ABSTRACT

Peening techniques like laser peening and shot peening were used to modify the surface of friction stir welded 7075-T7351 Aluminum Alloy specimens. The tensile coupons were machined such as the loading was applied in a direction perpendicular to the weld direction. The peening effects on the global and local mechanical properties through the different regions of the weld were characterized and assessed. The surface hardness levels resulting from various peening techniques were also investigated for both sides of the welds. Shot peening resulted in an increase to surface hardness levels, but no improvement was noticed on the mechanical properties. In contrast, mechanical properties were improved by laser peening when compared to the unpeened material.

The results on laser shot peening and its characterization for Inconel 600 are presented.Using an X-ray diffraction technique we show that the residual compressive stresses can besuccessfully induced in Inconel 600 with parameters of laser peening, which are close to that of 316L stainless steel. The on-line monitoring system involving the acoustic pulse measurements is described for quality control of laser shot peening process. We found that sample scanning during laser processing results in a system of column-like microstructure, which is tilted in direction of scanning. The features of optical properties of tilted microstructure are described.We revealed that the base material injected into the confining water due to laser ablation is transformed to spherical nanoparticles with diameter of 60 nm.

INTRODUCTION

Laser shot processing (peening) (LSP) is a surface treatment technology, which consists of irradiating a metallic target with a short and intense laser pulse in order to generate, through high-pressure surface plasma, a plastic deformation and a surface strengthening. Particular benefit is achieved for improving the fatigue behavior [1], and the stress corrosion cracking of various materials like austenitic stainless steel in power plants [2,3]. By now the theoretical aspects of LSP are well elaborated and are widely presented in many publications, which describe physical processes of laser-driven shock wave generation [4,5], models of pressure generation [6-8], and mechanics of a laser shock interaction with matter [9]. Meanwhile the experimental conditions of LSP allow one to create considerable changes of the surface morphology of the treated metal in a form of surface structures. Actually a large variety of the surface relief structures were observed with the laser energy higher than the melting threshold for semiconductors and metals. According to their shape these structures could be roughly divided into three groups. The first group of structures (ripples) is due to the development of the capillary wave instability due to nonuniform interference field. The presence of a strong correlation between the parameters of the structures and the characteristics of the laser irradiation allow one to speak about “laser-induced capillary waves” [10]. The second group (cellular structures) arises is resulted from instability of capillary waves due to thermal-capillary effects in laser-melted film [11]. The structures mentioned above were obtained at single-shot irradiation conditions. The third group is represented by conical and column microstructures, which are developed during multi-shot irradiation (up to 104 laser shots) [12,13]. In this paper the results are presented on the LSP technology application to induce the residual compressive stresses in Inconel 600. To provide reliability of our measurements the results obtained for Inconel 600 are compared to the results obtained under the same experimental conditions for the reference sample (stainless steel 316L), whose mechanical properties are very similar. We also report observation of column-like microstructures tilted in the direction of laser scanning. The light reflected from these conical microstructures is predominantly scattered at angles different from the incidence angle.

Since its invention by the Welding Institute in 1991 [1], friction stir welding (FSW) has emerged as a promising solid state process with encouraging results. FSW is considered a better technique than fusion welding for many aluminum alloys, and it surpasses other fusion welding processes in terms of the lack of solidification cracks, andporosity. This is particularly important when used on high strength aerospace aluminum alloys that are generally difficult to weld. The modified microstructure resulting from FSW is asymmetric about the weld centerline [2]. This is due to the advancing and retreating sides of the weld corresponding to maximum and minimum relative velocities between the tool and work-piece [3]. The FSW consists of a nugget, or the stirred zone, the thermo-mechanical affected zone (TMAZ), and a heat affected zone (HAZ). The use of FSW is expanding rapidly and is resulting in welded joints being used in critical load bearing structures. Therefore, it is important to investigate methods to improve the weakened mechanical properties produced from the welding process for components welded using FSW. Peening techniques like laser and shot peening has been reported to enhance mechanical properties in fusion welds [4, 5], however none of the investigations in literature assessed laser peening effects on the various regions of the FSW.In this study, the laser peening, shot peening, and a combination of both was used to introduce compressive residual stresses into FSW AA 7075-T7351. The influence of the different peening techniques on mechanical properties and hardness levels on both sides of the FSW specimens were characterized and assessed.

Figure 1. Severe erosive longitudinal cavitation

Figure 2. Schematic diagram of the development and collapse of longitudinal cavitation.

