激光冲击处理对6082-T651铝合金高周疲劳裂纹扩展速率和断裂韧性的影响外文文献翻译、中英文翻译
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激光冲击处理对6082-T651铝合金高周疲劳裂纹扩展速率和断裂韧性的影响外文文献翻译、中英文翻译,激光,冲击,处理,6082,T651,铝合金,疲劳,裂纹,扩展,速率,断裂韧性,影响,外文,文献,翻译,中英文
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外文题目 Effects of laser shock processing on high cycle fatigue crack growth rate and fracture toughness of aluminium alloy 6082-T651 译文题目 激光冲击处理对6082-T651铝合金高周疲劳裂纹扩展速率和断裂韧性的影响 外文出处 International Journal Of Fatigue 激光冲击处理对6082-T651铝合金高周疲劳裂纹扩展速率和断裂韧性的影响摘要： 研究无保护涂层的激光冲击加工对高周疲劳裂纹扩展和断裂韧性的影响。 激光冲击强化处理是在垂直于裂纹生长方向的两侧对紧凑拉伸试样进行，然后进行后续研磨。 疲劳裂纹扩展试验在116和146 Hz之间的频率下进行，在R = 0.1和疲劳裂纹萌生阶段和K减少期间的恒定应力强度范围内进行。疲劳预裂纹需要的循环次数较少，并且LSP处理的试样中的疲劳裂纹扩展更快。LSP处理后，裂纹增长值下降了60。LSP处理后断裂韧性下降28-33。断裂表面的疲劳延展性过渡边界显示未处理样品中的线性疲劳裂纹前沿和LSP处理后的曲线。2016 Elsevier Ltd.保留所有权利。1介绍激光冲击处理（LSP）是一种相对较新的表面处理技术，可延长动态加载元件的使用寿命。几项研究已经证明LSP通过压力冲击波引起的材料局部塑性变形产生大振幅压应力来降低疲劳裂纹扩展速率。 由于疲劳裂纹通常从表面开始，所以在外部载荷下，锁定的压缩应力将叠加在所施加的应力上，导致裂纹开端处的应力强度因子更小并且可能的裂纹闭合导致有效驱动力的降低。由于位错和显微组织的晶粒细化，LSP也可以提高疲劳强度和疲劳裂纹萌生寿命。 Huang等人证实了由于在Al-Mg-Si合金的LSP处理表面中高度缠结。致密的位错排列而降低的FCG速率。在另一项研究Lu等人报道了LSP后Al-Cu-Mg合金显微组织细化。 结果证实，LSP是获得高密度位错以提高耐疲劳性的有前途的方法，而多次LSP冲击后顶面中的最小晶粒尺寸大约是100-200纳米。此外，与传统的喷丸硬化相比，激光加工产生的残余应力可以显着更高，并且材料中更深的有效穿透。Huang等人报道了不同工艺设置下LSP对6061 T6铝合金疲劳裂纹扩展性能的有益影响。他们的结果表明FCG率随着激光能量的增加而降低和LSP覆盖区域特别是在疲劳裂纹扩展初期。然而，随着裂纹扩展，残余应力释放，疲劳裂纹扩展的最后阶段的强化效应较弱。在另一项研究中，Hatamleh 等比较了搅拌摩擦焊接的Al-Zn合金（7075-T7351）接头中FCG的激光和喷丸强化效果，报告与LSP处理试样的基材相比，FCG率降低，而喷丸硬化显示可忽略不计的改善的FCG速率。然而，尽管使用保护涂层的LSP工艺已经被证明是用于改善各种金属和合金的疲劳寿命，耐腐蚀性，硬度和耐磨性。但是存在一些限制这一过程广泛应用的负面因素。关于最后一次，即2015年辛辛那提第5届激光喷丸及相关现象国际会议的公开讨论10 指出在工业现场应用中需要新的策略。基本上，激光冲击处理实现可观的位错合并和压缩残余应力的产生是两种主要方法：（i）第一种方式，使用高能量激光脉冲，与保护涂层结合的脉冲持续时间长达100 ns。（ii）第二种方法使用较少的焦耳或更少的次序，增加与重叠。较小的斑点和无涂层（也称为LSPwC）的较低激光能量。尽管加工零件上的保护涂层可以防止表面烧蚀，同时保持较高的表面质量，但这是一件非常耗时的事情，因为在LSP过程中覆盖层受到严重损坏，并且需要频繁更换，从而使其在工业应用中变得缓慢并且昂贵。此外，由于部分喷丸处理可能无法应用保护性覆盖，因此人们非常关注LSP处理，而不需要保护性覆盖层。 然而，在该过程之后，由于局部表面烧蚀而获得具有增加的表面粗糙度和波纹的典型表面凹坑，其继而产生高压等离子体，由于在样品表面上流动的薄层水所包含，产生压力波传播到标本中。基于公开的文献，关于疲劳裂纹扩展的没有保护涂层的激光冲击加工已经进行了很少的全面研究，并且在随后的研究程序中没有评估疲劳特性以消除未涂保护层的LSP诱导的表面。Rubio-Gonzalez等人研究了双层非涂层LSP工艺对2205双相不锈钢和6061-T6铝合金的紧密张力试样的影响。结果证实，随着脉冲密度的增加，FCG速率下降而断裂韧性增加。 尽管有许多关于激光加工对材料疲劳裂纹扩展性能影响的研究，但是关于激光冲击处理的不利影响的文献是有限的。此外，所有上述研究都是在预先已知的疲劳预裂纹，低正弦波频率和低周疲劳状态下的高最大负荷之后进行的。考虑到这一点，并且早期认识到过早失效在工程部件中起着至关重要的作用，未涂层LSP对疲劳裂纹扩展速率影响的详细研究将得到评估。特别关注于确定高谐振频率下动态载荷下的疲劳裂纹扩展行为， 使用小负载和高循环疲劳方案测量裂纹生长速率。为了获得静态载荷下最终破裂阶段和未处理LSP试样的断裂韧性的定量比较。通过位错排列，残余应力和深度分布中的显微硬度，由于局部表面熔化，电离，和激光过程中的再凝固而可能产生的软化效应。所有正在研究的LSP样品都要在后续的研磨过程中进行分析，以获得合适的表面光洁度。2试验方法2.1材料和标本制备对商业锻制的Al-Mg-Si铝合金板用于疲劳裂纹扩展和断裂韧性测试。三个样本在未处理的条件下进行测试，并且三个样本用不同的参数集进行LSP处理：LSP 1，LSP 2和LSP 3.使用0.35mm电线的放电加工制造凹口形成一个半径为0.19毫米的凹口。用于Rumul Cracktronic装置和Krak-gauge AMF-5测试的试样的几何形状见图2。使用所施加的应力强度范围DK，根据该等式计算。2.2激光冲击处理使用具有1064nm的照射波长：YAG激光器进行激光冲击处理。最大激光束能量为2.8J/脉冲，而产生的高斯强度脉冲的FWHM（全宽度半最大值）为10ns。聚焦激光束直径通过会聚透镜的修改而变化。在LSP 1，LSP 2和LSP 3处理中使用的光束直径设定为1.5，X2.0和2.5毫米。在这项研究中，选择了1600脉冲/cm的预定重叠脉冲密度，其中激光前进方向平行于轧制方向（L）。使用激光冲击处理方法在没有任何保护涂层的情况下对样本进行照射，而激光脉冲重叠并且以锯齿形图案扫描（图3）。使用设置的水射流形成薄水层并避免形成水泡或由材料烧蚀产生的杂质浓度，因此，不断确保纯激光物质相互作用。C（T）标本上的治疗区域大约是10mm在样品的两侧。对于透射电镜微观结构分析，额外的LSP标本准备使用。图1.拉伸和C（T）标本取向铝表格1平均测量的6082-T651板的拉伸性能负载方向 L T拉伸强度 - R（MPa） 311 310屈服应力 - R（MPa） 296 295伸长率 - A（） 9.7 8.5图2 C（T）试样的几何形状和测量电阻计AMF-5 图3.激光冲击处理设置和处理的C（T）样品图4 C（T）试样（a）不处理，（b）LSP 1处理后，（c）研磨后的LSPLSP处理后（图4）。 在激光冲击处理之前，没有启动预裂纹。 然而，用于TEM分析的样本没有以任何方式磨碎以获得由激光喷丸处理产生的代表性微结构。2.3表面粗糙度在接近预加工切口的表面经过激光表面处理之后，使用Surtronic 轮廓测量仪（Taylor Hobson）和xy微米滑动台行程来测量表面轮廓和粗糙度。使用TalyProfile Lite v.3软件，采用2.5微米的微粗糙度滤波比率和高斯滤波器来提取粗糙度参数。用于评估表面粗糙度的所选参数是R，它是与测量长度上的平均线的平均算术偏差。 用于测量粗糙度参数R的测量轮廓长度的LSP处理表面为16mm，而用于测量裂纹表面粗糙度的测量长度为2.5mm。2.4射透电子显微镜使用在200kV电压下操作的JEOL 2000-FX透射电子显微镜（TEM）表征LSP对材料微结构的影响。为了获得对微观结构和位错配置的LSP效应的正确认识，我们选择不以任何方式研磨LSP表面.因为可能会丢失信息, 取而代之的是，我们在以下步骤中制备用于TEM分析的横截面薄箔：（i） 用Gatan胶（中间处理过的表面）面对面地粘成两块LSP处理的样品（ii） 将横截面标本切成（iii）仔细研磨（iv）冲出3mm直径的圆盘，（v）通过Ar 轰击离子稀化。平均而言，每个样品条件的80个微观结构图像从原始表面的深度0至约500lm取得。为了方便起见，选择具有最高位错密度的图像作为每个LSP处理条件的代表性图像。2.5显微硬度显微硬度用微量维氏（Vickers）测试，使用Leica Microsystems的Leitz Wetzlar量具，使用0.071N的小载荷进行测量。 通过试样的整个横截面在垂直方向上进行压痕。 为了获得关于LSP对硬度影响的正确信息.表面未被研磨，因为硬化材料可能已被完全去除。在地表以下的第一次和最后10次测量是使用32m的间距进行的，以增加感兴趣的地下区域中的测量次数。2.6残余应力残余应力分析采用钻孔松弛法,按照ASTM E 837-08钻孔应变计法测定残余应力的标准试验方法进行19。Vishay Measurements Group（Vishay Intertechnology Inc.，Malvern，PA，USA）的产品测量设备用电阻计CEA-06-062UM-120测定残余应力。使用带千分尺的深度进料器钻孔（分辨率0.01毫米）。钻孔后测量钻孔直径，并使用光学Olympus宏观显微镜，ColorView相机和AnalysisDocu软件在三个位置进行平均。使用H-drill v3.10软件和应力计算的积分方法计算最大和最小主残余应力和方位角，应变偏差误差估计值为3.2 lm。2.7疲劳裂纹增长试验疲劳测试是在Russen-bergerPrfmachinenAG的电磁驱动测试设备Cracktronic 8204上使用Fractomat进行的。 根据Cracktronic的可用测试空间和用于测量疲劳裂纹生长速率的标准测试方法ASTM E647进行测试。Cracktronic的基本模块具有单独的静态和动态驱动器。 静态负载由直流电机驱动的滚珠丝杠轴产生，并通过扭杆连接到摆动系统。 动态负载通过电磁驱动谐振器产生。电磁铁集成在一个封闭的并且以其自然谐振频率激励振荡系统，其中工作点位于谐振曲线的峰值处，典型频率在50和200Hz之间。 动态和静态部分由独立电路控制，与Credo控制单元和计算机软件相连，并允许任何高精度的应力比组合。使用最大载荷为8000N，分辨率为1N的Rumul称重传感器（压电式）测量交变力。此外，具有弹性的样本是系统的一部分。使用间接电势降法连续监测裂纹长度。在这种方法中，与“直接电位下降法”相比，电位值从薄的（5 lm）康铜电阻箔Krak-gages（型号RMF-A5）中取出，其中电位直接取出从标本。克拉克量规附着在样品的两侧以监测和平均裂纹长度。裂缝在量具中与试样的实际裂纹同时增长。以0.001mm的分辨率记录裂纹长度的测量结果，精度高于2。误差的主要来源是将试样上的应变计粘贴在距离实际切口尖端半径0.2毫米处时的标准偏差。使用基于氰基丙烯酸酯的胶（Vishay Measurements）在光学宏观范围下在15-20放大率下使用与应变计粘合相同的技术进行Krak-测量仪的胶合。定位误差通过空间校准显微镜测量。在实际电火花加工切口的定位误差长度在开始切口尺寸输入时考虑到了压痕温度。为了执行DK = const，以恒定的R比进行测试，方程(1) 被集成在测试机软件算法中以控制测试的动态力。使用两种方法计算裂纹增长率。该方法涉及Da / DN的直接点对点计算，其中与标称裂纹长度的小偏差导致大的散射。第二种方法涉及使用最小二乘法拟合适当的最佳拟合方程，这在第一节中描述3.5。使用具有测量能力为50000N并具有1N分辨率的压电力传感器的通用测试机BETA 50Messphysik在疲劳裂纹扩展测试之后在紧密张力试样上测量断裂韧性K。使用一种精度在2以内的光学激光散斑引伸计。在光学显微镜下从载荷线到裂纹前缘测量初始疲劳裂纹长度，其中测量误差为0.1mm。断裂韧度在厚度依赖的平面应力或过渡平面应力/平面应变条件下进行。使用Olympus立体显微镜SZX10和JEOL 5610扫描电子显微镜（SEM）进行断裂分析。这些图像是在疲劳表面上的预选位置以及从疲劳断裂表面到韧性断裂的过渡区域拍摄的。3.结果与讨论3.1LSP处理过的表面的粗糙度分析由于在激光冲击处理过程中不存在吸收涂层，与未处理样品相比，LSP处理样品的平均算术粗糙度参数R显着增加（表2）。 接收状态下的平均粗糙度为1.15 lm，而平均R粗糙度。在使用1.0,1.5,2.0mm束直径的LSP之后，分别增加到7.3,10.4和7.92lm的值。 随后在FCG测试之前将所有LSP样品研磨至R值为0.37-0.38lm。图5显示测量的轮廓，从未经处理的表面开始，继续到更粗糙的激光冲击处理表面。3.2透射电子显微镜LSP处理的试样的横截面显微结构显示在图6。强烈的应变硬化变形的证据在表面层中是明显的，具有各种位错排列。由于不同的地方使用不同的地区，地点之间的位错不均匀LSP处理以及由于特定晶粒内部和之间的塑性变形的非均质性质。对比图6b和d分别清楚地表明，试样LSP1和LSP3分别获得了最高和最低的位错密度。 由于激光束直径反映较小的激光强度和冲击波压力，因为强度与激光束面积成反比，所以这种结果是相当合理的。因此，LSP 1试样的显微组织（图6a和b）揭示随机排列的，具有位错缠结和位错壁的高密度位错结构。 此外，确认了形成超细晶粒的晶粒细化，其可以有效地减少多晶金属中的疲劳裂纹生长。看来，晶粒细化机制主要是通过形成和演化有助于形成的位错结构来实现的。图6. LSP处理样品的透射电子显微镜（TEM）明场图像; （a）LSP 1，（b）LSP 2和（c）LSP表2基材和LSP表面的平均R值 负载方向L T 拉伸强度 - R（MPa）311 310 屈服应力 - R（MPa）296 295 伸长率 - A（）9.7 8.5图7.显微硬度测量图8.残余应力状态的变化，（a）原样收到，（b）LSP 3处理后第19 页 共20 页3.3 LSP处理的表面下层的硬度分析由于应变硬化，与基础材料相比，LSP处理可能导致次表层硬度增加。例如，冈萨雷斯等人的研究。对于脉冲密度分别为900和1200脉冲/ cm，Al-Mg-Si合金的LSP处理后的维氏显微硬度在83 HV下分别增加到96和94 HV。图7显示位置和显微硬度压痕线。 相同合金系列上的载荷为0.071N的维氏硬度测量显示LSP之后的次表层中没有显着的硬度变化。维氏压痕的平均尺寸为11.16 lm（0.27 lm）。对于硬度的恒定函数，残差分析揭示了硬度的正态分布和108HV（7HV）的平均值;基本上，所有测试样本都具有相似的值。因此，LSP对地下和贯穿深度硬度的影响可以忽略不计。在我们的研究中获得的显微硬度结果与Peyre等报道的一致。Fabbro等人关于激光冲击喷丸在铝合金上的应用。 他们的结果也表明LSP处理试样的硬度性能几乎没有改善，与传统喷丸处理试样相比HV值更低。 据Fabbro等人称这种行为可以用三种方式来解释：图8残余应力状态的变化，（a）原样收到，（b）LSP 3处理后（i）冲击持续时间非常短，因此硬度不能在材料内部开始和传播，（ii）与喷丸过程相比，没有发生接触变形，即没有赫兹载荷和（iii）冲击压力通常低于喷丸处理的冲击压力。3.4残余应力选择处于接收状态的试样和研磨后的试样LSP 3，用于利用钻孔松弛方法分析残余应力。图8a表示冷轧后的母材的初期残余应力状态，在160下进行10小时的应力消除和时效处理。在钻孔深度达0.5毫米的表面上，残余应力的大小在0到50兆帕之间。 在0和-25MPa之间的压缩应力的短暂过渡发生在0.65mm的深度，然后是渐变恢复0.9毫米的张力。图8b显示了LSP 3处理后残余应力与钻孔深度的函数关系。在地表以下的较低深度处，缓解的应力对于最大应力分量稍微处于拉伸状态，对于最小主要应力分量略微压缩。LSP 3样品的小表面压缩RS很可能是由于局部表面熔化，电离和再固化引起的激光处理过程中的软化效应。此外，与TEM微观结构有很好的一致性，其中标本LSP3揭示了所有LSP处理的样本中最低的位错密度。尽管如此，从中可以看出图8b残余应力在rv0.05mm处变为纯粹的压缩，在0.25mm处达到-230MPa的最大压缩RS。残余应力完全保持在压缩高达0.85毫米的深度，然后过渡到平衡拉伸残余应力。应该指出的是，在1到9毫米深的板材区域内没有进行残余应力测量。然而，为了平衡表面的压应力，拉伸残余应力很可能出现在试样的中心。最近在Gill等人研究Inconel 718合金没有保护涂层的LSP效应中也报道了类似的结果。他们的结果也证实了小的表面压应力（50MPa），其本质上变成了拉伸，并且从表面持续大约80lm，然后变为压缩状态。通过TEM，同一作者提出，在激光表面烧蚀过程中，冲击波会将烧蚀的颗粒重新沉积到基体上，从而阻止晶粒细化。他们的建议与我们的结果部分一致，但仅限于LSP 3标本，因为LSP 1和LSP 2都显示高位错密度，位错单元和细化晶粒。3.5低应力强度水平下的疲劳裂纹扩展分析为了研究LSP处理前后的疲劳裂纹扩展速率，疲劳预裂的标准程序为:在R = 0.1和在DK =常数时执行。4兆帕/米裂纹长度为1.5毫米。预裂的长度是一致的，.使用ASTM E640-05，其中预裂纹长度应至少为0.1 B或至少1 mm。预裂后，应力强度范围DK从4MP/米增加到开始K减少测试值。K减少程序非常适用于低于10m /周期的疲劳裂纹扩展速率。在这种测试方法中，K值减小是通过开裂力来进行的，或者是连续的，或者是随着裂缝的增长而增加的一系列步骤。根据指数函数设定应力强度范围DK = Dk其中DK是应力强度的起始值，C是梯度值。