夾子沖壓件設(shè)計【落料沖孔-彎曲壓筋復(fù)合沖壓模具設(shè)計】【說明書+CAD】

夾子沖壓件設(shè)計【落料沖孔-彎曲壓筋復(fù)合沖壓模具設(shè)計】【說明書+CAD】,落料沖孔-彎曲壓筋,復(fù)合沖壓模具設(shè)計,說明書+CAD,夾子沖壓件設(shè)計【落料沖孔-彎曲壓筋,復(fù)合沖壓模具設(shè)計】【說明書+CAD】,夾子,沖壓,設(shè)計,沖孔,彎曲,曲折,復(fù)合,模具設(shè)計,說明書,仿單,cad 常州紡織服裝職業(yè)技術(shù)學(xué)院 畢業(yè)設(shè)計（論文）說明書 設(shè)計（論文）題目夾子沖壓件設(shè)計 系 專 業(yè) 班 級 姓 名 學(xué) 號 指 導(dǎo) 老 師 2008年5月31日 目錄 第一章、課題簡介3 第二章、工藝分析4 一、零件工藝分析4 二、工藝方案的確定4 三、工藝參數(shù)的確定5 第三章、工作力的計算及壓力機的選擇10 一、沖壓力的計算10 二、粗選壓力機12 三、機床壓力中心12 第四章、填寫沖壓工序卡14 第五章、模具結(jié)構(gòu)設(shè)計15 一、模具結(jié)構(gòu)形式的選擇15 二、模具結(jié)構(gòu)的分析與說明15 三、模具工作部分的尺寸和公差的確定16 四、模具結(jié)構(gòu)設(shè)計20 五、校核壓力機安裝尺寸23 第六章、 彎曲模具的設(shè)計24 一、制件彎曲工藝分析24 二、沖壓工藝參數(shù)的確定25 第七章、彎曲模的結(jié)構(gòu)設(shè)計27 一、模具結(jié)構(gòu)的分析說明27 二、彎曲模的卸料裝置的設(shè)計說明28 第八章、彎曲模的工作尺寸計算29 畢業(yè)設(shè)計小結(jié)33 參考文獻35 第一章 課題簡介 零件分析說明 1 零件形狀及其一般要求 制件如圖1-1所示，材料為不銹鋼，材料厚度為0.5mm，制件尺寸精度 按圖紙要求，未注按IT12級，生產(chǎn)綱領(lǐng)年產(chǎn)10萬件。 圖1-1 第二章 工藝分析 1、 零件工藝分析 本制件形狀簡單、尺寸、厚度適中，一般批量生產(chǎn)，屬于普通沖 壓件，但在設(shè)計冷沖壓模具時要注意以下幾點： 2 制件的外形輪廓、結(jié)構(gòu)都算簡單，但是要考慮幾個孔的加工 A、 兩個2的孔的位置要求，B、 由于要裝配，C、 兩孔必 須有一定的同D、 軸度要求，E、 其值為0.15mm。 3 此制件的加工難點主要在孔2的中心距的定位。 4 由于幾個孔的直徑都較小，并且有一定的批量，在設(shè)計時要重視 模具的材料和結(jié)構(gòu)的選擇，保證一定的模具壽命。 二、工藝方案的確定 根據(jù)制件的工藝的分析，其基本工序有落料、沖孔、彎曲三種。按其 先后順序組合以及合理的加工方案有以下幾種： 1、落料-沖孔-彎曲，單工序沖壓。 2、落料-彎曲-沖孔，單工序沖壓。 3、落料沖孔-彎曲壓筋，復(fù)合沖壓。 方案1)為單工序沖壓模具。由于此制件有一定的生產(chǎn)批量，過多的 工序，降低產(chǎn)品的精度，而且此方案生產(chǎn)效率底，不宜批量生產(chǎn)，故 不宜采用此方案。 方案2)也為單工序沖壓模具。它除有方案1的毛病外，還有孔的位 置精度難以保證，在并且在彎曲時也缺少定位精度難保證，故不宜采 用此方案。 方案3)復(fù)合沖壓模具。由于制件的結(jié)構(gòu)，材料的厚度較薄，沖孔與 落料一次沖壓完成。故最宜采用此方案 具體方案示意如下： A、零件的排樣方案圖2-1 圖2-1 三、工藝參數(shù)的確定 1 毛坯尺寸的計算 外形尺寸的長度計算 零件相對彎曲半徑為： R/t=2/0.5=40.5 式中 R彎曲半徑； T料厚。 可見，制件屬于圓角半徑較大的彎曲件，應(yīng)現(xiàn)求彎曲變形區(qū)的 中性層曲率半徑。 由課本p145中性層位置計算公式 =R+Xt 式中 X由實驗測定的應(yīng)變中性層位移系數(shù) 由課本p145 表4-5查出X取0.42 所以： =R+Xt =2+0.420.5 =2.21mm 圓角半徑較大（R0.5t）的彎曲件毛坯長度計算公式 L=L直+L彎 L彎=（180-a）/180* ABCDEFGHIJKL 圖2-2 所以制件長度為如圖2-2 L=LAB+LBC+LCD+LDE+LEF+LFG+LGH+LHI+LIJ+LJK+LKL L=3+3.0342+8.54+14+1+1+1.8028+13+1.8028+1+2 L=50.1798 DE段屬于工藝設(shè)置,目的是為了減少彎曲回彈,其半徑很大,可看成是直 線，所以DE段長度按直線計算。 1 外形輪廓寬度的計算 : 由于考考慮到板料的利用率和排樣的方便，此制件由CAD生成的 工藝尺寸為：外形輪廓寬度為L=53.74mm 2 排樣尺寸的計算 3 搭邊值的確定： 由課本上P46表3-14查得L50的工件間a1的值為2.2*0.8=1.76 側(cè)邊a的值為2.5*0.8=2 3 條料寬度的計算 在設(shè)計模具是為了方便，采用無側(cè)壓裝置送料方式條料寬度計算 公式如下： B=(D+2a +Z) 0- 式中 B為條料寬度的基本尺寸； D為條料寬度方向零件輪廓的最大尺寸 a側(cè)面搭邊 Z導(dǎo)料板與最寬條料之間的間隙 條料寬度的負向偏差 搭邊距a如上所示 間隙、剪切公差Z查課本p47表3-17查得=0.5mm、Z=0.5mm 由上面公式計算得B=(53.74+2*2+0.5 0-0.5=56.240-0.5mm 2 步距的計算 由制件的展開圖（見右圖2-3） 所生成的橫向有效尺寸為Ls=53.74mm 圖2-3 步間距計算公式為： L=Ls+a1 式中L為步間距 Ls為橫向有效尺寸 a1搭邊距 L=53.74+2=55.74mm 由此可得模具排樣圖如下： 圖 2-4 根據(jù)要求查模具設(shè)計指導(dǎo)史鐵梁主編 表4-1，選板的規(guī)格 15008000.5每塊板可剪150056規(guī)格條料14塊，材料利用率達80% 以上。 3 計算材料利用率 由課本p43式子3-19 =nA/BL100% 為材料利用率； A一個步距內(nèi)沖裁件的實際面積； B條料的寬度 L條料的長度 n一張板料上沖裁件的總數(shù)目 =28*1932.765/（56*1500 ）=65% 第三章 工作力的計算及壓力機的選擇 工作力的計算以落料沖孔模具為例計算： 一、沖壓力的計算 完成本制件所需的沖壓力由沖裁力、彎曲力、及卸料力、推料力、 頂料力和壓料力組成 1. 沖裁力F 沖 的計算 由本課本p49式子3-28 F沖 =KtL 式中： 為材料的抗剪強度（MPa） F沖 沖裁力（N） L沖裁周邊總邊長（mm） t材料厚度（mm） 說明：系數(shù)K是考慮到?jīng)_裁刃口的磨損、凸模與凹模的波動（數(shù)值 的變化或分布不均） 、潤滑情況、材料力學(xué)性能與厚度公差的變化等因 素而設(shè)置的安全系數(shù)，一般取1.3。當(dāng)查不到抗剪強度時，可用抗拉 強度 b代替，而取Kp=1的近似計算法計算。 查模具設(shè)計指導(dǎo)p90表4-12取=350 MPa 所以沖裁力F 沖 =1.30.5211.059350=48016 N =48.016KN 2. 卸料力F x 、 推料力FT 、 頂料力FD的計算 在實際生產(chǎn)中影響卸料力、推料力、頂料力的因素很多，要精確計算 很困難。在實際生產(chǎn)中常采用經(jīng)驗公式計算：（查課本p50公式3-30 3-31 3-32） 卸料力： F x=KxF 推料力： FT = nKTF 頂料力： FD =KDF 式中： F沖裁力（N） Kx卸料力系數(shù)，其值為0.045-0.055(薄料取大值、厚料取 小值） ； KT推件力系數(shù)，其值為0.063 KD頂件力系數(shù)，其值為0.08 n時卡在凹模內(nèi)的沖裁件數(shù)(或廢料)數(shù) n=h/t h凹模洞口的直刃壁高度 t板料厚度 卸料力和頂料力是設(shè)計卸料裝置和彈頂裝置的依據(jù)。 因此： Fx =0.04548.016=2.17 KN FT=0.06348.016=3.025008 KN FD=0.0848.016=3.84128 KN 總壓力F總的計算 F總= F 沖 F xFD=54.02728 KN 二、粗選壓力機 由于該制件是一普通制件，且精度要求不高，因此選用開式可傾 壓力機。它具有工作臺三面敞開，操作方便，成本低廉的優(yōu)點。由于 沖孔落料復(fù)合模的壓力行程的特點是在開始階段即 需要很大的壓力， 而在后面階段所需要的反倒要小的多。因此若按總的壓力來選取壓力 機，很可能出現(xiàn)雖然總的壓力滿足要求，但是在初始階段沖裁時已經(jīng) 超載。同時，選用拉深壓力機還應(yīng)該對沖裁功進行核算，否則會出現(xiàn) 壓力機在力的大小滿足要求，但是功率有可能過載，飛輪轉(zhuǎn)速降低， 從而引起電動機轉(zhuǎn)速降低過大，損壞電動機。因此精確確定壓力機壓 力應(yīng)當(dāng)根據(jù)壓力機說明書中給出的允許工作負荷曲線，并校核功率。 但是在一般條件下，可以根據(jù)生產(chǎn)車間的實際條件，在現(xiàn)有壓力機中 選取。在這里根據(jù)總壓力為54.02728KN,從模具設(shè)計指導(dǎo)史鐵梁主 編一書中表4-33提供的壓力機公稱壓力序列中選取100KN的壓力機，型 號為J23-10.由此可知，電動機的功率遠遠大于拉深所需壓力機的電動 機功率。故可以選用此電動機。 三、計算壓力中心 由于制件圖形規(guī)則，上下對稱，所以其壓力中心必在中心線上，用 實驗法測的該制件的壓力中心坐標為X24，Y0如下圖： 第四章、填寫沖壓工序卡 由上可知該制件共有兩道工序：即落料沖孔，彎曲壓筋。 