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EXPERIMENTAL PROCEDURE AND DETAILS

The 7075 aluminum alloy (AA) was used in this investigation. This high strength precipitation-hardened aluminum alloy is used extensively in aerospace applications. The base metal was supplied in a T651 temper which is solution heat treated and artificially aged, then stress relieved by stretching. The 7075-T651 was supplied as a 6.35mm thick plate with an ultimate and yield strength of 601 MPa, and 534 MPa respectively and an elongation of 11%. The FSW specimens for this investigation were made at the NASA- Johnson Space Center using a 5-axis milling machine, and were welded in a butt-weld configuration along rolling direction. The tool rotation was set at 350 rpm in the counterclockwise direction, and the traverse speed was set at a rate of 2.54 cm/min stirring the interface and producing a solid-state weld. The FSW panels were122 cm x 40 cm x 0.65 cm. The welded plates were heat treated to prevent the material from continuing to age at room temperature [6, 7]. Following the welding process, the welded plates were aged from the T651 condition to the T7351 condition in accordance with the SAE AMS-H-6088 requirements.The FSW plates were inspected using radiographic and penetrant inspections after the heat treatment process, with no defects being detected. Bending tests using strips specimens with dimensions of 17.8 cm x 2.54 cm were also performed. Both the root and the crown sides of the weld were tested to evaluate the quality of the weld. The samples were inspected visually afterward with no crack indications revealed.Laser peening (LP), as shown in Figure 1, is a technology that introduces a state of residual compressive stresses with the ability to develop deep, high compressive stresses in the areas treated. The samples of 316L steel and Inconel 600 of the rectangular cross-section of 25 by 25 mm with thickness of 5 mm were used as target materials. The sample surface was mechanically polished with no particular treatment. The samples were attached to a holder for rigid support and easy handling.

The holder was driven along (X, Y)-directions in the water jacket during laser irradiation using 3-D motorized actuator and multi-axis servo amplifier system. The treated area (15 by 15 mm) was created by sample displacement in X-Y plane. The displacement in Zdirection was used to determine the influence of defocusing of the laser beam (or the surfaceroughness) on the effectiveness of the LSP process. The sample was immersed in a water jacket with 3-10 mm water thickness between the sample surface and the input window

As a laser source the Nd:YAG Q-switched oscillator and amplifier (DCR family, SpectraPhysics) were used to produce laser pulse of 12 ns pulse duration with the energy of 0.8 J at the repetition rate 4 Hz. The focusing lens (f = 100 mm) was placed in front of the input window of the water jacket. The LSP of the sample was carried out during the sample scanning in X-Y plane (ladder trajectory) in such a way as to provide the laser spot two-dimensional overlapping more than 80% of the spot diameter. This means that laser pulses affected each point of the treated area approximately four times. The diameter of the single shot area was about 1 mm, which corresponds to the energy density of 102 J/cm2.

The acoustic sensor was mounted on a back surface of the sample. As the acoustic sensor,we used a narrowband sensor with a frequency resonance of 10 KHz (electro-acoustic DPA type 4006). The acoustic signal was rectified and was integrated by the diode-RC circuit and then was collected by the data acquisition system. The value measured was linearly proportional to the energy of the acoustic pulse.

The LP process uses high energy laser pulses (several GW/cm2) fired at the surface of a metal coated with an ablative film, and covered with a transparent layer (usually water). As the laser beam passes through the transparent layer and hits the surface of the material, a thin layer of the ablative layer is vaporized.The vapor continues to absorb the remaining laser energy and is heated and ionized into plasma. The rapidly expanding plasma is trapped between the sample and the transparent layer, creating a high surface pressure, which propagates into the material as a shock wave [8].

Figure 1 Laser peening process

When the peak pressure of the shock wave is greater than the dynamic yield strength of the material, it produces extensive plastic deformation in the metal. The actual depths of the LP induced stresses will vary depending on the type, intensity of the processing conditions chosen and the material properties [9].The laser peening was performed at the Metal Improvement Company in Livermore California. The surfaces of the specimens intended for peening were covered with an aluminum tape 0.22 mm thick. The aluminum tape was replaced between layers of peening. The tamping layer consisted of an approximately 1mm thick laminar layer of flowing water. Some laser peened specimens were peened using a single layer (100%), and others using a triple (300%) layers.

A squa re laser spot size of 4.72 x 4.72 mm 2 was used using a laser power density of 5 GW/cm2 and 18 ns duration. The spots within a layer were overlapped 3%Peening between layers had an off-set of 50% in each direction. A peening frequency of 2.7 Hz and a 1 micron wavelength laser was employed. The peening was applied on the total length of the gauge section on both faces and sides of the specimens.The shot peening process was optimized using “Peenstress” a software developed at Metal Improvement Company. Based on this evaluation, the samples were shot-peened with 0.0234” glass beads, with an Almen intensity of 0.008-0.012A and a 200% coverage rate. To investigate the effects of combining laser and shot peening on the mechanical properties, some of the laser coupons that were processed with a single layer of laser were also shot peened.

