Mechanisms of Bonding for Solid-State Welding Processes

Jerry E. Gould, Edison Welding Institute

SOLID-STATE WELDING PROCESSES the diffusion bonding literature. Resulting mod- Resistance processes encompass all the methods in which metallic els show the importance of the size and distri- a. Flash butt welding bonding occurs without the presence of resolidi- bution of the residual oxide particles, as well b. Resistance butt welding fied liquid metal. These processes range from cold as the role of the thermal cycle. c. Projection welding methods (cold-pressure welding) to hot upset pro- Finally, the third underlying mechanism of d. Mash seam welding cesses (forge processes) to diffusion processes forge welding processes is the decomposition of Friction processes (diffusion bonding). Typically, these processes the interfacial structure. Following the forging a. Inertia welding take advantage of applied strain and/or heat to required to displace contaminants, the bondline b. Continuous-drive friction welding facilitate joining. Joining is largely the result of can be characterized as a highly dislocated, c. Linear friction welding intimate intermetallic contact in the absence of high-energy structure. Improvements in weld d. Friction stir welding local protective films. performance can be made by decomposing this Arc processes This article focuses on the underlying mechan- structure and reducing the residual bondline a. Percussive welding isms of bonding for these processes, with particu- strain energy. Decomposition can occur either b. Magnetically impelled arc butt welding lar emphasis on mechanisms for the forge-type by recovery or recrystallization, depending on processes, while mechanisms for both the cold the thermal cycle employed. There is evidence These processes can be thought of as having and the diffusion processes are considered in to suggest that decomposition by recrystalliza- two generally separable stages. These include a other articles. Specific mechanisms for different tion gives better bond performance. Decomposi- heating stage and an upsetting stage. As such, stages of these processes are identified and quan- tion by recrystallization can be promoted by these welding methods can be generally classi- tified using best-available theory. Further, these appropriate thermal cycles as well as appropriate fied as heat and forge processes. Distinctions mechanisms are used to understand the roles of distributions of strain following upsetting. between these processes then are largely in temperature and strain in facilitating bonding how heat and forging are applied. Inevitably, with these classes of joining technologies. however, heat is first applied. Mechanistically, There are three categories of bonding Solid-State Welding Processes this heat is used for two purposes. First, heating mechanisms for the forge welding processes to the workpieces reduces the yield strengths of be considered: Solid-state welding processes are the oldest of the base materials and permits forging to occur welding processes, with the official American with high degrees of strain at reduced upsetting Contaminant displacement/interatomic Welding Society definition of forge welding forces. Second, if heating is properly applied, bonding requiring an anvil and a hammer (Ref 1). Solid- upsetting creates high degrees of strain over a Dissociation of retained oxides state welding processes have proliferated, particu- very localized area (at the bonding surface). Decomposition of the interfacial structure larly over the last several decades, as new power Once the appropriate heating has been accom- systems have developed. General classifications plished, forging (or upsetting) commences. Modeling of contaminant displacement/inter- of these processes include cold-pressure welding, Upsetting also has two major functions. These atomic bonding is largely taken from the cold- externally heated hot-pressure processes, resis- include collapsing asperities to create intimate pressure literature and adapted to the forge tance processes, friction processes, arc-heated contact, and displacing/dispersing protective welding processes. This modeling suggests that processes, and diffusion processes. In this article, oxides and films to facilitate metal-to-metal with increasing surface strain, bond strengths mechanisms of bonding are described for those bonding. Residual heat content/heating is also can asymptotically approach that of the base processes using both mechanically applied strain- considered advantageous, to further consoli- metal. The cold-pressure models can be adapted ing and heating. Detailed examinations of bonding date/homogenize the joint. to forge welding processes by considering the mechanisms of the other processes are available in role that the developing temperature field has the literature. These include the cold-pressure on the distribution of strain. Generally, how- welding processes (Ref 2–8) and the diffusion Mechanisms of Solid-State Bonding ever, after the surface strain is applied (upset), bonding processes (Ref 9–12). Specific variants there are still residual oxide particles trapped of the other bonding processes are as follows: To attempt to define the specific mechanisms in the bondline, and thermal dissolution of these of bonding for solid-state welding processes, it particles can further improve joint performance. Externally heated hot-pressure processes is first necessary to have an understanding Thermal dissolution modeling has been a. Forge welding about the microstructural and surface condi- adapted from the carbide dissolution (in steels) b. Gas pressure welding tions of the workpieces planned for joining. literature, with stability data collected from c. Induction hot-pressure welding On a microscopic scale, the surfaces for 172 / Fundamentals of Solid-State Welding bonding have been well categorized as irregular relatively unstable and is unlikely to yield an A more general plot showing the relationship and covered with various oxide and contami- adequate joint. As a result, the last stage of between oxide/metal hardness ratio and nant films (Ref 2–12). In addition, there may the process is to relieve these local bondline required deformation for bonding is presented be microstructural/compositional irregularities, stresses, typically with some sort of thermal in Fig. 3. One method of improving the charac- which further complicate the joining process. assist. Depending on the treatment, this local teristics of surface film fracture is to locally A typical representation of the prebond surface concentration of strain energy can result in a cold work the base metal. It has been demon- condition is presented in Fig. 1. This surface is final bond structure ranging from local recovery strated that cold working the surface using generally characterized as three layers: the base to recrystallization (Ref 15). scratch brushing