冷沖壓工藝卡片如下圖 產(chǎn)品型號 零（部）件 名稱 共 頁常州紡織 服裝職業(yè) 技術(shù)學(xué)院 沖壓工藝卡片 產(chǎn)品名稱 夾子 零（部）件 型號 第 頁 材料牌號及規(guī)格 材料技術(shù)要求 毛坯尺寸 每毛坯可制件數(shù) 毛坯質(zhì)量 輔助材料 1CrB 條料1500*800 29 工序 號 工序 名稱 工序內(nèi)容 加工簡圖 設(shè)備 工藝設(shè)備 工時 0 下料 剪床上撬板 1500*800 11-6 2500 1 落料 沖孔 落料與沖孔 J23-250 落料沖孔 復(fù)合模 2 彎曲 壓筋 彎曲制件并 壓筋 J23-250 彎曲模 3 檢查 4 5 6 7 8 9 第五章、模具結(jié)構(gòu)設(shè)計 根據(jù)確定的沖壓工藝方案和制件的形狀、特點、要求等因素確定沖 模的類型及結(jié)構(gòu)形式。 1 模具結(jié)構(gòu)形式的選擇 在模具設(shè)計中雖然單工序模具比較簡單也比較容易制造,但是制件 孔離制件邊緣尺寸較小在落料后沖孔勢必會影響模具的精度。且， 兩孔在單工序模具中很難保證兩孔的位置精度。所以考慮到用落料 沖孔復(fù)合模具來加工第一工序。又因為沖孔在前，落料在后，以凸 模插入材料和凹模內(nèi)進行落料，必然材料的切向流動的壓力，有可 能使4的凸模變形，因此考慮采用彈壓卸料裝置的復(fù)合沖壓。這 樣既提高了工作效率又提高了模具的壽命，這樣一來提高了模具的 使用價值。因為制件精度不是多高，采用兩副模具，一副是沖孔落 料復(fù)合模、一副是彎曲模。這樣就降低了模具的制造難度，且適合 生產(chǎn)條件不是很好的企業(yè)生產(chǎn)模具，給模具生產(chǎn)帶來一定的廣度。 給生產(chǎn)降低了成本，帶來了更大的經(jīng)濟利益。 二、模具結(jié)構(gòu)的分析與說明 1、沖孔落料模結(jié)構(gòu)的分析與說明 本道模具主要用來完成落料沖孔，目的明確簡單，看似易設(shè)計。但 對于本制件來說它是一個上下對稱的制件并且有兩個4的小孔， 為了保證兩孔在彎曲后的精度，所以要在本工序中做到兩孔的精度。 在本制件考慮到形狀有一定的復(fù)雜，且較?。?.5mm） ，為了保證制 件有較高的平直度本，故采用正裝式復(fù)合模。制件制造時還要考慮 到制件的定位，根據(jù)制件的特征，為了保證凹模的強度，故采用鉤 形當(dāng)料銷、和采用外形定位的定位銷。 2、彎曲模結(jié)構(gòu)的分析與說明 為了保證坯料在彎曲時不發(fā)生偏移，在設(shè)計時用5的孔為定位孔， 用定料銷定位。為了防止坯料轉(zhuǎn)動，采用左右定位銷定位。 三、模具工作部分的尺寸和公差的確定 沖裁件的尺寸精度主要決定于模具刃口尺寸精度，模具的合 理間隙值也主要靠刃口尺寸及制造精度來保證。正確決定模具刃口 尺寸及制造公差，是設(shè)計沖裁模具的主要任務(wù)之一。 在決定模具刃口尺寸及其制造公差時需考慮下列原則： 8 落料件尺寸由凹模尺寸決定。故設(shè)計落料精度時，以凹模為 基準，間隙取在凸模上；設(shè)計沖孔模時，以凸模為基準，間 隙在凹模上。 9 考慮到?jīng)_裁中凸、凹模的磨損，設(shè)計時，凹?；境叽鐟?yīng)取 制件尺寸公差范圍的較小尺寸；設(shè)計沖孔時，凸?；境叽?則應(yīng)取工件孔尺寸公差范圍的較大尺寸。這樣，在凸、凹磨 損到一定的程度的情況下，仍能沖出合格的制件。凸凹模間 隙則取最小合理間隙值。 10 確定沖模刃口制造公差時，應(yīng)考慮制件的公差要求。如果對 刃口尺寸精度要求過高（即制造公差過?。?，會使模具制造 困難，增加成本，延長生產(chǎn)周期；如果對刃口尺寸精度要求 過低（即制造精度公差過大） ，則生產(chǎn)出來的制件可能不合 格，會使模具的壽命降低。若制件沒有標注公差，則對于非 圓形件按國家標準“非配合尺寸的公差數(shù)值” IT14級處理，沖 模則可按IT11級制造；對于圓形件，一般可按IT7IT6級制 造模具。沖壓件的尺寸公差應(yīng)按“入體”原則標注為單向公差， 落料件上偏差為零，下偏差為負；沖孔件上偏差為正，下偏 差為零。 1） 、 計算落料凸、凹模的刃口尺寸： 本制件為簡單的軸對稱圖形，故按配作法計算凸、凹模刃口尺寸。 根據(jù)凸、凹模刃口尺寸計算公式，先計算出落料凹模刃口尺寸 課本p35式子3-3,3-4（落料） A0maxA)( D mina 0inT TZ Zp DA落料凹模的基本尺寸mm DT落料凸模的基本尺寸mm Dmax落料件最大極限尺寸mm Zmin凸凹模最小初始雙面間隙mm T凸模制造下偏差，可查表3-6,或取T0.4(ZmaxZmin) A凹模制造上偏差，可查表3-6,或取A0.6(ZmaxZmin) 沖裁件制造公差 x系數(shù)，為了避免沖裁件尺寸偏向極限尺寸(落料時偏向最 小尺寸，沖孔時偏向最大尺寸),X值在0.51之間,與工件精度有關(guān)。 可查課本p36表3-7或按下列關(guān)系取值： 當(dāng)制件公差為IT10以上時，取x=1； 當(dāng)制件公差為IT11IT13時，取x=0.75； 當(dāng)制件公差為IT14以下時，取x=0.5。 由課本p33表3-4 3-6 3-7得 Zmin=0.035mm,Zmax=0.045mm, T=0.02mm ， A=0.03mm X=0.75mm 則Z max-Zmin=(0.045-0.035)=0.01mm T A=(0.02+0.03) =0.05mm0.01mm 說明所取凸凹模公差不滿足T AZ max-Zmin條 件,此時可調(diào)整如下: T=0.4(Zmax-Zmin)=0.40.01=0.004 A =0.6(Zmax-Zmin)=0.60.01=0.006 校核： 0.004+0.006=0.01mm 由此可得該尺寸能保證間隙在合理的范圍內(nèi)，故可取 故有： =(50-0.750.1) A0maxA)( D =49.925 mm 06. =(49.925-0.035)minmax 0inAT TZ Zp = 49.89 mm 04. 由課本p36式子3-5 3-6（沖孔） 0minT)(pxd AZ0minin0iAAZ DA沖孔凹模的基本尺寸mm DT沖孔凸模的基本尺寸mm dmin沖孔件最小極限尺寸mm 沖裁件制造的公差mm Zmin凸凹模最小初始雙面間隙mm T凸模下偏差，可查表3-6,或取T0.4(ZmaxZmin) A凹模上偏差，可查表3-6,或取T0.6(ZmaxZmin) x系數(shù)，為了避免沖裁件尺寸偏向極限尺寸(落料時偏向最 小尺寸，沖孔時偏向最大尺寸),X值在0.51之間,與工件精度有關(guān)。 可查課本p36表3-7或按下列關(guān)系取值： 當(dāng)制件公差為IT10以上時，取x=1； 當(dāng)制件公差為IT11IT13時，取x=0.75； 當(dāng)制件公差為IT14以下時，取x=0.5。 計算孔4的凸、凹模尺寸 查課本p33表3-4得 Zmin=0.035mm,Zmax=0.045mm, T=0.02mm ， A=0.02mm X=0.75mm 則Z max-Zmin=(0.045-0.035)=0.01mm T A=(0.02+0.02) =0.04mm0.01mm 說明所取凸凹模公差不滿足T AZ max-Zmin條 件,此時可調(diào)整如下: T=0.4(Zmax-Zmin)=0.40.01=0.004 A =0.6(Zmax-Zmin)=0.60.01=0.006 校核： 0.004+0.006=0.01mm 由此可得該尺寸能保證間隙在合理的范圍內(nèi)，故可取 故 :d T=4+0.750.12= 04.9 DA=4.09+0.035= 6125 四、模具結(jié)構(gòu)設(shè)計 1、凹模周界尺寸計算: 因制件形狀簡單,尺寸不大,又是對稱零件?？紤]到為了便于加工，故 選用整體式凹模比較合理. 1 凹模厚度尺寸H的計算:由凹模的計算公式為 = =16mm 3沖FH016.48 又因為沖裁輪廓線全長為211.059mm,超過了50mm,故應(yīng)乘以修正系數(shù)K. 由模具設(shè)計指導(dǎo)表4-21可得凹模厚度的修正系數(shù)K的值為k=1.37 則 H凹=1.37*16=21.92mm 將凹模厚度圓整成22mm 2 落料凹模周界尺寸LB的計算:因為凹?？卓谳喞獮橹本€和圓弧形組 成所以： BS（2.54）H53.74（5588）109142mm LS12S253.74228109.74110mm 2、選擇模架及確定其他沖模零件的有關(guān)標準:根據(jù)凹模周界尺寸 B109142mm、L110mm, 查模具設(shè)計指導(dǎo)史鐵梁主編表5-8選 取典型結(jié)構(gòu)并結(jié)合實際125125150190I（GB/T2851.31990）,并 選用滑動導(dǎo)向后側(cè)導(dǎo)柱模架. 3、落料凸模的強度和剛度校核 凸模承載能力的校核： 凸模最小斷面承受的壓應(yīng)力 ，必須小于凸模材料強度允許的壓應(yīng)力， 即 minbcZAF， -凸模的最小斷面面積（ ）inA2m -凸模縱向總壓力包括沖裁力和推件力（N），ZF bc-凸模材料的許用抗彎強度 對于一般工具鋼，凸模淬火硬度HRC為5862時，取10001600MPa；如果有特 殊導(dǎo)向時，可取20003000 MPa 51.85728/50.2654824574371.032 MPa minA Z， 由已知 bc450 MPabc 即落料凸模承載壓力完全合格。 由于本產(chǎn)品的生產(chǎn)綱領(lǐng)為10萬件,即在沖裁過程中由于材料本身 屬于硬質(zhì)材料,為此需要經(jīng)常磨刃口,適當(dāng)給落料凸模適當(dāng)加厚。 