Tensile testing was performed at room temperature on a 200 KN servo-hydraulic universal testing machine using a constant crosshead speed of 0.1mm/min. The transverse tensile specimens consisted of conventional dog bone coupons and were 20 cm long with a gage length of 8.5 cm and a gage width of 1.27 cm in accordance with ASTM E8 standard. The coupons were oriented such that the weld was in the center of the specimen and the load was applied perpendicular to the weld direction. Prior to the peening process, the specimens were milled on the top side of the weld removing about 0.4 mm of material. Mechanical properties obtained in the transverse tensile test of the FSW weld generally represent the weakest region of the weld. In that configuration, the elongation constitutes an average strain over the whole gage length which includes the different weld region. This in return does not provide an insight into the correlation between the intrinsic tensile properties and localized microstructure [10]. Therefore the intrinsic tensile properties for various locations across the weld zone were also characterized by a tensile test using a set of strain gauges as illustrated in Figure 2. The local strain data was mapped to the corresponding global stress levels by assuming that the transversely loaded FSW specimens were considered a composite material loaded in an iso-stress configuration [9]. Using this assumption, local constitutive stress-strain relationships were obtained. The accuracy of the measured tensile properties is therefore determined by the degree of non-homogeneity and residual stress levels at all cross sections to which the load is applied.

RESULTS AND DISCUSSION

The standard X-ray diffraction procedure [14] was used to determine the residual stresses induced by LSP. The typical diffractograms of 316L steel and Inconel 600 samples before and after LSP were obtained using RIGAKU, MINIFLEX diffractometer (Cr Kα1 X-ray source, λ =2.29Å, relative angle accuracy 0.010 ) and the results are presented in Fig. 1. The diffracted peaks for steel 316L and Inconel 600 are very similar in position and most likely they originated from Ni (face-centered cubic) planes (111), (200), and (220) with lattice constant of 3.52 Å. It is seen from Fig. 1 that for both materials the LSP results in considerable shift of the X-ray diffractedpeaks to a larger Bragg’s angle.

For calculations of the residual compressive stresses we selected a highest possible diffraction peak (220) for stainless steel 316L and (200) for Inconel 600, which provide most accurate measurements [14]. The mechanical properties of the samples are as follows: (stainless steel 316L/Inconel 600: static compressive yield strength (GPa) = 0.48/0.31, Young’s modulus (GPa) = 195/207, Hugoniot Elastic Limit (GPa) = 1.1/1.3 Poisson’s ratio = 0.36/0.42). Using these values the induced residual compressive stresses were found to be 0.5 ± 0.1 and 0.3 ± 0.1 GPa for 316L and Inconel 600, respectively. One can conclude that the results obtainedevidently confirm the effectiveness of the LSP processing to impart surface compressive stresses on regular surface of the Inconel 600.

To exclude the uncertainty of the focal lens position we studied the dependence of the acoustic pulse energy as a function of the distance between treated surface and focal plane of the lens. It was found that the dependence of the energy of the acoustic pulse on the lens displacement has an extremely resonant character, so the focal length position can be exactly adjusted to most effective way of laser energy delivery to the target.

The asymmetrical shape of the dependence obtained is obviously connected with a shock wave generated by the impurity (microparticles) evaporation when the focal plane is located far before the irradiated surface. The connection was also established between the resonance curve of the acoustic pulse energy and the residual compressive stresses (RCS) obtained at the variable distance of the focal plane relative to the irradiated surface. We found that the maximum of RCS (XRD shift is about 0.30 ) takes place when the lens-target position corresponds to the maximum acoustic signal. When the lens position goes away from maximum of APE to its half-maximum the XRD shift is not found The surface morphology of the samples was checked by the atomic force microscopy (AFM). The AFM images of untreated and LSP treated area are shown in Fig. 2. The tilted column-like microstructure is clearly seen from Fig. 2 for laser treated sample. The microstructure has an approximate periodicity of 30 μm and protrudes above the surface for 2.5 –3 μm. The tilt of microstructure to the surface is about 250. The main feature of the microstructures observed in our experiments is their tilt. It must be emphasized that the nature of this tilt differs significantly from that of conical microstructures described in previous publication [13]. According to the analysis [13] conical microstructure grows towards the axis of laser beam. The authors showed that the tilt of this structure is determined by the incidence angle of laser beam. Contrary to their studies our experiments were performed at the normal incidence, and indicate different mechanism of the microstructure tilt arising under LSP conditions.

