FE modeling of surface stresses in erosion-resistant coating under single particle impact

FEM prediction of welding residual stresses in a SUS304 girth-welded pipe with emphasison stress distribution near weld start/end location Dean Deng a and Shoichi Kiyoshima ba College of Materials Science and Engineering, Chongqing University, Shazheng Street 174,Shapingba, Chongqing 400045, Chinab Research Center of Computational Mechanics Inc., Togoshi NI-Bldg, 1-7-1 Togoshi, Shinagawa-ku,Tokyo 142-0041, JapanReceived 2 July 2010;revised 31 August 2010;accepted 23 September 2010.Available online 16 October 2010.AbstractA finite element approach based on Quick Welder software is developed tosimulate welding temperature field and welding residual stress distribution in a 3D multi-pass girth-welded pipe model. The characteristics of welding residual stress distributions in a SUS304 stainless steel pipe induced by heating with a tungsten inert gas arc welding torch are investigated numerically.Meanwhile, an emphasis is focused on examining the welding residual stress distributions in and near the weld start/end location. Moreover, the residual stresses predicted by the present computational approach are compared with the measured data; and the comparison suggests that the numerical simulation method has basically captured the feature of welding residual stress distribution near the weld start/end region. The numerical simulation results show that both the hoop and the axial residual stresses near the weld start/end region have sharp gradients and are significantly different from those in the steady range.Keywords: Finite element analysis; Numerical simulation; Multi-pass welding; Residual stressArticle Outline1.Introduction2.Brief description on experimental procedure3.Finite element analysis3.1. Thermal analysis3.2. Mechanical analysis4.Results of thermal analysis5.Results of mechanical analysis5.1. Characteristics of residual stress distribution of girth-welded pipe5.2. Comparison of simulation results and measured data5.3. Welding residual stresses changing with weld pass6.ConclusionsAcknowledgementsReferences1. IntroductionResidual stress in a steel structure induced by thermal process such as welding is oftenup to or even over the yield strength at room temperature. Tensile residual stress is a mainfactor resulting in stress corrosion cracking, fatigue damage and brittle fracture. When therisk for growth of defects such as surface cracking at piping systems in nuclear power plantsis assessed, the welding residual stresses maybe give a large contribution to thetotal stress field. Therefore, it is very important to obtain the accurate informationon welding residual stress distribution in a welded joint.Generally, there are two methods for determining welding residual stress. One isexperimental method, and the other is numerical simulation technology. Limited by physicalconditions, experimental method is usually adopted to get welding residual stresses insmall-scale weldments [1]. With the development of computer hardware and software,numerical simulation technology has increasingly been employed to predict welding residualstress in various welded joints [2], [3] and [4]. In the past decades, many numerical modelswere developed to simulate welding residual stress in various welded joints. Among thesemodels, a number of models studied the welding residual stresses in girth-welded pipejoints through using thermal elastic plastic finite element method. However, most of themonly investigated welding residual stresses in the steady regions of a girth-welded pipe byusing axis-symmetric 2D models or 3D models [5], [6], [7] and [8]. Deng andMurakawa [9] studied welding residual stress in SUS304 pipe with 6 mm thickness using3D finite element model, however, they laid emphasis upon comparison between thesimulation results and the measured data, and the authors did not examine the welding residual stress distribution near weld start/end location in detail. Recently, Lee andChang[10] simulated welding residual stress distributions in a pipe model with considering 3D effect, but they only provided very limited information on residual stress distribution near the weld start/stop location. In actual pipe welds there is a start and end point of each weld pass, which always causes a discontinuity of stress state near this location. There is a tendency for the formation of cracking at the crater during welding[11]. In addition, fatigue fracture usually initiates from the surface especially at the weld start/stop position during service [11]. Therefore, it is necessary to obtain the accurate residual stress distribution in and near the weld start/end region.In this work, a finite element approach based on Quick Welder software [12] is developed to simulate welding residual stresses in a 3D SUS304 multi-pass girth-welded pipe model. In the developed approach, a moving heat source model is used to simulate the heat input, and the temperature-dependent thermo-physical and mechanical properties are taken intoaccount. In this study, an emphasis is focused on examining the characteristics of welding residual stress distributions in and near the weld start/end region. The simulation results are compared with the measured data cited from existing literature [13], and the effectiveness of the numerical model has been verified.2. Brief description on experimental procedureIn order to help readers to understand the comparison between simulation results obtained by the current study and the measured data, a brief description on the experimentalprocedure is provided in this section. It should be stated that the experiments were carried out by Japan power engineering and inspection corporation (JAPEIC) [13].Two mock-ups (mock-up A and mock-up B) were manufactured for measuring welding residual stresses in and near weld start/end location of girth-welded pipes. The materials of these two mock-ups (pipes) are austenitic stainless steel SUS304. The dimensions, the shape of groove and the weld passes are shown in Fig. 1. As a fundamental study, the complexity caused by filler metal for measuring residual stress was avoided when the experiment was designed. Tungsten inert gas (TIG) arc welding was used in theexperiments, and the TIG torch only heated the surface of the groove without filler metal. To investigate how welding residual stresses change with weld passes, only pass 1 was performed in mock-up A, while pass 1, pass 2 and pass 3 were sequentially heated in mock-up B. In mock-up B, the inter-pass temperature was lower than 50 °C.The welding direction of each pass is shown in Fig. 2. In the same figure, the central angle (θ) is also defined. As shown in this figure, the weld start/end location of mock-up A was 0°location. In mock-up B, the weld start/end of the first and the third weld passes was 0°location, while that of the second weld pass was 180° location. During welding, thewelding current, the arc voltage and the welding speed were 132A, 10 V and 90 mm/min, respectively. When the torch reached to the end of weld the welding mode was changed into the crater treatment mode; and the torch stopped 3 s with decreasingthe welding current. The welding current and weld speed are schematically shown in Fig.3.Full-size image (12K)Fig. 1.Dimensions of welded pipe, shape of groove and weld passes.Full-size image (29K)Fig. 2.Weld start/end position and welding direction of each pass.Full-size image (13K)Fig. 3.Welding current and welding speed used in mock-ups.After welding, neutron diffraction method was used to measure welding residualstresses near the weld start/end region of the two mock-ups. The measured locations are directly under the center of the second pass with 2.5 mm depth from the surface. The grey dots schematically show the measured locations in Fig. 4. In this figure, −20° and −10° correspond to 340° and 350° (central angle), respectively.Full-size image (20K)Fig. 4.Measured locations of welding residual stress.3. Finite element analysisIn this section, a computational approach based on Quick Welder software is developed to simulate temperature fields and residual stress distributions induced by tungsten inert gas (TIG) arc welding in a SUS304 pipe. The finite element model is shown in Fig. 5. The whole length of the pipe is 500 mm; the inside diameter is 255 mm; and the thickness of wall is25 mm. The groove shape is shown in Fig. 6. To roughly consider the penetration shapeinduced by TIG welding torch the heated zones of cross-section in three passes are also defined in Fig. 6. It should be stressed that the simulation results predicted by the finite element model are compared with the measurements cited from Ref. [13], so the groove shape of the numerical model is identical to that of the corresponding mock-up.Full-size image (38K)Fig. 5.3D finite element model.Full-size image (24K)Fig. 6.Shape of groove, weld passes and their sections used in finite element model.In order to reasonably simulate the temperature fields and the residual stress distributions induced by TIG welding process, a finer mesh is designed at the weld zone and its vicinity. To balance the accuracy of prediction and computing time, the number of divisions in the circumferential direction is 60 as shown in Fig. 5. In the finite element model, the number of nodes is 20,040; and that of elements is 23,400.In this study, the thermo-mechanical behavior is simulated using a sequentially coupled formulation because the dimensional change during the entire welding process is negligible and mechanical work is insignificant compared to the thermal energy from the arc. The heat conduction problem is solved independently from the stress–strain problem to obtain temperature histories. However, the formulation takes into account the contributions of the transient temperature field to the mechanical analysis through thermal expansion coefficient, as well as temperature-dependent properties such as Young’s modulus, yield strength and Poisson’s ratio. The solution procedure consists of two steps. At the first step, thetemperature histories of all nodes in the finite element model are computed according to the given welding conditions and the thermal boundary conditions. At the second step, the temperature histories obtained at the first step are taken as thermal loads in the mechanical analysis.3.1. Thermal analysisIn this work, only the torch is used to heat the surface of SUS304 pipe and no filler metal is