3、卸料.壓邊彈性元件的確定 沖壓工藝中常見的彈性元件有彈簧和橡膠等,但是由于這副模具的 結(jié)構(gòu)和結(jié)合實際生產(chǎn),因此我們選用橡膠作為卸料的彈性元件. 1 確定卸料橡膠 確定橡膠的自由高度H自,有資料查模具設(shè)計指導(dǎo)史鐵梁主編表 3-9得: H自=L工/(0.25-0.30)+h修磨 式中的L工為模具的工作行程再加1-3mm.本模具的工作行程為沖孔落料, 故L工1mm,h修磨的取值范圍為4-6mm,在這里取中間值5mm. H自=(1/0.3+5)mm9mm 確定L預(yù)和H裝.由表3-9可得如下計算公式: L預(yù)=(0.1-0.15)H自=0.15*9=1.35mm H裝=H自-L預(yù)=(9-1.35)mm=7.65mm 確定橡膠橫截面積A A=F/q F由前可知為F=2.17KN,q=0.26-0.5MPa.在這里,由于該模具的工作行程 比較小,因此取q=0.4MPa 則 A=2170/0.4mm=5425mm 核算橡膠的安裝空間: 可以安裝橡膠的空間可按 凹模外形表面積與凸凹模底部面積之差 的80估算.經(jīng)計算為S=13692.235mm,則可以安裝橡膠的面積 S=10953.788mm,大于所需的橡膠面積,因此滿足安裝橡膠的需要. 五、校核壓力機安裝尺寸 模座的外形尺寸為125125,閉合高度為160mm,由資料查模具設(shè)計 指導(dǎo)史鐵梁主編表4-33查得J23-10型壓力機的工作臺尺寸為 360mm240mm,最大閉合高度為180mm,連桿調(diào)節(jié)長度為50mm,故符合安 裝要求. 第六章 彎曲模具的設(shè)計 彎曲是使材料產(chǎn)生塑性變形，形成有一定角度或一定曲率形狀零件 的沖壓工序。彎曲的材料可以是板材、型材，也可以是棒料、管料。 彎曲工序除了使用模具在普通壓力機上進行外，還可以使用其他專門 的彎曲設(shè)備進行，以下幾例可看出： 圖6-1 1 制件彎曲工藝分析 本制件彎曲為綜合的折彎，折彎角度為165度和90度，彎曲半 徑為2mm，并且制件為對稱件的，彎曲形狀為U和V形彎曲的綜合彎 曲。在彎曲過程中要考慮彎曲回彈。 最小彎曲半徑計算r/t=2/0.5=458，彎曲半徑變化一般很小， 可不考慮，而僅考慮彎曲角度的回彈變化?；貜椊且詮澢昂蠊ぜ?彎曲角度的變化量表示?；貜椊? 0-t，其中 0為工件彎 曲后的實際彎曲角度， t為回彈前的彎曲角度（即凸模的彎曲角） 。 可以運用相關(guān)手冊查取回彈角修正經(jīng)驗系數(shù)值。 本制件在彎曲時帶有壓筋，在減小回彈上有一定的作用，即 它在結(jié)構(gòu)上改良了彎曲，使制件更容易達到要求。 2 沖壓工藝參數(shù)的確定。 1、沖壓力的計算 (1)彎曲力的計算 為了選擇合適的壓力機需要對彎曲力進行計算，影響彎曲力的因 素很多，如材料的性能，工件的形狀尺寸，板料厚度，彎曲方式，模 具結(jié)構(gòu)等。此外模具間隙和模具的工作表面質(zhì)量也會影響彎曲力的大 小。因此，理論分析的方法很難精確計算彎曲力。在生產(chǎn)實際中，通 常根據(jù)板料的機械性能以及厚度和寬度，按照經(jīng)驗公式計算彎曲力 （查課本p147式子） 。 彎曲力的計算F彎 (為了有效控制回彈,采用校正彎曲) 由模具設(shè)計指導(dǎo)史鐵梁主編一書中文獻(2)知 F自=0.7Kbt b/（r+t） b為彎曲件的寬度mm t為彎曲件的厚度mm r為內(nèi)圓彎曲半徑mm b為彎曲材料的抗拉強度MPa，查表 4-12得b =450MPa K為安全系數(shù)，一般取1.3 故有： F自=0.71.353.740.25450/(2+0.5)= 2.200653KN 3 頂料力和壓料力的計算，計算公式： FQ =（0.30.8）F自 =0.66019591.7605224KN 4 總壓力F總 F自+ FQ= 2.200653+1.76052243.9605224653K N 5 F壓機1.3F總=5.14867920489 根據(jù)F壓機可選壓力機類型為J23-4 第七章 彎曲模的結(jié)構(gòu)設(shè)計 彎曲模具的結(jié)構(gòu)設(shè)計是在彎曲工序確定后的基礎(chǔ)上進行的，設(shè)計 時應(yīng)考慮彎曲件的形狀要求、材料性能以及生產(chǎn)批量等因素。 一、模具結(jié)構(gòu)的分析說明 本制件為U形彎曲和V形彎曲的綜合，并且在下面兩側(cè)還有加 強壓筋，假如把模具設(shè)計成整體的，凸凹模難以加工制造，而且也 會增加成本，即使能制造出來也會在以后的生產(chǎn)中修模也帶來了困 難，不易修理。故可以把模具設(shè)計成相拼的。在外面用一凹?？蚶?面用凹模鑲塊，這樣有利于模具的加工和修理。其結(jié)構(gòu)如圖7-1 注：鑲塊主要用于限位的作用，凹模主要用于成型作用。 壤 塊凹 模 圖7-1 上圖的結(jié)構(gòu)是模具工作時的部位是這副模具的核心，這樣的結(jié)構(gòu)可以 適用于批量生產(chǎn)，即使壞了也可以再制造一個新的重新加工，這大大 縮短了工作周期，提高了生產(chǎn)率，而且卸料方便。給生產(chǎn)產(chǎn)家?guī)砹?更好的經(jīng)濟利益。 二、彎曲模的卸料裝置的設(shè)計說明 本制件在彎曲后將包在凸模上，在模具開模時用采用剛性卸料，在 開模時利用卸料板把制件從凸模上卸下來。在此同時下面也有一個 彈性卸料裝置把料和凹模從壤塊中頂出來其結(jié)構(gòu)如圖6-3 圖7-2 圖中在卸料螺釘下面有個彈性裝置在合模時它處于壓縮狀態(tài)，在開模 時要恢復(fù)原狀而釋放力起了卸料作用。 第八章 彎曲模的工作尺寸計算 工作尺寸設(shè)計的幾點說明： 1、 凸模圓角半徑 當(dāng)彎曲件的相對彎曲半徑r/t較小時，取凸模圓角半徑等于或 略小于工件內(nèi)側(cè)的圓角半徑r，但不能小于材料允許的最小彎曲半 徑。若彎曲件的r/t小于最小相對彎曲半徑，則應(yīng)取凸模圓角半徑 rtrmin，然后增加一道整形工序，使整形模圓角半徑rt=r0 當(dāng)彎曲件的相對彎曲半徑r/t較大時（大于10） ，并且精度要求 較高時，必須考慮回彈的影響，根據(jù)回彈值的大小對凸模圓角半 徑進行修正。 2、凹模圓角半徑 凹模入口處的圓角半徑的大小對彎曲力以及彎曲件的質(zhì)量均 有影響。過小的凹模圓角半徑會使彎距的彎曲力臂減小，毛坯沿 凹模圓角滑入時的阻力增大，彎曲力增加，并容易使工件表面擦 傷甚至出現(xiàn)壓痕。 在生產(chǎn)中，通常根據(jù)材料的厚度選擇凹模圓角半徑： 當(dāng)t2mm時，ra=（3）t 當(dāng)tmm時，ra（）t； 當(dāng)tmm時，rat。 本制件凹模圓角半徑應(yīng)取ra=2.5mm 、凹模的深度 凹模的深度要適當(dāng)，若過小則彎曲件的兩端自由部分太長，工 件回彈大，不平直；若深度過大則凹模增加高度，多耗材并需要較 大的壓力機工作行程。 對于彎曲件，若直邊高度過大或要求兩邊平直，則凹模深度 應(yīng)大于工件深度。 、彎曲凸凹模的間隙 行件的彎曲時必須合理確定凸、凹之間的間隙，間隙過大則 回彈，工件形狀和尺寸誤差增大。間隙過小會增加彎曲力，使工件 厚度減薄，增加摩擦，擦傷工件并降低模具的壽命。形件的凸凹 模具的單面間隙值一般可按下式計算： ct+kt 式中： c為凸凹模的單面間隙mm t板料厚度的基本尺寸mm 板料厚度的正偏差mm k根據(jù)彎曲件的高度和寬度而決定的間 隙系數(shù) 當(dāng)工件精度高的時候，間隙值應(yīng)適當(dāng)減小，可以取ct 、形件彎曲模工作部分的尺寸計算 ） 、彎曲件外形尺寸的標注應(yīng)以凹模為基準，先確定凹模的尺寸， 然后再減去間隙值確定凸模尺寸。 當(dāng)彎曲為雙向?qū)ΨQ偏差時，凹模尺寸為： d（/） d0 當(dāng)彎曲件為單向偏差時，凹模尺寸為： dLd0)43( 凸模尺寸為： )2(pcdp 或者凸模尺寸按凹模實際尺寸配制，保證單面間隙值c式中 d、p為凹模、凸模的制造公差。 ） 、彎曲件內(nèi)形尺寸的標注應(yīng)以凸模為基準件，先確定凸模尺寸， 然后再增加間隙值確定凹模尺寸。 當(dāng)彎曲件為雙向?qū)ΨQ偏差時，凸模尺寸為： 0)21(pLp 當(dāng)彎曲件為單向偏差時，凸模尺寸為： 0)43(pp 凹模尺寸為： dcLd0)2( 或者凹模尺寸按凸模實際尺寸配制，保證單面間隙值c 式中 p、 d為凸模、凹模的制造公差，選用IT7IT9級精度mm。 經(jīng)過上述幾點說明可得凹凸模尺寸為如圖8-1 凸模 c 凹模 圖8-1 畢業(yè)設(shè)計小結(jié) 時光飛逝，轉(zhuǎn)眼間我們就要大學(xué)畢業(yè)了。剛結(jié)束畢業(yè)設(shè)計的我有 很多的感想要對大家說，在此我就利用這個機會來發(fā)表我的長篇大論 了。通過這次的畢業(yè)設(shè)計真的讓我學(xué)到了很多東西。以前在學(xué)理論的 時候，我以為自己學(xué)到了很多專業(yè)方面知識，曾經(jīng)還一度的認為自己 不錯，但是在實習(xí)的這段時間以后，我才發(fā)現(xiàn)自己的缺陷和不足，而 且還非常的缺乏經(jīng)驗，因為在學(xué)校的時候我們接觸的都是些純理論的 專業(yè)知識，并沒有與實踐聯(lián)系起來，所以我們對模具的認識只局限在 表面上的感性的認識，而沒有上升到理性認識的高度，換句話說就是 還沒有真正的認識模具這一名詞的真正含義。但是只有實踐還是不夠 的，經(jīng)驗才是最寶貴的。而對與我們這些剛剛走出大學(xué)校園的畢業(yè)生 來說最缺乏的就是經(jīng)驗了。所以在實習(xí)的這段時間我吃了不少苦頭。 