The compressive residual stresses can be responsible for the formation of the tilt. It is clear that due to overlapping of laser pulses the areas that were previously shocked (compressed) and the original surface response differently. The compressive state of the shocked area is caused by the reaction of the surrounding region. During laser irradiation both of these area (compressed and uncompressed) are melted creating thin liquid film.

In a liquid state the energy of the compressed area is released in a form of mechanical pulse towards the uncompressed area. Possibly that this released pulse energy results in a tilt of the column microstructure created by capillary waves. Our observation is supported by publication [15] in which a decrease in hardness between two neighboring laser scanned regions was found.

Figure 2. The AFM images of the metal surface after LSP treatment (left) and original surface(write.)

During LSP treatment we have found that a certain amount of target material is injected into the water due to laser ablation. We analyzed the structure of the ablated material using TEM technique.

The image obtained is shown in Fig. 3. It is seen that the ablated material is presented in a form of spherical nanopartricles with typical dimension of 60 nm. Such a phenomenon was observed early in the experiments on laser ablation of metals in liquid media [16,17]. Possible explanation of nanoparticle formation can be given in terms of laser ablation process.

Figure 3. The TEM image of the nanoparticles resulted from LSP. One division is 60 nm

WELD MICROSTRUCTURE

The Microstructure of the weld zone was assessed using digital, optical, and scanning electron microscopes. The specimen used for metallographic investigation was cut and sectioned in a direction normal to the welding direction, and then subjected to several successive steps of grinding and polishing. After that, the specimen was etched using a Keller’s reagent that consists of 190ml of H2O, 5ml of HNO3, 2ml of HF, and 3ml of HCL.

A weld cross section showing different regions of the weld is shown in Figure 3. The cross section revealed the classical formation of the elliptical onion rings structure in the center of the weld. These rings were caused by the rotational flow of the welding tool, and have been attributed to the incremental advance of the tool per revolution [11]. The FSW sample showed no evidence of porosity, or other kinds of defects. The weld nugget seems wider on the crown region of the weld because the upper surface contact with the tool shoulder.

Another cross section of the weld (Figure 4) illustrates the transition from the nugget- TMAZ-HAZ microstructure on the retreating area of the weld. The grain structure at the nugget is fine and equiaxed grains typical of a recrystallized structure.

The grain sizes in this region are of the order of 5-12 μm, and are significantly smaller than the parent material grain due to the higher temperature and extensive plastic deformation. The grain structure at the TMAZ region is elongated, with some considerable distortions that may be attributed to mechanical action from the welding tool. resembles the parent material grain structure. Even though the grain size in this region resemble the base material, previous work by [12] showed the strengthening precipitates in this region have grown in size and were several times larger than in the parent material The HAZ is unaffected by the mechanical effects from the tool, and has a grain structure that

The difference between the base material and the nugget grain structure is also shown in Figure 5. The base material in these figures exhibits elongated grain or pan-caked type morphology typical of that resulting from cold rolling compared to the fine equiaxed grain of the dynamically recrystallized zone (nugget).

Base material1000x

Nugget at 1000x

MICROHARDNESS PROFILES

To quantify the changes in hardness at different regions across the surface of the weld, Micro-hardness test measurements were taken on the top and bottom surface of the specimens as shown in Figures 6 & 7. The measurements were taken using a Struers microhardness machine using a 300g for 3 seconds. The figures show a softened region corresponding to the weld nugget. The aluminum alloy sheets which are generally cold rolled tend to increase the mechanical properties of the produced sheets by increasing the dislocation density. As the grains undergo recrystallization in the weld nugget, the strain induced dislocations will annihilate resulting in a decrease to the mechanical properties of the weld. The soft regions noted throughout the weld could also be attributed to coarsening and dissolution of strengthening precipitates during the thermal cycle of the FSW. It was noticed that the lowest hardness levels were outside of the weld nugget and close to the edge of the TMAZ. The variations in hardness can be correlated to the microstructure developed after the welding process. Previous work by [11, 12] has indicated that strengthening precipitates in the HAZ have grown in size and were several times larger than in the parent material, hence resulted in a reduction in hardness. The local coarsening and growth process for strengthening precipitates is a function of temperature, which is a function of distance from the weld nugget [13]. Accordingly, the hardness levels increased with increasing distance from the weld as precipitation hardening became more effective.