在初步测试和Shelton等人的研究基础上，脱落率的影响，梯度值C被设置为0.8毫米。初始值DK也是根据未经处理和LSP处理的试样的初步测试结果选定的。其目的是获得最大DK间隔的数据（达到DK），用于选择脱落参数和可用5 mm疲劳裂纹长度（Krak-gage AMF-5限制）。对于未处理的样品1和未处理的样品2.D从12MPam达到. 因此，本研究提出的LSP处理试件从较低的起始值7 MPam进行测试。在疲劳试验的两个阶段中，试样两侧均采用间接电位下降法监测预裂过程和K减少试验，试验共振频率和裂纹长度。图9a表示疲劳裂纹长度作为预裂化和K减少试验期间许多循环的函数，其中每条曲线代表单个试样。 在该图中，每个样本都经历相同的测试参数制度。此外，裂纹方向没有偏离对称平面超过2。从图9a，很明显，LSP处理过的试样中的疲劳裂纹早于未处理过的材料。 周长的数量，需要增长一个precrack（的总长度1.5毫米）是最大的未处理标本和最低的LSP 3样本。在1.21.5mm范围内的N-N曲线的线性部分，用正割法计算裂纹扩展速率，表3根据这些数据，计算出LSP 1，LSP 2和LSP 3试样预裂前1.5倍所需的周期数分别减少了28，33和47。谐振测试频率随裂纹长度而减小，或者在缺口或裂纹尖端的应变硬化情况下增加。测试共振频率是样品弹性常数的函数。图9b显示了作为疲劳裂纹长度函数的测试频率f。 总体而言，在未处理材料的测试过程中记录的频率最高。 由于缺口尖端的应变硬化效应，疲劳裂纹开始前，测试频率略有增加。 破解之后生长稳定，频率随裂纹长度而减小。 在达到1.5mm的预裂纹长度后，进行测试。 机器正在增加应力强度比范围，直至达到7MPa/m。未处理的spec-imen在测试标本中处于最高水平和最凸面。从n = 5次基材测试中获得的起始频率标准偏差达到2.1Hz。假设LSP处理标本的相似标准偏差是正确的，LSP处理之间测试频率的差异不显着。在应力强度因子范围增加之后立即发生裂纹增长的短暂初始加速。未处理试样的初始疲劳预裂纹长度从1.5mm增加到1.64mm，LSP 1试样的初始疲劳预制裂纹长度从1.64mm增加到LSP 2试样的1.76mm和LSP 3试样的1.81mm。首先，使用点对点割线方法来生成Da / DN（DK）图（图10）来自（N）个数据。用这种方法疲劳裂纹扩展曲线中的连续实际a（N）的小误差偏差导致更大的“散射”。测量的可变性可以观察到三个测试未处理的样本。数据显示在图10，可以得出结论，对于所有的标本数据显示在图10，可以得出结论，对于所有的标本经过激光冲击处理后，LSP 1，LSP 2和LSP 3显示出较高的裂纹扩展速率，这支持了较早的结论。 相比之下，只有很小的差异之间的曲线。图9.监测的疲劳裂纹长度和频率随循环数的变化而变化。 （a）疲劳裂纹长度，（b ）测试（共振）频率据Broek介绍，数据集（如图所示）图10）必须解释，然后才能用于分析。为了获得散点较少且没有实际裂纹扩展长度的有意义变化的图，根据最佳拟合方程计算新的da / dN数据。a（N）数据使用具有不同函数的最小二乘法拟合。例如，未处理标本的最佳拟合函数是Hocket-Sherby 2D函数。图11代表da / dN（DK）纯化的疲劳裂纹扩展曲线，实际上没有散射（由于连续函数的da / dN计算）以及Paris和NASGRO的拟合模型。模型系数以英寸表示表4。用于LSP处理标本中，FCG曲线不是S形，其斜率在4.6-4.9MPa/m处逐渐增加。对于巴黎区域，曲线拟合了LPS处理样本的斜率变化。实验数据显示了斜率（膝盖）的变化，其中使用了从DK到膝盖以及从膝盖到K的一个功能。LSP 1和LSP 2的FCG曲线几乎重叠，表明等裂纹增长率。在LSP 3试样中发现了稍高的裂纹扩展速率。 因此，如果不进行多次重复，就不能得出关于LSP参数对疲劳裂纹扩展的影响的肯定结论。在3.36MPa/m的阈值水平下，基材中的疲劳裂纹不再传播（da / dN0），而在同一个SIF处的LSP 3试样中的裂纹正在传播.约3m /循环。因此，在最低阈值的LSP 3试样中发现最高的疲劳裂纹扩展速率。值为1.1兆帕/米。虽然，LSP过程引起高密度的位错，晶粒细化甚至是近表面的压缩残余应力，显然这仅仅是不能有效地降低FCG速率。 很明显，与未处理的试样相比，所有LSP试样都表现出较高的裂纹扩展速率，这是由于试样中心处的平衡拉伸应力，这降低了裂纹闭合的影响。很明显，残余压应力并没有有效地降低裂纹扩展速率。Chahardehi等人他研究了由LSP产生的残余应力的影响，他指出观察到的裂纹的增长率受拉伸核心的影响要大于压缩的表面应力。 在这项研究中，相对于循环次数测量的裂纹长度没有显示任何疲劳裂纹的延迟，这被认为是由于材料中的拉伸核心作为激光冲击强化的“副产物”而出现的。图10.裂纹增长曲线da / dN与应用的DK，用直接割线方法获得图11.裂纹增长曲线da / dN与应用的DK，从拟合方程中获得表4.PARIS和NASGRO方程的未处理和LSP 1，LSP 2，LSP 3处理标本的常量 不处理8.37 . 10-93.733.369.56 . 10-112.750.931.2LSP 13.4 . 10-113.721.41.05 . 10-102.700.450.93.5 . 10-114.104.5 0.6LSP 23.4 . 10-113.731.41.15 . 10-102.700.450.91.5 . 10-103.576 0.6LSP 37.4 . 10-103.521.11.80 . 10-102.700.450.99 . 10-102.506 0.6还应该指出的是图10和11为被绘制为应用的DK数据的函数。在低应力下裂纹发生，并且在临界值附近影响增加。 据Borego等人和Gavras等人塑性诱导和粗糙诱导闭合机制是铝合金的主要研究对象（其中6082-T6也是研究的主题）。考虑到这一点，校正的da / dN与DK曲线将从da / dN与DK数据向左偏移。另一个需要注意的因素是LSP治疗的顺序。与大多数调查相反在预先存在的疲劳预制裂纹之后进行LSP，我们的标本在LSP之后进行预裂。其次，LSP在EDM切口的边缘进行，这可能在切口边缘上产生拉伸热点。以前Ivetic等报道了类似的观察结果。研究“未镀膜”LSP工艺对开孔薄铝7075-T73试样的影响。他们的结果表明操作顺序（第一个LSP和裸眼或反之亦然）对疲劳裂纹扩展行为具有巨大影响。此外，有人建议激光束通过孔内侧的“溢出”等离子体聚焦，产生局部拉伸热点，这大大降低了疲劳寿命。 同一作者认为，这种效应仅影响裂纹萌生阶段，因为一旦裂纹形成，残余应力就会逐渐松弛。3.6断裂韧性在FCG试验后对C（T）试样进行断裂韧性K试验以评价LSP处理的效果。为了满足断裂韧性测试中的平面应变条件，样品的厚度B应该大于2.5（K/ rY）。所研究的6082-T651板的屈服应力和断裂韧性为295MPa。此外，根据文献综述，6082-T6的断裂韧性的预期值取决于时效硬化，在35-40MPa/m，因此，按照平面应变线弹性断裂韧性试验时，试样厚度应为46 mm。由于厚度非常大的铝板实际上很少有用，因此过渡平面应变,平面应力厚度相关测试非常普遍。对于较薄的铝板，铝合金B 646-12的断裂韧性的标准实践（a）厚度小于6.3mm的薄产品，（ii）中间板厚度，对于有效的平面应变断裂韧性测试来说太薄，但对于处理为薄板，厚度太厚范围从6.3到50毫米，和（三）相对厚的标本。B = 10mm的6082-T651钢板划分成（ii）中间厚度，其中断裂韧性也应根据试验方法E399（作为对铝合金B645的线弹性平面应变断裂韧性测试的标准实践的补充。使用三个未经处理的样本来测量平均断裂韧度，而每个LSP处理使用单个样本。表5 显示了起始裂纹长度，a / W比，最大力P，参考力P和断裂韧性K的计算值。裂纹长度（即缺口和疲劳裂纹长度的总长度）在0.67和0.65之间（由ASTM E399推荐的介于0.45和0.55 W)。图12显示起始裂纹长度为0.71-0.74W的试样的P-v图。未处理试样的平均断裂韧度基于三个独立测试被发现为37.8兆帕/米。LSP处理后，断裂韧性为28。与未经处理的材料相比降低了33。考虑到与基础材料测试的标准差相似的值（0.86 MPa/m）也适用于LSP处理，可以得出结论，在LSP处理的试样中对断裂韧性值没有显着影响（LSP 1 = 29.01，LSP 2 = 29.34和LSP3 = 28.21MPa/m）。表5.LSP处理前后的平面应力断裂韧性K测试结果未处理（3次重复）#119.880.712849262038.7237.8 0.86#217.600.634175395037.68#318.650.673784343037.00激光冲击治疗后LSP 120.300.722358196029.01LSP 220.650.742175184029.34LSP 320.190.722058195028.21图12.载荷 - 位移曲线3.7分形分析使用光学和扫描电子显微镜在不同放大倍数下比较未处理和LSP 3样品的宏观断裂表面。图13a在代表性未处理的样品中显示出相对线性的形状裂纹前缘。相反，图13b显示宏观可以观察到疲劳裂纹前缘的曲率。 因此，残余应力的影响影响LSP处理试样的裂纹扩展，疲劳裂纹长度随深度增加而增加; 它是标本中间最大的。裂纹前缘的曲率可以通过在存在表面压缩残余应力的情况下两侧的裂纹延迟来解释。相反，拉伸残余应力具有不利影响，因为它们降低了裂纹闭合的影响。图13b和d显示了在断裂韧性试验期间发生的接近阈值应力强度和过渡到断裂表面的疲劳裂纹表面的细节。 进行静态断裂韧性测试的表面的断裂分析观察显示典型的凹痕破裂，其由称为微孔聚结的过程引起。在裂纹扩展的近阈值范围内，疲劳断口表面没有表现出典型的疲劳条纹，这对于阶段I疲劳断裂表面是常见的。影响疲劳裂纹表面视觉外观的重要现象是裂纹闭合。疲劳期间接触的裂纹表面限定了纹理和裂纹表面粗糙度。如前所述，在铝合金应力强度范围接近阈值时，主要发生粗糙引起的裂纹闭合。通常使用粗糙感应裂纹闭合来描述尺寸小的多面表面特征的接触，例如，颗粒尺寸的顺序。基于SEM显微照片上的疲劳裂纹表面的视觉外观，图13b和d在接近阈值时，该基材具有刻面和圆形的乳沟飞机，随后在40英尺处裂开。DK为3.3MPa/m。 相比之下，LSP 3标本更精细和平滑的粗糙表面，的相对较低阈值处的裂纹扩展停止区域的裂纹扩展。 该表面更光滑表明能量更低，更易碎。图13.疲劳和断裂韧性测试后的断裂表面 在LSP处理的样本中沿着结晶平面发生断裂。 此外，视觉宏观检查（光学宏观奥林巴斯）证实，在LSP 3标本中，裂纹表面的中心部分看起来更平滑，特别是裂纹扩展方向。使用接触轮廓仪（Taylor Hobson）和微米滑动台测量裂纹表面的粗糙度。 如表1所示，粗糙度测量区域是裂纹表面中心部分2.5 mm的矩形区域图14一个。 单个样品上每个方向的测量线数量为8个，测量线偏移增量约为350 lm，测量线长度为2.5 mm。 根据这些值计算平均粗糙度值R和标准偏差并以直方图表示，图14湾 从结果可以看出，z方向的平均粗糙度是相似的（z方向的平均R在16.2和17.5lm之间，并且具有平均标准偏差2.45lm）。 此外，在z方向上的整体表面粗糙度R.比在x方向上高33-54，这主要是由于在板制造期间轧制方向上的晶粒伸长的结果。直方图显示裂纹表面粗糙度的平均值在未处理的样本中x方向上的R高于全部。表六裂纹表面粗糙度测量Rlm在（x方向），n = 8用于测试样品的相等性i12345678不处理7.8213.27.989.3711.011.3012.714.7LSP 15.47.355.010.57.8111.105.688.2LSP 25.95.988.479.24.029.044.277.2LSP 38.8512.38.047.898.856.296.428.4在所有三种情况下，H假设被拒绝，这意味着未处理试样的x方向裂纹粗糙度显着高于LSP处理试样的x方向粗糙度。但是，人们意识到，接触轮廓仪的测量轮廓并不完全遵循裂纹表面的微粗糙度。然而裂纹表面粗糙度分析证实了视觉定性SEM和宏观检查结果。所有上述结果表明，在部件表面上引入压缩残余应力时应该格外小心，该部件表面将在HCF状态下受到动态负载。此外，尽管表面残余应力测量技术通常用于预测疲劳寿命，但这不能作为最终的标准。为此，应该注意的是，虽然LSP已被广泛认，用于改善金属部件的疲劳寿命的先进性，且非常有效，但是工艺参数需要针对每种应用进行优化，并且为了防止处理材料的任何有害影响操作应始终以最佳顺序进行4结论LSP是增加动态加载组件寿命的创新新方法。尽管预计近表面的残余应力和细化晶粒会延缓LSP后的FCG速率。但在此情况并非完全如此。 根据所进行的研究，可以得出以下结论：1. 在LSP表面粗糙度处理之后，与未处理材料相比，参数Ra增加了7倍以上。2透射电子显微镜分析证实了位错细胞和细化晶粒的纠缠错位。在LSP处理的C（T）试样中发现LSP诱导的近表面压应力。3未经处理的试样的裂纹增长 值高于LSP处理后。4在相同的测试条件下，与LSP处理的试样相比，在未处理的试样的测试中共振测试频率更高。5在LSP两种情况下裂纹增长速率都显着增加。6LSP试样获得的断裂韧性平均比未处理试样低28-33。7在近临界测试范围内的断口分析显示，LSP处理试样的疲劳裂纹前缘和弯曲前缘的 裂纹表面粗糙度测量结果表明裂纹扩展的x方向的粗糙度低于z方向的粗糙度。 此外，未处理试样的平均裂纹表面粗糙度（x方向）在试样中心部分显着高于选定试验领域LSP 1，LSP 2和LSP 3试样的平均裂纹表面粗糙度。 第20 页 共20 页Effects of laser shock processing on high cycle fatigue crack growth rateand fracture toughness of aluminium alloy 6082-T651Zoran Bergant, Uro Trdan, Janez GrumFaculty of Mechanical Engineering, University of Ljubljana, Askerceva 6, 1000 Ljubljana, Sloveniaa r t i c l ei n f oArticle history:Received 25 November 2015Received in revised form 10 February 2016Accepted 16 February 2016Available online 3 March 2016Keywords:Laser shock processingFatigue crack growthAluminium alloyResidual stressesFracture toughnessa b s t r a c tThe effects of laser shock processing without protective coating on high-cycle fatigue crack growth andfracture toughness were investigated. Laser shock peening treatment was performed on compact tensionspecimens from both sides perpendicular to the crack growth direction, followed by subsequent grinding.Fatigue crack growth tests were performed at frequencies between 116 and 146 Hz, at R = 0.1 and a con-stant stress intensity range during the fatigue crack initiation phase and K-decreasing test. A lower num-ber of cycles was required to initiate a fatigue precrack, and faster fatigue crack growth was found intensile residual stress field of LSP-treated specimens. The crack growth threshold decreased by 60% afterLSP treatment. The fracture toughness decreased by 2833% after LSP treatment. The fatigue-to-ductiletransition boundary on fractographic surfaces show linear fatigue crack fronts in non-treated specimensand curves after LSP treatment.? 