這次我設(shè)計的是書夾，這個零件的形狀不是很復(fù)雜，是個軸對稱 零件，對于工廠來說，是個比較簡單的零件，不過對于我們這些剛要 畢業(yè)的學(xué)生來說還是比較困難的。 這個零件共分兩道工序，落料沖孔、彎曲壓筋，兩道工序中有一道 是復(fù)合的，特別是一道彎曲壓筋難度較大，難的不是彎曲而是壓筋這 樣來實現(xiàn)，我看到了廠的的設(shè)計他們設(shè)計了很巧妙。把彎曲外形和壓 筋鑲塊分開制造，這樣即省了錢又提高了效率，也在制造上降低了難 度是很可取的辦法。我開始是打算用個比較精密的沖側(cè)孔模具把之加 工出來，但是是師傅看到了問為什么這樣而不用金加工的方法加工出 來，我的回答另他們很不滿意。在后來我才發(fā)覺到我這樣加工太浪費 了不切合實際。理論就是理論要把它和實際結(jié)合起來才能算是真正的 有用理論。我那樣加工固然是個很不錯的加工方法，但是在生產(chǎn)是太 浪費達不到原有的經(jīng)濟利益，在制造模具時要浪費大量的錢，并且在 模具裝夾定位時也很困難，很難不易不讓它變形。最后不得不向?qū)嶋H 低頭采納師傅們的意見。確定了該制件的工藝方案以后，通過力的計 算，進行各套模具的壓力機的選擇；通過凸凹模尺寸的計算，選擇所 需要的模架和導(dǎo)柱導(dǎo)套，并對其進行校核；最后畫各套模具的零件圖 和裝配圖。 在設(shè)計的時候經(jīng)常會遇到問題，比如說，設(shè)計打桿的時候，由于不 知道打桿在模具中靠什么工作的，所以就不知道它的長度該怎么確定， 每次遇到這樣的問題時，都是師傅不厭其煩地教我，直到我懂為止。 這段時間師傅教了我很多在學(xué)校里沒學(xué)過的東西，使我對模具的認識 進一步加深了，對模具的興趣也越來越濃了。 通過這次畢業(yè)設(shè)計我的收獲很多，最主要的是實踐方面的收獲， 因為通過畢業(yè)設(shè)計，讓我了解了沖壓模具的主要結(jié)構(gòu)組成，同時還了 解了設(shè)計一套模具的主要流程。但學(xué)會這些還是不夠的，在以后的工 作學(xué)習(xí)中我將更加努力，從而彌補自己的不足。 由于水平所限，設(shè)計中的錯誤再所難免，懇請各位老師批評指正。 參考文獻 1王新華 袁聯(lián)富編.沖模結(jié)構(gòu)圖冊.北京：機械工業(yè)出版社.2003 2機械設(shè)計手冊 3沖壓手冊 4模具設(shè)計手冊 5 金滌塵，宋放之. 現(xiàn)代模具制造技術(shù)M 機械工業(yè)出版社，2001 6 胡石玉.模具制造技術(shù)M.東南大學(xué)出版社.1997. 7 李云程. 模具制造工藝學(xué)M. 機械工業(yè)出版社.1998. 8 孫鳳勤. 模具制造工藝與設(shè)備M. 機械工業(yè)出版社.1999. 9 孫鳳勤. 沖壓與塑壓設(shè)備M. 機械工業(yè)出版社.1997. 10 模具標準匯編M. 機械工業(yè)出版社.1999. 11 標準件手冊M. 機械工業(yè)出版社.1999. 12 機械加工工藝手冊M. 機械工業(yè)出版社.1999. Annals of the CIRP Vol. 56/1/2007 -269- doi:10.1016/j.cirp.2007.05.062 Design of Hot Stamping Tools with Cooling System H. Hoffmann 1 (2), H. So 1 , H. Steinbeiss 1 1 Institute of Metal Forming and Casting, Technische Universitt Mnchen, Garching, Germany Abstract Hot stamping with high strength steel is becoming more popular in automotive industry. In hot stamping, blanks are hot formed and press hardened in a water-cooled tool to achieve high strength. Hence, design of the tool with necessary cooling significantly influences the final properties of the blank and the process time. In this paper a new method based on systematic optimization to design cooling ducts in tool is introduced. The optimization procedure was coupled with FE analysis and a specific evolutionary algorithm. Through this procedure each tool component was separately optimized. Subsequently, the hot stamping process was simulated both thermally and thermo-mechanically with the combination of optimized solutions. Keywords: Hot Stamping, Finite element method (FEM), Optimization 1 INTRODUCTION In recent years, weight reduction while maintaining safety standards has been strongly emphasized in the automotive industry for building new models. Hot stamping of high strength steels for automotive inner body panels offers the possibility of fuel saving by weight reduction and enhances passenger safety due to its higher strength. In order to achieve high strength by hot stamping with high strength steels, blanks should be heated above austenitic temperature and then cooled rapidly such that the martensitic transformation will occur. Normally, the tools are heated up to 200C without active cooling systems in serial production 1. However, in hot forming processes, the tool temperature must maintain below 200C to achieve high strength. So far, very few studies have been conducted regarding the design of cooling systems in a hot stamping tool. This paper presents a systematic method to design hot stamping tools with cooling systems in optimal and fast manners, in which the cooling system is optimized with the help of FE analysis and a specific evolutionary algorithm. Subsequently, a series of hot forming processes was simulated thermally as well as thermo-mechanically to observe the heat transfer and the cooling rate through the optimized cooling system. In the hot stamping process the tool motion requires relatively short time compared to the whole process time. Therefore, thermal analysis of a series of hot stamping processes without considering the tool motion could be sufficient with reasonable accuracy and shorter computation time for quick design of the hot stamping tools with cooling system. However, thermo- mechanical analyses that include the motion of the punch and the forming process are necessary to enhance the accuracy of the predictions. In this paper, a crash relevant hot stamped component of a vehicle and its corresponding prototype of hot stamping tool are introduced in chapter 2. And the optimization procedure with FE analysis and evolutionary algorithm is introduced in chapter 3. Subsequently, the results of thermal and thermo-mechanical analyses with the optimized hot stamping tool are presented. 