1213錘式破碎機(jī)結(jié)構(gòu)設(shè)計(jì)與三維建模含proe三維及8張CAD圖

1213錘式破碎機(jī)結(jié)構(gòu)設(shè)計(jì)與三維建模含proe三維及8張CAD圖,破碎,結(jié)構(gòu)設(shè)計(jì),三維,建模,proe,cad 燒結(jié)廠反擊式破碎機(jī)的AISI A2工具鋼打漿機(jī)頭的開發(fā) 摘要 介紹了一臺沖擊式破碎機(jī)的工具鋼（AISI A2）打漿機(jī)頭的故障分析，以及開發(fā)合適的熱處理工藝以改善其性能。打手頭因其針孔位置的脆性斷裂而過早失效。該研究包括目視檢查，斷層成像，化學(xué)分析，使用光學(xué)和掃描電子顯微鏡（SEM）表征微結(jié)構(gòu)，EDS分析以及顯微硬度分布的測定。使用SEM和EDS分析的微觀結(jié)構(gòu)表征揭示了馬氏體基體中的大量連續(xù)的粗碳化物網(wǎng)狀物。它增加了硬度（64HRC）以及基體的不均勻性（如微觀硬度分布圖所示），并降低了韌性（3J），因?yàn)榇痔蓟锞W(wǎng)絡(luò)非常硬且脆。在制造商方面，奧氏體化溫度和熱處理的回火溫度都較低。新推薦的熱處理導(dǎo)致較少量的不連續(xù)Cr-碳化物以及大量均勻分布在整個(gè)基體中的細(xì)小沉淀物，從而使硬度（59 HRC）和韌性（6.5 J）達(dá)到應(yīng)用所需的最佳組合。按照推薦的熱處理工藝生產(chǎn)的打漿機(jī)機(jī)頭表現(xiàn)出比以前更好的性能（壽命增加4倍）。 關(guān)鍵詞反擊式破碎機(jī)；錘子；AISI A2；工具鋼；網(wǎng)狀碳化物；脆性斷裂；韌性；熱處理 介紹 燒結(jié)礦被用作高爐煉鐵的原料。燒結(jié)是一種通過初期熔合將基礎(chǔ)混合物凝聚成多孔物質(zhì)的過程?；A(chǔ)混合材料是通過按比例混合鐵礦粉，地面助熔劑材料，磨碎的焦炭粉和還原混合物，然后在堆場中逐層堆積來制備的?；匣旌显系念w粒是燒結(jié)礦質(zhì)量的重要因素。通過在沖擊式破碎機(jī)中破碎，然后以適當(dāng)?shù)谋壤Y分和混合不同的顆粒來實(shí)現(xiàn)正確的造粒。在燒結(jié)礦中用作助熔劑的石灰石和純橄欖石通過吊桿重新加熱輪回收，并送到三臺150 tph的錘磨機(jī)中，稱為初級破碎機(jī)，然后將它們送到二級破碎機(jī)（Hammer Mills）中，它們被研磨成3.2毫米的大小。 粉碎后，將這些材料篩分并儲存在配料箱中。 該沖擊式破碎機(jī)廣泛用于粉碎石灰石，白云石和純橄欖石等原料，用于燒結(jié)配料的床上用品和混合設(shè)備。它使用多個(gè)旋轉(zhuǎn)攪拌頭（錘子）通過撞擊固定在機(jī)器沖擊壁上的磨削襯套來粉碎原材料[1]。打漿機(jī)頭在鋸齒形周向排列的攪拌臂幫助下連接到轉(zhuǎn)子軸（圖1）。通過調(diào)整旋轉(zhuǎn)錘頭與沖擊壁之間的距離，可以改變破碎機(jī)的顆粒產(chǎn)品的細(xì)度。 圖1 (a)沖擊式破碎機(jī)，顯示了轉(zhuǎn)子軸比頭總成，(b)更近地觀察一個(gè)打頭。 沖擊式破碎機(jī)的攪拌頭在脆性斷裂下經(jīng)常失敗（在15天內(nèi)）。打擊面上打擊頭的逐漸磨損是一種正常的失效模式，而其突然的脆性斷裂是一個(gè)值得關(guān)注的問題。沖擊式破碎機(jī)中的錘頭或錘頭不僅會(huì)受到磨損，還會(huì)受到?jīng)_擊或沖擊載荷的作用。因此，打擊頭材料的韌性是延長錘子硬度或耐磨性能的重要因素[2-6]。每個(gè)月都需要更換大量的打漿機(jī)頭，因?yàn)樗鼈兘?jīng)常發(fā)生破損，導(dǎo)致維護(hù)成本高和生產(chǎn)損失。在現(xiàn)有的文獻(xiàn)[2-10]中提供了關(guān)于合金鑄鐵和AISI H11，H13，D2和M2等工具鋼的幾項(xiàng)工作，這些文獻(xiàn)涉及微觀結(jié)構(gòu)和機(jī)械特性。但是在空氣淬火工具鋼AISI A2及其作為反擊式破碎機(jī)（錘式破碎機(jī)）中的打漿機(jī)頭材料的應(yīng)用在公開文獻(xiàn)中很少見。此外，在現(xiàn)有文獻(xiàn)中沒有關(guān)于AISI A2工具鋼作為粉碎錘的性能的工業(yè)或應(yīng)用數(shù)據(jù)。這項(xiàng)工作提出了對AISI A2攪拌頭過早失效的根本原因進(jìn)行分析，并開發(fā)合適的熱處理工藝以改善其性能。 實(shí)驗(yàn)程序和結(jié)果 實(shí)地考察和視覺觀察 燒結(jié)廠有兩臺原料床上用品和混合（RMBB）單元，分別是RMBB＃1和RMBB＃2，其中有6臺沖擊式破碎機(jī)和7臺沖擊式破碎機(jī)。 沖擊式破碎機(jī)粉碎燒結(jié)中使用的助熔劑材料，如石灰石，白云石，輝石巖等。以約675rpm的轉(zhuǎn)速旋轉(zhuǎn)的初級破碎機(jī)首先將原材料從-50mm的網(wǎng)孔破碎成 -15毫米。然后這些再次在具有約830rpm的二次破碎機(jī)的幫助下破碎成大小-3.15毫米。破碎機(jī)的容量大約為125噸/小時(shí)，它包含40-52個(gè)打漿機(jī)頭。 圖1（a）顯示了轉(zhuǎn)子軸攪拌頭組件。 打漿機(jī)頭在幫助下附在打漿機(jī)臂上。如圖1（b）所示，銷子插入打擊頭內(nèi)的孔中。 圖2（a）和（b）顯示了不合格的攪拌器頭從其針孔位置斷裂。 失效攪拌器頭部的斷裂表面（圖2c和d）顯示明亮的顆粒外觀，表明脆性斷裂。 打擊頭的另一端被發(fā)現(xiàn)磨損（圖2a和b），原因是在輸入助焊劑材料磨損時(shí)受到磨損或撞擊。 雖然攪拌器頭表面的逐漸磨損是所需的失效模式，但在服務(wù)期間從針孔位置斷裂是一個(gè)必須解決的問題。 圖2 (a，b) 打樣頭樣品在打孔位置處失敗(c，d)靠近打頭針孔附近的裂縫面 斷口 包含斷裂表面的小樣品在斷裂發(fā)生位置的針孔附近被切割。 將樣品進(jìn)行超聲波清洗以檢查裂隙表面。 使用在15kV的加速電壓下操作的掃描電子顯微鏡（SEM）（型號：JXA6400，JEOL，Japan）檢查斷裂表面。 在各種放大率下記錄斷裂表面各個(gè)位置的顯微照片。 如圖3所示，檢查裂紋萌生區(qū)附近的斷裂表面顯示斷裂表明脆性斷裂。 攪拌頭的脆性斷裂表明其在使用過程中的沖擊載荷下失效。 圖3分形圖(92000)失敗的擊劍頭顯示了脆性斷裂的分裂。 物料 從失敗的打漿機(jī)機(jī)頭上切下一小塊樣品，并準(zhǔn)備進(jìn)行化學(xué)分析。樣品的化學(xué)分析使用X射線熒光光譜法（XRF）進(jìn)行;碳（C）和硫（S）含量采用紅外燃燒技術(shù)測定。打擊頭的化學(xué)分析結(jié)果匯總在表1中?；瘜W(xué)分析發(fā)現(xiàn)更接近AISI A2（ASTM A681）等級的空氣淬火中等合金冷作工具鋼，其鉻（Cr）含量稍高，降低釩量（V）。錳，鉻和鉬是這種等級鋼中的主要合金元素，其賦予高淬透性并且鋼可在空氣中硬化。 AISI A2提供了在使用中原材料粉碎操作所需的耐磨性和韌性的最佳組合。圖4給出了從制造商那里得到的攪拌機(jī)頭的熱處理周期。鑄造材料在850-870℃下熱處理1h / in。隨后空氣冷卻。風(fēng)冷材料在175-200℃的溫度范圍內(nèi)以兩個(gè)階段回火3h / in。 表1錘頭頭樣品化學(xué)分析(wt%)。 表1 beater頭樣品化學(xué)分析(wt%)。Sample C Mn Si S P Cr Mo V W Beater head 1.01 0.67 0.50 0.030 0.026 5.7 1.02 0.12 0.010 ASTM A681 type A2 0.95–1.05 0.4–1 0.1–0.5 0.03 max 0.03 max 4.75–5.5 0.9–1.4 0.15–0.5 … （a） （b） 圖4 (a) 錘頭制造商所給出的典型熱處理時(shí)間表，(b)推薦的錘頭熱處理周期 微觀結(jié)構(gòu)檢查 從失敗的攪拌頭上切下樣品用于橫截面的微觀結(jié)構(gòu)檢查。然后將樣品安裝在樹脂中，研磨并使用標(biāo)準(zhǔn)金相技術(shù)拋光。用Villela's試劑（1g苦味酸，5mL鹽酸和100mL乙醇）蝕刻后，在光學(xué)顯微鏡（Leica，型號：DMRX，德國）上檢查橫截面。在圖5（a）和（b）中顯示了破碎打擊頭的橫截面的典型顯微照片。打擊頭橫截面的微觀檢查揭示了馬氏體基體和晶界上的鏈狀塊狀碳化物網(wǎng)絡(luò)（圖5a和b）。碳化物表現(xiàn)為明亮的相，它們優(yōu)先沿著晶界聚集。使用圖像分析軟件在各個(gè)領(lǐng)域測量碳化物的面積分?jǐn)?shù)。發(fā)現(xiàn)碳化物的平均面積分?jǐn)?shù)為7.2±0.45％。 打擊頭元件所需的最重要的材料特性是硬度，韌性和耐磨性。隨著碳化物數(shù)量的增加，硬度和耐磨性也增加，但在熱處理時(shí)必須小心，以避免韌性損失[2,3,6,7]。粗大的碳化物團(tuán)簇對韌性產(chǎn)生不利影響;在使用過程中，碳化物易碎易于引發(fā)裂紋[2,8,9]。 