了使用商業(yè)有限元程序PAMSTAMP2G/PAMSTAMP2000年的模擬過程： （a） 對制造的超高頻三維?12米反射鏡的可行性論證 （b） 預測在高壓成形過程?12米后反射鏡的回彈 （c） 確定的工藝參數影響/材料特性（即各向異性片材，板材初始厚度，界面摩擦條件下，重力和模具的幾何尺寸對回彈）形成的反射器 （d） (d) 計算在回彈形成的反射可以很容易地通過修改模具的補償的一個幾何尺寸。 這項研究得出的結論是： 形成反射超高頻- D的可行性過程：通過有限元模擬，用模具（超高頻- D）的內高壓成形過程中形成大的反光板使用的可行性證明。有一個表面精度要求0.2毫米的RMS上的內形成反射朗讀，在超高頻- D過程是有厚度的變化而變形， 還有的由于殘留在反射形成的反射回彈應力，因此，為了獲得所需的關于最終形成準確地表反射剖面，對模具/模具幾何體必須進行修改。因此需要在設計上與Z軸方向高硬度模具結構，盡量慎重考慮減少上層偏轉死于高勢力。觀察附近的頂點?12米天它是很好的資產負債表內的材料最高為14％，變薄，線 斷裂極限為AA3003 – O。 影響工具之間的摩擦片：工藝條件，即下垂的空白由于重力作用，模具的幾何形狀（法蘭角）和初始板厚影響了部分形成的細化。然而，在部分變薄分布發(fā)生的變化僅± 2％，有限元模擬關于在液壓成形零件變薄界面摩擦條件的影響被認為是微不足道的。 各向異性對板料回彈的影響：由于AA 3003 – O為非各向異性數據，假設的數據被用來量化回彈各向異性的影響，在板材各向異性導致圓周的一部分非均勻回彈。因此，需要考慮材料的各向異性估計和補償在模具的幾何形狀的回彈，回彈是受a）板材各向異性，B）初始板厚，C）下垂的空白，由于重力和d）模具的幾何形狀的相互影響。 模法蘭角對回彈的影響：入模腔片材下垂導致更多的物質流入模腔，從而減少了減薄形成的部分，增加成型后的回彈。因此，模具的幾何形狀（法蘭角）顯著影響了內高壓成形零件的回彈。高等法蘭角允許在空白以及在初始階段的夾緊下垂更多的物質流進模腔，因此，較大的模具幾何法蘭角導致更多的回彈。 然而，更大的法蘭角導致回彈更加均勻分布，這是可取的。統(tǒng)一回彈可以更容易地通過修改模具幾何尺寸進行補償。其中模具幾何尺寸（法蘭角）在本研究中考慮，75°法蘭角的發(fā)現使得了初步板材厚度6.35毫米（0.25英寸）的材料回彈分布均勻。 板材厚度對回彈的影響：對于較大的法蘭角，入模腔更多的物質流。此外，小薄床單的抗彎曲和更容易流動到模具型腔。太多的材料進入模腔流導致皺紋和折疊。 - 模具結構的“最佳“回彈補償：較低的值幾何模法蘭角（30°，45°和60 °）導致了非均勻回彈這是很難彌補。高等法蘭角導致過度進入模具型腔物質流的重力及控股階段，從而導致皺紋和折疊。因此，“最優(yōu)“模具結構（法蘭角）取決于初始片厚度，負債表和厚度的減小。據估計，在形成?12米反射鏡厚度為4.75毫米（0.185英寸）的初始表的“最佳“死幾何（法蘭角）將在60°?75 °。 顯示對應的拉丁字符的拼音 PROCESS ANALYSIS AND DESIGN IN STAMPING AND SHEET HYDROFORMING DISSERTATION Presented in Partial Fulfillment of the Requirements for the Degree Doctor of Philosophy in the Graduate School of The Ohio State University By Ajay D. Yadav, M.S. * * * * * The Ohio State University 2008 Dissertation Committee: Professor Taylan Altan, Adviser Associate Professor Jerald Brevick Professor Gary L. Kinzel Approved by: - Adviser Industrial and Systems Engineering Graduate Program ii ABSTRACT This thesis presents initial attempts to simulate the sheet hydroforming process using Finite Element (FE) methods. Sheet hydroforming with punch (SHF-P) process offers great potential for low and medium volume production, especially for forming (a) lightweight materials such as Al- and Mg- alloys and (b) thin gage high strength steels (HSS). Sheet hydroforming has found limited applications and is thus still a relatively new forming process. Therefore, there is very little experience-based knowledge of process parameters (namely forming pressure, blank holder tonnage) and tool design in sheet hydroforming. For wide application of this technology, a design methodology to implement a robust SHF-P process needs to be developed. There is a need for a fundamental understanding of the influence of process and tool design variables on hydroformed part quality. This thesis addresses issues unique to sheet hydroforming technology, namely, (a) selection of forming (pot) pressure, (b) excessive sheet bulging and tearing at large forming pressures, and (c) methods to avoid leaking of pressurizing medium during forming. Through process simulation and collaborative efforts with an industrial sponsor, the influence of process and tool design variables on part quality in SHF-P of axisymmetric punch shapes (cylindrical and conical punch) is investigated. In stamping and sheet hydroforming, variation in incoming sheet coil properties is a common problem for stamping plants, especially with (a) newer light weight materials for automotive applications (aluminum-, magnesium- alloys) and (b) thin gage high strength steels. Even though incoming sheet coil may meet tensile test specifications, high scrap rate is often observed in production due to inconsistent