2016 Elsevier Ltd. All rights reserved.1. IntroductionLaser shock processing (LSP) is a relatively new surface treat-ment technology to extend the lifetime of dynamically loadedcomponents. Several studies 13 have already demonstrated thatLSP treatment decreases the fatigue crack growth (FCG) ratethrough the generation of large amplitude compressive stressesas a result of material local plastic deformation induced by thepressure shock wave. Since the fatigue cracks normally initiatefrom the surface, locked compressive stresses will, under externalload, be superimposed on the applied stress, leading to smallerstress intensity factors at the crack tip and possible crack closureto incur the reduction of effective driving force for the FCG 1,2.LSP can also improve fatigue strength and fatigue crack initia-tion life due to a considerable densification of dislocations andmicrostructural grain refinement. Huang et al. 3 confirmed areduced FCG rate due to the highly tangled and dense dislocationarrangements in the LSP-processed surface of an AlMgSi alloy(6061-T6). In another study Lu et al. 4 reported obviousmicrostructure refinement in AlCuMg alloy (LY2) after LSP. Theresults confirmed LSP to be a promising method to obtain a highdensity of dislocations to improve fatigue resistance, whereas theminimum grain size in the top surface after multiple LSP impactswas about 100200 nm. Furthermore, residual stress resultingfrom laser processing can be significantly higher with much deepereffective penetration in the material compared to conventionalshot peening 5.Huang et al. 1,3 reported beneficial effects of LSP under differ-ent process setups on the fatigue crack growth properties of 6061T6 aluminium alloy. Their results indicate that the FCG ratedecreases with increasing laser energies 1 and LSP coverage areas3, especially in the initial fatigue crack growth stage. However,the strengthening effects are weak in the final period of fatiguecrack growth since the residual stresses release as the crack grows.In another study, Hatamleh et al. 5 compared laser and shotpeening effects on FCG in friction-stir-welded AlZn alloy (7075-T7351) joints, reporting reductions of the FCG rate in comparisonwith the base material for the LSP-treated specimen, whereas shotpeening revealed negligible improvement of the FCG rate.However, despite the fact that the LSP process in which a pro-tective coating is used has already been proven as an advanced,competitive surface treatment for improving the fatigue life, corro-sion resistance, hardness and wear resistance of various metals andalloys 69, several negative points exist that limit the widespreadapplication of this process. Open discussion on the last, i.e. 5thinternational conference on laser peening and related phenomenain Cincinnati, 2015 10 pointed out that new/combined strategiesare needed in the industrial field applications. Basically, two mainapproaches laser shock treatment to achieve considerable mergingof dislocations and generation of compressive residual stresses are/10.1016/j.ijfatigue.2016.02.0270142-1123/? 2016 Elsevier Ltd. All rights reserved.Corresponding author. Tel.: +386 1 477 1203; fax: +386 1 477 1225.E-mail addresses: zoran.bergantfs.uni-lj.si (Z. Bergant), uros.trdanfs.uni-lj.si(U. Trdan), janez.grumfs.uni-lj.si (J. Grum).International Journal of Fatigue 87 (2016) 444455Contents lists available at ScienceDirectInternational Journal of Fatiguejournal homepage: /locate/ijfatigueused; (i) first regime, uses high-energy laser pulses (100 J), longerpulse duration up to 100 ns combined with the protective coating,(ii) the second approach uses lower laser energies in the order offew joules or fewer, increased overlap, smaller spots and no coat-ing (also known as LSPwC). A detailed description of both princi-ples was reported previously 11.Although, the protective coating on the processed part preventssurface ablation, while maintaining high surface quality, it is atime-consuming affair since the overlay is damaged severely dur-ing the LSP process and requires frequent replacements, makingit slow and expensive in industrial applications 12. Furthermore,since some parts being peened might not be accessible for theapplication of protective overlay, much attention was focused onLSP process without a need for a protective ablative overlay. How-ever, after this process typical surface craters are obtained withincreased surface roughness and waviness due to the local surfaceablation, which in turn creates a high-pressure plasma that, con-tained by a thin layer of water flowing on the specimen surface,generates a pressure wave propagating into the specimen 13.Based on the open literature review, very few comprehensive stud-ies have been conducted on laser shock processing without protec-tive coating on fatigue crack growth, and none of them haveevaluated fatigue characteristics after subsequent grinding proce-dure to eliminate the downward effect on uncoated LSP-inducedsurface topography.Rubio-Gonzalez et al. 14,15 investigated the effects of a two-sided uncoated LSP process on compact tension (CT) specimensof 2205 duplex stainless steel and 6061-T6 Al alloy. The resultsconfirmed that the FCG rate decreased while fracture toughnessincreased with the increased pulse density. Although there arenumerous investigations of laser processing effect on material fati-gue crack growth properties 1,5,14, there is limited literature onthe possible detrimental effect of laser shock processing. Further,all the above studies have been performed after pre-existing fati-gue pre-crack, under low sine wave frequencies ( 106) 16. To obtain a quantitativecomparison between the fracture toughness of untreated and LSPspecimens in the final rupture stage under static load will be deter-mined as well. In addition, the possible softening effect due to thelocal surface melting, ionization, and re-solidification during laserprocess will be carefully investigated by means of dislocationarrangements, residual stress, and microhardness in depth distri-bution. All LSP specimens under investigation were analyzed aftersubsequent grinding procedure in order to obtain a proper surfacefinish.2. Experimental methods2.1. Material and specimen preparationPlates of commercial wrought AlMgSi aluminium alloy (ENAW 6082-T651) were machined to obtain the specimens. The over-all heat treatment T-651 procedure (homogenization, solutiontreatment, aging, etc.) is detailed in Ref. 17. Its chemical compo-sition in wt% was 0.87 Si, 0.72 Mg, 0.42 Mn, 0.35 Fe, 0.15 minor ele-ments (Cu, Cr, Ni, Zn, Ti) and balance Al. In order to obtainmechanical properties for the given material in the specific rollingdirection, three cylindrical tensile specimens were machined inlongitudinal and three in the transversal direction of rolling witha test diameter of 5 mm and a measuring length of 20 mm(Fig. 1). The average measured tensile strength, yield stress, andelongation are given in Table 1. A total of six C(T) specimens weretested from the plate in the LT direction for fatigue crack growthand fracture toughness tests. Three specimens were tested in not-treated conditions, and three specimens were LSP-treated with dif-ferent parameter sets: LSP 1, LSP 2 and LSP 3. Notches were madeusing electrical discharge machining (EDM) using a 0.35 mm wire,which resulted in a notch with a radius of 0.19 mm. The geometryof the C(T) specimen, adopted for testing with Rumul Cracktronicunit and Krak-gages AMF-5, is given in Fig. 2.The applied stress intensity rangeDK was used, calculatedaccording to the equation 18:DK DPBffiffiffiffiffiffiWp? Y;1whereDP = Pmax? Pminis the alternating force, B is the specimenthickness, W is the length from the load centerline to the edge ofspecimen, Y is a geometrical factor, which is derived for linear elas-tic material for compact-tension specimens:Y 2 a1 ?a3=20:886 4:64a? 13:32a2 14:72a3? 