2 COOLING OF HOT STAMPING TOOL 2.1 Motivation To enhance the economical production procedure and good characteristics of the formed parts, hot stamping tools need to be designed optimally. Therefore, the main objective of this study is the optimal designing of an economical cooling system in hot stamping tools to obtain efficient cooling rate in the tool. So far, very few researches have been conducted regarding the design of cooling systems in hot stamping tools. Therefore, an advanced design method is required. Also, an adequate simulation model is required to perform the optimization and investigation of tools and products as fast and accurate as possible. 2.2 Characteristics of 22MnB5 In direct hot forming process, the quenchable boron- manganese alloyed steel 22MnB5 is commonly used. Also, 22MnB5 is one of the representative materials of ultra high strength steels. Therefore, in this study, aluminium pre-coated 22MnB5 sheet (Arcelors USIBOR) was considered as the blank material. The material 22MnB5 has a tensile strength of 600MPa approximately at the delivery state, and the tensile strength can be significantly increased by hot stamping process. Higher tensile strength is achieved in the hot stamping process by a rapid cooling at least at the rate of 27C/s 2. The initial sheet of 22MnB5 consisting of ferritic-perlitic microstructure must be austenitized before forming process in order to achieve a ductility of blank sheet. As the austenite cools very fast during quenching process martensite transformation will occur. This microstructure with martensite provides the hardened final product with a high tensile strength up to 1500 MPa. 2.3 Tool component and test part The components of the prototype hot stamping tool and its kinematics are shown in Figure 1 and the initial blank and the proposed test part in Figure 2. The initial blank has the dimension of 170mm x 430mm x 1.75mm and the draw depth of the proposed test part is 30mm. -270- faceplate counter punch blank holder punch faceplate table table blank distance bolts die barells plunger Figure 1: Schematic of a test hot stamping tool. Initial thickness: 1.75mm 4 3 0 m m1 7 0 m m 4 0 0 m m 1 0 0 m m Draw depth: 30mm Figure 2: Initial blank and drawn part. 2.4 Cooling systems in stamping tools The tool must be designed to cool efficiently in order to achieve maximum cooling rate and homogeneous temperature distribution of the hot stamped part. Hence, a cooling system needs to be integrated into the tools. The cooling system with cooling ducts near to the tool contour is currently well known as an efficient solution. However, the geometry of cooling ducts is restricted due to constraints in drilling and also the ducts should be placed as near as possible for efficient cooling but sufficiently away form the tool contour to avoid any deformation in the tool during the hot forming process. To guarantee good characteristics of the drawn part, the whole active parts of the tool (punch, die, blank holder and counter punch) need to be designed to cool sufficiently. 3 DESIGNING OF COOLING SYSTEMS 3.1 Optimization with Evolutionary Algorithm x s a boring position minimum distance between loaded contour and cooling duct (x) between unloaded contour and cooling duct (a) between cooling ducts (s) loaded contour unloaded contour coolant bore Constraints sealing plug input parameters of cooling system number of cooling channels and coolant bores diameter of cooling duct evaluation criteria cooling intensity and uniform cooling Optimization (Evolutionary Algorithm) 1 solution per given input separate optimization Solution Figure 3: Optimization procedure for each tool. The optimization procedure for design of a cooling system is presented in Figure 3. In this procedure, cooling channels can be optimized in each tool by a specific Evolutionary Algorithm (EA), which was developed at ISF (Institut fr Spannende Fertigung, Universitt Dortmund, Germany) for