added to the groove. Welding heat transfer analysis with given welding conditions is performed in the 3D finite element model. At this step, temperature histories at all nodes are computed during the multi-pass welding process. In the course of welding, the governing equation for transient non-linear heat transfer analyses is:(1)where λ, T, q, ρ and c arei thermal conductivity, temperature, rate of internal heat generation, density and specific heat capacity, respectively. For TIG welding process, the heat input due to moving arc is often modeled by using surface heat source model such as typical Gaussian heat source [14].When a Gaussian heat source is chosen, correspondingly a fine mesh division especially in the circumferential direction is required. To reduce the number of element and savecomputing time, the moving heat source is treated as volumetric heat flux with uniform density instead of Gaussian surface heat source in the present analysis.In the present finite element model, each weld pass is divided into a number of small parts (mesh blocks) with a same length, and each mesh block is sequentially heated to simulate the movement of welding torch. The volume of the moving heat source is equal to that of those elements composing the corresponding weld bead in one mesh block. The image of heat source is shown in Fig. 7.Full-size image (52K)Fig. 7.Heat source defined in the finite element model.After each weld bead and the welding conditions are defined, the volume of the heat sourceand the heat flux are automatically calculated in the finite element model. In the finiteelement model, the welding direction, the welding sequence and the welding speed arethe same as those used in the mock-up [13].The volumetric heat flux of each weld pass is determined using the following equation:(2)q=ηUI/V H where U is the arc voltage, I is the welding current, V H is the volume of heat isource, and η is the arc efficiency.The value of arc efficiency is assumed to be 0.7 for TIG welding process.Because the number of divisions in circumferential direction is 60, the heating time (t h) foreach mesh block is calculated as follows:(3)t=L/(60v w)where L is the total length of each weld pass, and v w is the welding speed of heach weld pass. In this simulation, the arc voltage, the current and the welding speed areidentical to those used in the experiments.Besides considering the moving heat source, heat losses due to convention and radiationare also taken into account in the finite element models. In the thermal analysis, all surfacesof the finite element model are assumed to lose heat by convention to the surrounding air.Heat loss due to convection (q c) is taken into account using Newton’s law:(4)q=-h c(T s-T0)where h c is the heat transfer coefficient; T s is the surface temperature of the cweldment; and T0 is the ambient temperature. The ambient temperature is assumed to be18 °C.In the present simulation, the heat transfer coefficient (h c) is set to be a15 × 10−6 W/(mm2 °C) [14].In addition, the heat loss due to radiation is modeled using Stefan–Boltzmann’s law:(5)q=-εσ[(T s+273)4-(T o+273)4]where ε is the emissivity, and σ is the Stefan–Boltzmannrconstant [15].In the present study, the emissivity (ε) is assumed to be 0.8 [14].Both the thermal effect due to solidification of the weld zone and the influence of fluid flow ofthe weld pool on the welding residual stress are insignificant, so these two factors areneglected in the thermal analyses.To consider the crater treatment at each weld end, the heat input used in this location isassumed to be twice as that used in the rest range. The inter-pass temperature is strictlycontrolled to be lower than 50 °C during the multi-pass welding process through setting anappropriate interval (30 min) between two passes. The temperature-dependent thermalproperties such as thermal conductivity, specific heat capacity and density are used in the finite element model. The temperature-dependent thermo-physical properties of SUS304 used in this study were tested by experiment [9] as shown in Fig. 8.Full-size image (36K)Fig. 8.Temperature-dependent thermal physical properties of SUS304.3.2. Mechanical analysisIn the mechanical analysis, the same finite element mesh used in the thermal analysis is employed. The simulation is conducted using the temperature histories computed by the thermal analysis as the input information. Generally, during the welding process, besides the elastic, plastic and thermal strains, the strains due to solid-state phase transformation and creep potentially give some contributions to the total strain. Because SUS304 stainless steel has no solid-state phase transformation during cooling and the heating time is relatively short, it can be expected that the strains due to phase change and creep can be neglected in the present simulation. The total strain increment {d} at a material point can be expressed as the summation of the elastic, plastic and thermal strain.(6){dε}={dεe}+{dεp}+{dεth}where {d e}, {d p} and {d th} is the elastic strain, plastic and thermalstrain increment, respectively.In the mechanical simulation, the elastic behavior is modeled using the isotropic Hooke’s rule with temperature-dependent Young’s modulus. The thermal strain is considered usingthermal expansion coefficient. For the plastic behavior, a rate-independent plastic model is employed. The yield criterion is the Von Mises yield surface. In the present model, the strain hardening is taken into account using the linear isotropic hardening law. Meanwhile, the annealing effect is also included in the numerical model.When the temperature of material point (integrated point in an element) is higher than the annealing temperature the material will lose its hardening memory. In the finite elementmodels, the effect of prior work hardening is removed by setting the equivalent plastic strain() to zero. The annealing temperature is assumed to be 800 °C for SUS304 stainless steel [14]. If the temperature of material point falls below the annealing temperature at a subsequent point in time, the material point can work harden again.Similar to the thermal analysis, the temperature-dependent mechanical properties are usedin the mechanical model. Fig. 9shows the Young’s modulus, the yield strength, thePoisson’s ratio and the thermal expa nsion coefficient of SUS304 as functions of temperature,and Fig. 10 shows the strain hardening coefficients at different temperatures [9].Full-size image (40K)Fig. 9.Temperature-dependent mechanical properties of SUS304.Full-size image (29K)Fig. 10.Strain hardening coefficients of SUS304 at different temperature.The mechanical boundary conditions are constrained only for preventing rigid body motionwith considering symmetry. The arrows towards points A–C in Fig. 5 show the restraintconditions. These three points locate in the Y–Z plane.The simulations are conducted on a personal computer with a 3.2 GHz CPU and a 4.0 GB memory. The total computational time used in the thermo-mechanical analyses is about 100 h.4. Results of thermal analysisTemperature field induced by moving heat source is shown in Fig. 11. From this figure, the image of weld pool and the maximum temperature can be seen. Thermal cycles at three points defined in Fig. 12 during the entire welding process are shown in Fig. 13. Because the second weld pass started at 180° location, we can see that each thermal cycle has two peak temperatures and point A was molten twice during the second welding.Full-size image (12K)Fig. 11.Temperature field due to the second weld pass.Full-size image (19K)Fig. 12.Points A–C at the section with 180° central angle.Full-size image (22K)Fig. 13.Temperature histories of points A–C.Fig. 14 shows the penetration shapes (molten zones) with peak temperature higher than 1400 °C (the melting point of SUS304) in three typical sections. Fig. 14a shows thepenetration shape of the section with 0° central angle. Because the first and the third weld passes started/ended at 0° location and the crater treatment was considered in the thermal analysis, the penetration depths (D p1 and D p3) of both sides are larger than that (D p2) of the center. Fig. 14c shows the penetration shape of the section with 180° central angle. Contrast to Fig. 14a, the penetration depth of the center is larger than those of both sides. Fig. 14b shows the penetration shape of the section with 90° central angle. This figure tells us that the penetration depths of three weld passes have no significant difference because the heat input used in each pass was identical at this location.Full-size image (22K)Fig. 14.Penetration shapes of different sections: (a) 0° location, (b) 90° location, (c) 180° location.5. Results of mechanical analysis5.1. Characteristics of residual stress distribution of girth-welded pipeFig. 15 shows the contours of hoop residual stress on four typical sections with different central angles. A common feature on the four sections is that large tensilehoop stresses generated in and near the weld zone and compressivehoop stresses produced at the regions away from the weld zone. Carefully comparing the four sections we can find that the hoop stress distribution on the section with 0° central angle is different from those on the other three sections. In the same figure, we can also obverse that the maximum hoop stress on the groove surface of the section with 0° central angle is larger than those of the other three sections. Fig. 16 qualitatively compares the hoop stress distributions along the axial direction on the inside surfaces of four sections. This figure suggests that the hoop stress distribution shape on the inside surface of the section with 0° central angle is similar to those on the other three sections, but both the peak tensile stress and the maximum compressive stress (absolute value) is smaller than those on the other three sections. Carefully observing this figure, we can find that thehoop stress distributions on these two locations with 90° and 270° central angle almost have no difference, while the hoop stress distribution on the surface with 180° central angle has a small difference with the former two locations. After the second welding, which started and ended at the location with 180° central angle, stress distributions with sharp variations generated near this location. However, the stress distribution induced by the previous passes (the second and the first passes) was significantly changed by the third (final) welding. The location with 180° central angle is a steady range during thethird welding, so it is not strange that the difference of stress distribution among the sections with 90°, 180° and 270° is small. In addition, Fig. 16 indicates that thehoop stress distribution have an asymmetric shape with respect to the weld center because the three weld passes were performed using a sequential welding procedure.Full-size image (22K)Fig. 15.Contours of hoop stress of four sections with different central angles.Full-size image (33K)Fig. 16.Hoop stress distributions on the inside surfaces.Fig. 17 shows the contours of axial residual stress on four typical sections with different central angles. It is clear that the axial stress distribution on the section with 0° central angle is significantly different from those on the other three locations. Fig. 18 qualitatively compares the axial stress distributions along the axial direction on the inside surfaces of four sections. This figure indicates that although the shapes of axial stress distribution on the inside surfaces of the four locations are similar, however the axial stresses on the surface with 0° location are significantly higher than those on the other three locations. From the same figure, one can find that the axial stress distributions on the locations with 90° and 270° central angle are very similar both in shape and in magnitude, while theaxial stress distribution on the inside surface with 180° central angle has a bit difference with the former two locations.Full-size image (26K)Fig. 17.Axial stress distributions of four sections with different central angles.Full-size image (36K)Fig. 18.Axial stress distributions on the inside surfaces.In this study, a main task is to investigate the welding residual stress distribution in and near the weld start/end location. To clarify the feature of welding residual stress distribution, the middle section of the welded pipe is defined in Fig. 19. The final hoop andaxial stress distributions of the middle section are shown in [Fig. 20] and [Fig. 21], respectively. It is very clear that both the hoop and axial stresses in and near the weldstart/end location are much different from those in the rest region. Fig. 22 shows the hoop and axial stress distributions on the outside surface around the pipe. From this figure, we can see that both the hoop and axial stresses within the range from −30° (330°) to +30° have large gradients. In the same figure, we can see another interesting feature that the hoop and axial stresses slightly change with central angle near 180° location. The reason resulting in the variation is that the second weld pass started and ended at 180° location and the third welding could not completely cancel out the residual stresses induced by the previous passes especially by the second pass. Fig. 23 shows the hoop andaxial stress distributions on the inside surface around the welded pipe, and the features reflected by this figure are much similar to Fig. 22. Comparing [Fig. 22] and [Fig. 23], one can find that although these two figure are similar, however both the hoop and axial stresses on the inside surface are higher than those on the outside surface.Full-size image (63K)Fig. 19.Definition of middle section.Full-size image (21K)Fig. 20.Hoop stress distribution on the middle section.Full-size image (23K)Fig. 21.Axial stress distribution on the middle section.Full-size image (25K) Fig. 22.Residual stress distribution on the outside surface of mid-section.Full-size image (24K)Fig. 23.Residual stress distribution on the inside surface of mid-section.5.2. Comparison of simulation results and measured dataBased on the above simulation results, we have known that the welding residualstress distributions near the weld start/end location are significantly different from those on the rest region. In this section, the measured data [13] are compared with the simulation results, and the conclusions obtained from the present numerical simulation are further verified. It should be stated that the results for pass 1 are from mock-up A and pass 3 from mock-up B.Fig. 24 compares the hoop stresses predicted by the finite element model and theexperimental values measured by neutron diffraction method. Similar to Fig. 4, −30°, −20° and −10° used in Fig. 24 represents 330°, 340° and 350° central angle, respectively. This figure tells us that the hoop stresses (solid curve) after the third welding predicted by the finite element model generally match the corresponding measured data (broken curve with dots). Meanwhile, Fig. 24 indicates that the hoop stresses(broken curve) after thefirst welding computed by the numerical model are significantly higher than the measured data (broken curve with square marks), however, the shape of the broken curve is similar to that of the broken curve with square marks. In the same figure, both the simulation results and the measured data show that the hoop stress sharply changes with central angle near the weld start/end region.Full-size image (35K)Fig. 24.Hoop stress distributions near the weld start/end region.Here, we give some explanations on why the large discrepancy producedbetween predictions and the measurements. Generally speaking, both the inaccuracies of finite element model and the errors caused by measuring method are the origins of discrepancy. On the aspect of finite element model, the current meshes (especially the number of division in the circumferential direction) and the material model (e.g. the strain hardening rule and the assumed annealing temperature) potentially have influence on the accuracies of predictions. Each strain component value measured by the neutron diffraction method represents the average one in a sampling gauge volume (e.g.2.5 mm × 2.5 mm × 2.5 mm). From the simulation results as shown in Fig. 15, we can see a relatively large stress gradient near the measurement locations (2.5 mm depth under the surface of groove). On the aspect of the experiment, it can be inferred that the gauge volume and the large stress gradient near the measurement locations possibly affect the accuracy of measured data.Fig. 25 compares the axial stresses predicted by the finite element model and the measurements obtained by neutron diffraction method. Similar to Fig. 24, the solid curve and the broken curve represent the axial stresses predicted by the finite element model after the third welding and the first welding, respectively; while the broken curve with dots and the broken curve with square marks express the axial stresses measured by experiment after the third welding and the first welding,respectively. This figure suggest that even though there are some differences between the predictions and the measurements, however, the simulation results match the corresponding measured data on the whole. In other words, the current numerical model has basically captured the feature of axial residual stress distribution near the weld start/end region. Similar to the hoop stress, both the simulation results and the measured data indicate that the axial stress distribution has large gradients near the weld start/end region.。

材料研究与应用 2024，18（1）：157‐165Materials Research and ApplicationEmail ：clyjyyy@http ：//mra.ijournals.cn 激光熔覆涂层对高锰钢耐磨性能的影响韩纪鹏1,2,赵德刚1*,刘国祥2,张文彬2,陈小伟2（1.济南大学材料科学与工程学院，山东 济南 250022； 2.浙江武精机器制造有限公司，浙江 金华 321200）摘要： 使用FeCrVSi 和Ni+WC 涂层粉末，在高锰钢材料表面成功制备了激光熔覆涂层，并对涂层组织形貌、显微硬度和耐磨性进行了研究。

结果表明，两种涂层均可提升高锰钢基体的耐磨性和显微硬度，Fe‐CrVSi 涂层对基材性能的提升更佳，添加FeCrVSi 和Ni+WC 涂层的材料表面磨损量分别降低9.5%和6.3%，硬度分别为470—550 HV 和500—630 HV ，高于基体的250 HV ，这主要源于合金元素的固溶强化作用和激光熔覆过程的激冷效果。

在高应力载荷冲击过程中，涂层为高锰钢提供了第一层防护，以高硬度质点抵抗磨料破坏；同时，表层基材发生塑性变形和强化，产生形变诱导马氏体和栾晶硬化，提供了很高的硬化效应，在协同强化的作用下为高锰钢提供了更高的强度和硬度，提升了其耐磨性能。

关键词： 激光熔覆；高锰钢；涂层；耐磨性；碳化物；硬化效应；协同作用；激冷效果中图分类号：TG156；TB114 文献标志码： A 文章编号：1673-9981（2024）01-0157-09引文格式：韩纪鹏，赵德刚，刘国祥，等.激光熔覆涂层对高锰钢耐磨性能的影响［J ］.材料研究与应用，2024，18（1）：157-165.HAN Jipeng ，ZHAO Degang ，LIU Guoxiang ，et al.Influence of Laser Cladding Coating on Wear Resistance Property of High Mn Steel ［J ］.Materials Research and Application ，2024，18（1）：157-165.0　引言高锰钢具有优异的耐磨性能及较低的成本，广泛应用于矿山、机械、冶金、铁路、电力等行业中，如破碎机的衬板、动锥、定锥、锤头等。

Trans. Nonferrous Met. Soc. China 22(2012) 1064í1072Correlation between welding and hardening parameters offriction stir welded joints of 2017 aluminum alloyHassen BOUZAIENE, Mohamed-Ali REZGUI, Mahfoudh AYADI, Ali ZGHALResearch Unit in Solid Mechanics, Structures and Technological Development (99-UR11-46),Higher School of Sciences and Techniques of Tunis, TunisiaReceived 7 September 2011; accepted 1 January 2011Abstract: An experimental study was undertaken to express the hardening Swift law according to friction stir welding (FSW) aluminum alloy 2017. Tensile tests of welded joints were run in accordance with face centered composite design. Two types of identified models based on least square method and response surface method were used to assess the contribution of FSW independent factors on the hardening parameters. These models were introduced into finite-element code “Abaqus” to simulate tensile tests of welded joints. The relative average deviation criterion, between the experimental data and the numerical simulations of tension-elongation of tensile tests, shows good agreement between the experimental results and the predicted hardening models. These results can be used to perform multi-criteria optimization for carrying out specific welds or conducting numerical simulation of plastic deformation of forming process of FSW parts such as hydroforming, bending and forging.Key words: friction stir welding; response surface methodology; face centered central composite design; hardening; simulation; relative average deviation criterion1 IntroductionFriction stir welding (FSW) is initially invented and patented at the Welding Institute, Cambridge, United Kingdom (TWI) in 1991 [1] to improve welded joint quality of aluminum alloys. FSW is a solid state joining process which was therefore developed systematically for material difficult to weld and then extended to dissimilar material welding [2], and underwater welding [3]. It is a continuous and autogenously process. It makes use of a rotating tool pin moving along the joint interface and a tool shoulder applying a severe plastic deformation [4].The process is completely mechanical, therefore welding operation and weld energy are accurately controlled. B asing on the same welding parameters, welding joint quality is similar from a weld to another.Approximate models show that FSW could