both minimizes the extent of material, a layer of mechanically and/or chemi- As mentioned, these mechanisms collectively contaminant films and creates a local layer of cally affected metal, and surface oxides/con- permit solid-state bonding between metallic heavily cold-worked material, which, on strain- taminant films. materials, although not all mechanisms are used ing, can fracture and carry more ductile oxide There are a number of mechanisms that can by all solid-state welding processes. Generally, films (Ref 6). proceed to form a bond between such surfaces. these mechanisms, particularly as they operate For most conventional forge welding pro- The most important of these is that asperities on within the group of forge welding processes, cesses, extension of the contacting surfaces is the surfaces must be collapsed to form intimate can be classified in three general areas. These done with a combination of heat and force. contact between materials. In forge welding include surface deformation mechanisms, con- For this stage of the process, local strain is the processes, creation of this intimate contact is taminant dissolution mechanisms, and interfa- most important factor. However, how that strain done mechanically; that is, local yield stresses cial structure homogenization mechanisms. is distributed is a strong function of how the are exceeded on the contacting surfaces, and These are described in detail, including their thermal field is applied. Figure 4 shows some surface deformation is used to create the con- direct relationship to the forge welding pro- results from numerical simulations of the tact. For diffusion bonding processes (not cov- cesses, in subsequent sections. flash-butt welding process (Ref 16). This plot ered extensively in this article), such surface Contaminant Displacement/Interatomic shows how contact surface strain is affected collapse is done under relatively low forces Bonding. As briefly described previously, sur- both by the amount of upset used and the level and relies on creep and surface diffusional face deformation mechanisms have two func- of flashing acceleration employed. Flashing mechanisms to consolidate the surfaces. tions: to collapse surface asperities and to acceleration directly controls the heat distribu- Once the surfaces have come under intimate displace surface contaminants. It is of interest tion in the workpiece, with higher flashing contact, bonding still cannot initiate until a num- that the best information on the role of surface accelerations indicating steeper, higher thermal ber of other criteria are met. The most important strain and its effect on the extent of solid-state gradients (Ref 17, 18). For conductivity materi- of these is how oxide and surface contaminant bonding is available in the cold-pressure weld- als, particularly aluminum and copper, strain films can be affected to allow intimate contact ing literature (Ref 2–8). Several authors have location provided simply by the thermal gradi- of the underlying virgin materials. Generally, examined the roles of surface condition, ent is difficult. In such cases, pinch-off dies there are two mechanisms for this. For forge mechanisms of interfacial breakdown, and are recommended (Ref 19). Pinch-off dies sim- welding processes, contaminant films can be bro- degrees of subsequent bonding for cold-pres- ply use the constraint of the die (rather than the ken up as a result of mechanical action. In addi- sure welding applications. Collectively, initial thermal profile) to create localization of tion, it is also possible to break down metal bonding, related to surface straining, appears forging. The function of pinch-off dies is shown oxides by dissolution into the matrix. This is a to progress through the following stages. schematically in Fig. 5. mechanism particularly important in diffusion bonding (Ref 13) but also plays a role in other thermally assisted forge processes. Extension of the Contacting Surfaces Even when base materials are in intimate Separation of the Contaminated contact, there are additional changes that must For any bonding to occur, it is essential that Areas occur to facilitate an adequate joint. First, crys- contaminating oxides/films be disrupted. This tallographic matching across the boundary must is, of course, accomplished by application of It is established that the onset of bonding occur (Ref 14). Obviously, most forge welding contact surface strains. It is equally important, occurs with applied surface strain as surface applications are between parent materials with however, that surface oxides/films be in a con- contaminants are separated and virgin base randomly oriented grain structures, and so, this dition in which they can be readily broken up materials are allowed to contact. Considerable bond surface must take on the characteristic of when the surface strain is applied. Table 1 lists work has been done, again largely in the cold- a series of high-angle grain boundaries. Gener- hardnesses for some common oxides. Of inter- pressure welding area, attempting to quantify ation of this dislocation structure can occur est here is the difference between the hard- the separation of these contaminants and the mechanically (Ref 6) thermally (Ref 9, 10), or nesses of the aluminum or copper oxides. resulting bond quality. In examining the role by a combination of the two. At this stage of Figure 2 shows some fractographs of bond sur- of contaminated surfaces, all workers agree that the process, intimate solid-state bonding has faces for cold-pressure welds below critical a critical strain must first be achieved that this undoubtedly occurred; however, the localized bonding deformation for aluminum, copper, surface ruptures (Ref 2–8). There are discrepan- high-angle grain-boundary structure is and silver. It is clear from these results that cies, however, on how this rupture occurs. the aluminum oxide fails in a brittle manner, Mohamed and Washburn (Ref 6) suggest that while the copper oxide fails in a shear manner. separations of the two contaminated surfaces are unrelated, while Wright et al. (Ref 7) and later Bay (Ref 8) suggest that surface contami- Table 1 Representative hardnesses of nants impinge on either side of the joining and some metal oxides at room temperature therefore separate as pairs. Local bonding is Metal Hardness, HV Oxide Hardness, HV then accomplished by extrusion of virgin material into the spaces between the separated con- Al 15 Al2O3 1800 taminated surfaces. Each set of authors has Cu 40 Cu2O 160 developed models based on their assumption Ag 26 Ag2O 135 Au 20 ...... of interfacial breakdown. In each case, the underlying assumption is that the strength of Fig. 1 Schematic representation of workpiece surface Source: Ref 6, Table 1 conditions in the prebonded state the joint is a direct function of the fraction of Mechanisms of Bonding for Solid-State Welding Processes / 173