the optimization of injection molding tools and adapted for design of cooling systems in hot stamping tools 3,4. As constraints for optimization, the available sizes of connectors and plugs, the minimum wall thicknesses as well as the nonintersection of drill holes were considered. The admissible minimal distance between cooling duct and unloaded/loaded tool contour (a/x) and the minimal distance between cooling ducts (s) were determined through FE analyses. Parameters of the cooling system such as the number of channels (a chain of sequential drill holes), drill holes per channel and the diameter of the holes for each tool component were also provided as input parameters to the optimization. These input parameters can be obtained from existing design guidelines or through FE simulations. Based on the input parameters initial solution is generated randomly by EA or manually by the user. From the initial solution, the EA generates new solutions by recombination of current solutions and modifying them randomly. The defined constraints were subsequently used for the correction of the generated solutions and the elimination of inadmissible solutions. All the generated solutions were evaluated by optimum criteria such as efficient cooling rate and uniform cooling. Finally, the best solution was selected as optimized cooling channels for a selected tool component. 3.2 Optimized cooling channels In our research, the selected diameters of ducts were 8mm and 12mm for punch, 8mm, 12mm and 16mm for die, 8mm and 10mm for counter punch and 8mm for blank holder. EA was used to place the cooling channels optimally according to the given input and constraints for each tool component. The optimized profiles of the channels for duct diameter of 8mm are shown in Figure 4. c a b 4 0 0 m m 100mm 145 mm pu n c h cou n ter p un ch die b l an k h o ld er a b a b c a b 5 1 0 m m 260 mm a b c 70mm 510mm ab 260 mm a 110mm cooling medium plug 380mm a 70mm 250 mm b c b direction of cut view Figure 4: Optimized cooling channels with 8mm duct diameter. 4 EVALUATION OF THE OPTIMUM COOLING CHANNEL DESIGNS The design of cooling channels was generated by EA for each tool component with different bore diameters and their cooling performances were evaluated by using FE simulations. 4.1 Thermal analysis In the design and development phase of hot stamping tools, it is important to estimate the hot stamping process qualitatively and quantitatively within a short time for -271- economic manufacturing of tools. For this purpose, two transient thermal simulations were carried out with ABAQUS/standard, which uses an implicit method. In this analysis steel 1.2379 was selected as a tool material. The simulation model comprises 4 tool components: punch, die, blank holder and counter punch. In Table 1, the selected combinations of tool components with optimized cooling channels are presented. The variant V1 is the combination of optimized tools with small cooling duct diameters, whereas variant V2 with large cooling duct diameters. V1 V2 punch counter punch blank holder 8mm 8mm 8mm 8mm 12mm 10mm 16mm 8mm diameter of cooling duct die Table 1: Combinations of designed tools for FE analysis. In order to represent a series of production processes, a number of cycles of the hot stamping processes were simulated as a cycle heat transfer analysis. The Figure 5 shows the FE model including boundary conditions. cooling duct (c) T c = 20C h c = 4700W/m 2 C tool (t) T t,0 = 20C environment (e) T e = 20C h e = 3.6W/m 2 C counter punch blank holder punch blank die blank (b) T b,0 = 850C blank - tool D c = f (d,P) Figure 5: FE model and boundary conditions. This hot forming process for the prototype part was designed such that the cycle time is 30 sec. In a cycle, the punch movement for forming requires 3 sec, the tool is closed for 17 sec for quenching the blank and it takes another 10 sec for opening the tool and locating the next blank on the tool. However, in this thermal analysis, the tool motion and deformation of the blank was not considered to reduce the computation time. Hence, only heat transfer analysis was performed in a closed tool. In thermal analysis, the quenching process takes places 20 sec instead of 17 sec, because the motion