be successfully modeled as a forging and extrusion process [5]. The plastic deformation field in FSW is compared with that in metal cutting [6í8]. The predominant deformation during FSW, particularly in vicinities of thetool, is expected to be simple shear, and parallel to the tool surface [9]. When the workpiece material sticks to the tool, heat is generated at the tool/workpiece contact due to shear deformation. The material becomes in paste state favoring the stirring process within the thermomechanically affected zone, causing a large plastic deformation which alters micro and macro structure and changes properties in polycrystalline materials [10].The development of the mechanical behavior model, of heterogeneous structure of the welded zone, is based on a composite material approach, therefore it must takes into account material properties associated with the different welded regions [11]. The global mechanical behavior of FSW joint was studied through the measurement of stress strain performed in transverse [12,13] and longitudinal [14] directions compared with the weld direction. Finite element models were also developed to study the flow patterns and the residual stresses in FSW [15]. B ased on all these models, numerical simulations were performed in order to investigate the effects of welding parameters and tool geometry on welded material behaviors [16] to predict the feasibility of the process on various shape parts [17].Corresponding author: Mohamed-Ali REZGUI; E-mail: mohamedali.rezgui@ DOI: 10.1016/S1003-6326(11)61284-3Hassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í1072 1065 However, the majority of optimization studies of theFSW process were carried out without being connectedto FSW parameters.In the present study, from experimental andmodeling standpoint, the mechanical behavior of FSWaluminum alloy 2017 was examined by performingtensile tests in longitudinal direction compared with theweld direction. It is a matter of identifying the materialparameters of Swift hardening law [18] according to theFSW parameters, so mechanical properties could bepredicted and optimized under FSW operating conditions.The strategy carried out rests on the response surfacemethod (RSM) involving a face centered centralcomposite design to fit an empirical models of materialparameters of Swift hardening law. RSM is a collectionof mathematical and statistical technique, useful formodeling and analysis problems in which response ofinterest is influenced by several variables; its objective isto optimize this response [19]. The diagnostic checkingtests provided by the analysis of variance (ANOV A) suchas sequential F-test, Lack-of-Fit (LoF) test, coefficient ofdetermination (R2), adjusted coefficient of determination(2adjR) are used to select the adequacy models [20].2 Experimental2.1 Welding processThe aluminum alloy 2017 chosen for investigationhas good mechanical characteristics (Table 1), excellentmachinability and formability, and is mostly used ingeneral mechanics applications from high strengthsuitable for heavy-duty structural parts.Table 1 Mechanical properties of aluminum alloy 2017Ultimate tensile strength/MPaYieldstrength/MPaElongation/%Vickershardness427 276 22 118 The experimental set up used in this study was designed in Kef Institute of Technology (Tunisia). A 7.5 kW powered universal mill (Momac model) with 5 to 1700 r/min and welding feed rate ranging from 16 to 1080 mm/min was used. Aluminum alloy 2017 plate of6 mm in thickness was cut and machined into rectangular welding samples of 250 mm×90 mm. Welding test was performed using two samples in butt-configuration, in contact along their larger edge, fixed on a metal frame which was clamped on the machine milling table.To ensure the repeatability of the FSW process, clamping torque and flatness surface of the plates to be welded are controlled for each welding test. At the end of welding operation, around 80 s are respected before the withdrawal of the tool and the extracting of the welded parts. In this experimental study, we purpose to screen theeffects of three operating factors, i.e. tool rotational speed N, tool welding feed F and diameter ratio r, on hardening parameters from Swift’s hardening law such as strength coefficient (k), initial yield strain (İ0) and hardening exponent (n). The ratio (r=d/D) of pin diameter (d) to shoulder diameter (D), is intended to optimize the tool geometry [21í23]. The welding tool is manufactured from a high alloy steel (Fig. 1).Fig. 1 FSW tool geometry (mm)Preliminary welding tests were performed to identify both higher and lower levels of each considered factors. These limits are fixed from visual inspections of the external morphology and cross sections of the welded joints with no macroscopic defects such as surface irregularities, excessive flash, and lack of penetration or surface-open tunnels. However, among these limits one is not sure to have a safe welded joint so often, but they show great potential on defect avoidance. Figure 2 shows some external macroscopic defects observed beyond the limit levels established for each factor. Table 2 lists the processing factors as well as levels assigned to each, and Table 3 shows the fixed levels for other factors needed to success the welding tests.A face centered central composite design, which comes under the RSM approach, with three factors was used to characterize the nature of the welded joints by determining hardening parameters. In this design the star points are at the center of each face of the factorial space (Į=±1), all factors are run at three levels, which are í1, 0, +1 in term of the coded values (Table 4). The experiment plan has been run in random way to avoid systematic errors.2.2 Tensile testsThe tensile tests are performed on a Testometric’s universal testing machines FSí300 kN. The tensile test specimens (ASME E8Mí04) proposed for characterizing the mechanical behavior of the FSW joint, were cut inHassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í10721066Fig. 2 Types of macroscopic defectsTable 2 Levels for operating parameters for FSW processFactorLow level (í1) Center point (0) High level(+1)N /(r·min í1) 653 910 1280 F /(mm·min í1)67 86 109r /%33 39 44Table 3 Welding parametersPin height/ mm Shoulder diameter/ mm Small diameter pin/mm Tool’s inclination angle/(°) Penetrationdepth of shoulder/mm5.3 18 4 30.78longitudinal direction compared with the weld direction, so that active zone is enclosed in the central weld zone (Fig. 3). Figure 4 shows the tensile specimens after fracture.Ultimately, it is a matter of experimental evaluation of hardening parameters of the behavior of FSW joints (k , İ0, n ) according to Swift’s hardening law:n k )(p 0H H V (1)These parameters are required to identify the plastic deformation aptitude of the FSW joints. They are also needed for numerical simulations of forming operations on welded plates. The hardening parameters have been calculated by least square method (LSM) from the stressüstrain curves data. Table 4 shows the experimental design as well as dataset performance characteristics according to the FSW parameters of aluminum Alloy 2017.3 Experimental results3.1 Development of mathematical modelsAlthough the basic principles of FSW are very simple, it involves complex phenomena related to thermo-mechanical and metallurgical transformation that causes strong microstructural heterogeneities in the welded zone. From an energy standpoint, welding process is generated by converting mechanical energy provided by FSW tool into other types of energy such as heat, plastic deformation and microstructural transformations. The nonlinear character of these different dissipation forms can justify research for nonlinear prediction models whose accuracy generally depends on the order of the models relating the responses to welding parameters. For this reason, we chose the RSM which is helpful in developing a suitable approximation for the true functional relationships between quantitative factors (x 1, x 2, Ă, x k ) and the response surface or response functions Y (k , İ0, n ) that may characterize the nature of the welded joints as follows:r 21),,,(e x x x f Y k (2)Hassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í10721067Table 4 Face centered central composite design for FSW of aluminum alloy 2017Factors levelCoded Actual Hardening parameterTypeStandard orderN F r N /(r·m í1)F /(mm·min í1)r /% k /MPan İ0/%1 í1 í1 í165367 33629.7 0.3296 0.00202 1 í1 í1 1280 67 33 654.7 0.4514 0.0035 3 í1 1 í1 653109 33 587.8 0.3712 0.0025 4 1 1 í1 1280 109 33 689.2 0.4856 0.00555 í1 í1 1 653 67 44 642.3 0.4524 0.00256 1 í1 1128067 44 218.6 0.2447 0.0015 7 í1 1 1 653 109 44 685.5 0.4885 0.0035 Factorialdesign8 1 1 1 1280 109 44 332.5 0.3405 0.00209 0 0 0 91086 39 624.9 0.4257 0.0025 10 0 0 0 910 86 39 639.9 0.4292 0.0025 11 0 0 0 910 86 39 640.9 0.4011 0.0020 Center point12 0 0 0 910 86 39 598.6 0.3960 0.0023 13 í1 0 0 653 86 39 690.6 0.4748 0.0027 14 1 0 0 128086 39 505.6 0.3909 0.0030 15 0 í1 091067 39499 0.3317 0.001716 0 1 0 910 109 39 545.6 0.4157 0.0026 17 0 0 í1 910 86 33 672.1 0.4385 0.0027 Star point18 0 019108644 509.7 0.41750.0019Fig. 3 Tensile test specimens (ASME E8Mí04) cut in longitudinal direction compared with weld direction (mm)Fig. 4 Tensile specimens after fractureThe residual error term (e r ) measures theexperimental errors. Such relationship was developed as quadratic polynomial under multiple regression form [19,20]:¦¦¦ r 20e x x b x b x b b Y j i ij i ii i i (3)where b 0 is an intercept or the average of response; b i , b ii , and b ij represent regression coefficients. For the three factors, the selected polynomial could be expressed as:2332222113210r b F b N b r b F b N b b YFr b Nr b NF b 231312 (4)In applying the RSM, the independent variable Y was viewed as surface to which a mathematical model was fitted. The adequacy of the developed model was tested using the analysis of variance (ANOV A) which quantifies the amount of variation in a process and determines if it is significant or is caused by random noise.3.2 Mathematic model of hardening parametersTable 5 lists the coefficients of the best linear regression models. All selected parameters (N , F , r ) for k and İ0 are statistically significant (P-value less than 0.05) at the 95% confidence level. However, for the response n , the term b 3r having a P-value=0.0654>0.05 is not statistically significant at the 95% confidence level even though the term b 