material behavior. Thus, tensile test specifications may not be adequate to characterize sheet material behavior in production stamping/hydroforming iii operations. There is a strong need for a discriminating method for testing incoming sheet material formability. The sheet bulge test emulates biaxial deformation conditions commonly seen in production operations. This test is increasingly being applied by the European automotive industry, especially for obtaining reliable sheet material flow stress data that is essential for accurate process simulation. This thesis presents a new inverse-analysis methodology for calculating flow stress curves at room temperature, using the biaxial sheet bulge test. This approach overcomes limitations of previously used closed-form membrane theory equations and exhibits great potential for elevated temperature bulge test application. To verify the developed methodologies presented in this thesis, selected case studies are presented, to (a) demonstrate the successful application of finite element (FE) simulation in tool design, process sequence design and springback reduction in stamping and sheet hydroforming and, (b) validate the developed methodology for automation/standardization of tool and process sequence design procedure and recording of existing design guidelines in transfer die stamping. iv ACKNOWLEDGMENTS I am immensely grateful to my graduate adviser, Dr. Taylan Altan, for his guidance and encouragement during my doctoral studies at the Engineering Research Center for Net Shape Manufacturing (ERC/NSM). This dissertation work would not have been possible if not for his continuous support and enthusiasm for applied research. I sincerely appreciate the suggestions and comments of my candidacy and dissertation committee: Dr. Jerald Brevick, Dr. Gary L. Kinzel and Dr. Jose Castro. I am grateful to sponsors of the Center for Precision Forming (CPF) for partially supporting this research. I am also thankful to ERC/NSMs industrial sponsors for supplying us with an adequate number of metal forming problems. I acknowledge all students and visiting scholars of ERC/NSM and CPF (years 2001 - 2008) for their contributions. Special thanks to Dr. Hariharasudhan Palaniswamy, Mr. Lars Penter (Dresden), Mr. Parth Pathak, Mr. Dario Braga (Brescia) and Mr. Gianvito Gulisano for their co- operative efforts. Thanks are also due to Dr. Manas Shirgaokar, Dr. Ibrahim Al-Zkeri, Mr. Thomas Yelich, Mr. Giovanni Spampinato, Dr. Yingyot Aue-u-lan, Dr. Hyunjoong Cho, Dr. Hyunok Kim, Mr. Shrinidhi Chandrasekharan and Dr. Gracious Ngaile (during my Masters) for moral and intellectual support. I wish to thank all my friends for their encouragement, advice and support. Finally, I dedicate my graduate study efforts to my parents for their everlasting support, patience and personal sacrifice. v VITA Oct 7, 1979 Born Pune, India. Jul 1997 Jun 2001 B.E. Mechanical Engineering, All India Shri Shivaji Memorial Societys College of Engineering (AISSMS COE), University of Pune, Maharashtra, India. Aug 2001 May 2003 M.S. Industrial and Systems Engineering, The Ohio State University, Columbus, Ohio, USA. Aug 2001 Jun 2008 Graduate Research Associate, NSF I/UCRCs Center for Precision Forming (CPF) and Engineering Research Center for Net Shape Manufacturing (ERC/NSM). vi PUBLICATIONS Yadav A. D., Palaniswamy H., and Altan, T., (2008), Sheet Hydroforming: Room and Elevated Temperature Chapter for ASM Sheet Metal Forming, Editor: T. Altan (in progress). Yadav, A.D., Pathak P., Altan, T., (2008), Progressive