5:6a4;2wherea= a/W, a is the crack length.2.2. Laser shock processingLaser shock processing was performed using a Spectra-PhysicsQ-switched Nd:YAG laser with an irradiation wavelength of1064 nm. The maximum laser beam energy was 2.8 J/pulse,whereas the FWHM (Full Width Half Maximum) of the generatedGaussian intensity pulses was 10 ns. Focused laser beam diameterswere varied via modification of convergent lens. Beam diameters,used during LSP 1, LSP 2 and LSP 3 treatments, were set to 1.5,2.0 and 2.5 mm, respectively. In this study, a predefined overlap-ping pulse density of 1600 pulses/cm2was chosen, with thelaser-advancing direction parallel to the plate-rolling direction (L).Specimens were irradiated using laser shock processing methodwithout any protective coating in a water confinement regime,whereas laser pulses were overlapped and scanned in a zig-zagpattern (Fig. 3). A water jet set up was employed to create a thinwater layer and to avoid the formation of water bubbles or the con-centration of impurities resulting from the material ablation, thus,constantly ensuring a pure lasermatter interaction.ThetreatedareaontheC(T)specimenswasapprox.10 ? 10 mm2on both sides of the specimens. For TEM microstruc-tural analysis, additional LSP specimens were prepared using theFig. 1. Tensile and C(T) specimen orientation in Al-plate.Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455445same processing conditions as for C(T) specimens. Prior to fatiguecrack growth testing, C(T) specimens were subsequently groundto the final thickness of 9.8 mm to remove the rough asperitiesafter LSP treatment (Fig. 4). Prior to laser shock processing, nopre-crack was initiated. However, specimens for TEM analysis werenot ground in any way to obtain the representative microstruc-tures produced by the laser peening treatment.2.3. Surface roughnessAfter the surface near pre-machined notch was laser surfacetreated, the surface profile and the roughness were measured usingSurtronic 3+ profilometer (Taylor Hobson) and the x-y micrometersliding table stroke. TalyProfile Lite v.3 software was used withmicro-roughness filtering ratio of 2.5lm and a Gaussian filter toextract the roughness parameter. The selected parameter for eval-uation of surface roughness is Ra, which is the average arithmeticdeviation from the mean line on the measuring length. The mea-suring profile length for the roughness parameter RadeterminationTable 1Average measured tensile properties of 6082-T651 plate.Direction of loadLTTensile strength Rm(MPa)311310Yield stress Rp0.2(MPa)296295Elongation A (%)9.78.5Fig. 2. Geometry of C(T) specimens and measuring resistance gage AMF-5.Fig. 3. Laser shock processing setup and treated C(T) specimen.Fig. 4. C(T) specimens (a) no treatment, (b) after LSP 1 treatment, (c) LSP 1 aftergrinding.446Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455of LSP-treated surfaces was 16 mm, while the measuring lengthused for measuring roughness on crack surfaces was 2.5 mm.2.4. Transmission electron microscopyThe LSP effect on the material microstructure was characterizedusing a JEOL 2000-FX transmission electron microscope (TEM)operated at a voltage of 200 kV. In order to obtain proper insightinto LSP effects on microstructure and dislocation configurations,we chose not to grind the LSP surfaces in any way, due to a possibleloss of information. Instead, we prepared cross-sectional thin foilsfor TEM analysis in the following steps: (i) bonding two pieces ofthe LSP-treated specimens face-to-face with GatanTMglue (treatedsurfaces in the middle), (ii) cutting a cross-sectional specimen into1.0 mm thick sheets, (iii) grinding it carefully to about 100lmthick; (iv) punching out to 3 mm diameter discs, (v) ion thinningby Ar+ bombardment. On average, 80 microstructural images foreach specimen condition were taken from the depth 0 to about500lm from the original surface. For the sake of convenience,images with the highest dislocation densities were chosen as therepresentative ones for each LSP-processing condition.2.5. MicrohardnessThe microhardness was measured with a micro-Vickers testusing a small load of 0.071 N, Leitz Wetzlar gage, Leica Microsys-tems. Indentations were made in the vertical direction throughthe entire cross section of C(T) specimen. In order to obtain properinformation about LSP effects on hardness, the surface was notground, since the hardened material could have been completelyremoved. The first and last 10 measurements below the surfacewere made using a spacing of 32lm to increase the number ofmeasurements in the subsurface region of interest. Then, the spac-ing over the rest of the material thickness increased to approx.214lm in order to reduce the experimental time.2.6. Residual stressesThe analysis of residual stresses was performed using the hole-drilling relaxation method in accordance with the Standard TestMethod for Determining Residual Stresses by the Hole-DrillingStrain-Gage Method ASTM E 837-08 19. Residual stresses weredetermined with measurement equipment, a product of the VishayMeasurements Group (Vishay Intertechnology Inc., Malvern, PA,USA) with the resistance gage CEA-06-062UM-120. The hole wasdrilledusingadepthfeederwithmicrometer(resolution0.01 mm). The hole diameter was measured after drilling and aver-aged on three locations using an optical Olympus macroscope, aColorView Camera, and AnalysisDocu software. Maximum andminimum principal residual stresses and the orientation anglewere calculated using H-drill v3.10 software and the integralmethod of stress calculation with a strain deviation error estima-tion of 3.2lm.2.7. Fatigue crack growth testFatigue testing was performed on the electromagnetically dri-ven testing device, Cracktronic 8204 with Fractomat, by Russen-berger Prfmachinen AG. Testing was performed according to theavailable testing space on Cracktronic and Standard Test Methodfor Measurement of Fatigue Crack Growth Rates ASTM E647 20.The basic module of Cracktronic has separate static and dynamicdrives. The static load is generated by a ball screw spindle drivenby a DC-motor and couples over the torsion rod into the oscillatingsystem. The dynamic load is generated by means of electromag-netic driven resonator. The electromagnet is integrated in a closedloop system and excites the oscillating system in its natural reso-nant frequency, where the operating point is situated at the peakof the resonance curve with typical frequencies between 50 and200 Hz. The dynamic and static parts are controlled by separatecircuits, connected with the Credo control unit and computer soft-ware, and allow any combination of stress ratios with high accu-racy. The alternating force is measured using Rumul load cells(piezoelectric) with the maximum load of 8000 N and a resolutionof 1 N. Furthermore, the specimen with its elasticity is part of thesystem 21. The crack lengths were continuously monitored usingthe indirect electric potential drop method. In this method, theelectric potential value is taken off the thin (5lm) constantanresistance foil Krak-gages (type RMF-A5), as opposed to the directpotential drop method”, in which the potential is taken off directlyfrom the specimen. The Krak-gages were attached on both sides ofthe specimen to monitor and average the crack lengths. The crackin the gage is growing simultaneously with the real crack of thespecimen. The measurement of crack length is recorded in a reso-lution of 0.001 mm and the accuracy is higher that 2%. The mainsource of error is gage misalignment when gluing the gage on spec-imen, which amounts to 0.2 mm from the actual notch tip radius.The gluing of Krak-gages was performed using cyanoacrylate-based glue (Vishay Measurements) under optical macro-scope at1520? magnification with the same technique as strain gage glu-ing. The error of positioning was measured by spatially calibratedmicroscope. The error length of positioning from actual EDM notchwas taken into account at starting notch length value input in Frac-tomat. In order to performDK = const. tests at a constant R-ratio,the relationship in Eq. (1) is integrated in testing machine softwarealgorithm to control the