13Nr is statistically significant. Consequently, b 3(r ) is kept in the model to improve the Lack-of-Fit test (Table 6). Furthermore, only theHassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í10721068Table 5 Coefficients of regression models for hardening parametersStrength coefficient (k) Hardeningcoefficient(n) Initial yield strain (İ0) CoefficientEst. SEP-value Est SE P-valueEst/10í4 SE/10í4 P-value b0 610.39,48<10í4 0.422 0.0073 <10-4 22.8 1.010 <10-4 b1 í83.58.48<10í4 í0.020 0.0065 0.0091 2.30 0.912 0.0267 b2 19.68.480.0410.0290.00650.00084.900.9120.0002b3 í84.58.48<10í4 í0.013 0.0065 0.0654 í4.80 0.912 0.0002 b11 5.561.3670.0009b22 í61.812.720.0005í0.0310.00980.009b33b12b13 í112.99.48<10í4 í0.074 0.0073 <10-4 -8,75 1.010 <10-4 b23R2 95.90% 92.38% 92.84%2adjR 94.19% 89.21% 89.86% SE of est. 30.7 0.021 2.9×10í4Est: Estimate; SE: Standard Error; SE of est.: Standard error of estimateTable 6 ANOV A for hardening parametersk n İ0Source of variationSS Df P-Value SS Df P-Value SS/10í7 Df P-Value Model 263946.0 5 <10í4 0.062357 5 <10í4 129.324 5 <10í4Residual 11296.4 12 0.005140912 9.97 12Lack-of-Fit 10130.4 9 0.2065 0.00428669 0.3678 8.295 9 0.3723 Pure error 1166.07 3 0.0008543 3 1.675 3 Total correction 275243.017 0.06749817 139.294 17 DW-value 1.31 1.42 2.26DW: Durbin-Watson statistic; SS: Sum of squares; D f: Degree of freedominteraction (Nír) is statistically significant on the three responses (Fig. 5). According to the adjusted R2 statistic, the selected models explain 94.19%, 89.21% and 89.86% of the variability in k, n and İ0 respectively.The ANOV A (Table 6) for the hardening parameter shows that all models (k, n, İ0) represent statistically significant relationships between the variables in each model at the 99% confidence level (P-value<10í4). The Lack-of-Fit test confirms that these models (k, n, İ0) are adequate to describe the observed data (P-value>0.05) at the 95% confidence level. The DW statistic test indicates that there is probably not any serious autocorrelation in their residuals (DW-value>1.4). The normal probability plots of the residuals suggest that the error terms, for these models, are indeed normally distributed (Fig. 6). The response surface models in terms of coded variables (Eqs. (5)í(7)) are shown in Fig. 7.k=610.3–83.5 N+19.6 F–84.5 r –61.8 F2–112.9 Nr(5) n=0.422–0.020 N+0.029 F–0.013 r–0.031 F2–0.074 Nr(6) İ0=22.8+2.3 N+4.90 F–4.80 r+5.56 N2–8.75 Nr(7) Fig. 5 Interaction plots of Nír (rotational speedídiameter ratio): (a) Strength coefficient k; (b) Hardening coefficient n;(c) Initial yield strain İ0Hassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í1072 1069Fig. 6 Normal probability plots for residual: (a) Strength coefficient k; (b) Hardening coefficient n; (c) Initial yield strain İ04 Validation of identified modelsValidation tests of the identified models were performed through comparative study between the experimental models (EM) of tensile tests and the computed responses given by numerical simulations of the same tests (Fig. 8). The computed responses, expressed in the form of tension and elongation, wereFig. 7 Response surfaces plots: (a) Strength coefficient k;(b) Hardening coefficient n; (c) Initial yield strain İ0 established by examining welded joints having an elastoplastic behavior in accordance with the Swift hardening law (Eq. (1)). These computed responses were deduced from the numerical simulations using the finite element code Abaqus/Implicit, in which the introduced elastoplastic behavior was obtained from the least square hardening models (LSHM) (Table 4) and the response surface hardening models (RSHM) (Table 5). The highest deviations (<10%), between EM and computed response, were recorded with the RSHM. Increasing deviations, as shown in Fig. 8, is due to the effect of combining damage with plastic strains accumulatedHassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í10721070Fig. 8 Relationship between tension and elongation: Confrontation between experimental model (EM), and computed responses (LSHM, RSHM) for three experimental testsduring the onset of localized necking.The relative average deviation criterion (EM/LSHM ]) between the experimental data and the numerical predictions of tensions, was used to assess the quality of the identified models.¦¸¸¹·¨¨©§'' '2exp num exp exp/num )()()(1i i i L F L F L F N] (8)where N is the number of experimental measurements,F exp (ǻL i ) and F num (ǻL i ) are respectively the experimental and predicated tensions relating to the i-th elongation ǻL i . Figure 9 illustrates that the relative average deviation of EM/LSHM (EM/LSHM ]) ranges between 1.64% and 6.75% while the relative average deviation of EM/RSHM (EM/RSHM ]) ranges between 4.52% and 9.32%.Fig. 9 Distribution of relative average deviations for most representative experimental testsFor the deviation within limits fluctuating between 4.52% and 6.75% the estimated models (LSHM and RSHM) are comparable. This applies particularly to welded joints characterized by a strength coefficient (k ), ranging from 520 to 610 MPa and a hardening exponent (n ) ranging between 0.30 and 0.45.5 DiscussionIn this study we evaluated, using RSM, the effect of FSW parameters such as tool rotational speed, welding feed rate and diameter ratio of pin to shoulder on the plastic deformation aptitudes of welded joints. The performed analysis highlights the incontestable significant effects of rotational speed (N ), welding feed rate (F ) and the interaction (Nír ) between rotational speed and diameters ratio on hardening parameters (k , n , İ0) according to Swift law. The established models show that tool diameter ratio has a linear effect only on (k ) and (İ0), it does not have any quadratic effect. They also show that rotational speed has a quadratic effect solely on (İ0); while welding feed rate has a quadratic effect on both (k ) and (n ).In addition, numerical simulation of tensile tests of welded joints has been made possible through the predictive models (LSHM and RSHM) of Swift’s hardening parameters. To judge whether the models represent correctly the data, a comparative study between the experimental response and the computed response, expressed in terms of tension-elongation, was carried out. It was found that the relative average deviation betweenHassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í1072 1071experimental model and numerical models is less than 9.5% in all cases.Moreover, correlation between welding and hardening parameters provided has many benefits. The correlation relationships can solve inverse problem relating to optimal choice of parameters linked up with the desired welded joints properties to produce welds having tailor-made mechanical properties. The correlation predictions offer the possibility to identify the behavior of friction stir welded joints necessary for finite element simulations of various forming processes while minimizing experimental cost and time. Ultimately, understanding correlations can be useful for studies on reliability of welded assemblies in service life expectancy.6 Conclusions1) Rotational speed and welding feed rate are the factors that have greater influence on hardening parameters (k, n, İ0), followed by diameter ratio that has no influence on the hardening coefficient (n).2) The numerical models RSHM were compared with those through LSHM and confronted to the experimental results. Indeed, within the limit of a relative average deviation of about 9.3%, between the experimental model and numerical models expressed in terms of tension-elongation, the validity of these models is acceptable.3) The predictive models of work-hardening coefficients, established taking into account the FSW parameters, have made possible the numerical simulation of tensile tests of FSW joints. These results can be used to perform multi-criteria optimization for producing welds with specific mechanical properties or conducting numerical simulation of plastic deformation of forming process of friction stir welded parts such as hydroforming, bending and forging.References[1]THOMAS W M, NICHOLAS E D, NEEDHAM J C, MURCH M G,TEMPLE-SMITH P, DAWES C J. Friction stir butt welding,PCT/GB92/ 02203 [P]. 1991.[2]XUE P, NI D R, WANG D, XIAO B L, MA Z Y. Effect of friction stirwelding parameters on the microstructure and mechanical propertiesof the dissimilar AlíCu joints[J]. Materials Science and EngineeringA, 2011, 528: 4683í4689.[3]LIU H J, ZHANG H J, YU L. Effect of welding speed onmicrostructures and mechanical properties of underwater friction stirwelded 2219 aluminum alloy [J]. Materials and Design, 2011, 32:1548í1553.[4]MISHRA R S, MA Z Y. Friction stir welding and processing [J].Materials Science and Engineering R, 2005, 50: 1í78.[5]ARB EGAST W J. A flow-partitioned deformation zone model fordefect formation during friction stir welding [J]. Scripta Materialia,2008, 58: 372í376.[6]LEWIS N P. Metal cutting theory and friction stir welding tooldesign [M]. NASA Faculty Fellowship Program Marshall SpaceFlight Center, University of ALABAMA, NASA/MSFC Directorate:Engineering (ED-33), 2002.[7]ARB EGAST W J. Modeling friction stir welding joining as ametalworking process, hot deformation of aluminum alloys III [C].San Diego: TMS Annual Meeting, 2003: 313í327.[8]ARTHUR C N Jr. Metal flow in friction stir welding [R]. NASAmarshall space flight center, EM30. Huntsville, AL 35812.[9]FONDA R W, B INGERT J F, COLLIGAN K J. Development ofgrain structure during friction stir welding [J]. Scripta Materialia,2004, 51: 243í248.[10]NANDAN R, DEBROY T, BHADESHIA H K D H. Recent advancesin friction stir weldingüProcess, weldment structure and properties[J]. Progress in Materials Science, 2008, 53: 980í1023.[11]LOCKWOOD W D, TOMAZ B, REYNOLDS A P. Mechanicalresponse of friction stir welded AA2024: Experiment and modeling[J]. Materials Science and Engineering A, 2002, 323: 348í353. [12]SALEM H G, REYNOLDS A P, LYONS J S. Microstructure andretention of superplasticity of friction stir welded superplastic 2095sheet [J]. Scripta Materialia, 2002, 46: 337í342.[13]LOCKWOOD W D, REYNOLDS A P. Simulation of the globalresponse of a friction stir weld using local constitutive behavior [J].Materials Science and Engineering A, 2003, 339: 35í42.[14]SUTTON M A, YANG B, REYNOLDS A P, YAN J. B andedmicrostructure in 2024–T351 and 2524-T351 aluminum friction stirwelds, Part II. Mechanical characterization [J]. Materials Science andEngineering A, 2004, 364: 66í74.