Fig. 3 Relationship between oxide/metal hardness ratio and the critical deformation for bonding during cold-pressure welding. Source: Ref 3

Fig. 4 Thermomechanical modeling results showing the relationship between ﬂashing acceleration, upset distance, and contact surface strain for ﬂash-butt welding mild steel

the bond area converted by actual base material/base material bonds. The simplest of these models is that developed by Mohamed and Washburn (Ref 6). This model assumes a completely brittle contamination layer and no coordination of contaminants on either side of the bondline. The physical representation of this model is presented in Fig. 6. The resulting equation for strength is: R 2 f ¼ C (Eq 1) R þ 1

where f is the ratio of the joint strength to the parent material strength, R is the surface strain, and C is a constant to incorporate contaminant mismatches and contaminant hardness. Fig. 2 Fractographs showing the faying surfaces of cold-pressure-welded aluminum, copper, and silver at subbonding strains. (a) Aluminum, 3% deformation. (b) Aluminum, 6% deformation. (c) Copper, 5% deformation. (d) Copper, 23% deformation. (e) Silver, 25% deformation. Source: Ref 6, Fig. 1 174 / Fundamentals of Solid-State Welding

"# The model proposed by Wright et al. is slightly 2 for bonding (extrusion pressure), and s is the 1 Rf o more complex. This model was generated for f ¼ C 1 (Eq 2) yield strength of the base material. Y and Y0 ðÞ1 R 2 roll-bonding applications, so largely plain-strain t are the surface exposure and threshold surface conditions exist. The physical representation of exposure, where the surface exposure is de- this model is presented in Fig. 7. This model In this case, C is considered an empirical fined by: assumes matchup of contaminants across the hardening factor, Rt is the threshold deforma- bondline and attempts to account for a degree of tion for bonding, and Rf is the total deformation 1 Y ¼ 1 (Eq 4) prebonding deformation. The resulting equation of the process. Equations 1 and 2 are similar, 1 þ X for joint strength is: asymptotically approaching a maximum bond strength as the total deformation (R or Rf) where X is the degree of expansion of the con- approaches 1. The most complex analysis is tact area. provided by Bay. This model includes the These models, of course, show a greater effects of contaminant films and subsurface degree of complexity, because a greater number hardened layers and is diagramed in Fig. 8. of bonding factors are included. It is important The resulting equation for joint strength is: to recall, however, that these models have been developed for cold-pressure processes, and 0 ¼ð bÞ p pe þ b Y Y p these complexities may be more or less relevant f 1 Y s 0 s (Eq 3) o 1 Y o for conventional forge welding processes. One factor of note is the extrusion pressure where f is now the ratio of weld tensile strength (described as pe in Eq 3). This factor is included to base material tensile strength, b is the frac- either directly or indirectly in each of these tion area covered by contaminant films, p is models. However, for conventional forge weld- the applied pressure, pe is the threshold pressure ing processes, extrusion pressures will fall dra- matically with temperature and may be less of a factor. Also to be questioned is the role of subsurface cold-worked layers, which, in conventional forge welding processes, will probably anneal substantially before any macroscopic deformation occurs.

Realignment of the Grain Structures for Bonding

There is considerable evidence that crystallographic matchup across the bondline is also Fig. 5 Use of pinch-off dies in upset welding processes important at this stage of bonding (Ref 3, 6, to localize strain 14). Detailed work (Ref 3) suggests that contact between similarly oriented close-packed or Metal near-close-packed planes most readily bonds. Fig. 7 Schematic illustration of interfacial breakup as Oxide proposed by Wright et al. Source: Ref 7 For aluminum, (111) to (111) and (110) to (110) were found to bond readily, while (111) Oxide to (100) were found difficult to bond. However, Metal most structural materials are polycrystalline, (a) so such ideal crystallographic matchups are relatively uncommon. To accomplish bonding requires some localized crystallographic reorientation. The model here is the one of a series of grain boundaries. Grain boundaries (b) can be thought of as a complex dislocation pileup, accommodating the misorientation Metal between grains over a very small distance. The types of macroscopic surface strains and Metal local intercontaminant extrusion described here (c) are ideal sources for dislocation generation and undoubtedly contribute to the generation of this bondline structure. An example of this dis- (d) located structure is presented in Fig. 9. This particular example is a resistance butt weld Fig. 6 Schematic illustration of interfacial breakup as on steel, showing evidence of a residual bond- proposed in the Mohamed and Washburn model. (a) Original interface. (b) Fracture of brittle oxide line. This region is characterized by relatively film. (c) First requirement for welding: formation of high internal strain energy and may be a overlapped oxide-free metallic areas. (d) Second quality concern. Reactions of this region to requirement for welding: extrusion of metal through the Fig. 8 Schematic illustration of interfacial breakup the applied thermal fields typical of the forge gaps created in the oxide and some relative shear as proposed by Bay. (a) Interfacial surfaces. displacement at the points of contact of oxide-free (b) Onset of extrusion and thinning of contaminant film. welding processes are described in a subsequent metal. Source: Ref 6 (c) Welds. Source: Ref 8 section. Mechanisms of Bonding for Solid-State Welding Processes / 175