圖5失敗錘頭的微觀結(jié)構(gòu):(a)微觀結(jié)構(gòu)(950)顯示馬氏體基體和碳化物網(wǎng)絡(luò)在晶界處 (b)放大后的微觀結(jié)構(gòu)(9200)顯示在前奧氏體晶界的連續(xù)碳化物網(wǎng)絡(luò)。 掃描電子顯微鏡和EDS分析 在加速電壓為15kV的掃描電子顯微鏡（SEM）的幫助下，對樣品的蝕刻橫截面進(jìn)行了檢查，以了解其微觀結(jié)構(gòu)以及元素表征。顯微照片顯示在晶界有粗大的碳化物網(wǎng)狀結(jié)構(gòu)，還有一些帶有馬氏體基體的細(xì)小球狀沉淀物（圖6a和b）。進(jìn)行晶界網(wǎng)絡(luò)的能量色散譜（EDS）以及細(xì)小沉淀物（如圖6a和b所示），進(jìn)行元素化表征，分析結(jié)果匯總在表2中。 EDS分析表明，基體內(nèi)的晶界網(wǎng)絡(luò)和細(xì)小沉淀物是碳化鉻。碳化物化學(xué)計(jì)量不能通過EDS微觀分析確定。但是對類似材料[2,10,11]的文獻(xiàn)調(diào)查，如AISI H11，H13，M2，D2等，以及早期文獻(xiàn)[5,12-15]討論的類似碳化物的電子探針顯微分析表明，邊界碳化物網(wǎng)絡(luò)為M7C3型（原碳化物）和細(xì)小球狀沉淀物為M23C6（其中M = Cr）型（次碳化物）。 如圖6（b）所示，在晶界處的大塊碳化物網(wǎng)狀骨架狀形態(tài)顯然表明它的共晶起源（一次碳化物），即起始于凝固階段[2,5,8- 10，13]。 另一方面，基體中細(xì)小的球狀析出物表明在鑄件熱處理過程中形成的二次碳化物[2,5,8,11]。 （a） (b) 圖6 錘頭樣品截面的SEM顯微圖 (a)微圖(91000)，在馬氏體基體中，在晶界處顯示粗碳化物網(wǎng)絡(luò)，并在馬氏體基體內(nèi)呈現(xiàn)細(xì)析出物(b)顯微圖(91500)顯示了粗碳化物網(wǎng)絡(luò)中EDS分析的位置，以及矩陣內(nèi)的細(xì)析出物。 表2 EDS的結(jié)果在不同的分析(wt %)如圖6所示 位置 C 附注 1 11.71 0.40 16.68 0.68 64.5 6.03 二次碳化物 2 19.35 … 40.92 … 30.63 8.71 主要的硬質(zhì)合金 3 … 1.00 4.83 0.50 92.71 0.96 矩陣 4 9.64 0.48 12.40 … 73.72 3.76 二次碳化物 5 17.21 … 31.67 … 45.08 6.03 主要的硬質(zhì)合金 6 … 1.05 5.33 … 92.23 1.39 矩陣 硬度測量 在標(biāo)準(zhǔn)ASTM E384后，從失敗的錘頭制備的樣品的截面上測量了宏觀和顯微硬度值。在Vickers硬度試驗(yàn)機(jī)中，用30kgf的負(fù)載測量了宏觀硬度值。在橫截面隨機(jī)位置進(jìn)行了5次測量，得到平均宏觀硬度值;macro-hardness值被發(fā)現(xiàn)781±12 HV30(相當(dāng)于&64 HRC)。AISI A2工具鋼材料的宏觀硬度較高，由于材料的硬度較高，對其韌性有不利影響。顯微硬度值在500 lm的一個(gè)固定間隔內(nèi)，由一組50gf的微維氏硬度試驗(yàn)機(jī)和15 s的壓痕持續(xù)時(shí)間進(jìn)行測量;顯微硬度剖面如圖7(a)所示。硬度值在600-1000 HV0.05的范圍內(nèi)變化。除硬度剖面外，還分別測量了一些晶界碳化物網(wǎng)絡(luò)的顯微硬度值。電石網(wǎng)絡(luò)的平均顯微硬度測定為1375 15 HV0.05;在現(xiàn)有文獻(xiàn)[12,16]中，發(fā)現(xiàn)碳化物網(wǎng)絡(luò)的微硬度值與M7C3型主要碳化物的值(1400 HV)相似。 圖7在橫截面上測量的顯微硬度剖面錘頭 (a)破壞的錘頭 (b)在推薦的熱處理之后 測量沖擊韌性 在標(biāo)準(zhǔn)沖擊試驗(yàn)機(jī)（Striking Energy：300±10 J）的幫助下，在環(huán)境溫度下，根據(jù)IS 1757：1988 [17]，使用V型缺口夏比試樣進(jìn)行樣品沖擊試驗(yàn)。 至少進(jìn)行了三次測試以獲得打擊頭樣品的平均沖擊能量值。 樣品的平均沖擊能量值為3±0.2 J。 在實(shí)驗(yàn)室中對打頭的熱處理進(jìn)行改進(jìn) 攪拌頭制造商給出的熱處理產(chǎn)生了具有顯著粗粒的微觀結(jié)構(gòu)邊界碳化物網(wǎng)絡(luò)。 粗晶界碳化物網(wǎng)絡(luò)不利地影響材料的韌性，使其在沖擊下容易斷裂[2,8]。 在目前的調(diào)查中，對一些鑄型打頭材料(由同一制造商提供)進(jìn)行了適當(dāng)?shù)臒崽幚怼崽幚硎窃诳煽貧夥諣t中進(jìn)行的。材料被預(yù)熱到788攝氏度，并保持在這個(gè)溫度直到被浸泡。然后,加熱到954°C和1 h /in。最大的橫截面。奧氏體化后,材料從熔爐中取出并在空氣中冷卻，然后在兩個(gè)階段立即回火，溫度為200-250攝氏度。 熱處理后的顯微結(jié)構(gòu)表征 在實(shí)驗(yàn)室中，用“顯微結(jié)構(gòu)檢查”一節(jié)所述的方法，制備了一種樣品，并在實(shí)驗(yàn)室中對其進(jìn)行了熱處理。光學(xué)顯微結(jié)構(gòu)如圖8(a)和(b)所示，微觀結(jié)構(gòu)主要表現(xiàn)為馬氏體基體，在晶界處有少量初級共晶碳化物。但這些碳化物的形式并不是連續(xù)的網(wǎng)絡(luò),而他們是不連續(xù)的,出席一些晶界上的離散位置聯(lián)合國——就像觀察圖5(a)和(b)。掃描電子mi - croscopy(SEM)以及碳化物的EDS分析進(jìn)行了研究其形態(tài)和類型。SEM和EDS分析表明，在晶界和許多球狀精細(xì)碳化物(副碳- m23c6型)的晶界和大量球狀精細(xì)碳化物的分布均勻分布在整個(gè)基質(zhì)中，如圖9和表3所示。 圖8 (a)在建議熱處理后，錘頭的微結(jié)構(gòu)(950)(b)微觀結(jié)構(gòu)(9200)在前奧氏體晶界的離散位置上顯示出具有不連續(xù)碳化物的馬氏體基體。 圖9在前奧氏體晶界處的不連續(xù)碳化物網(wǎng)絡(luò)(92500)以及在基體內(nèi)的精細(xì)球狀碳化物析出物。 實(shí)驗(yàn)室熱處理后的機(jī)械特性 在實(shí)驗(yàn)室中對鑄坯頭塊進(jìn)行熱處理后，對試樣進(jìn)行硬度和沖擊韌性值的測試，分別在“硬度測量”和“沖擊韌性測量”各部分進(jìn)行分析。macro-hardness值測量是670±10高壓(相當(dāng)于59 HRC)。如圖7(b)所示，在樣品的橫截面上測量的mi- cror硬度值的剖面如圖7(b)所示。在實(shí)驗(yàn)室處理的beater頭塊熱處理的硬度值(600-750 HV)比在制造商端處理的600 - 1000hv更低。目前的硬度值與典型AISI A2工具鋼的(60 HRC)一致[10,18]。此外，測量的顯微硬度剖面(圖7b)與較早的測量值相對平滑或均勻。用v形缺口試樣進(jìn)行了熱處理后的錘頭試樣的沖擊試驗(yàn)。能量值的影響被測量是6.5±0.3 J顯示顯著改善(117%)相比,韌性的制造商。這兩種樣品的硬度剖面和沖擊韌性的差異(即:由于采用了新的熱處理計(jì)劃，生產(chǎn)出的熱處理樣品和現(xiàn)有的實(shí)驗(yàn)室熱處理樣品都是由原始碳化物的粗化網(wǎng)絡(luò)相對自由的。碳化物的粗連續(xù)網(wǎng)絡(luò)增加了硬度，但在沖擊荷載作用下，粗碳化物管網(wǎng)破壞的韌性增加了[2,8]。 討論 通過分析，找出了沖擊破碎機(jī)錘頭失效的根本原因，并通過適當(dāng)?shù)臒崽幚矸椒▽ζ溥M(jìn)行了開發(fā)。失敗的打頭樣品的微觀結(jié)構(gòu)和力學(xué)特性均表明熱處理不當(dāng)。在鑄造過程中形成的粗硬碳化物網(wǎng)絡(luò)是有害的;在熱處理過程中，由粗連續(xù)碳化物網(wǎng)絡(luò)形成均勻分布在基體上的次生碳化物的細(xì)球狀沉淀物。 使它們在某些位置不連續(xù)或分離，而不是減少它們的數(shù)量[5]。由制造商給出的熱處理時(shí)間表(圖4)顯示在淬火過程中，在淬火過程中，奧氏體化溫度較低[2,18]。在熱處理過程中，由于奧氏體化溫度較低，產(chǎn)生精細(xì)次生碳化物的mas-主要碳化物并沒有完全溶解在基體中;這產(chǎn)生了粗碳化物網(wǎng)絡(luò)，少量的精細(xì)碳化物析出，使材料脆性或韌性較差[2,5,19]。 建議制造商采用“改善實(shí)驗(yàn)室中攪拌器頭的熱處理”一段適當(dāng)?shù)臒崽幚碛?jì)劃，以改善攪拌器頭的微觀結(jié)構(gòu);奧氏體化tem -perature預(yù)熱后增加到954°C在788°C和回火溫度增加到200 - 200°C。新的熱處理保證了顯著的效果。 