[15]ZHANG H W, ZHANG Z, CHEN J T. The finite element simulationof the friction stir welding process [J]. Materials Science andEngineering A, 2005, 403: 340í348.[16]ZHANG Z, ZHANG H W. Numerical studies on controlling ofprocess parameters in friction stir welding [J]. Journal of MaterialsProcessing Technology, 2009, 209: 241í270.[17]B UFFA G, FRATINI L, SHIVPURI R. Finite element studies onfriction stir welding processes of tailored blanks [J]. Computers andStructures, 2008, 86: 181í189.[18]SWIFT H W. Plastic instability under plane stress [J]. Journal of theMechanics and Physics of Solids, 1952, 1: 1í18.[19]MONTGOMERY D C. Design and analysis of experiments [M].Fifth Edition. New York: John Wiley & Sons, 2001: 684.[20]MYERS R H, MONTGOMERY D C, ANDERSON-COOK C M.Response surface methodology: Process and product optimizationusing designed experiment [M]. 3rd Edition. New York: John Wiley& Sons, 2009: 680.[21]VIJAY S J, MURUGAN N. Influence of tool pin profile on themetallurgical and mechanical properties of friction stir welded Al–10% TiB2 metal matrix composite [J]. Materials and Design,2010, 31: 3585í3589.[22]ELANGOV AN K, BALASUBRAMANIAN V. Influences of tool pinprofile and tool shoulder diameter on the formation of friction stirprocessing zone in AA6061 aluminum alloy [J]. Materials andDesign, 2008, 29: 362í373.[23]PALANIVEL R, KOSHY MATHEWS P, MURUGAN N.Development of mathematical model to predict the mechanicalproperties of friction stir welded AA6351 aluminum alloy [J]. Journalof Engineering Science and Technology Review, 2011, 4(1): 25í31.Hassen BOUZAIENE, et al/Trans. Nonferrous Met. Soc. China 22(2012) 1064í107210722017䪱 䞥 ⛞ ⛞⹀ ⱘ ㋏Hassen BOUZAIENE, Mohamed-Ali REZGUI, Mahfoudh AYADI, Ali ZGHALBesearch Unit in Solid Mechanics, Structures and Technological Development (99-UR11-46),Higher School of Sciences and Techniques of Tunis, Tunisia㽕˖ 2017䪱 䞥䖯㸠 ⛞ ˈ㸼䗄Swift⹀ 㾘 DŽ䞛⫼䴶 䆒䅵 ⊩䖯㸠⛞ ⱘ Ԍ 偠䆒䅵DŽ䞛⫼ Ѣ ѠЬ⊩ 䴶⊩ⱘ2⾡῵ 䆘Ԅ ⛞ ⛞ ㋴ ⹀ ⱘ DŽ䞛⫼ 䰤 ⿟ Abaqus ῵ ⛞ Ԍ⌟䆩㒧 DŽⳌ 㒧 㸼 ˈ 偠㒧 ῵ 㒧 䕗 DŽ䖭ѯ㒧 㛑⫼Ѣ 偠 Ⳃ Ӭ ˈ 㸠 ԧ⛞ ⛞ 䳊ӊ 䖛⿟Ё ⱘ ῵ ˈ ⎆ ǃ 䬏䗴DŽ䬂䆡˖ ⛞ ˗ 䴶 ⊩˗䴶 Ё 䆒䅵˗⹀ ˗῵ ˗Ⳍ(Edited b y LI Xiang-qun)。

Designation:A262−15Standard Practices forDetecting Susceptibility to Intergranular Attack in Austenitic Stainless Steels1This standard is issued under thefixed designation A262;the number immediately following the designation indicates the year of original adoption or,in the case of revision,the year of last revision.A number in parentheses indicates the year of last reapproval.A superscript epsilon(´)indicates an editorial change since the last revision or reapproval.This standard has been approved for use by agencies of the U.S.Department of Defense.1.Scope*1.1These practices cover the followingfive tests:1.1.1Practice A—Oxalic Acid Etch Test for Classification of Etch Structures of Austenitic Stainless Steels(Sections4to 13,inclusive),1.1.2Practice B—Ferric Sulfate-Sulfuric Acid Test for De-tecting Susceptibility to Intergranular Attack in Austenitic Stainless Steels(Sections14to25,inclusive),1.1.3Practice C—Nitric Acid Test for Detecting Suscepti-bility to Intergranular Attack in Austenitic Stainless Steels (Sections26to36,inclusive),1.1.4Practice E—Copper–Copper Sulfate–Sulfuric Acid Test for Detecting Susceptibility to Intergranular Attack in Austenitic Stainless Steels(Sections37to46,inclusive),and 1.1.5Practice F—Copper–Copper Sulfate–50%Sulfuric Acid Test for Detecting Susceptibility to Intergranular Attack in Molybdenum-Bearing Austenitic Stainless Steels(Sections 47to58,inclusive).1.2The Oxalic Acid Etch Test is a rapid method of identifying,by simple etching,those specimens of certain stainless steel grades that are essentially free of susceptibility to intergranular attack associated with chromium carbide precipitates.These specimens will have low corrosion rates in certain corrosion tests and therefore can be eliminated (screened)from testing as“acceptable.”The etch test is applicable only to those grades listed in the individual hot acid tests and classifies the specimens either as“acceptable”or as “suspect.”1.3The ferric sulfate-sulfuric acid test,the copper–copper sulfate–50%sulfuric acid test,and the nitric acid test are based on weight loss determinations and,thus,provide a quantitative measure of the relative performance of specimens evaluated.In contrast,the copper–copper sulfate–16%sulfuric acid test is based on visual examination of bend specimens and,therefore, classifies the specimens only as acceptable or nonacceptable.1.4The presence or absence of intergranular attack in these tests is not necessarily a measure of the performance of the material in other corrosive environments.These tests do not provide a basis for predicting resistance to forms of corrosion other than intergranular,such as general corrosion,pitting,or stress-corrosion cracking.N OTE1—See Appendix X1for information regarding test selection.1.5The values stated in SI units are to be regarded as standard.The inch-pound equivalents are in parentheses and may be approximate.1.6This standard does not purport to address all of the safety concerns,if any,associated with its use.It is the responsibility of the user of this standard to establish appro-priate safety and health practices and determine the applica-bility of regulatory limitations prior to use.Some specific hazards statements are given in10.1,20.1.1,20.1.9,31.3,34.4, 53.1.1,and53.1.10.2.Referenced Documents2.1ASTM Standards:2A370Test Methods and Definitions for Mechanical Testing of Steel ProductsA380/A380M Practice for Cleaning,Descaling,and Passi-vation of Stainless Steel Parts,Equipment,and Systems D1193Specification for Reagent WaterE3Guide for Preparation of Metallographic Specimens 2.2ASME Code:3ASME Boiler&Pressure Vessel Code,Section IX2.3ACS Specifications:4Reagent Chemicals,Specifications and Procedures1These practices are under the jurisdiction of ASTM Committee A01on Steel, Stainless Steel and Related Alloys and are the direct responsibility of Subcommittee A01.14on Methods of Corrosion Testing.Current edition approved Sept.1,2015.Published September2015.Originally approved st previous edition approved in2014as A262–14.DOI: 10.1520/A0262-15.2For referenced ASTM standards,visit the ASTM website,,or contact ASTM Customer Service at service@.For Annual Book of ASTM Standards volume information,refer to the standard’s Document Summary page on the ASTM website.3Available from American Society of Mechanical Engineers(ASME),ASME International Headquarters,Two Park Ave.,New York,NY10016-5990,http:// .4Available from American Chemical Society(ACS),1155Sixteenth Street,NW, Washington,DC20036,*A Summary of Changes section appears at the end of this standard Copyright©ASTM International,100Barr Harbor Drive,PO Box C700,West Conshohocken,PA19428-2959.United States2.4ISO Standard:5ISO 3651-2Determination of Resistance to Intergranular Corrosion of Stainless Steels—Part 2:Ferritic,Austenitic,and Ferritic-Austenitic (Duplex)Stainless Steels—Corrosion Test in Media Containing Sulfuric Acid 3.Purity of Reagents3.1Purity of Reagents—Reagent grade chemicals shall be used in all tests.Unless otherwise indicated,it is intended that all reagents conform to the specifications of the Committee on Analytical Reagents of the American Chemical Society 6where such specifications are available.Other grades may be used,provided it is first ascertained that the reagent is of sufficiently high purity to permit its use without lessening the accuracy of the test result.3.2Purity of Water—Unless otherwise indicated,references to water shall be understood to mean reagent water as defined by Type IV of Specification D1193.PRACTICE A—OXALIC ACID ETCH TEST FOR CLASSIFICATION OF ETCH STRUCTURES OFAUSTENITIC STAINLESS STEELS (1)74.Scope4.1The Oxalic Acid Etch Test is used for acceptance of wrought or cast austenitic stainless steel material but not for rejection of e of A262Practice A as a stand-alone test may reject material that the applicable hot acid test would find acceptable;such use is outside the scope of this practice.4.2This test is intended to be used in connection with other evaluation tests described in these practices to provide a rapid method for identifying qualitatively those specimens that are certain to be free of susceptibility to rapid intergranular attack in these other tests.Such specimens have low corrosion rates in the various hot acid tests which require from 15to 240h of exposure.These specimens are identified by means of their etch structures,which are classified according to the criteria given in Section 11.4.3The Oxalic Acid Etch Test may be used to screen specimens intended for testing in Practice B—Ferric Sulfate-Sulfuric Acid Test,Practice C—Nitric Acid Test,Practice E—Copper-Copper Sulfate–16%Sulfuric Acid Test,and Prac-tice F—Copper-Copper Sulfate–50%Sulfuric Acid Test.4.4Each of these other practices contains a table showing which classifications of etch structures on a given stainless steel grade are equivalent to acceptable or suspect performance in that particular test.Specimens having acceptable etch structures need not be subjected to the hot acid test.Specimens having suspect etch structures must be tested in the specified hot acid solution.4.5There are two classes of specimens to be considered:base metal,and process-affected metal.4.5.1Process-affected metal contains any condition that affects the corrosion properties of the material in a non-uniform way,such as (but not limited to)welds;carburized.nitrided,or oxidized surfaces;mechanical deformation;and areas affected by heat.Base metal has none of these conditions.4.5.2Because Practices B,C,and F involve immersing the entire specimen and averaging the mass loss over the total specimen area,and because welding,carburization,mechanical deformation,and the like affect only part of a specimen,the presence of process-affected metal in a specimen can affect the test result in an unpredictable way depending on the propor-tions of the area affected.4.5.3If the presence of these or other localized conditions is a concern to the purchaser,then tests that do not average the mass loss over the total specimen surface area,such as Practice A,the Oxalic Acid Etch Test,or Practice E,the Copper–Copper Sulfate–Sulfuric Acid Test for Detecting Susceptibility to Intergranular Attack in Austenitic Stainless Steels,should be considered.5.Summary of Practice5.1A specimen representative of the material to be evalu-ated is polished to a specified finish and over-etched using oxalic acid electrolytically.The etched