Figures 6 & 7 indicate that hardening was higher in shot peened specimen compared to laser peened ones. This is consistent with the findings of Peyre et al [14] on unwelded 7075-T7351 aluminum plates. The investigation revealed that Vickers hardness increased from 160 HV(25g) to 170 for 7075-T7351 in the laser peened specimens, but Shot peening produces greater surface hardening than laser peening over the first 0.2 mm from the surface. This was due to differences in the pressure duration produced by the peening process where laser shock interaction times are smaller than those for conventional shot peening. The longer interaction time in the shot peening process results in higher dislocation generation and motion.

The higher hardening in shot peening was also attributed to the number of slip planes activated by multiaxial surface loading [15]. The surface hardening in 7075-T7351 caused by laser peening was also demonstrated by Clauer and Fairand [16] who showed that that hardness properties were significantly improved as compared to the unpeened properties. 10

It was also noticed from the hardness profiles that the soften region at the top portion of the weld was relatively wider than the one on the bottom. That is due to the specimen heating characteristics during welding. For example the top region generally exhibits higher temperature compared to the bottom region due to the heating from the tool shoulder, while the bottom surface is in contact with the backing plate which acts as a heat sink. The lower heat input at the bottom can significantly reduce the extent of metallurgical transformations such as re-precipitation and coarsening of precipitates that take place during welding. Therefore, the local strength of individual regions across the weld zone is improved [17]. This is also reflected in Table 1 where the average hardness levels at the weld nugget are represented. The average values on the top region were lower than their correspondent values on the bottom region for the conditions tested. The hardness levels using a single layer of laser peening did not have an effect of the hardness, while using a triple layer resulted in hardness levels comparable to the shot peened ones.

Table 1. Mechanical properties for the various peening configurations Specimen

Hardness HV 300 gf (Top of weld)

Hardness HV 300 gf (Bottom of weld)

Unpeened 117 133 Laser Peened (100%) 118 136 Laser Peened (300%) 129 143 Shot Peened 134 144 Combination 136 145

MECHANICAL PROPERTIES

The global tensile properties for the different peening configurations are shown in Table 2. While laser peening resulted in a 17% increase to yield stress, no improvement was obtained to ultimate strength. All tensile specimens fractured at or near the interface between the weld nugget and the TMAZ on the retreating side of the weld at a 45 degree angle. That distance corresponded roughly to the radius of the tool shoulder.

This is normally explained by strain localization within the minimum hardness region of the weld boundary on the retreating side of the weld. This usually happens in areas softened by the welding process which results in a comparatively low overall strain. The interface between the weld nugget and TMAZ correspond to low hardness region because the original structure in this region is over aged and there is not enough solute left in the material. Therefore, this area of the weld will be relatively ineffective in inhibiting dislocation motion and the strain localization in the softened area of the weld will result in a degradation of the mechanical properties. Since the yield strength of the transversely loaded FSW specimens were less than the yield strength of the base metal, the base metal experienced predominantly elastic strain throughout the test [18]. In all cases, mechanical properties for FSW specimens where significantly lower than the base material. 14

Table 2. Mechanical properties for the various peening configurations

SPECIMEN

YIELD STRESS(MPA)

ULTIMATE STRESS (MPA)

ELONGATION (%)

Laser Peening(one layer)

266 322 4.47

Laser Peening(three layers)

266 323 5.02

Laser & Shot Peening 248 320 4.47Shot Peening 228 320 4.46

Unpeened 227 319 4.57Base 534 601 11

The consistent failure at the retreating side of the weld also suggests that the intrinsic tensile properties of the welded joints are not symmetric on the two interfaces of the weld. Tensile properties on the retreating side were weaker than the advancing side. Generally the global yield strength in FSW is measured using the 0.2% offset. This could result in inaccuracies because the strain is not uniform along the gauge length in an under matched weld specimen [2].

Figure 8 shows an example of the mechanical properties at different regions of the weld for a peened specimen. The results were obtained with the aid of strain gages using the iso-stress condition. For the iso-stress condition to be valid the different weld regions are assumed to be arranged in series with a homogeneous cross section at any location in the specimen [19, 20]. Because of the through the weld thickness property gradients that exists in FSW, this assumption may not be very accurate. To verify the iso-stress assumption, Lockwood and Reynolds [20] conducted a series of tensile tests on reduced thickness specimens at different regions across the weld thickness. Overall, the thick specimen properties closely match the thin specimen properties and seem to justify the iso-stress approximation for friction stir welds. It can be seen from the Figure 8 that the lowest properties corresponded to the interface between the weld nugget and the TMAZ. The highest properties in the FSW took place in the nugget region. It was also noticed that the different weld zones resulted in different resistances to deformation due to differences in grain and precipitate size and distribution

It should be noted that the iso-stress assumption used to determine the local property determination assumes an initially stress free material. Although residual stresses in the unpeened sample were eliminated or greatly reduced by the act of cutting the specimens from the welded plate, substantial residual stresses are expected from the peened specimens. Therefore, the results for the peened specimens may have a higher uncertainty associated with it.