Thermal Dissolution of Oxides/ the solubility product for non-base-material for modeling oxide dissolution during diffusion Contaminants metal oxides relative to the base metal (as well bonding (Ref 21). However, this analysis was as the temperature dependence) can also be esti- largely based on continuous oxide films and mated from the phase diagram. Here, the solubil- focused on the maximum thickness of these The preceding discussion indicates the ity product is estimated from the maximum films for relatively long (diffusion bonding) degrees to which bondline strain can be used solubilities of the secondary metal and oxygen heating cycles. Such an analysis does not take to create a solid-state bond. However, implicit in the base material. into account the breakup of the oxide film into in that discussion and related modeling were If Raoultian behavior of oxygen in the base discrete particles caused by the applied surface two related facts: some level of contamination material is assumed, the proportionality con- strain during forge welding processes, or the was always present in the joint area, and joint stants in Eq 5 and 6 become equal to 1. Then, relatively short heating times. strengths could only asymptotically approach knowing the stoichiometry of the oxide present A better analysis can parallel that done by parent material strengths. The relationship here and using the appropriate phase diagram, Ashby and Easterling for the dissolution of car- is straightforward. As long as contaminants approximate solubility products for some dif- bide particles during welding (Ref 22). That exist in the joint, they reduce effective bond ferent metals can be done directly. Table 2 lists analysis examines the dissolution of discrete area and act as initiation sites for subsequent approximate solubility products for some stan- particles. The approach used attempts to esti- mechanical failures (Ref 20). dard engineering materials with their most mate the roles of both the solubility of the car- To achieve improved joint properties, partic- common oxide. These solubility products are bide constituents into the matrix and diffusion ularly in industrial applications, some further calculated for approximately bonding tempera- of these constituents away from the decompos- reduction in the residual bondline contaminant tures estimated for forge welding processes ing carbide. The approach is based on the content is advantageous. Fortunately, for many (0.9 T ). Materials shown include aluminum, assumption that distinct spherical particles can metallic systems, oxides are soluble in the m iron, and titanium. These solubility products be dissociated completely into a volume of matrix at elevated temperatures. The degree of cover approximately 30 orders of magnitude, matrix with radius l. Further, the particle will solubility of a specific oxide in M xOy in a indicating, on one extreme, the difficulty of dis- dissociate into this volume at a locus of times matrix of metal “A” at equilibrium can be solving aluminum oxide into an aluminum and temperatures defined by: defined by the solubility product: matrix as well as the relative ease with which 1=2 ¼ ð Þxð Þy titanium can dissolve its own oxide. l ¼ðD t Þ (Eq 7) Keq Z CM Co (Eq 5) Similar calculations can be done for nonmatrix metal oxides. The approximately solubility prod- where D* and t* are the combinations of the where Keq is the equilibrium solubility product, uct for Al O on iron is calculated and compared C and C are the compositions of the metal 2 3 diffusivities (D, a function of temperature, M O to the similar calculation for the oxide of iron defined at T*) and times (t*) over which the “M” and oxygen in the matrix metal “A,” and (Fe O ) in Table 3. These calculations were done Z is a proportionality constant relevant to the 2 3 particle can be completely dissolved into the at the approximate bonding temperature for forge volume matrix defined by l. Combining this activity coefficient. If the oxide is of the matrix welding iron. In this case, the stoichiometry for metal, this expression reduces to: approach for examining the role of diffusion the two oxides is similar (x =2,y = 3), so the dif- can be combined with an expression for the ference between the two solubility products is K ¼ ZðC Þy (Eq 6) temperature dependence of the solubility prod- eq o directly related to the solubility of the aluminum uct of the particle, to examine particle dissolu- in the steel. This fact appears to account for the tion behavior. The discussions on solubility with Z a different proportionality constant. This relatively low solubility of Al O in iron. suggests that the solubility product is similarly a 2 3 products for oxides detailed previously can be Such solubility products and diffusivities of used to adapt these equations for oxide particle power function of the maximum soluble oxygen oxygen in the matrix have been used as a basis content in the base material. For oxides of the dissolution. The resulting governing equations base materials, Eq 6 suggests that the solubility include for base-metal oxide particles: product can be estimated from the maximum solubility of oxygen in the matrix as taken from the Bhi Table 2 Estimated solubility products for Ts ¼ (Eq 8) appropriate phase diagram. In addition, the shape ðÞO y oxides present on some common A ln of the oxygen solvus provides some indication of f the temperature dependence for the solubility engineering materials assuming Raoultian product. In a similar manner, Eq 5 suggests that behavior of oxygen in solution in the parent and for non-base-metal oxides: material B Metal Oxide Keq hi Ts ¼ (Eq 9) ðÞM xðÞO y 29 A ln Al Al2O3 3 10 f 15 Fe Fe2O3 1 10 1 Ti TiO2 1 10 where Ts is the dissolution temperature, A and B Calculations are done for temperatures approximately representing are the temperature coefficients for the appro- bonding temperatures for forge welding processes (0.9 Tm) (internally generated). priate solubility product, and f is the matrix volume fraction affected by the decaying oxide, defined by: Table 3 Comparison of solubility products ¼ 1 for Al2O3 and Fe2O3 in an iron matrix at the f hi= (Eq 10) 3 2 t Q2 1 1 appropriate bonding temperature for forge 1 þ exp T t R T Ts welding iron (0.9 m) Base material Oxide K eq In this expression, Q2 is the activation 19 Fe Al2O3 1.8 10 energy associated with the appropriate diffusiv- 15 Fe Fe2O3 1.0 10 ity (oxygen or metal + oxygen), and R is the Fig. 9 Resistance butt weld on mild steel indicating a Internally generated ideal gas constant. With some estimate of t* highly deformed zone down the bondline and T*, Eq 10 combined with either Eq 8 or 9 176 / Fundamentals of Solid-State Welding