dynamic force of testing. The crack growthrates were calculated using two methods. First, the point-to-pointsecant method:dadNffiDaDNai1? aiNi1? Ni3This method involves a direct point-to-point calculation ofDa/DN, in which small deviations from nominal crack length lead tolarge scatter. The second method involves fitting an appropriatebest-fitting equation using the least squares method, which isdescribed in Section 3.5.The fracture toughness Kcwas measured on compact-tensionspecimens after fatigue crack growth tests using universal testingmachine BETA 50 Messphysik with a piezoelectric force sensorwith measuring capacity of 50,000 N and with a resolution of1 N. The displacement was measured using an optical laser-speckle extensometer with accuracy within 2%. Initial fatigue cracklength was measured under the optical microscope from the line ofload to the crack front, where the measurement error is 0.1 mm.Fracture toughness was conducted under thickness-dependentplane stress or transition plane stress/plane strain conditions.Fractographic analysis was conducted using an Olympus stereomacroscope SZX10 and JEOL 5610 scanning electron microscope(SEM). The images were taken on preselected locations on the fati-gued surface and at transition zone from fatigue fracture surface toductile fracture.3. Results and discussion3.1. Roughness analysis of LSP-treated surfacesAs a consequence of the absence of an absorbent coating duringlaser shock processing, the mean arithmetic roughness parameterRaof LSP-treated specimens increased significantly in comparisonwith untreated specimens (Table 2). The average roughness inthe as-received state was 1.15lm, while the average RaroughnessZ. Bergant et al./International Journal of Fatigue 87 (2016) 444455447after LSP using 1.0, 1.5, 2.0 mm beam diameters, increased to val-ues of 7.3, 10.4 and 7.92lm, respectively. All LSP specimens weresubsequently ground to Ravalues of 0.370.38lm before the FCGtests.Fig. 5 shows measured profiles, starting on a non-treated sur-face continuing to a rougher laser shock-treated surface.3.2. Transmission electron microscopyCross-sectional microstructures of the LSP-treated specimensare shown in Fig. 6. Evidence of intense, strain-hardening deforma-tion is obvious in the surface layer, with various dislocationarrangements. Dislocations are inhomogeneous in terms of depthfrom place to place, due to different regimes used with specificLSP treatment and due to the heterogeneous nature of the plasticdeformation within and between specific grains.Comparing Fig. 6b and d clearly indicates that the highest andthe lowest dislocation density was achieved with the specimensLSP 1 and LSP 3, respectively. Such results are rather logical sincelarger a laser beam diameter reflects a smaller laser intensity andshock wave pressure, since the intensity is inversely proportionalto the laser beam area. Thus, the microstructure of the LSP 1 spec-imen (Fig. 6a and b) revealed randomly arranged, high-density dis-location structures with dislocation tangles and dislocation walls.Moreover, a grain refinement with the formation of ultra-finegrains was confirmed, which can effectively reduce fatigue crackgrowth in the polycrystalline metals 3. It seems that a grainrefinement mechanism is primarily achieved via formation andevolution of dislocation structures contributing to the formationTable 2Average Ravalues for base material and LSP surface.SpecimenStateRoughness as receivedRoughness after LSP Ra(lm)Roughness after grinding Ra(lm)W/O treatmentAs received1.15LSP 11.167.300.37LSP 2LSP surface1.1410.470.37LSP 31.157.920.38Fig. 5. Roughness profile (a) LSP-treated C(T) specimen with detail at the vicinity of the notch, (b) roughness profiles at transition.Fig. 6. TEM bright field images of LSP-treated specimens; (a) LSP 1, (b) LSP 2 and (c) LSP 3.448Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455of dislocation cells (Fig. 6c). With further hardening, dislocation,cells convert to ultra-fine equiaxed grains in order to minimizethe total energy state due to repeated shockwave impacts on thetarget surface.3.3. Analysis of hardness below the LSP-treated sub-surface layerThe LSP-treatment could lead to an increase in hardness in thesub-surface layer compared with the base material, due to strainhardening. For example, the study of Gonzales et al. 14 showedthat Vickers microhardness after LSP treatment of Al-Mg-Si alloyis increased in the sub-surface layer from 83 HV to values 96 and94 HV for pulse densities 900 and 1200 pulses/cm2, respectively.Fig. 7 shows location and the line of micro-hardness indentations.Vickers hardness measurements with loads of 0.071 N on samealloy series revealed no significant hardness changes in the subsur-face layer after LSP. The average Vickers indentation size amountsto 11.16lm (0.27lm). For a constant function of hardness, theanalysis of residuals reveals a normal distribution of hardnessand a mean value of 108 HV (7 HV); essentially, all tested speci-men had similar values. Thus, the effect of LSP on subsurface andthrough-depth hardness was negligible.Microhardness results obtained in our study are consistent withthose reported by Peyre et al. 22 and Fabbro et al. 23 concerningthe application of laser shock peening on aluminium alloys. Theirresults also revealed little improvement in the hardness propertiesof LSP-treated specimens, with lower HV values in comparison tothe specimens treated by the conventional shot-peening process.According to Fabbro et al. 23 this behaviour may be explainedin three ways: (i) shock duration is very small, so that hardnesscannot initiate and propagate inside the material, (ii) comparedto the shot-peening process, no contact deformation occurs, i.e.no Hertzian loading, and (iii) impact pressures are usually lowerthan those from shot-peening process.3.4. Residual stressesThe specimen in an as-received state and specimen LSP 3 aftergrinding was selected for analysis of residual stress with thehole-drilling relaxation method. Fig. 8a shows the initial residualstress state of the base material after cold rolling, stress-relievingand aging at 160 ?C for 10 h. The magnitude of residual stress isbetween 0 and 50 MPa in tension at the surface up to 0.5 mm ofdrilling depth. The short transition to compressive stress between0 and ?25 MPa occurs at 0.65 mm of depth, followed by a gradualreturn to tension at 0.9 mm. Fig. 8b shows the residual stresses as afunction of drilling depth after LSP 3 treatment. At lower depthsbelow the surface, relieved stresses are slightly in tension for themaximum stress component and slightly compressive for the min-imum principal stress component.Small surface compressive RS with the LSP 3 specimen mostlikely originate from the softening effect during laser treatmentdue to the local surface melting, ionisation, and re-solidification.Moreover, good agreement with TEM microstructures can be seen,where specimen LSP 3 revealed the lowest dislocation densityamong all LSP-treated specimens. Nevertheless, as can be seenfrom Fig. 8b the residual stresses became purely compressive atthe depth of ?0.05 mm and reached a maximum compressive RSof ?230 MPa at 0.25 mm. The residual stresses remain fully inFig. 7. Location of Vickers micro-hardness measurements.Fig. 8. Change of residual stress state, (a) as-received, (b) after LSP 3 treatment.Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455449compression up to 0.85 mm of depth, followed by the transition tobalancing tensile residual stresses. It should be noted that therewere no residual stress measurements made in the region of theplate from 1 to 9 mm deep. However, in order to balance the com-pressive stresses at the surface, tensile residual stresses are mostlikely present in the centre of the specimen. Similar results wererecently reported in the investigation of LSP effects without protec-tive coating on Inconel 718 alloy by Gill et al. 12. Their resultsalso confirmed small surface compressive stresses (?50 MPa),which became tensile in nature, and persisted for approximately80lm from the surface that afterwards changed to the compres-sive state. Through TEM, the same authors proposed that duringlaser surface ablation, shock waves re-deposit ablated particlesonto the matrix, which prevented grain refinement. Their proposalis partially consistent with our results, but only for the LSP 3 spec-imen, since both LSP 1 and LSP 2 exhibit high dislocation densities,dislocation cells, and refined grains.3.5. Analysis of fatigue crack growth at low-stress intensity levelsTo investigate the fatigue crack growth rate before and after LSPtreatment, the standard procedure of the fatigue precracking wasperformed at R = 0.1 and atDK = const. = 4 MPapm to reach a fati-gue crack length of 1.5 mm. The length of pre-crack is in agreementwith ASTM E640-05, where the precrack length should be at a min-imum length of 0.1 ? B or at least 1 mm. After precracking, thestress intensity rangeDK increased from 4 MPapm to startingvalue for K-decreasing test. The K-decreasing procedure is well sui-ted for fatigue crack growth rates below 10?8m/cycle 20. In thistest method, K-decreasing is performed by shedding force, eithercontinuously or by a series of incremental steps as the crack grows.The stress intensity rangeDK was set according to the exponentialfunction:DK DK0? eCpa?a0?;4whereDK0is the starting value of stress intensity and Cpis the gra-dient value. On a basis of preliminary tests and the research of Shel-ton et al. 24 who investigated the influence of shed-rate onthreshold values, the gradient value Cpwas set to ?0.8 mm?1. Thestarting valueDK0was also selected on the basis of findings frompreliminary tests of not-treated and LSP-treated specimens. Theaim was to obtain the data for the largest possibleDK interval (toreachDKth) for the shedding parameter selected and the availablefatigue crack length of 5 mm (Krak-gage AMF-5 limitation). Fornot-treated specimen #1 and not-treated #2 specimen,DKthwasreached from 12 MPapm. In contrast, due to much faster crackgrowth rate in LSP-treated specimens in our preliminary tests,DKthwas not established before crack length reached the limiting valueof 5 mm. Therefore, the LSP-treated specimens presented in thisresearch were tested from lower starting value of 7 MPapm. How-ever, to establish equality in test and adequate comparison, addi-tional not-treated specimen #3 was tested from 7 MPapm.During both stages of fatigue testing, the pre-cracking proce-dure and K-decreasing test, the testing resonant frequency, andthe crack length was monitored with the indirect electric potentialdrop method on both sides of the specimen. Fig. 9a presents thefatigue crack length a0as a function of a number of cycles duringpre-cracking and K-decreasing test, where each curve represent asingle specimen. In this plot, each specimen was subjected to thesame testing parameter regime. Moreover, the crack direction didnot deviate from the plane of symmetry for more than 2?. FromFig. 9a, it is evident that fatigue crack initiated earlier in LSP-treated specimens than in not-treated material. The number ofcycles, required to grow a precrack (of the total length of1.5 mm) is largest for the non-treated specimen and lowest forthe LSP 3 specimen. In the linear part of aN curve that is withininterval from 1.2 to 1.5 mm, the crack growth rate was calculatedwith secant method, Table 3. From this data, it was calculated thatthe number of cycles required to grow 1.5 of pre-crack for LSP 1,LSP 2 and LSP 3 specimens decreased by 28%, 33% and 47%,respectively.The resonant testing frequency reduces with the crack length orincreases in case of strain hardening at the tip of the notch or crack.Testing resonant frequency is a function of the elastic spring con-stant of the specimen. Fig. 9b shows testing frequency as a functionof fatigue crack length, f(a0). Overall, the highest frequency wasrecorded during the test of non-treated material. The testing fre-quency was increasing slightly before the fatigue crack initiateddue strain hardening effects in the tip of the notch. After the crackFig. 9. Monitored fatigue crack length and frequency as a function of number ofcycles; (a) fatigue crack length, (b) testing (resonant) frequency.Table 3Fatigue crack growth values during pre-cracking procedure at constant R = 0.1 andconstantDK = 4 MPapm.Specimenda/dNa(m/cycle)Number of cyclesbTimec(min)No treatment4.4 ? 10?93.60 ? 10541LSP 16.9 ? 10?92.57 ? 10525LSP 28.8 ? 10?92.40 ? 10524LSP 312.0 ? 10?91.90 ? 10522aFatigue crack growth rate, estimated between 1.21.5 mm of crack length oflinear aN part.bNumber of cycles required to reach pre-crack length of 1.5 mm.cTime of testing to reach pre-crack length of 1.5 mm.450Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455growth stabilized, the frequency was decreasing with the cracklength. After reaching the pre-crack length of 1.5 mm, the testingmachine was increasing the stress intensity ratio range until7 MPapm was reached. The frequency curve for not-treated spec-imen was on the highest level and most convex among tested spec-imens. The starting frequency standard deviation, as obtained fromn = 5 tests of base material, amounts to 2.1 Hz. Assuming thatsimilar standard deviation is true for LSP-treated specimens, differ-ences of testing frequencies among LSP treatments are notsignificant.A short initial acceleration of crack growth occurs immediatelyafter the increase of stress intensity factor range. The initial fatiguepre-crack lengths increased from 1.5 to 1.64 mm for not-treatedspecimen, to 1.74 mm for the LSP 1 specimen, to 1.76 mm for theLSP 2 specimen and to 1.81 mm for the LSP 3 specimen.First, the point-to-point secant method was used to generateDa/DN(DK) graphs (Fig. 10.) from a(N) data. With this method,small error deviations from continuous real a(N) lead to largerscatter” in fatigue crack growth curves. Variability of measure-ments can be observed for three tested not-treated specimens,(#1, #2 and #3), where #1 and #2 were tested fromDK0= 12 -MPapm and #3 fromDK0= 7 MPapm. The source of variabilityis the inhomogeneity of material as well as the fatigue crackgrowth rate measuring error. From the fatigue crack growth ratedata presented in Fig. 10, it can be concluded that for all specimensafter being treated with laser shock processing, LSP 1, LSP 2 and LSP3 show higher crack growth rate, which supports the earlier con-clusions. In contrast, only small differences are between curves ofindividual LSP-treatments.According to Broek 36, data sets (as shown in Fig. 10) must beinterpreted before they can be used in analysis. In order to obtainplots with less scatter without any meaningful change of the actualcrack growth length, new da/dN data was calculated from best-fitting equations. The a(N) data was fitted using with least squaresmethod with different functions. For example, best fitting functionfor not-treated specimen is the HocketSherby 2D function:aNLSP 1 b ? b ? a ? e?c?xd5where a = 1.43026, b = 3.10645, c = 0.101526, d = 0.48909.The experimental a(N) data for LSP 1 specimen was fitted usingHoerl function:aNLSP 1 abxNc;6where a = 1.11347187, b = 1 and c = 0.125. The a(N) data for LSP 2specimen was fitted using Dr. Hills function:aNLSP 2ahNgjg Ng;7wherea= 4.286326, h = ?2.82718,g= ?0.481,j= 1.8E5. The a(N)data for LSP 3 specimen was fitted using the same function witha= 4.579980126, h = ?3.220466252,g= ?0.459,j= 1.24E5.The data for stable II. region growth was fitted using the Parisequation:dadN C ?DKm8where