Figure 9 represent the tensile properties for different peening methods in the weld nugget region. It is evident that peening the surface resulted in an increase to the yield strength, with the laser peening (300%) exhibiting the most pronounced effect.The mechanical properties at the weld interface are represented in Figure 10. This region of the weld resulted in the lowest resistance to plastic deformation due to the microstructure associated with this region as discussed earlier. The results were also compared to the base unwelded material as shown in Figure 11 for comparison. It is evident that significant reduction in strength has resulted from the welding process.Figure 12 illustrates the same effects for the HAZ region. Except for the laser peening using three layers, other peening techniques did not result in a significant increase to tensile properties. The increase in mechanical properties from the laser peening was mainly attributed to the increase in dislocation density by the laser peening process, and the high level of compressive residual stresses introduced during the high energy peening. 19

FRACTOGRAPHY

An example of the fractured surface is shown in Figure 13. The microscopically ductile shear fracture took place at 45 degrees by slip along the slip planes that were favorably orientated with respect to the planes of maximum shear stress. Mahoney [13] also attributed the fracture path in FSW to the configuration of the temperature profile through the thickness of the sheet, corresponding to a location where strengthening precipitates were coarsened

Fractographic examinations of the broken tensile samples revealed characteristic features like dimples indicative of ductile failure. This process takes place in parts containing inclusions or precipitates. Under increased strain microvoids grow, coalescence and finally voids coalesce into a crack until rupture occurs.

Figure 13 FSW Fracture surface

ANALYSIS Laser peening is similar to shot peening but imparts compressive stresses much deeper into components with mnimal surface deformation. The process replaces the stream of tiny metal or ceramic balls with short blasts of intense laser light, which generates high-pressure plasma, resulting in consistent and deeper compressive stresses in the material near surface, significantly improving performance and fatigue life.

Laser peening technology has been under development since the 1970s at research facilities such as Battelle Laboratories, but only recently has new laser technology developed at the Lawrence Livermore National Laboratory allowed the development of systems that can peen fast enough for industrial use. With the advent of higher output systems, laser peening is now being used in a wide range of industries.

APPLICATIONS

Now that laser peening technology and processes have been refined for practical use, commercial applications range from large aircraft components to power generation system parts to knee replacements—anywhere that critical, high-cost components require greater depths of compressive stress.

Mobile systems are also advancing the widespread acceptance and use of laser peening. Transportable laser peening systems, currently in routine operation, facilitate reaching components of arbitrarily large size, such as naval vessels in a shipyard. They are completely self-contained and allow quick setup and teardown on site anywhere laser peening is needed.

Industrial applications either in production or detailed development include components for:

Commercial wide body aircraft engine blades and discs, drive train components on U.S. Army

helicopters, engine components for automobiles and biomechanics

Energy systems such as steam turbines, power reactors and others Naval systems including structural components; propulsion system and hull applications; arresting and launching components; and lightweight alloys

Laser peenforming of wing skin components of aircraft, especially large wide body aircraft.

The Navy Metalworking Center (NMC) is working with Metal Improvement,Company NAVAIR-sponsored

project that will evaluate the potential benefits of laser peening on selected aircraft components in the Navy inventory. The NMC project will develop and optimize the laser peening process for specific Navy components through material evaluation, demonstration and validation tasks. The project will evaluate the residual stress level and compressive layer depth as a function of laserbeam parameters of intensity, duration and number of applications; develop a model that predicts residual stress distribution to include location and distribution of positive stress profiles; and conduct metallurgical evaluations of specimens to determine characteristics after peening andthen after surface finishing on critical aircraft components.

In addition to increasing the fatigue strength of metals, laser peenin is being used to form large components with complex contours for commercial aircraft. The 8 x 3-foot wing panel represented an initial

demonstration of the unique processing capability. The forming process is going into production in the 2nd quarter of 2008 on wing skin panels for the new Boeing 747-8. Photo courtesy of Metal Improvement

Company

Laser peening processing produces an intense plasma on a protective layer placed on the surface of metals, resulting in a high pressure wavethat imparts compressive residual stress deep into the near surface of