(as appropriate) defines an implicit relationship energy, this structure can quickly decompose and temperatures are almost always advanta- between the time/temperature profile for the to a lower energy variant. Parks (Ref 15) has geous, permitting maximum homogenization process and the degree of oxide dissolution. done considerable work to understand the of the joint microstructure. In these expressions, t* and T* are direct breakdown of contacting interfaces. In his functions of the oxide particle size and distribu- work, he suggests two regimes for breakdown tion. Values for these can presumably be esti- of this interfacial structure. These parallel the Comparison of Solid-State Bonding mated from the original distribution of oxides concepts of recovery and recrystallization. Processes on the bonding surfaces and estimations of sur- Recovery of the interface implies a realignment face strain, as described previously. of the dislocated structure to reduce the overall The above discussion suggests that, in sum- From these equations, some qualitative esti- strain energy of the system. This is typically mary, solid state processes employ one or more mate can be made of the role of both the degree done at relatively low temperatures, permitting of three mechanisms to accomplish bonding. of forging and time-temperature profile on bond only local movement of the dislocations that These include disruption of contacting interfaces quality. With increasing strain applied to the make up the boundary structure (Ref 23–25). for nacient metal contact, diffusion related disso- contacting surface, both particle size and den- During recovery, these dislocations realign ciation of residual contaminants, and breakdown sity will inevitably fall. These factors reduce themselves into dislocation cells. An example of the remaining interfacial structure. These fun- amounts of oxygen (and potentially second of such cells is shown in Fig. 11 (Ref 26). Parks damental mechanisms, are driven by process metal) that must be diffused and increase the found that very little effective bonding occurred mechanisms, specifically temperature, time, and kinetics of oxide dissolution. Extended heating if interfacial decomposition was limited to recov- deformation. Obviously, temperature and time (welding) times are important in that, again, ery. Rather, substantial bond strengths were both drive diffusion reactions (promoting both diffusion is promoted. Increasing welding tem- found if higher annealing temperatures were contaminant dissolution and breakdown of the peratures are not only advantageous for increas- used, resulting in bondline recrystallization. interfacial structure), while deformation pro- ing rates of diffusion but also for increasing the This is shown in Fig. 12. Recrystallization is motes interfacial disruption. This approach was solubility product for the dissolution reaction. essentially the nucleation and growth of new used by Fenn (Ref 27). In the developed con- grains. Provided activation energies are high struct, Fenn created a ternary diagram with axes enough, this mechanism of interfacial decompo- of temperature, time, and deformation. For this Breakdown of the Interfacial sition shows the greatest reduction in bondline diagram, the axes are represented as conceptual Structure energy and is suggested by Parks as essential for fractional values. Here, temperature can be con- forming high-integrity bonds. sidered as the fraction of the absolute melting A third mechanism of bonding results from During welding, residual stored energy (as temperature, time as a dimensionless fraction, the decomposition of any interfacial structure. local deformation) can play a role in the kinet- and deformation as the relative collapse of the As described previously, straining of the bond ics of recovery and/or recrystallization. Parks two components. On this diagram, Fenn then surface, extrusion of material around residual has demonstrated that actual bonding tempera- placed hypothetical ranges for a number of solid oxide particles, and matching crystallographic tures can be reduced depending on the degree state processes. The resulting diagram is shown structures across the bondline result in a highly of deformation in the material. Required recrys- in Fig. 14. Limits of the diagram include diffu- dislocated bondline structure. This highly dislo- tallization temperatures as a function of the sion bonding (all time and temperature) and, cold cated structure is of relatively high energy as degree of deformation for a range of materials welding (all deformation). Of note, most conven- well as planar. An example of such a highly are shown in Fig. 13. tional solid state welding processes (flash weld- dislocated bond is shown in Fig. 10. This struc- Obviously, the extent to which this interfa- ing, friction welding, upset (or forge) welding) ture obviously develops during straining the cial structure can decompose is a function of fall toward the middle of the diagram, utilizing contact surface. However, the presence of vari- both the amount of strain applied and the tem- components of all the mechanisms described ous particulates from the contaminated bond perature cycle experienced. Increasing amounts above. Of note, the diagram is not material spe- surfaces may also stabilize this structure. of strain (upset) obviously increase the amount cific. As a result, how individual processes fall Decomposition of this structure is largely a of work in the material and promote subsequent on this diagram will be strongly affected by the thermally assisted process. To develop this breakdown of the interfacial structure. Time at substrate welded, shifting to match the specific highly dislocated structure, considerable energy temperature, however, provides the activation combination of mechanisms most advantageous for deformation is required. Much of this energy to allow this aspect of the bonding pro- to individual material systems. energy is stored in the interfacial structure cess to proceed. It is interesting from this dis- itself. With varying degrees of activation cussion that greater levels of upset may permit bonding at shorter times and lower tempera- Summary tures. However, in practice, extended times This article provides a systematic look at mechanisms of bond formation during solid-state (forge) welding processes. Discussions have been limited to those processes that can be characterized as having two stages: heating and forging. Explicitly excluded were those processes that do not use heating (cold-pressure welding processes) or forging (diffusion bonding processes). For the forge welding processes, three distinct mechanisms of bonding have been discussed. These include contaminant displacement/interatomic bonding, dissociation of retained oxides, and decomposition of the interfacial structure. Contaminant Displacement/Interatomic Bonding. This mechanism of bonding relates Fig. 10 Interfacial structure on resistance-projection- Fig. 11 Dislocation cells in a dynamically recovered to displacement of contaminants by local strain welded mild steel iron microstructure. Source: Ref 26 at the contacting surface. Displacement of these Mechanisms of Bonding for Solid-State Welding Processes / 177