C and m are material constants. Table 4 gives C and m param-eters with the fitting correlation coefficient R2, which are higherthan 0.98. The NASGRO model 25,26 was used to predict a crackgrowth in other regions. The equation for positive R ratios was used:dadNC ?DKn1 ? DKth=DKp1 ? Kmax=Kcq9whereKmaxDK=1 ? R;10where C, n, p, and q are empirical constants.Fig. 11 represents the da/dN(DK) purified fatigue crack growthplots with practically no scatter (due to the da/dN calculation fromcontinuous function) and with fitted models of Paris and NASGRO.The model coefficients are presented in Table 4. For LSP-treatedspecimens, the FCG curve was not sigmoidal where the slope grad-ually increases at values from 4.6 to 4.9 MPapm. For the Parisregion, the curve was fitted up to change of slope of LPS-treatedspecimens. The experimental data shows change of slope (knee),where one function fromDKthto knee and second from knee toKcwas used. The FCG curves for LSP 1 and LSP 2 almost overlap,indicating the equal crack growth rate. A slightly higher crackgrowth rate was found in the LSP 3 specimen. Therefore, withoutperforming many repetitions, no firm conclusions can be madeabout the impact of LSP-parameters on fatigue crack growth.The fatigue crack in the base material was no longer propagat-ing (da/dN ? 0) at a threshold level of 3.36 MPapm, while thecrack in the LSP 3 specimen at the same SIF was propagating atapprox. 3 ? 10?9m/cycle. Hence, the highest fatigue crack growthrate was found in the LSP 3 specimen with the lowest thresholdvalue of 1.1 MPapm. Although, LSP process induced high densityof dislocations, grain refinement and even compressive residualstresses in the near surface, it is obvious that this alone cannoteffectively reduce the FCG rate. It is clear that all LSP specimensexhibited higher crack propagation rates in comparison to theuntreated specimen due to the equilibrium tensile stresses in thecentre of specimen, which reduced the effect of crack closure. Itis obvious that the compressive residual stresses have not effec-tively decreased the crack growth rate. Chahardehi et al. 2, whostudied the effect of residual stresses arising from LSP, stated thatthe growth rates of the crack observed are more affected by thetensile core than by the compressive surface stresses. In this study,crack length against number of cycles measurements did not showany retardation of fatigue cracks, which is believed to be due to thetensile core in the material that arises as a by-product” of lasershock peening.Fig. 10. Crack growth curves da/dN vs. appliedDK, obtained with direct secantmethod.Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455451It should also be noted that Figs. 10 and 11 are plotted as a func-tion of appliedDK data. At low-stress intensities, crack closureoccurs, and the influence is increasing near the threshold level.According to Borego et al. 27 and Gavras et al. 35, plasticityinduced and roughness induced closure mechanisms are dominantstudied aluminium alloys (among them the 6082-T6 was also thesubject of the research). Considering this, corrected da/dN vs.DKeffcurves would be shifted to the left from da/dN vs.DK data.Another factor to be noted is the sequence of LSP treatment. Incontrast to the majority of the investigations 3,14 in which LSPwas performed after pre-existing fatigue precrack, our specimenswere precracked after LSP. Secondly, LSP were performed overthe edge of the EDM notch, which possibly created tensile hotspots over the notch edge. Similar observations were reported pre-viously by Ivetic et al. 28, investigating the effects of an un-coated” LSP process on open-holed thin Al 7075-T73 specimens.Their results indicate that the sequence of operations (first LSPand than open-hole or vice versa.) has an enormous influence onfatigue crack propagation behaviour. Moreover, it was suggestedthat laser beam, being focused by spilled” plasma by the innerside of the hole, creates a local tensile hot-spot, which drasticallyreduces fatigue life. The same authors argued that this effect isinfluencing only the crack initiation phase, since the progressiverelaxation of residual stresses occurred once the crack has formed.3.6. Fracture toughnessThe fracture toughness Kctest was performed on C(T) speci-mens after FCG tests to evaluate the effect of LSP treatments. Inorder to satisfy plane-strain conditions in fracture toughness test-ing, the thickness B of the specimen should be greater than 2.5(Kc/rY)2. The yield stress and fracture toughness of investigated 6082-T651 plate is 295 MPa. Moreover, based on the literature review,the expected value of fracture toughness of 6082-T6, whichdepends on age-hardening, is between 35 and 40 MPapm 29.Thus, to perform a plane-strain linear elastic fracture toughnesstesting according to the B 2.5(Kc/rY)2, the specimen thicknessshould be 46 mm. Because an aluminium plate with very largethickness is seldom practically useful, transition plane-strain/plane-stress thickness-dependent tests are quite common. In thecase of thinner aluminium plates, the Standard Practice for Frac-ture Toughness of Aluminium Alloys B 646-12 30 provides guid-ance for the testing of aluminium alloys with (i) thin products,with thicknesses of less than 6.3 mm, (ii) intermediate plate thick-nesses, too thin for valid plane-strain fracture toughness tests buttoo thick for treatment as a sheet, thickness range from 6.3 to50 mm, and (iii) relatively thick specimens. The investigated6082-T651 plate with B = 10 mm classify into (ii) intermediatethicknesses, where the fracture toughness shall also be determinedaccording with the test method E399 18 (as supplemented byStandard Practice for Linear-Elastic PlaneStrain Fracture Tough-ness Testing of Aluminium Alloys B645 31).Three specimens without treatment were used to measure theaverage fracture toughness, while a single specimen was used foreach LSP treatment. Table 5 shows the starting crack lengths, a/W ratio, maximum force Pmax, reference force Pq, and the calculatedvalues of fracture toughness Kc. The crack lengths (i.e. the totallength of notch and fatigue crack length) were between 0.67 and0.74 W (recommendation by ASTM E399 is between 0.45 and0.55 W). Fig. 12 shows Pv plots for specimens with starter cracklength of 0.710.74 W. The average fracture toughness of non-treated specimen, based on three separate tests, was found to beTable 4Constants for Paris and NASGRO equation for not-treated and LSP 1, LSP 2, LSP 3 treated specimens.DesignationParisNasgroFromDKthto slope changeFrom slope change to KcCmDKth(MPapm)CnpqCnpqNot-treated8.37 ? 10?93.733.369.56 ? 10?112.750.931.2LSP 13.4 ? 10?113.721.41.05 ? 10?102.700.450.93.5 ? 10?LSP 23.4 ? 10?113.731.41.15 ? 10?102.700.450.91.5 ? 10?103.5760.6LSP 37.4 ? 10?103.521.11.80 ? 10?102.700.450.99 ? 