在基體中粗碳化物的粗化網(wǎng)絡(luò)的溶解，然后是均勻分布在整個(gè)基體(圖8和圖9)的細(xì)小的球狀次生汽車-等待粒子的沉淀。推薦的熱處理時(shí)間安排導(dǎo)致了顯微結(jié)構(gòu)的改善，其硬度和韌性都得到了最佳的組合。材料的沖擊韌性從3增加到3.6.5 J。，根據(jù)a級AISI A2的規(guī)格要求，117%的硬度值。硬度剖面(圖7b)相對于較早的一個(gè)確保更均勻的矩陣，也表現(xiàn)出相對平穩(wěn)的趨勢。按照推薦的熱處理時(shí)間表生產(chǎn)的新噴頭，事實(shí)上，比早期的更適合于人工。服務(wù)期限為15天至2個(gè)月，即通過4次。新錘頭的失效模式也從突然的脆性斷裂轉(zhuǎn)變?yōu)橹饾u磨損在突出在表面上。 結(jié)論 從上述分析可以得出以下結(jié)論: (1)對AISI A2刀具鋼錘頭的斷口形貌和斷口形貌的觀察，表明其在沖擊載荷作用下，在脆性模式下的針孔位置失效。 (2)微觀結(jié)構(gòu)檢查發(fā)現(xiàn)了大量粗顆粒的粗晶粒邊界網(wǎng)絡(luò)。粗碳鋼網(wǎng)的硬度高(64 HRC)，韌性較低(3 J)，使其易碎。 (3)對beater head的熱處理進(jìn)度表進(jìn)行了分析，表明其不適當(dāng)?shù)奈⒔Y(jié)構(gòu)是由于低奧氏體化(850-870 C)的結(jié)果，以及在制造商端熱處理時(shí)的溫度(175-200 C)。 (4)新推薦的熱處理時(shí)間表(預(yù)熱788°C,奧氏體化在954°C符合低下由兩個(gè)階段回火200 - 500°C)導(dǎo)致少量的不連續(xù)和大量的細(xì)碳化物網(wǎng)絡(luò)分布式pre - cipitates回火馬氏體內(nèi)部的二次碳化物矩陣。改進(jìn)后的顯微結(jié)構(gòu)產(chǎn)生了硬度(59 HRC)和韌性(6.5 J)的最佳組合，使錘頭的使用壽命提高了4倍。 參考文獻(xiàn) [1] Y. Umucu, V. Deniz, S. Cayirli, A new model for comminution behavior of different coals in an impact crusher. Energy Sources A 36(13), 1406–1413 ,2014. [2]M. Pellizzari, A. Molinari, The application-oriented heat treat- ment of tool steels. Mater. Manuf. Process. 24(7–8), 723–728 ,2009. [3] B. Smoljan, An analysis of relationships between behavior and microstructure constitution of hot-work tool steel. Mater. Manuf. Process. 24(7–8), 786–790 ,2009. [4]P. Hansson, Modern prehardened tool steels in die-casting applications. Mater. Manuf. Process. 24, 824–827 ,2009 [5]A.E. Karantzalis, A. Lekatou, H. Mavros, Microstructural modifications of as-cast high-chromium white iron by heat treatment. J. Mater. Eng. Perform. 18, 174–181 ,2009. [6]J.O. Agunsoye, V.S. Aigbodion, O.S. Sanni, Effect of heat treatment on microstructure and mechanical properties of NF 6357A cast alloy for wear resistance application. J. Miner. Mater. Charact. Eng. 10(11), 1077–1086 ,2011. [7]C.R. Gagg, Premature failure of thread rolling dies: material selection, hardness criteria and case studies. Eng. Fail. Anal. 8, 87–105,2001. [8]M.F.Mirjana,Iron-chromium-carbon-vanadiumwhitecastirons—microstructure and properties. Chem. Ind. 68(4), 413–427 ,2014. [9] G. Mukhopadhyay, S. Bhattacharyya, Cracking of a tool steel guide roller in a Bar Mill. J. Mater. Eng. Perform. 12, 594–603 ,2012. [10]D. Viale, J. Be′guinot, F. Chenou, G. Baron, Optimizing microstructure for high toughness cold-work tool steels, in Pro- ceedings of the 6th International Tooling Conference, Use of Tool Steels: Experience and Research, ed. by J. Bergstrom, G. Fre- driksson, M. Johansson, O. Kotik, F. Thuvander ,Karlstad University, Karlstad, 2002. [11] N. Mebarki, P. Lamesle, D. Delagnes, C. Levaillant, F. Delmas, Relationship between microstructure and mechanical properties of a 5%Cr hotwork tool steel, in Proceedings of the 6th Inter- national Tooling Conference, Use of Tool Steels: Experience and Research, ed. by J. Bergstrom, G. Fredriksson, M. Johansson,O. Kotik, F. Thuvander ,Karlstad University, Karlstad, 2002. [12] A. Ray, D. Mukherjee, B. Sarkar, S. Mishra, Influence of mi- crostructure on the premature failure of a second-intermediatesendzimir mill drive roll. J. Mater. Eng. Perform. 3(5), 649–656 ,1994. [13] K.M. Ibrahim, M.M. Ibrahim, Heat treatment in high chromium white cast iron Ti alloy. J. Metall. 2014, Article ID 856408 ,2014. doi:10.1155/2014/856408 [14]N.F. Garza-Montes-de-Oca, R. Cola′s, W.M. Rainforth, On the damage of a work roll grade high speed steel by thermal cycling. Eng. Fail. Anal. 18, 1576–1583 ,2011. [15] M.M. Dhobe, I.K. Chopde, C.L. Gogte, Investigations on surface characteristics of heat treated tool steel after wire electro-dis- charge machining. Mater. Manuf. Process. 28(10), 1143–1146 ,2013. [16] F.J. Belzunce, A. Ziadi, C. Rodriguez, Structural integrity of hot strip mill rolling rolls. Eng. Fail. Anal. 11, 789–797,2004. [17]Bureau of Indian Standards, Method for Charpy Impact Test ,V notch) for Metallic Material, IS 1757 (Bureau of Indian Standards, New Delhi, 1988. [18] ASTM International, Standard Specification for Tool Steels Alloy, ASTM A681-08 ,ASTM International, West Conshohocken, 2008. [19] R. Schneider, J. Perko, G. Reithofer, Heat treatment of corrosion resistant tool steels for plastic moulding. Mater. Manuf. Process. 