ADVANTAGES

Laser peening exhibits substantial advantages over conventional shot peening relative to fatigue strength improvement, depth of compression layer and process control. Each laser pulse creates an intense shock wave over a roughly 5 x 5-millimeter area that drives a residual compressive stress approximately 1- to 2-millimeters deep into the base metal. In conventional peening, this ompressed layer is approximately 0.25 millimeters deep. The added depth is key to laser peening’s superior ability to prevent cracks from initiating and propagating, which extends the life of parts three to 10 times over that provided by conventional treatments. Because the fatigue strength of laser-peened parts is significantly increased, components that are processed with laser peening technology often can be produced thinner and lighter, allowing for greater flexibility in the design and operation of systems. Among the other advantages of laser peening over conventional shot peening, laser peening does not require physical contact with treated components and is not limited by surface finish or geometry

The application of laser peening to resolve the FOD tolerance problem on the F119 4th stage IBR will avoid a redesign effort, estimated to cost $10M while preventing a potential delay in the introduction of a full flight envelope qualified F119 engine to the fleet (F-22,JSF) – a major benefit to the warfighter. Compared to the previous state of the art, reductions in processing costs associated with this program will result in savings of more than $10M for the 4th stage IBR alone, over the entire F-22 engine delivery program (738 engines). Since LSP may be applied to as many as four other F119 rotors as well as to all of the F119 engines produced for the JSF, the total savings are likely to be many times higher. LSP surface treatment will also decrease the cost per flight hour and increase mission readiness for the F-22 Raptor by decreasing the frequency of required maintenance inspections on this IBR by 30 to 50 percent

F119 Integrally Bladed Rotors

SUMMARY AND CONCLUSION

The effects from laser peening, shot peening, and a combination of both on the mechanical properties on Friction Stir Welds AA 7075-T7351 were investigated. The peening effects on the global and local mechanical properties through the different regions of the weld were characterized and assessed. The tensile coupons were machined such as the loading was applied in a direction perpendicular to the weld direction. The surface hardness levels resulting from various peening techniques were also investigated for both sides of the welds. The bottom surface indicated higher hardness levels when compared to the top surface. That was attributed to the lower heat input at the bottom side of the weld, which can significantly reduce the extent of re-precipitation and coarsening of precipitates that take place during welding. Although shot peening resulted in a high increase to hardness it did not improve the tensile properties of the FSW. In contrast, single layer laser peening did not improve the surface hardness but resulted in higher tensile properties. The highest increase in tensile properties resulted from using three layers of laser peening. The increase was mainly attributed to the increase in dislocation density and the high level of compressive residual stresses introduced during the high energy peening.

We developed an experimental prototype of the LSP system using commercially available optical and mechanical equipment and Nd:YAG laser. Our experiments have demonstrated the LSP abilities for the improvement of several fundamental characteristics of stainless steel 316L and Inconel 600. The X-Ray analysis was made for 316L steel and Inconel 600, which showed that the residual compressive stresses up to 0.4-0.6 GPa could be produced by LSP. We found that overlapping of the laser spots during scanning of the surface results in tilted column microstructure of the treated surface. The tilt of columns is probably caused by the energy of the compressed area, which is released due to laser pulse melting. We found that during LSP the target material injected into the water due to laser ablation is transformed is spherical nanoparticles with diameter of 60 nm.

REFERENCES

1. Thomas W. Metal. Friction stir butt welding. Int Patent App PCT/GB92/02203, and GB Patent App 9125978.8, December1991.US patent No.5,460,317,October 1995

2. Kroninger, H.R, Reynolds, A.P.R-curve behavior of friction stir welds in aluminum-lithium alloy 2195. Fatigue Fract Engineering Mater Struct, (2002) 25, 283-290

3. Sutton, M. A, Reynolds, A. P. et al. A study of residual stresses and microstructure in 2024-T3 aluminum friction stir butt welds. Journal of Engineering Materials and Technology (2002) Vol. 124. 215-221 Clauer, A. et al. Metall Transactions. A, 1976, 8, 1871-1876

4. Montross C. et al. Subsurface properties of laser peened 6061-T6 Al weldments. Surface Engineering 2000, Vol. 16, No2. 116-121.

5. John, R., et al. Residual stress effects on near-threshold fatigue crack growth in friction stir welds in aerospace alloys. International Journal of Fatigue 25 (2003) 939-948

6. Mahoney, M. W. et al. Properties of friction –stir-welded 7075 T651 aluminum. Metallurgical and Materials Transactions A. Volume 29A, 1998. pp 1955-1964.

7. Tan Y. et al. Laser shock peening on fatigue crack growth behavior of aluminum alloy. Fatigue and Fracture of Engineering Materials. 27, 649-656.