specimen is then examined using a metallurgical microscope.The etched struc-ture is compared with reference photographs to determine whether the material is acceptable or suspect.Suspect material is then subjected to the specified hot acid immersion test.6.Significance and Use6.1Use of the etch test allows rapid acceptance of specific lots of material without the need to perform time-consuming and costly hot acid immersion tests on those lots.7.Apparatus7.1Etching Cell:7.1.1An etching cell may be assembled using components as described in this section.Alternatively,a commercial electropolisher/etcher (as used for metallographic sample preparation)may be used for small specimens provided the current density requirement of 10.2is met.7.1.2Source of Direct Current—Battery,generator,or rec-tifier capable of supplying about 15V and 20A.7.1.3Ammeter—For direct current;used to measure the current on the specimen to be etched.7.1.4Variable Resistance—Used to control the current on the specimen to be etched.7.1.5Cathode—A stainless steel container,for example,a 1-L (1-qt)stainless steel beaker.7.1.5.1Alternate Cathode—A piece of flat stainless steel at least as large as the specimen surface.7.1.6Electrical Clamp—To hold the specimen to be etched and to complete the electrical circuit between the specimen and the power source such that the specimen is the anode of the cell.5Available from International Organization for Standardization (ISO),1,ch.de la V oie-Creuse,CP 56,CH-1211Geneva 20,Switzerland,.6For suggestions on the testing of reagents not listed by the American Chemical Society,see Analar Standards for Laboratory Chemicals ,BDH Ltd.,Poole,Dorset,U.K.,and the United States Pharmacopeia and National Formulary ,U.S.Pharma-copeial Convention,Inc.(USPC),Rockville,MD.7The boldface numbers in parentheses refer to a list of references at the end of thisstandard.--`,`,``,``,`,,`,,,,,,```,`,,```-`-`,,`,,`,`,,`---7.1.7The power source,resistor,and ammeter must be sized appropriately for providing and controlling the current as specified in10.2of this practice.7.1.8As described,the electrolyte container is the cathode; it may be a stainless steel beaker or fabricated from stainless steel such as by welding a section of tube or pipe to aflat plate or sheet.Alternatively,the electrolyte container may be glass (or other non-conducting,corrosion resisting material)in lieu of a stainless steel container,and the cathode may be aflat plate or sheet of a corrosion resisting alloy.In this latter case,theflat surface of the cathode must be at least as large as,facing,and approximately centered on,the prepared surface of the speci-men.Other configurations of the electrodes might not provide uniform etching over the specimen surface.In any case,the size and shape of the specimen dictate the size and construction of the etching cell and of the power source and controls.The overriding principle is that the etch needs to be uniform over the surface to be examined.7.2Metallurgical Microscope—For examination of etched microstructures at250to500diameters.8.Reagents and Materials8.1Etching Solution(10%)—Dissolve100g of reagent grade oxalic acid crystals(H2C2O4·2H2O)in900mL of reagent water.Stir until all crystals are dissolved.8.1.1Alternate Etching Solution(See10.7)—Dissolve100g of reagent grade ammonium persulfate((NH4)2S2O8)in 900mL of reagent water.Stir until dissolved.9.Sampling and Test Specimens9.1The specified hot acid test provides instructions for sampling and for specimen preparation such as a sensitization heat treatment.Additional instructions specific to Practice A follow:9.2The preferred specimen is a cross-section including the product surface to be exposed in service.Only suchfinishing of the product surface should be performed as is required to remove foreign material.9.3Whenever practical,use a cross-sectional area of1cm2 or more.If any cross-sectional dimension is less than1cm, then the other dimension of the cross-section should be a minimum of1cm.When both dimensions of the product are less than1cm,use a full cross section.9.4Polishing—On all types of materials,polish cross sec-tional surfaces through CAMI/ANSI600[FEPA/ISO P1200]in accordance with Guide E3prior to etching and examination. Not all scratches need to be removed.10.Procedure10.1(Warning—Etching should be carried out under a ventilated hood.Gas,which is rapidly evolved at the electrodes with some entrainment of oxalic acid,is poisonous and irritating to mucous membranes.)10.2Etch the polished specimen at1A/cm2for1.5min.10.2.1To obtain the correct specified current density: 10.2.1.1Measure the total immersed area of the specimen to be etched in square centimetres.10.2.1.2Adjust the variable resistance until the ammeter reading in amperes is equal to the total immersed area of the specimen in square centimetres.10.3A yellow-greenfilm is gradually formed on the cath-ode.This increases the resistance of the etching cell.When this occurs,remove thefilm by rinsing the inside of the stainless steel beaker(or the steel used as the cathode)with an acid such as30%HNO3.10.4The temperature of the etching solution gradually increases during etching.Keep the temperature below50°C. This may be done by alternating two containers.One may be cooled in tap water while the other is used for etching.10.4.1The rate of heating depends on the total current (ammeter reading)passing through the cell.Therefore,keep the area to be etched as small as possible while at the same time meeting the requirements of desirable minimum area to be etched.10.5Avoid immersing the clamp holding the specimen in the etching solution.10.6Rinsing—Following etching,rinse the specimen thor-oughly in hot water and then in acetone or alcohol to avoid crystallization of oxalic acid on the etched surface during drying.10.7It may be difficult to reveal the presence of step structures on some specimens containing molybdenum(AISI 316,316L,317,317L),which are free of chromium carbide sensitization,by electrolytic etching with oxalic acid.In such cases,an alternate electrolyte of ammonium persulfate may be used in place of oxalic acid.(See8.1.1.)An etch for5or10 min at1A/cm2in a solution at room temperature readily develops step structures on such specimens.11.Classification of Etch Structures11.1Examine the etched surface on a metallurgical micro-scope at250×to500×for wrought steels and at about250×for cast steels.11.2Examine the etched cross-sectional areas thoroughly by complete traverse from inside to outside diameters of rods and tubes,from face to face on plates.11.2.1Microscopical examination of a specimen shall be made on metal unaffected by cold-working,carburization, welding,and the like.If any of these conditions are found,note their presence in the report.11.3Classify the etch structures into the following types (Note2):11.3.1Step Structure(Fig.1)—Steps only between grains, no ditches at grain boundaries.11.3.2Dual Structure(Fig.2)—Some ditches at grain boundaries in addition to steps,but no single grain completely surrounded by ditches.11.3.3Ditch Structure(Fig.3)—One or more grains com-pletely surrounded by ditches.11.3.4Isolated Ferrite(Fig.4)—Observed in castings and welds.Steps between austenite matrix and ferrite pools. 11.3.5Interdendritic Ditches(Fig.5)—Observed in castings and welds.Deep interconnectedditches.--` , ` , ` ` , ` ` , ` , , ` , , , , , , ` ` ` , ` , , ` ` ` -` -` , , ` , , ` , ` , , ` ---11.3.6End-Grain Pitting I (Fig.6)—Structure contains a few deep end-grain pits along with some shallow etch pits at 500×.(Of importance only when the nitric acid test is used.)11.3.7End-Grain Pitting II (Fig.7)—Structure contains numerous,deep end-grain pits at 500×.(Of importance only when nitric acid test is used.)N OTE 2—All photomicrographs were made with specimens that were etched under standard conditions:10%oxalic acid,room temperature,1.5min at 1A/cm 2.11.4The evaluation of etch structures containing only steps and of those showing grains completely surrounded by ditches in every field can be carried out relatively rapidly.In cases that appear to be dual structures,more extensive examination is required to determine if there are any grains completely encircled.If an encircled grain is found,classify the steel as a ditch structure.11.4.1On stainless steel castings (also on weld metal),the steps between grains formed by electrolytic oxalic acid etching tend to be less prominent than those on wrought materialsorFIG.1Step Structure (500×)(Steps Between Grains,No Ditchesat GrainBoundaries)FIG.2Dual Structure (250×)(Some Ditches at Grain Boundaries in Addition to Steps,but No One Grain CompletelySurrounded)FIG.3Ditch Structure (500×)(One or More Grains CompletelySurrounded byDitches)FIG.4Isolated Ferrite Pools (250×)(Observed in Castings and Welds.Steps Between Austenite Matrix and FerritePools)--`,`,``,``,`,,`,,,,,,```,`,,```-`-`,,`,,`,`,,`---are entirely absent.However,any susceptibility to intergranular attack is readily detected by pronounced ditches.11.4.2Some wrought specimens,especially from bar stock,may contain a random pattern of pits.If these pits are sharp and so deep that they appear black (Fig.7)it is possible that the specimen may be susceptible to end grain attack in nitric acid only.Therefore,even though the grain boundaries all have stepstructures,specimens having as much or more end grain pitting than that shown in Fig.7cannot be safely assumed to have low nitric acid rates and should be subjected to the nitric acid test whenever it is specified.Such sharp,deep pits should not be confused with the shallow pits shown in Figs.1and e of Etch Structure Classifications12.1The use of these classifications depends on the hot acid corrosion test for which stainless steel specimens are being screened by etching in oxalic acid and is described in each of the practices.13.Precision and Bias13.1Precision and Bias—No information is presented about either the precision or bias of Practice A—Oxalic Acid Etch Test for classification of Etch Structures of Austenitic Stainless Steels since the test result is nonquantitative.PRACTICE B—FERRIC SULFATE–SULFURIC ACID TEST FOR DETECTING SUSCEPTIBILITYTO INTERGRANULAR ATTACK IN AUSTENITIC STAINLESS STEELS (2)14.Scope14.1This practice describes the procedure for conducting the boiling 120-h ferric sulfate–50%sulfuric acid test which measures the susceptibility of austenitic stainless steels to intergranular attack.14.2The presence or absence of intergranular attack in this test