and Transfer Die Stamping Chapter for ASM Sheet Metal Forming, Editor: T. Altan (in progress). Yadav, A. D., Kaya S., Altan, T., (2008), Servo Drive Presses Stamping Technology Chapter for ASM Sheet Metal Forming book, Editor: T. Altan (in progress). Yadav A. D., Penter, L., Pathak, P., Altan, T., (2008/2009), Flow stress determination with biaxial sheet bulge test using inverse analysis approach (Manuscript in preparation) for peer review journal. Yadav A. D., Palaniswamy H., and Altan, T., (February, March, April 2006), Three-part series on “Sheet Hydroforming”, Stamping Journal R Initial sample thickness for ERCs VPB bulge test = 1.3 mm Mirtsch 2006 . 65 xvi Figure 4.9: Chart showing the comparison of the thickness measured at dome apex with predictions from Excel macro. Calculated values of thickness are 4% lower than measured thickness values. . 66 Figure 4.10: Schematic shows springback error (exaggerated) in dome height measurement . 67 Figure 4.11: Comparison of flow stress curves, with and without springback correction . 69 Figure 5.1: Inverse analysis methodology to determine flow stress curve (for materials following the power law fit) using the dome geometry evolution from the biaxial bulge test. . 75 Figure 5.2: Two dimensional schematic of FE model in LS-DYNA v9.71 . 76 Figure 5.3: Schematic illustration of the objective function. (a) Changing dome height for five time increments (t = 1 through 5). Comparison of measured and simulated dome height at final time t = 5. . 78 Figure 5.4: (a) 2D view of design space describing the two variables in flow stress determination problem Design variable 1: K (in MPa), design variable 2: n- value (b) 3D view of design space, showing objective function (response) obtained for each K and n-value from simulation. The objective of this study is to find the minimum on the design surface created from response values. . 79 Figure 5.5: Schematic of a full factorial set of FE simulations for a given input range for K and n-value. For K = 400 to 500, a set of 11 values is chosen (namely, 400, 410, 420 500). Similarly, for n-value = 0.1 to 0.5, a set of 11 points is chosen. The number of points is user defined. Thus, for a full factorial set of simulations, 11 x xvii 11 = 121 simulations are planned by LS-OPT. This is computationally expensive. . 80 Figure 5.6: (a) Selective FE simulations using a D-optimal design (b) An example response polynomial surface (second order) obtained from the first 10 FE simulations conducted using the D-optimal design. . 82 Figure 5.7: Minimization of objective function E using RSM. In each iteration (set of 10 FE simulations), the margins for K and n-value shrink to a smaller design space, until objective function is minimized. . 83 Figure 5.8: Results for bulge tests for AA 5754-O (t = 1.01 mm) conducted at industrial sponsor: Convergence history of strength coefficient K at end of eight iterations, K converged to 432 MPa (b) Convergence history of the strain hardening exponent n-value. At the end of eight iterations, n-value converged to 0.3. . 84 Figure 5.9: Flow stress results using inverse analysis methodology at room temperature: Aluminum alloy AA5754-O, sheet thickness = 1.01 mm. Biaxial hydraulic bulge tests were conducted at industrial sponsor (seen in Figure 4.3). . 85 Figure 5.10: Flow stress results using inverse analysis methodology at room temperature: Aluminum alloy AA5754-O, sheet thickness = 1.3 mm. Biaxial bulge experiments were conducted at ERC/NSM using the VPB test tool geometry. . 