24, 903–908 ,2009. 12 TECHNICAL ARTICLE Development of AISI A2 Tool Steel Beater Head for an Impact Crusher in a Sinter Plant Goutam Mukhopadhyay ? Piyas Palit ? Sandip Bhattacharyya Received: 15 December 2014 / Revised: 23 February 2015 / Accepted: 23 February 2015 / Published online: 7 March 2015 C211 Springer Science+Business Media New York and ASM International 2015 Abstract Failure analysis of a tool steel (AISI A2) beater head of an impact crusher and development of suitable heat treatment process to improve its performance have been presented. The beater heads were failing prematurely by brittle fracture from its pin-hole locations. The investigation consisted of visual inspection, fractography, chemical ana- lysis, characterization of microstructures using optical and scanning electron microscopes (SEM), EDS analysis, and determination of micro-hardness profile. Microstructural characterization using SEM and EDS analysis revealed significant amount of coarse continuous Cr-carbide networks in the martensite matrix. It increased hardness (64 HRC) as well as heterogeneity of the matrix as depicted by the micro- hardness profile, and decreased the toughness (3 J) since coarse carbide networks are very hard and brittle. The austenitizing temperature as well as tempering temperature of heat treatment was found lower at the manufacturer’s end. The new recommended heat treatment resulted in lower amount of discontinuous Cr-carbides along with significant amount of fine precipitates uniformly distributed throughout the matrix which led to an optimum combination of both hardness (59 HRC) and toughness (6.5 J) required for the application. The beater heads manufactured following the recommended heat treatment exhibited better performance (life increased by 4 times) compared to the earlier ones. Keywords Impact crusher C1 Hammers C1 AISI A2 C1 Tool steel C1 Carbide network C1 Brittle fracture C1 Toughness C1 Heat treatment Introduction Sinter is used as a raw material in blast furnace for iron making. Sintering is a process of agglomeration of base mix into a porous mass, by incipient fusion. The base-mix materials are prepared by proportionate mixing of iron ore fines, ground flux materials, ground coke breeze, and revert mix, followed by layer wise stacking of the same in the storage yard. The granulation of raw materials for base mix is an important quality factor for sinter. The correct granulation is achieved by crushing in an impact crusher, followed by sieving and mixing different granulations in appropriate ratios. Limestone and dunite, used as fluxes in sinter, are reclaimed by means of a wheel on boom re- claimer and sent to three 150 tph Hammer Mills, called primary crushers from where they are then sent to sec- ondary crushers (Hammer Mills), where they are ground to a size of 3.2 mm. After crushing, these materials are screened and stored in proportioning bins. The impact crusher is extensively used for crushing the raw materials like limestone, dolomite, and dunite in the bedding and blending plant for sinter preparation. It crushes the raw materials using multiple rotating beater heads (hammers) by impact against the grinding liners fixed on the impact walls of the machine [1]. The beater heads are connected to a rotor shaft with the help of beater arms arranged circumferentially in a zig-zag fashion (Fig. 1). The fineness of the granulated product of the crusher can be altered by adjusting the distance between the rotating hammer heads and the impact walls. The beater heads of the impact crusher were failing very frequently (within 15 days) under brittle fracture. Gradual wear of the beater head at the striking face is a normal mode of failure while its sudden brittle fracture is a con- cern. The hammer heads or beater heads in an impact G. Mukhopadhyay ( goutam2007@yahoo.co.in 123 Metallogr. Microstruct. Anal. (2015) 4:114–121 DOI 10.1007/s13632-015-0192-6 crusher are exposed not only to wear, but also to impact or shock loads. Therefore, toughness of the beater head material is an important factor for extended service life of the hammers apart from its hardness or wear resistance properties [2–6]. A huge number of beater heads had to be replaced every month because of their frequent breakage which led to high maintenance cost and loss in production. Several works on alloy cast irons and tool steels like AISI H11, H13, D2 and M2 are available in the existing lit- erature [2–10] which deals with the microstructural and mechanical characterizations. But work on air hardening tool steel AISI A2 and its application as a beater head material in impact crusher (hammer mill) is rare in the open literature. Further, no industrial or application data on the performance of AISI A2 tool steel as a crushing ham- mer is available in the existing literature. This work has presented the analysis of root cause for the premature failure of the AISI A2 beater heads and develop- ment of suitable heat treatment process to improve their performance. Experimental Procedure and Results Site Visit and Visual Observation The sinter plant has two raw materials bedding and blending (RMBB) units, namely, RMBB#1 and RMBB#2, which have 6 and 7 impact crushers. The impact crushers crush the flux materials used in sintering like lime stone, dolomite, pyroxenite, etc. The primary crushers which ro- tate at an rpm of around 675 first crush the raw materials from a mesh size of -50 mm into a size of -15 mm. These are then again crushed with the help of secondary crushers having an rpm of around 830 into a size of -3.15 mm. The capacity of a crusher is approximately 125 t/h and it contains 40–52 beater heads. Figure 1(a) shows a rotor shaft-beater head assembly. Beater heads are attached to the beater arms with the help of pins inserted through the holes present in the beater heads as shown in Fig. 1(b). Figure 2(a) and (b) shows failed beater heads which fractured from their pin-hole locations. The fracture surfaces (Fig. 2c and d) of the failed beater head reveal bright granular appearance suggesting brittle fracture. The other end of the beater head was found to be worn out (Fig. 2a and b) due to abrasion or impact with input flux materials while crushing. While gradual wear of the beater head surface is a desired mode of failure, fracture from the pin-hole locations during service is a concern which must be addressed. Fractography A small sample containing fracture surface was cut near the location of pin-hole from where the fracture initiated. The sample was ultrasonically cleaned for examining the frac- ture surface. The fracture surface was examined using a scanning electron microscope (SEM) (model: JXA6400, JEOL, Japan) operated at an accelerating voltage of 15 kV. Micrographs at various locations of the fracture surface were recorded at various magnifications. Examination of the fracture surface near crack initiation region as shown in Fig. 3 revealed cleavages indicating brittle fracture. Brittle fracture of the beater head suggests its failure under impact load during service. Materials A small piece of sample was cut from the failed beater head and prepared for its chemical analysis. Chemical analysis of the sample was carried out using x-ray fluorescence spectroscopy (XRF); carbon (C) and sulfur (S) content of the sample were determined using combus- tion infrared technique. The chemical analysis of the beater head is compiled in Table 1. The chemical analysis is found to be closer to AISI A2 (ASTM A681) grade of air hardening medium-alloy cold-work tool steel with marginally higher amount of chromium (Cr) and lower (a) (b) Beater Head Beater Arm Rotor shag332 Beater Head Beater Arm Pin Fig. 1 (a) Impact crusher showing rotor shaft-beater head assembly, (b) closer view of a beater head Metallogr. Microstruct. Anal. (2015) 4:114–121 115 123 amount of