of punch was not considered. It was assumed that the blank has an initial homogeneous temperature (T b,0 ) of 850C due to free cooling from 950C during the transfer in environment. The initial tool temperature (T t,0 ) was assumed as 20C at the first cycle and changes from cycle to cycle. The temperature of the cooling medium (T c ) was assumed as room temperature. Beside the boundary conditions, the required material properties of 22MnB5 were obtained from hot tensile test conducted at LFT (Lehrstuhl fr Fertigungstechnologie, Universitt Erlangen-Nrnberg, Germany), with whom a joint research on hot stamping is being conducted 2. In this analysis, convection from blank and tools to the environment (h e ), conduction within each tool, convection from tool into cooling channels (h c ) and heat transfer from hot blank to tool (D c ) were considered. Here, D c , is the contact heat transfer coefficient (CHTC) which describes the amount of heat flux from blank into tools. This coefficient usually depends on the gap d between tool and blank and the contact pressure P. It increases usually as the contact pressure increases. However, in thermal analysis the pressure dependent CHTC was not available, but a gap dependent coefficient was used. CHTC was assumed as 5000W/m 2 C at zero distance between blank and tool (gap) and keeps constant until the gap increases beyond critical value. 4.2 Thermo-mechanical analysis Simulation of hot forming is different from conventional sheet metal forming simulation, in which the distribution of temperatures or stresses in tools is neglected. For fast and easy way to analyze the hot forming process the tool and the blank were modelled with shell elements as in other studies 5,6. In these studies, the temperatures could be distributed along the thickness of the shell element with the user-defined function of temperature, but the temperature within the tool was not considered. Also, in this simulation model the heating of tools in a series of hot stamping processes were not considered. Furthermore, the shell model for thermal contact problems is just adequate for relatively short contact time 6. Therefore, in our studies the tools and the blank were modelled with volume elements to simulate the sequential heat transfer in a series of processes. The thermo- mechanical simulation was conducted with ABAQUS/explicit. In comparison to the thermal analysis, the whole forming and quenching process were modelled and the dynamic temperature and stress responses of tools in contact with hot blank were simulated by using time-temperature dependent flow stress curves. The heat transfer could be more accurately expressed using pressure dependent CHTC at contact places which change during forming process. In addition, temperature dependent thermal conductivity and specific heat were also considered. However, in thermo-mechanical analysis, as the number of elements increases, the complexity of the FE problem significantly increases. In conventional forming simulation an adaptive mesh can be normally used to spare the simulation time and to obtain a more accurate solution in the contact area. However, adaptive mesh refinement causes instability during computation in thermo- mechanical analysis. Therefore, a refined mesh with higher punch velocity was considered to reduce the simulation time. The heat transfer coefficients were scaled accordingly to obtain the same heat flux 7. 5 SIMULATION RESULTS AND DISCUSSION 5.1 Thermal analysis Figure 6 shows the temperature changes in the tool components for 10 cycles at tool combination V1 and V2. T C 400 300 100 0 030100 0 300100 die punch t s t s V1 V2 Figure 6: Temperature changes in heat transfer analysis. The results show that the hottest temperatures of the tools at the end of each cycle do not change almost after some cycles. The obtained cooling rates of the blank at the hottest point from 850C to 170C are 40C/s with V1 and 33C/s with V2 at 10th cycle and these are greater than the required minimum cooling rate of 27C/s. Furthermore, V1 leads to a more efficient cooling performance than V2. Better cooling performance for V1 compared to V2 can be explained with the geometric restrictions and the minimal wall thickness. A cooling duct with