8. J. K. Gregory, H. J. Rack, and D. Eylon (eds.) Surface Performance of Titanium, TMS, Warrendale, PA. (1996) pp. 217-230

9. Mishra R. S. et al. Friction Stir Welding and Processing. Materials Science and engineering R 50 (2005) 1-78.

10. Lee, W. et al. Effects of the local microstructures on the mechanical properties in FSWed joints of a 7075-T6 Al alloy. Zeitschrift fuer Metallkunde/Materials Research and Advanced Techniques, v 96, n 8, August, 2005, p 940-947

11. Jata K et al. Metall Materials Transactions. 2000, 31A. 2181-2192.

12. Mahoney M.W. et al. Properties of friction stir welded 7075 T651 aluminum. Metallurgical and Materials Transactions A. Volume 29A (1998) 1955-1964.

13. Peyre P. et al. Laser induced shock waves as surface treatment for 7075-T7351 aluminum alloy. Surface Engineering. Vol 11, No 1 (1995) 47-52

14. Peyre P et al. Laser shock processing of aluminium alloys. Application to high

15. cycle fatigue behaviour. Materials Science and Engineering A210 (1996) 102-113

16. Fairand B. et al. Laser generates stress waves: their characteristics and their effects on materials. American Institute of Physics Conference. Laser Solid Interactions and Laser Processing. 1978.

17. Lomolino S. et al. On the fatigue behaviour and design curves of friction stir butt-welded Al alloys. International Journal of Fatigue 27 (2005) 305–316

18. Peel M. et al. Microstructure, mechanical properties and residual stresses as a function of welding speed in aluminum AA5083 friction stir welds. Acta Materialia 51 (2003) 4791-4801

19. Lockwood W.D et al. Mechanical response of friction stir welded AA2024: Experiment and modeling. Materials Science and Engineering A, v 323, n 1-2, Jan 31, 2002, p 348-353

20. Lockwood W.D, Reynolds, A.P. Simulation of the global response of a friction stir weld using local constitutive behavior. Materials Science and Engineering A, v 339, n 1-2, Jan 2, 2003, p 35-42

21. M. Gerland, M. Hallouin, H. N. Presles, Material Science and Engineering 156A, 175(1992).

22. Y. Sano, N. Mukai, K. Okasaki, M. Obata, Nucl. Instr. and Meth. in Phys. Res. B121, 432(1997).

23. M. Obata, Y. Sano, N. Mukai, M. Yoda, A Shima, and M. Kanno, Proc. of the 7th International Conference on Shot Peening (ICSP7), Warsaw, Poland, 1987, p.387

24. P. Peyre, R. Fabro, L. Berthe and C. Dubouchet, J. of Laser Application 8, 135 (1996).

25. P. Peyre, L. Berthe, X. Scherpereel, R. Fabbro, and E. Bartnicki, J. of Appl. Phys. 84,5985 (1998).

26. R. Fabbro, J. Fournier, P. Ballard, D. Devaux, and J. Virmont, J. Appl. Phys. 68, 775(1990).

27. R. Fabbro, P. Peyre, L. Berthe, and X. Scherpereel, J. of Laser Application 10, 265(1998).

28. D. Devaux, R. Fabbro, L. Tollier, and E. Bartnicki, J. Appl. Phys. 74, 2268 (1993).

29. P. Peyre, R. Fabbro, Optical and Quantum Electronics 27, 1213 (1995).

30. S. A. Akhmanov, V. I. Emelyanov, N. I. Koroteev, and V. N. Seminogov, Usp. Fiz. Nauk 147, 675 (1985) [Sov. Phys. Usp. 28, 1084 (1985)].

31. A. A. Bugayev, V. A. Lukoshkin, V. A. Urpin, and D. G. Yakovlev, Zh. Tekh. Fiz. 58,908 (1988) [Sov. Phys. Tech. Phys. 33, 550 (1988)].

32. F. Sanchez, J. L. Morenza, R. Aquiar, J. C. Deldago, M. Varela, Appl. Phys. A 60, 83(1998)

33. S. I. Dolgayev, S. V. Lavrishev, A. A. Lyalin, A. V. Simakin, V. V. Voronov, G. A.shafeev, Appl. Phys. A 73, 177 (2001)

34. H. P. Klug, L. E Alexander, X-ray Diffraction Procedures, (John Wiley & Sons, 1973),New York, Chichester, Brisbane, Toronto, Singapur, 966 p.

35. D. Grevey, L. Maiffredy, A. Vannes, Journal of Material Science 27, 2110 (1992)

36. H. Cai, N. Chandhary, J. Lee, M. F. Becker, J. R. Brock, J. W. Keto. Journal of AerosolScience and Technology 29 627 (1998).

37. A. Miotello, M. Bonetti, G. De Marchi, G. Mattei, P. Mazzoldi, C. Sada, F. Conella,Appl. Phys. Lett. 79, 2456-2458 (2001)