vanadium (V). Manganese, chromium, and molybdenum are the principal alloying elements in this grade of steel, which impart high hardenability and the steel can be hardened in air. AISI A2 provides an optimum combination of wear resistance and toughness required for crushing operation of raw materials in service. The heat treatment cycle given to the beater head as received from the manufacturer is presented in Fig. 4. The as-cast mate- rial was heat treated at 850–870 C176C for 1 h/in. followed by air cooling. The air-cooled material was tempered in two stages in the temperature range of 175–200 C176C for 3 h/in. Microstructural Examination A sample was cut from the failed beater head for microstructural examinations at the cross-section. The sample was then mounted in resin, ground, and polished using standard metallographic technique. The cross-section was examined under optical microscope (Leica, model: DMRX, Germany) after etching using Villela’s reagent (1 g picric acid, 5 mL hydrochloric acid, and 100 mL ethanol). Typical micrographs at the cross-section of the broken beater head are shown in Fig. 5(a) and (b). Mi- crostructural examination at the cross-section of the beater head reveals martensite matrix with networks of chain-like massive primary carbides at the grain boundaries (Fig. 5a and b). The carbides appear as bright phases which are preferentially clustered along grain boundary. The area fractions of carbides were measured at various fields using image analysis software. The average area fraction of carbides was found to be 7.2 ± 0.45%. The most significant material properties required for the beater head component are hardness, toughness, and wear resistance. As the amount of carbides increases, the hardness and wear resistance also increase but care has to be taken in heat treatment to avoid loss of toughness [2, 3, 6, 7]. The clusters of coarse carbides detrimentally affect the toughness; the carbides being brittle are sus- ceptible to initiate cracks during an impact in service [2, 8, 9]. (a) (b) (c) (d) A A B B Wear Pin hole Pin hole Fig. 2 (a, b) Beater head samples failed from the pin-hole locations, (c, d) closer view of fracture surfaces near the pin- hole of the beater heads Fig. 3 Fractography (at 92000) of failed beater head shows cleav- ages indicating brittle fracture 116 Metallogr. Microstruct. Anal. (2015) 4:114–121 123 Scanning Electron Microscopy and EDS Analysis The etched cross-section of the sample was examined with the help of the scanning electron microscope (SEM) operated at an accelerating voltage of 15 kV for its microstructural as well as elemental characterizations. Micrograph showed coarse carbide network at the grain boundary along with some fine globular precipitates with the martensitic matrix (Fig. 6a and b). Energy dispersive spectroscopy (EDS) of the grain boundary network as well as fine precipitates (as marked in Fig. 6a and b) was carried out for their elemental charac- terization and the results of the analyses were compiled in Table 2. The results of EDS analysis indicate that the grain boundary network and the fine precipitates within the matrix are chromium carbides. The carbide stoichiometry could not be determined by EDS micro-analysis. But literature survey on similar materials [2, 10, 11] like AISI H11, H13, M2, D2 and so on, and electron probe micro analysis of similar car- bides discussed in earlier literature [5, 12–15] suggests the grain boundary carbide network to be of M 7 C 3 type (Primary Carbide) and the fine globular precipitates as M 23 C 6 (where M = Cr) type (Secondary Carbide). The skeleton-like mor- phology of massive carbide network at the grain boundary as shown in Fig. 6(b) is apparently indicative of its eutectic origin (primary carbide), i.e., origin at the solidification stage [2, 5, 8–10, 13]. On the other hand, fine globular precipitates within the matrix indicate secondary carbides which form during heat treatment of the casting [2, 5, 8, 11]. Measurement of Hardness Both macro- and