86 Figure 5.11: Flow stress results using inverse analysis methodology at room temperature: Stainless steel SS 304, sheet thickness = 0.86 mm. Biaxial bulge experiments were conducted at ERC/NSM using the VPB test tool geometry. . 87 xviii Figure 5.12: (a) Elevated temperature hydraulic bulge test set up Erlangen, Germany (b) 3D and 2D view of a deformed Mg-alloy AZ31-O dome (measured by CCD camera) Hecht et al. 2005. . 88 Figure 5.13: Elevated temperature hydraulic bulge tests at Virginia Commonwealth University, USA Koc et al. 2007 . 90 Figure 5.14: (a) Schematic of the elevated temperature bulge test tools at University of Darmstadt, Germany (die opening = 115 mm, die corner radius = 4 mm) (b) Sample Mg-alloy specimen at 25C and 225C Kaya et al. 2008. . 91 Figure 5.15: Deviation of the dome surface from spherical shape (at elevated temperature) Kaya et al. 2008. . 92 Figure 5.16: (a) Definition of objective function (b) Proposed inverse analysis methodology for elevated temperature bulge test, to determine flow stress curve for materials following the power law fit = K () n () m using dome geometry evolution. . 96 Figure 5.17: Two dimensional schematic of the axisymmetric of the biaxial sheet bulge test for elevated temperature flow stress determination. Die diameter = 4 inches (101.8 mm), die corner radius = 0.25 inch (6.35 mm) . 98 Figure 5.18: Methodology for selecting lockbead dimensions for elevated temperature bulge test tools. . 99 Figure 5.19: (a) 2D schematic of the designed axisymmetric elevated temperature bulge test tool. (b) Effective strains in the clamped sheet (within lockbeads) are lesser than uniform elongation value for AA5754-O material. Thus, design is safe. . 101 xix Figure 5.20: Locating the three LVDTs for measuring the dome curvature. One LVDT will measure the dome apex, while the other two are designed for flexibility in measuring location. . 102 Figure 6.1: Schematic shows top view of example non-axisymmetric (rectangular) part. Due to asymmetry, the state of stress in the deforming sheet metal is different along the perimeter. . 107 Figure 6.2: Pot pressure P a is not sufficient to lift the sheet metal off the die corner at Section 1. Thus, insufficient pot pressure causes rubbing of sheet with the die corner leading to excessive thinning. . 107 Figure 6.3: Sheet metal flow in sections 1 and 2 (from Figure 1) for the pot pressure of P b (Note: P b P a ). An increased pot pressure P b ensures that sheet metal does not rub at the die corner at Section 1. However, this increased pot pressure P b causes the sheet metal to “bulge against drawing direction” at Section 2. . 108 Figure 6.4: Process limits on pot pressure are limited by sheet bulging in straight sections (for high pot pressure) and sheet rubbing at die corner (for low pot pressure). 109 Figure 6.5: Schematic of the 90 mm cylindrical cup tooling at Schnupp Hydraulik. 110 Figure 6.6: Photograph of the 800 kN (80 ton) hydraulic press used for experiments at Schnupp Hydraulik, Germany . 111 Figure 6.7: (a) Pot pressure and blank holder force (BHF) curve estimated by Schnupp Hydraulik through trial and error experiments. (b) Picture of 90 mm cup hydroformed cup using the estimated process parameters. . 113 Figure 6.8: (a) The optimum BHF curve was estimated using numerical techniques coupled with FE simulation. The “ERC/NSM pot pressure curve” was estimated by xx Contri, et al. 2004 through trial-and-error FE simulation (b) Part formed successfully, using ERC/NSM pot pressure curve and optimum BHF curve estimated by Braedel, et al. 2005 . 