is not necessarily a measure of the performance of the material in other corrosive environments.The test does not provide a basis for predicting resistance to forms ofcorrosionFIG.5Interdendritic Ditches (250×)(Observed in Castings andWelds.Deep InterconnectedDitches)N OTE 1—To differentiate between the types of pits,use a magnification of 500×and focus in the plane of etched surface.The pits which now appear completely black are end grain pits.FIG.6End Grain Pitting I (500×)(A Few Deep End Grain Pits(See 1in Figure)and Shallow Etch Pits(2))N OTE 1—This or a greater concentration of end grain pits at 500×(using standard etching conditions)indicates that the specimen must be tested when screening is for nitric acid test.FIG.7End Grain Pitting II(500×)--`,`,``,``,`,,`,,,,,,```,`,,```-`-`,,`,,`,`,,`---other than intergranular,such as general corrosion,pitting,or stress-corrosion cracking.15.Summary of the Ferric Sulfate-Sulfuric Acid PracticeB15.1A specimen representative of the material to be evalu-ated is immersed in a boiling solution of ferric sulfate and sulfuric acid for a specified time.The resulting mass loss is converted to a corrosion rate,which is compared to a specified maximum value to determine whether the material has the resistance to attack expected of the grade of material being tested.16.Significance and Use16.1The ferric sulfate-sulfuric acid test detects susceptibil-ity to intergranular attack associated primarily with chromium carbide precipitate in unstabilized austenitic stainless steels, and to intergranular attack associated with sigma phase. 16.2The corrosion potential of the ferric sulfate-sulfuric acid test has been reported as0.6V versus a standard calomel electrode(SCE),as compared with0.75to1.0V for Practice C, and0.1V for Practices E and F.(3)N OTE3—A higher corrosion potential indicates more severely oxidizing conditions.17.Rapid Screening Test17.1Before testing in the ferric sulfate-sulfuric acid test, specimens of certain grades of stainless steels(see Table1) may be given a rapid screening test in accordance with procedures given in Practice A,Oxalic Acid Etch Test for Classification of Etch Structures of Austenitic Stainless Steels. Preparation,etching,and the classification of etch structures are described therein.The use of etch structure evaluations in connection with the ferric sulfate-sulfuric acid test is specified in Table1.17.2Heat treat the material in accordance with22.1prior to performing the etch test.17.3Ignore“process-affected”areas(see Section21);ap-plication of the ferric sulfate-sulfuric acid test to process-affected areas is currently outside the scope of Practice B. 17.4Corrosion test specimens having acceptable etch struc-tures in the Oxalic Acid Etch Test will be essentially free of intergranular attack in the ferric sulfate-sulfuric acid test.Such specimens are acceptable without testing in the ferric sulfate-sulfuric acid test.All specimens having suspect etch structures shall be tested in the ferric sulfate-sulfuric acid test.18.Apparatus18.1The apparatus is illustrated in Fig.8.N OTE4—Other ground glass joints,such as the45/40joint may also be used.18.1.1An Allihn condenser with a minimum of four bulbs and with a ground glass joint to match that of theflask. 18.1.1.1Substitutions for this condenser orflask are not allowed.Specifically,the cold-finger type of condenser with standard Erlenmeyerflasks shall not be used.Corrosion rates obtained using the cold-finger type of condenser are lower than those obtained using the Allihn type of condenser whether due to loss of vapor or to higher oxygen content in the solution or both.Such lower corrosion rates lead to acceptance of material that should be rejected.18.1.2A1-L Erlenmeyerflask with a ground glass joint to match that of the condenser.Theflask opening limits the size of the specimen;a larger opening is desirable.TABLE1Use of Etch Structure Classifications from the Oxalic Acid Etch Test with Ferric Sulfate-Sulfuric Acid Test AGrade Acceptable EtchStructuresSuspect Etch Structures B304Step,dual,end grain,I&II Ditch304L Step,dual,end grain,I&II Ditch316Step,dual,end grain,I&II Ditch316L Step,dual,end grain,I&II Ditch317Step,dual,end grain,I&II Ditch317L Step,dual,end grain,I&II DitchCF-3Step,dual,isolated ferrite pools Ditch,interdendritic ditches CF-8Step,dual,isolated ferrite pools Ditch,interdendritic ditches CF-3M Step,dual,isolated ferrite pools Ditch,interdendritic ditches CF-8M Step,dual,isolated ferrite pools Ditch,interdendritic ditches A Grades not listed in this table either have not been evaluated for use of Practice A with Practice B or have been found to give acceptable results in the etch test while giving unacceptable results in Practice B.In the latter case Practices A would pass material that should have been subjected to the ferric sulfate-sulfuric acid test.B Specimens havingthese structures shall be tested in the ferric sulfate-sulfuricacid test.FIG.8Apparatus for Ferric Sulfate-Sulfuric Acid Test--`,`,``,``,`,,`,,,,,,```,`,,```-`-`,,`,,`,`,,`---18.1.3Glass cradle(Note5)—Can be supplied by a glass-blowing shop.It must be sized so as tofit,with the specimen, through theflask opening.It must be designed to allow free flow of the testing solution around the specimen.N OTE5—Other equivalent means of specimen support,such as glass hooks or stirrups,may also be used.18.1.4Boiling Chips—Used to prevent bumping.18.1.5High Vacuum Silicone Grease—For the ground glass joint.18.1.6Hot plate,capable of providing heat for continuous boiling of the solution.18.1.7An analytical balance capable of weighing to the nearest0.001g.N OTE6—During testing,there is some deposition of iron oxides on the upper part of the Erlenmeyerflask.This can be readily removed,after test completion,by boiling a solution of10%hydrochloric acid in theflask.18.1.8Desiccator—For storage of prepared specimens prior to testing.19.Reagents and Materials19.1Ferric Sulfate Hydrate(Fe2(SO4)3·xH2O),about75% (Fe2(SO4)3)by mass.19.1.1Ferric sulfate is a specific additive that establishes and controls the corrosion potential.Substitutions are not permitted.19.2Sulfuric Acid(H2(SO)4),95.0to98.0%by mass.20.Ferric Sulfate-Sulfuric Acid Test Solution20.1Prepare600mL of50%(49.4to50.9%)solution as follows:20.1.1(Warning—Protect the eyes and use rubber gloves for handling acid.Place the testflask under a hood.)20.1.2First,measure400.0mL of Type IV reagent water and pour into the Erlenmeyerflask.20.1.3Then measure236.0mL of reagent-grade sulfuric acid.Add the acid slowly and with constant stirring to the water in the Erlenmeyerflask to avoid boiling by the heat evolved.N OTE7—Loss of vapor results in concentration of the acid.20.1.4Weigh25g of reagent-grade ferric sulfate to the nearest0.1g and add to the sulfuric acid solution.20.1.5Drop boiling chips into theflask.20.1.6Lubricate ground glass joint with silicone grease.20.1.7Coverflask with condenser and circulate cooling water.20.1.8Boil the solution until all ferric sulfate is dissolved (see Note7).20.1.9(Warning—It has been reported that violent boiling resulting in acid spills can occur.It is important to ensure that the concentration of acid does not increase and that an adequate number of boiling chips(which are resistant to attack by the test solution)are present.)21.Sampling21.1Obtain and prepare only base metal samples.21.1.1There are two classes of specimens to be considered: base metal,and process-affected metal.Process-affected metal contains any condition that affects the corrosion properties of the material in a non-uniform way,such as(but not limited to) welds;carburized.nitrided,or oxidized surfaces;mechanical deformation;and areas affected by heat.Base metal has none of these conditions.21.1.2The Practice B test involves immersing the entire specimen and averaging the mass loss over the entire surface of the specimen.Welding,carburization,mechanical deformation, and the like,affect only part of a specimen.21.1.3The mass loss rate from process-affected metal is expected to differ from that from base metal;the presence of process-affected metal in a specimen will affect the calculated test result in an unpredictable way.21.1.4If the presence of these or other localized conditions is a concern to the purchaser,then tests that do not average the mass loss over the total specimen surface area,such as Practice A,the Oxalic Acid Etch Test,or Practice E,the Copper–Copper Sulfate–16%Sulfuric Acid Test for Detecting Susceptibility to Intergranular Attack in Austenitic Stainless Steels,should be considered.Details of the test and acceptance criteria shall be as agreed by the purchaser and producer.21.2Unless otherwise specified by the purchaser,the pro-cedures for obtaining representative base metal samples,for removing the specimens from the samples,and the number of specimens shall be at the discretion of the producer.22.Preparation of Test Specimens22.1Heat treat extra-low carbon and stabilized grades at 650to675°C(1200to1250°F),which is the range of maximum carbide precipitation,prior to testing.The length of time of heating,and the method of subsequent cooling used for this sensitizing treatment together with the corresponding maximum permissible corrosion rate shall be as agreed be-tween the material producer and purchaser.N OTE8—The most commonly used sensitizing treatment is1h at 675°C(1250°F).22.2Prepare the specimens,each having a total surface area of5to20cm2.22.3Where feasible for the product form,grind all the specimen surfaces using CAMI/ANSI120[FEPA/ISO P120] paper-backed,wet or dry,closed coated abrasive paper,with water as a coolant.If abrasive paper is used dry,polish slowly to avoid overheating.Do not use abrasives with grinding aids; some grinding aids containfluorides that can affect the measured corrosion rate.22.4Remove all traces of oxide scale and heat tint formed during heat treatments.Any scale that cannot be removed by grinding(for example,in stamped numbers)may be removed by using one of the pickling solutions described in Practice A380/A380M,Table A1.1.(Residual oxide scale causes gal-vanic action and consequent activation in the test solution.)22.5Measure the specimens,including the inner surfaces of any holes,to the nearest0.05mm(0.001in.)and calculate the total exposed area.22.6Degrease the specimens using suitable nonchlorinated agents,such as soap and lukewarm water,or acetone.Drythe --` , ` , ` ` , ` ` , ` , , ` , , , , , , ` ` ` , ` , , ` ` ` -` -` , , ` , , ` , ` , , ` ---。