Breakdown of the Interfacial Structure. The side result of the ﬁrst two mechanisms is a highly dislocated interfacial bond structure. This structure results largely from the application of bondline strain but can be stabilized by the presence of discrete oxide particles. This structure is of relatively high energy and can be a detriment to weld quality. Decomposition of this structure does improve bond quality. The mechanism of decomposition, however, depends on the thermal cycle employed. For relatively short or low- temperature cycles, the structure may only recover, resulting in a series of dislocation cells. At higher temperatures and longer times, recrystallization of the metal at the bondline can also occur. Some results suggest that recrystallization of the bondline structure results in better weld properties. Increasing stored energy at the bondline (caused by higher levels of strain) also appears to aid the kinetics of recrystallization and improve weld quality.

REFERENCES 1. “Standard Welding Terms and Defini- tions,” ANSI/AWS A3.0-94, American Welding Society, Miami, FL, 1994 2. D.R. Miller and G.W. Rowe, Fundamentals of Solid Phase Welding, Metall. Rev., Vol 28 (No. 7), 1962, p 433–480 3. R.F. Tylecote, Investigations on Pressure Welding, Br. Weld. J., Vol 1 (No. 3), 1954, p 117–135 4. R.F. Tylecote, D. Howd, and J.E. Fur- midge, The Influence of Surface Films on the Pressure Welding of Metals, Br. Weld. J., Vol 5 (No. 1), 1958, p 21–38 5. L.R. Vaidyanath, M.G. Nicholas, and D.R. Milner, Pressure Welding by Rolling, Br. Weld. J., Vol 6, 1959, p 13–38 6. H.A. Mohamed and J. Washburn, Mechan- isms of Solid State Pressure Welding, Weld. J. Res. Suppl., Vol 54 (No. 9), 1975, p 302s–310s 7. P.K. Wright, D.A. Snow, and C.K. Tay, Fig. 12 Weld strengths as a function of annealing temperature for a range of materials. Source: Ref 15 Interfacial Conditions and Bond Strength in Cold Pressure Welding by Rolling, Metals Technol., Vol 1, 1978, p 24–31 contaminants allows exposure of clean surfaces these particles can be thermally dissolved in 8. N. Bay, Mechanisms Producing Metallic for direct interatomic bonding. The basics for the matrix. The relative solubility of specific Bonds in Cold Welding, Weld. J. Res. modeling this mechanism were largely taken types of particles can be assessed directly by Suppl., Vol 62 (No. 5), 1983, p 137s–142s from the cold-pressure literature. Although sev- examining solubility products between the con- 9. K. Inoue and Y. Takashi, Recent Void eral models are available with increasing levels stituent elements of the particle compared with Shrinkage Models and Their Applicability of complexity, all predict that bond strengths solution in the base material. For base-material to Diffusion Bonding, Mater. Sci. Technol., asymptotically approach that of the base metal oxides, this solubility product is only a function Vol 8 (No. 11), 1992, p 953–964 with increasing surface strain. For the forge of the solubility limit of oxygen in the matrix. 10. Y. Takashi, K. Inoue, and K. Nishiguchi, welding processes, the developed temperature This analysis was used to indicate relative stabil- Identifications of Void Shrinkage Mechan- distribution also plays a role, increasing metal ity of a range of base-metal oxides. This exami- isms, Acta Metall. Mater., Vol 41 (No. plasticity, assisting in localizing strain at the nation was extended, using previous work done 11), 1993, p 3077–3084 bondline, and reducing required upset loads. for dissolution of carbide particles in steel, to 11. T. Enjo, K. Ikeuchi, and N. Akikawa, Thermal Dissolution of Oxides/Contami- examine the kinetics of dissolution. This analysis Effect of Oxide Film on the Early Process nants. The applied surface strains described incorporates both solubility product and diffusiv- of Diffusion Welding, Trans. JWRI, Vol previously permit considerable bonding but ity factors. The results indicate the effects of 10 (No. 2), 1981, p 45–53 leave a residue of oxide/contaminant particles residual oxide particle size as well as the role 12. T. Enjo, K. Ikeuchi, and N. Akikawa, dispersed over the bond surface. As a mecha- of the severity of the thermal cycle for dissol- Effect of the Roughness of the Faying Sur- nism for further facilitating bonding, many of ving these oxide particles. face on the Early Process of Diffusion 178 / Fundamentals of Solid-State Welding