micro-hardness values were measured at the cross-section of the sample prepared from the failed beater head following the standard ASTM E384. Macro-hardness values were measured in a Vickers hardness testing machine with a load of 30 kgf. Five measurements were taken at random locations of the cross-section to get the average macro-hardness value; the macro-hardness value was found to be 781 ± 12 HV30 (equivalent to the micro-hardness profile is shown in Fig. 7(a). The hardness values are found to be varied in the range of 600–1000 HV0.05. Apart from hardness profile, micro-hardness values on some grain boundary carbide network were also measured separately. The average micro-hardness of carbide network was measured to 1375 ± 15 HV0.05; the measured micro-hard- ness value of carbide network is found to be similar to the value (1400 HV) for M 7 C 3 type of primary Cr-carbide reported in other existing literature [12, 16]. Measurement of Impact Toughness The impact tests of the samples were carried out using V-notch Charpy specimens in accordance with IS 1757:1988 [17] with the help of a standard impact testing machine (Striking Energy: 300 ± 10 J) at ambient temperature. At least three tests were carried out to get the average impact energy values of the beater head sample. The average impact energy value of the samples was 3 ± 0.2 J. Heat Treatment of Beater Head in Laboratory for Improvement The heat treatment given by the manufacturer of the beater head yielded a microstructure with significant coarse grain Table 1 Chemical analysis (wt%) of beater head sample Sample C Mn Si S P Cr Mo V W Beater head 1.01 0.67 0.50 0.030 0.026 5.7 1.02 0.12 0.010 ASTM A681 type A2 0.95–1.05 0.4–1 0.1–0.5 0.03 max 0.03 max 4.75–5.5 0.9–1.4 0.15–0.5 … Temperature ( °C) Time 850-870°C (1h/inch) 175-200°C (3h/inch) 175-200°C (3h/inch) Hardening Tempering 780° - 790°C (soaking) Temperature ( °C) Time 950-960°C (1h/inch) 250°C (3h/inch) 200°C (3h/inch) Hardening Tempering Room Temperature (a) (b) Room Temperature Fig. 4 (a) Typical heat treatment schedule given by the manufacturer of beater head, and (b) recommended heat treatment cycle of beater head Metallogr. Microstruct. Anal. (2015) 4:114–121 117 123 boundary carbide networks. The coarse grain boundary carbide network adversely affects the toughness of the material making it prone to fracture under impact [2, 8]. During present investigation, a suitable heat treatment was given to some as-cast beater head material (supplied by the same manufacturer). The heat treatment was carried out in a controlled-atmosphere furnace. The material was pre- heated to 788 C176C and held at this temperature until thor- oughly soaked. Then, it was heated to 954 C176C and held for 1 h/in. of greatest cross-section. After austenitization, the material was removed from the furnace and cooled in air followed by immediate tempering in two stages at a tem- perature of 200–250 C176C. Microstructural Characterization After Heat Treatment in Laboratory A sample was prepared form the beater head block heat treated in the Laboratory for examination of the microstructure fol- lowing the procedure as described in ‘‘Microstructural Examination’’ section. The optical microstructures are shown in Fig. 8(a) and (b). The microstructures reveal predominantly Fig. 5 Microstructures of failed beater head: (a) Microstructure (at 950) shows martensite matrix along with carbide networks at the grain boundaries, and (b) microstructure at magnified view (at 9200) shows continuous carbide network at the prior austenite grain boundary Fig. 6 SEM micrographs at the cross-section of beater head sample: (a) micrograph (at 91000) shows coarse carbide network at the grain boundary along with fine precipitates within martensite matrix, and (b) micrograph (at 91500) shows locations of EDS analysis on the coarse carbide network as well as fine precipitates within the matrix Table 2 Results of EDS analysis (wt%) at different locations as shown in Fig. 6 Locations C Si Cr Mn Fe Mo Remarks 1 