114 Figure 6.9: FE-Model (quarter geometry) in PAMSTAMP 2000 (AQUADRAW) for the 90 mm cylindrical cup SHF-P process. . 116 Figure 6.10: Flow stress of St1403 sheet material of thickness 1 mm estimated by Viscous Pressure Bulge (VPB) test at ERC/NSM. Uniaxial tensile tests were conducted to obtain the anisotropy coefficients (r 0 , r 45 and r 90 ). . 117 Figure 6.11: In investigating influence of punch-die clearance in SHF-P of cylindrical cups, the previously estimated optimum pot pressure curve (Max. value 400 bar) was too high. Thus, a reduced pot pressure curve (Max. value 180 bar) was used. . 119 Figure 6.12: Optimum blank holder force (BHF) curves for experiments 1 and 2 for pot pressure 180 bar (see Figure 6.11), for punch corner radius = 5 mm. . 120 Figure 6.13: Optimum blank holder force (BHF) curves for experiments 3 and 4 for pot pressure 180 bar (see Figure 6.11), for punch corner radius = 10 mm. . 120 Figure 6.14: Formed cylindrical cups for the four experimental cases. . 121 Figure 6.15: Comparison of (a) material draw-in and (b) thinning distribution (along rolling direction of sheet) between FE simulation and experiment, for Experiment 1. . 122 Figure 6.16: Comparison of (a) material draw-in and (b) thinning distribution (along rolling direction of sheet) between FE simulation and experiment, for Experiment no. 2. . 123 Figure 6.17: Comparison of (a) material draw-in and (b) thinning distribution (along sheet rolling direction) between FE simulation and experiment, for Experiment 3. . 124 xxi Figure 6.18: Comparison of (a) material draw-in and (b) thinning distribution (along rolling direction) between FE simulation and experiment, for Experiment 4. . 125 Figure 6.19: Sidewall wrinkles in the part were observed for punch-die clearance 5.5 mm (a) Formed part from Experiment no. 2 (b) Formed part from Experiment no. 4. . 128 Figure 6.20: Pot pressure curve (180 bar) was insufficient to form the cup completely against the punch. Therefore, pot pressure curve was modified by Schnupp Hydraulik and experiments were repeated. . 129 Figure 6.21: Blank holder force (BHF) curves by trial-and-error, for experiments 1 and 2 (punch corner radius = 5 mm) using Schnupp pot pressure curve (Figure 6.20). . 129 Figure 6.22: BHF curves obtained by trial-and-error, for experiment nos. 3 and 4 (punch corner radius = 10 mm) using Schnupp pot pressure curve (Figure 6.20). . 130 Figure 6.23: Four cups were formed with the modified pot pressure curve (one cup for each experimental case). . 130 Figure 6.24: Thinning distribution comparison (along sheet rolling direction) for Experiments. 3 and 4. . 131 Figure 6.25: Schematic shows three punch stroke locations using conical punch. Punch-die clearance changes with punch stroke. . 134 Figure 6.26: (a) Conical punch geometry selected for the study (b) Tool and blank dimensions for conical punch simulations/experiments . 135 Figure 6.27: FE model (quarter geometry) and input parameters used in FE simulations in PAMSTAMP 2000 (AQUADRAW) for conical cup SHF-P process. . 137 Figure 6.28: Pot pressure curves for conical punch FE simulations. . 138 xxii Figure 6.29: Sheet bulging observed for the selected conical punch geometry, at punch stroke 3.5 mm for Simulation no. 1. This bulge was observed for pot pressure = 20 bar and punch-die clearance 28 mm. . 139 Figure 6.30: Data points indicating bulging/no bulging for the planned FE simulation matrix. With increasing punch stroke (reducing punch-die clearance) higher values of pot pressure are needed to bulge the sheet against drawing direction. . 