Welding, Trans. JWRI, Vol 11 (No. 2), 1981, p 49–56 13. A. Nied, General Electric Company, Research and Development Center, Sche- nectady, NY, private communication, 1991 14. V.M. Zalkin, Theoretical Problems of Cold Pressure Bonding of Metals, Svar. Proiz., Vol 11, 1982, p 41–42 15. J.M. Parks, Recrystallization Welding, Weld. J. Res. Suppl., Vol 32 (No. 5), 1953, p 209s–222s 16. J.E. Gould and T.V. Stotler, “An Examina- tion of Morphological Development during Flash Butt Welding,” EWI Cooperative Research Report MR9602, 1996 17. E.F. Nippes, W.F. Savage, J.J. McCarthy, and S.S. Smith, Temperature Distribution during the Flash Welding of Steel, Weld. J. Res. Suppl., Vol 30 (No. 12), 1951, p 585s–601s 18. E.F. Nippes, W.F. Savage, S.S. Smith, J.J. McCarthy, and G. Grotke, Temperature Distribution during the Flash Welding of Steel—Part II, Weld. J. Res. Suppl., Vol 32 (No. 3), 1953, p 113s–122s 19. Resistance Welding Manual, 4th ed., Resis- tance Welding Manufacturers Association, Philadelphia, PA 20. W.F. Savage, Flash Welding—Process Variables and Weld Properties, Weld. J. Res. Suppl., Vol 41 (No. 3), 1962, p 109s–119s 21. Z.A. Munir, A Theoretical Analysis of the Stability of Surface Oxides during Diffusion Welding of Metals, Weld. J. Res. Suppl., Vol 62 (No. 12), 1983, p 333s–336s 22. M.F. Ashby and K.E. Easterling, A First Report on Diagrams for Grain Growth in Welds, Acta Metall., Vol 30, 1982, p 1969–1978 23. J.D. Embury, A.S. Keh, and R.M. Fisher, Fig. 13 Recrystallization temperatures as a function of degree of deformation for a range of materials. Source: Ref 15 Substructural Strengthening of Materials Subject to Large Plastic Strains, Trans. Metall. Soc. AIME, Vol 236 (No. 9), 1966, p 1252–1260 24. J.H. Cairns, J. Clough, M.A.P. Dewey, and J. Nutting, The Structure and Mechanical Properties of Heavily Deformed Copper, J. Inst. Met., Vol 99, 1971, p 93–97 25. A.L. Wingrove, Some Aspects of Relating Structure to Properties of Heavily Deformed Copper, J. Inst. Met., Vol 100, 1972, p 313–314 26. J.E. Pratt, Dislocation Substructure in Strain Cycled Copper as Inﬂuenced by Temperature, Acta Metall., Vol 15 (No. 2), 1967, p 319–327 27. R. Fenn, “Solid phase welding-an old answer to new problems?” Metallurgist and Materials Technologist, Vol. 16 (No. 7): 1984, pp 341–342.

Fig. 14 Construct of time-temperature-pressure regimes of solid state welding processes proposed by Fenn (Ref 27). This diagram includes dimensionless values for temperature, time, and deformation, and contains suggested ranges for speciﬁc solid state processes. AN OVERVIEW OF WELDING IN SOLID STATE Mihaela Iordachescu , Elena Scutelnicu

Caminos, Canales y Puertos, Dep. Ciencia de Materiales, Universidad Politecnica de Madrid, Espana Dunarea de Jos University of Galati, Romania [email protected]

ABSTRACT The importance of the Solid State Processes (SSP) has increased in the last decade due to the industry demands of improved properties of joined/surfaced materials, combined with cost reduction and energy saving. New and/or micro-scale solid state processed materials are used by aerospace, automotive and electrotechnics industry. Nowadays, classic SSP are mainly applied to light materials, but progresses were also reported in steels. In this field, the tools design, the technology and practical techniques surpassed the fundamental approach of the materials solid state processing. The SSP parameters evaluation is based on different experiments, approaching the material flow in the large plastic deformation domain. The paper approaches the solid state welding/joining and surface processing. The envisaged SSP are solid state joining processes as Cold Welding (butt and spot welding), Friction Stir Welding - FSW, and surface processing, Friction Stir Processing - FSP. Therefore, the investigation targeted the deformation and flow pattern of the parent metal in case of cold welding and FSW/FSP, processes parameters evaluation and correlation, local analysis of the material structural transformations, and material hardening.

KEYWORDS: solid state processes, microstructure, hardening, material flow

1. INTRODUCTION welding. Of significant importance was the ability to produce a product in solid state. The use of friction In the beginning of the XXth century, together for welding came in the 1950's, when Bishop [9] with the development of resistance welding process, it reported many applications of Russian origin. On a was noticed the decisive influence of the pressure in more worldwide scale, the process gained joints achievement, leading up oriented researches in acceptability for high volume production and its the field of cold pressure welding . Nowadays, ability to join a wide range of materials from the the process is used to achieve the joints of the high 1960's. The automotive industry adopted the process voltage networks' wires, as well as for joining several to weld bimetallic valves, rear axle and front wheel parts of the cryogenic equipment. drive shafts, while the electrical industry was welding Cold welding (CW) process can be easily and copper/aluminium connectors in large scale comfortably achieved, being practically the result of Another major milestone was reached in 1991 the pressing force applied between two metal sheets when Thomas Wayne from The Welding Institute appropriately and carefully cleaned. This process (TWI) in UK patented FSW, extending the requires important materials deformation, obtained by opportunities to use friction heating and material flow using high pressing forces, able to generate upsetting to join sheets and plates in solid state . The pressures 10 times greater than the maximum process principle is illustrated in fig. 2a. material's yield strength. As fig. 1 presents, CW can FSP is a new solid state processing technique be achieved mainly by two methods: spot, (fig. 2b), which can locally eliminate casting defects respectively butt CW and refine microstructures, thereby improving The complexity of the cold weld formation strength and ductility, increasing resistance to reported by W. Zhang and N. Bay was studied, corrosion and fatigue, enhancing formability and in case of aluminium bars butt colds welding, by M. improving other properties . FSP can also produce Iordachescu fine-grained microstructures through the thickness to In 1891, Bevington realised the opportunity impart superplasticity. FSP provides the ability to to use friction to generate heat for both forming and thermomechanically process selective locations on the structure's surface and to some considerable depth to accomplished by adapting the technology developed enhance specific properties. This was mainly for FSW