11.71 0.40 16.68 0.68 64.5 6.03 Secondary carbide 2 19.35 … 40.92 … 30.63 8.71 Primary carbide 3 … 1.00 4.83 0.50 92.71 0.96 Matrix 4 9.64 0.48 12.40 … 73.72 3.76 Secondary carbide 5 17.21 … 31.67 … 45.08 6.03 Primary carbide 6 … 1.05 5.33 … 92.23 1.39 Matrix 0 200 400 600 800 1000 1200 1 2 3 4 5 6 7 8 9 10111213141516171819 Micro-vickers hardness (Hv) Micro-Hardness Profile Failed Beater Head (heat treated at Manufacturer's end) Ag332er recommended heat treatment (a) (b) Fig. 7 Micro-hardness profiles measured at the cross-section of the beater heads: (a) Failed beater head, and (b) after recommended heat treatment 118 Metallogr. Microstruct. Anal. (2015) 4:114–121 123 martensitic matrix with a small amount of primary eutectic carbides at the grain boundary. But these carbides are not in the form of continuous network rather they are discontinuous and present at some discrete locations on the grain boundaries un- like that observed in Fig. 5(a) and (b). Scanning electron mi- croscopy (SEM) as well as EDS analysis of the carbides was carried out to study their morphology and type. SEM and EDS analyses show discontinuous primary Cr-carbides (M 7 C 3 type) at the grain boundary and numerous globular fine carbide par- ticles (secondary carbides-M 23 C 6 type) uniformly distributed throughout the matrix as illustrated in Fig. 9 and Table 3. Mechanical Characterization After Heat Treatment in Laboratory After carrying out the heat treatment of the as-cast beater head block in the Laboratory, the sample was tested for hardness and impact toughness values following the pro- cedure described in ‘‘Measurement of Hardness’’ and ‘‘Measurement of Impact Toughness’’ sections, respec- tively. The macro-hardness values were measured to be 670 ± 10 Hv (equivalent to during heat treatment fine globular precipitates of secondary carbides which were uniformly distributed throughout the matrix were formed from the coarse continuous carbide network Fig. 8 (a) Microstructure (at 950) of beater head after suggested heat treatment, (b) Microstructure (at 9200) shows martensite matrix with discontinuous carbides at discrete locations on the prior austenite grain boundary Fig. 9 Discontinuous carbide network (at 92500) at prior austenite grain boundary along with fine globular Cr-carbide precipitates within the matrix Metallogr. Microstruct. Anal. (2015) 4:114–121 119 123 making them discontinuous or isolated at certain locations apart from reducing their amount [5]. The heat treatment schedule (Fig. 4) given by the manufacturer showed a lower austenitizing temperature during hardening as well as temperature during tempering [2, 18]. Because of lower austenitizing temperature during heat treatment, the mas- sive primary carbides which generated fine secondary carbides afterwards were not fully dissolved in the matrix; this yielded coarse carbide network with a little amount of fine secondary carbide precipitates imparting brittleness or poor toughness to the material [2, 5, 19]. A proper heat treatment schedule as described in ‘‘Heat treatment of Beater Head in Laboratory for improvement’’ section was recommended to the manufacturer to improve the microstructure of beater head; austenitizing tem- perature was increased to 954 C176C after preheating at 788 C176C and tempering temperature was increased to 200–250 C176C. The new heat treatment ensured significant dissolution of coarse primary carbide network in the matrix followed by precipitation of fine globular secondary car- bide particles uniformly distributed throughout the matrix (Figs. 8 and 9). The recommended heat treatment schedule resulted in improved microstructure having an optimum combination of both hardness and toughness. The impact toughness of the material was found to increase from 3 to 6.5 J, i.e., by 117% with a hardness value desired as per the specification of the grade AISI A2. The hardness profile (Fig. 7b) also exhibited a relatively smooth trend compared to the earlier one ensuring a more uniform matrix. The new beater he