141 Figure 6.31: State of stress in example rectangular (non-symmetric) part geometry. . 143 Figure 6.32: Example non-symmetric punch geometry. This punch geometry will be used for (a) prediction of optimum pot pressure and BHF curve, (b) an investigation on the effect of punch-die clearance on sheet deformation. . 144 Figure 7.1: (a) Large reflector (b) Schematic of the reflector geometry (2D section view) . 146 Figure 7.2: 2D schematic of sequence of SHF-D forming operation for large reflectors. . 147 Figure 7.3: 3D finite element model (quarter geometry) in PAMSTAMP. . 149 Figure 7.4: Plastic strain ratio along rolling direction (0), diagonal direction (45) and transverse direction (90) used in FE simulation . 150 Figure 7.5: Thinning distribution predicted by FE simulation for anisotropic material (Inset: Schematic of the curvilinear length of deformed 50 inch reflector) . 152 Figure 7.6: Springback in the formed 50 inch reflector predicted by FE simulation for anisotropic sheet material (r 0 = 1.4, r 45 = 1.6 and r 90 =1.2) . 152 Figure 7.7: Schematic of the assumed tooling for FE simulations . 153 Figure 7.8: Comparison of Z displacement due to springback in the hydroformed part for the die geometry with different flange angle (for initial sheet thickness = 0.25 in) . 156 xxiii Figure 7.9: Comparison of radial displacement due to springback in the hydroformed part for the die geometry with different flange angle (for initial sheet thickness = 0.25 in) . 156 Figure 7.10: Comparison of springback predictions in 12 m reflector for different interface friction conditions, for flange angle 60 (for initial sheet thickness = 0.25 in). . 157 Figure 7.11: Cross sectional view of a typical formed axisymmetric part. . 162 Figure 7.12: Methodology for automating the process sequence and tool design. . 164 Figure 7.13: Steps followed in tool and process sequence design . 165 Figure 7.14: User interface to accept final part dimensions (in inches). . 167 Figure 7.15: Surface area calculation of the final part. Source: Progressive Dies Certificate program, Society of Manufacturing Engineers, at Dearborn, MI 2004. . 168 Figure 7.16: Deliverable of developed methodology: Process sequence for an axisymmetric part (with 4 drawing stages). . 170 Figure 7.17: Deliverable of developed methodology: Tool design and drawings, generated using the developed Visual Basic methodology. . 172 Figure 7.18: (a) Stainless steel case for a flow measuring device. Sketches of conceptualized process sequences to form possible final part geometries, namely (b) donut- shaped part (c) half donut-shaped part. The donut-shaped part was selected for process/tool design using commercial finite element (FE) code. . 174 Figure 7.19: Material properties (flow stress data) for SS304 and SS304L Source: Kalpakjian, S., Manufacturing Processes for Engineering Materials, Addison Wesley 1991 . 175 xxiv Figure 7.20: Finite element (FE) model. (a) Two dimensional view of the 1st stage forming process, inner curvature forming. (b) Two dimensional view of the 2nd stage forming process outer curvature forming. . 176 Figure 7.21: (a) Quarter geometry view showing thinning distribution in formed part, maximum value 30% (b) Thinning distribution along curvilinear length of the part. . 178 Figure 7.22: Quarter geometry view showing “springback” in the formed part. Maximum springback predicted is 2.5 mm (0.1 inches) (a) Back view (b) Front view . 178 Figure