small diameter can be placed closer to the tool surface in a convex area and the amount of the cooling channels can be increased additionally. Usually, the heat dissipation in the convex area is slower than in concave area 6. The result shows also that the temperature of convex area in the punch -272- cools down slower than the concave areas in the die. Due to this fact, it can be concluded that the efficient cooling is most desired at convex area. 5.2 Thermo-mechanical analysis The heat transfer with optimized tool components was simulated thermally at first. However, there was a simplification of a hot stamping process in thermal analysis. Therefore, a thermo-mechanical analysis for V1 was performed to observe the differences and the significance of modelling the punch movement. Temperature change curves at the hottest point from the end of the first cycle in the blank, die and punch are shown in Figure 7. The tool cooled further 10 sec after quenching and the temperature changes in the die and punch were presented for 30 sec. A coupled thermo- mechanical analysis was done using gap-pressure dependent CHTC. The results from thermal analysis shows a cooling rate of 92C/s from 850C to 170C in comparison to 75C/s from thermo-mechanical analysis. 400 300 100 0 die punch 05 20 1000 800 400 T C 200 Thermal analysis Thermo-mechanical analysis t s 15 blank 0 0 5 30 0 5 25 30t s10 202510 20t s T C Figure 7: Temperature changes in thermal and thermo- mechanical analysis (1th cycle). To verify the accuracy of a thermal analysis or to predict a serial production process more accurately a series of thermo-mechanical analysis was done. For this analysis the punch velocity was increased 10 times and 10 hot stamping processes were simulated. In Figure 8, the temperature change curves at the hottest point of the die and punch from a thermal and thermo-mechanical analysis are compared for 10 cycles. Finally, the temperature distributions in the blank at the end of the 10th cycle are shown in Figure 9. 400 300 100 0 TC 030ts100 030ts100 die punch thermal thermo-mechanical Figure 8: Temperature changes for 10 cycles. (b) T C (a) 130 60 102 74 88 116 T C 140 70 112 84 98 126 Figure 9: Temperature fields of blanks at the end of 10th cycle: (a) thermal and (b) thermo-mechanical analysis. In Figure 8, the temperature differences at the end of 10th cycle between the thermal and thermo-mechanical analyses were 7C in the die and 3C in the punch. Subsequently, the Figure 9 indicates that the maximum temperature of the blank from the thermal analysis is slightly greater than that of the thermo-mechanical about 10C. Nonetheless, the temperature fields of blanks from both analyses are very similar. As a consequence, the thermal analysis for a series of hot stamping processes is relatively accurate compared to the thermo-mechanical analysis. Furthermore, a thermal heat transfer analysis could be used to design and develop the hot stamping tools in the early phase due to its timesaving computation. 6 CONCLUSION AND FUTURE WORK A systematic method has been developed for optimizing the geometrical design of the cooling systems of hot stamping tools. This methodology was successfully applied to design of cooling channels in a prototype tool for efficient cooling performance. This indicates that the method can be used for designing cooling systems in other stamping tools as well. This paper presented both thermal and thermo- mechanical simulations to represent a series of hot stamping processes. The thermal analysis could be used for an optimization and investigation of hot stamping processes especially in the developing stage. However, a thermo-mechanical analysis is needed to predict more accurately but it is still time consuming to analyze the processes within adequate time period. To resolve this problem, an alternative simulation model will be further studied. Also, a more accurate contact condition for thermo-mechanical analysis remains to be studied. To validate this proposed method and its corresponding FE model, a prototype tool is currently being built and experiments will be carried out for validation. 7 ACKNOWLEDGMENTS We extend our sincere thanks to all joint project researchers of LFT and ISF. 8 REFERENCES 1 Sik