10?102.5060.6Fig. 11. Crack growth curves da/dN vs. appliedDK, obtained from fitting equations.Table 5Results of plane-stress fracture toughness Kctest before and after LSP treatments.TreatmentDesignationa (mm)aWPmax(N)PQ(N)KC(MPapm)KC(MPapm)Not-treated (3 repetitions)#119.880.712849262038.7237.8 0.86#217.600.634175395037.68#318.650.673784343037.00After laser shock treatmentLSP 120.300.722358196029.01LSP 220.650.742175184029.34LSP 320.190.722058195028.21452Z. Bergant et al./International Journal of Fatigue 87 (2016) 44445537.8 MPapm. After LSP treatment, the fracture toughness was 2833% lower compared to the untreated material. Considering thatsimilar value of standard deviation from base material tests(0.86 MPapm) also apply for LSP treatments, it can be concludedthere is not a significant effect on fracture toughness values amongLSP-treated specimens (LSP 1 = 29.01, LSP 2 = 29.34 and LSP3 = 28.21 MPapm).3.7. Fractographic analysisThe macro-fracture surfaces of the untreated and LSP 3 speci-mens were compared under different magnifications using opticaland scanning electron microscopes.Fig. 13a shows a relatively linear shape crack front in represen-tative not-treated specimen. In contrast, Fig. 13b shows macro-fractograph of the LSP 3 specimen, where the curvature of the fati-gue crack front can be observed. Therefore, the effects of residualstresses influence the crack growth in LSP-treated specimens,where the fatigue crack length increases with the depth; it is thelargest in the middle of the specimen. The curvature in crack frontcan be explained through crack retardation at the sides in the pres-ence of surface compressive residual stresses. In contrast, tensileresidual stresses have an adverse effect, because they reduce theeffect of crack closure.Fig. 13b and d shows the detail of fatigue crack surface at near-threshold stress intensity and transition to fracture surface, whichoccured during fracture toughness test. Fractographic observationof the surface where the static fracture toughness test was con-ducted showed typical dimple rupture, caused by a process knownas microvoid coalescence 32.At the near-threshold regime of crack propagation, the fatiguefractographic surfaces exhibited no typical fatigue striation, whichis common for stage I fatigue fracture surfaces 32. The importantphenomenon, which in turns influences the visual appearance offatigue crack surface, is the crack closure. The crack surfaces incontact during fatigue define the texture and crack surface rough-ness. As mentioned earlier, mainly roughness-induced crack clo-sureoccursatnear-thresholdofstressintensityrangeinaluminium alloys 27. Roughness-induced crack closure is gener-ally used to describe contact of faceted surface features which aredimensionally small, e.g. the order of the grain size 33. Based onthe visual appearance of fatigued crack surfaces on SEM micro-graphs, Fig. 13b and d at near threshold, that base material hasfaceted and rounded cleavage planes, followed by crack arrest atDKthof 3.3 MPapm. In contrast, LSP 3 specimens has finer andsmoother asperities with fatigue tear ridges oriented in the direc-tion of crack growth in the region where crack growth stopped atcomparatively much lower threshold value of 1.1 MPapm. Thesmoother surface suggests that lower energy and more brittleFig. 12. Loaddisplacement curves.Fig. 13. Fracture surfaces after fatigue and fracture toughness test.Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455453cleavage fracture occurred along crystallographic planes in LSP-treated specimens. Moreover, visual macro-scope examination(optical macroscope Olympus) confirmed that on LSP 3 specimenthe central part of crack surface appears to be smoother, especiallyin the direction of crack growth. In order to support the qualitativevisual examinations, the roughness on crack surface was measuredusing a contact profilometer (Taylor Hobson) and a micrometersliding table stroke. The area of measurement of roughness is arectangular area of 2.5 ? 2.5 mm in the central part of the cracksurface, as shown in Fig 14a. The number of measurement linesfor each direction on an individual specimen was eight, measuredwith offset increments of around 350lm and a measuring linelength of 2.5 mm. From these values, the average roughness valueRaand standard deviation were calculated and presented in a his-togram, Fig. 14b. From the result it can be depicted that the aver-age roughness in z-direction is similar (average Rafor z-direction isbetween 16.2 and 17.5lm and with an average standard deviationof 2.45lm). Moreover, overall surface roughness Rain z-directionis 3354% higher than in the x-direction, mainly as the conse-quence of grain elongation in the rolling direction during platefabrication.Histograms reveal that the average of crack surface roughnessRain x-direction is higher in the non-treated specimen than all inthree LSP-treated specimens. However, the standard deviation ishigh and, therefore, it is not sufficient to conclude before perform-ing a statistical analysis. The double-sampling hypothesis test withStudents t-distribution was used to determine whether the x-direction roughness of the crack surface in the not-treated speci-men is statistically different from LSP-treated specimens. The basisfor the hypothesis test is roughness measurement data, given inTable 6. To reject hypothesis of sample equality mean, that isH0:l1l2the value of t0in and ta=2;n1n2?2is compared. Ifjt0j ta=2;n1n2?2, the H0is rejected and if jt0j ta=2;n1n2?2the H0hypothesis is accepted. The value of statistic was calculated as:t0y1? y2Sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi1=n1 1=n2p;11where y1and y2are average value of the samples, n1and n2is num-ber of measurements, S is the variance estimator for both samples;it is calculated:S2n1? 1S21 n2? 1S22n1 n2? 2;12where S21and S22are variances of samples. Furthermore, for the sig-nificancelevelofa= 0.05andfromStudentst-distribution,ta=2;n1n2?2 t0:025;14 2:1448 is obtained 34. Thus, from Eq. (11)statistic for LSP 1 amounts t0= 2.83, for LSP 2 t0= 3.58 and, asexpected lowest value in LSP 3 specimen, t0= 2.389. In all threecases, the H0hypothesis is rejected, which means that the crackroughness in x-direction in the not-treated specimen is significantlyhigher than the roughness in the x-direction of LSP-treated speci-mens. However, one should be aware that the measured profile ofcontactprofilometerdoesnotperfectlyfollowthemicro-asperities of cracked surface. Nevertheless, the crack surface rough-ness analysis confirms the visual qualitative SEM and macro-scopeexamination findings.All the above results indicate that extreme caution should bemade when introducing compressive residual stresses on the sur-face of the component, which will be subjected to dynamic loadingunder an HCF regime. Moreover, despite the fact that surface resid-ual stress-measuring techniques are commonly used in practice forthe prediction of fatigue life, this cannot serve as any ultimate cri-teria. To this end, one should note that although LSP has beenwidely approved as an advanced, very efficient treatment forimproving fatigue life of metallic components, process parametersneed to be optimized for every application and the operationsshould always be performed in the optimum sequence in orderto prevent any detrimental effect of the treated material.4. ConclusionsThe LSP process is an innovative new method to increase thelifetime of dynamically loaded components. Although it wasexpected that compressive residual stresses and refined grains inthe near surface would retard the FCG rate after LSP, that wasnot the case here. Based on the research conducted, the followingcan be concluded: After LSP surface roughness, the parameter Ra was increased bya factor of more than 7 in comparison to the untreated material. TEM analysis confirmed tangled dislocations, with dislocationcells and refined grains. LSP-induced near-surface compressive stresses was found inLSP-treated C(T) specimens.Fig. 14. Crack surface roughness measurement (a) location of random measure-ments in x and z direction in rectangular area of crack surface, (b) histograms ofaveraged roughness results Rawith standard deviation error plot.Table 6Crack surface roughness measurement Ralm in (x-direction), n = 8 for the test ofthe equality of samples.i12345678Not-treated7.8213.27.989.3711.011.3012.714.7LSP 15.47.355.010.57.8111.105.688.2LSP 25.95.988.479.24.029.044.277.2LSP 38.8512.38.047.898.856.296.428.4454Z. Bergant et al./International Journal of Fatigue 87 (2016) 444455 The crack growth threshold value for untreated specimen washigher than after LSP treatment. The resonant tes
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