BUTT COLD :JZf^ r^^F^*"'

WELDING ' ' ••• \ SPOT COLD WELDING 2 _ --J

a) b) Fig. 1. Cold welding process variants: a) spot cold welding; b) butt cold welding; 8 - material displacement; F - upsetting force; 1- pressing/clamping devices; 2 - samples to be welded 1/

FRICTION STIR PROCESSING

a) b) Fig.2. FSW/FSP variants: a) FSW; b) FSP; F - pressing force; sr - rotating speed; Sj- advancing speed

2. COLD WELDING PROCESS technological factors: clamping dies selection, preparation of bar contact surfaces, the initial standoff 2.1. Cold Welding process parameters value, up-setting force, clamping force, bars deformation, and welding equipment adequate selection. Cold welding process can be obtained as result of applying a pressing force on two metal sheets, The selection of clamping dies have to be appropriately cleaned. The process requires important adapted to the bar cross-section shape. Furthermore, a material deformations. Easy deformable metals as large contact surface between bars and clamping dies Aluminium or Copper (or their alloys) can be cold- must be ensured, to avoid the bars sliding at up welded, but the process can be also achieved between setting. dissimilar metals (Aluminium-Stainless Steel, etc.). Figure 3a presents a cross-section through the Butt cold pressure welding rises very interesting clamping die. It can be noticed that this is made of theoretical and practical issues relate to joint three distinct pieces, machined inside for allowing achievement, material deformation, material flow and contact with the bar exterior surface. The space cold hardening during deformation, and also to between these pieces makes possible the initial material thermal response during the up-setting clamping of the bar. During up-setting, some of the process. bar material fills this space, creating longitudinal burrs. The cogged active surface of the clamps The butt cold pressure welding procedure of the prevents the bar sliding in the clamps during up aluminium bars depends on the following determinant setting (Fig.3b). The clamping length used during the experiments was LB = 40 mm (an empirical clamping/squeezing force (Fs = 8,650 N), before technological prescription indicates as minimum upsetting. During upsetting, the bars material actuates value LB = 4 • d, where 'd' is the bar initial diameter towards the clamping dies, developing forces of [1,2]). similar magnitude as the upsetting force. Thus, The geometry of the clamp active side (d! reaction forces of important values are generated in =1.4 • d, a = 5°, (3 = 60°) was designed to ensure the clamping device, without increasing the squeezing appropriate material flow and joint strength. Figure 3c force. In conclusion, whilst the necessary squeezing presents the initial position of the clamps during up force is about 7.30 times smaller than the upsetting setting, and final one, respectively. one, the reaction forces have the same order of The bars contact surfaces preparation, their magnitude as the pressing force. Consequently, the smoothness, alignment and perpendicularity are actuation of the squeezing devices must be designed necessary for preventing their eventual relative sliding to provide the 'Fs' value, whilst the clamping device and compressive buckling. Furthermore, the cleaning itself should be able to carry out the bigger loads and degreasing of the bars extremities were necessary generated by the reaction forces. before welding. The cleaning of the bars extremities The bar deformation, '8', is the ratio between with a rotating wire brush followed the mechanical the one-bar standoff variation (during up-setting) and cutting of the bars samples. its initial standoff. Previous research consider that cold pressure welding process by single up-setting can be obtained only if a minimum deformation (Smin — 0,7 for aluminium) is exceeded during pressing. The adequate selection of the welding equipment depends mainly on the necessary up setting force value, capable to ensure the achievement of cold welded joint.

2.2. Cold Welding material flow

Different tests on butt pressing of aluminium bars of 10mm diameter were performed. The up setting process was stopped at different values of bars deformation, for better understanding the material deformation process and the cold welded joint formation. Fig.3. Clamping device: a) cross-section; Experiments have confirmed that bars cold b) longitudinal section of the clamping area; welding occurs at the deformation S = 0.75. c) longitudinal section of the joint area, before (left) Continuous pressing at higher deformation values led and after (right) applying the pressing force, 'F'; Fs - to corresponding decrease of standoff, with an squeezing force increased certitude of obtaining good quality joint.

The initial standoff represents the initial length of the non-clamped end of the bar to be welded. An optimum positioning of the samples in the clamps is described by an initial standoff capable to ensure accurate up-setting that produces welded joint of good quality. The bars standoff is experimentally determined according to the base materials qualities. The correct standoff value experimentally determined is of 10mm, equal with the diameter of the aluminium bars. An excessive standoff doesn't lead to a correct Fig.4. Cold weld in case of 8 = 0.68 pressing, causing the bars buckling occurs. The up-setting force, 'F', is the actuation force The study intended also to determine the produced by the hydraulic motor of the toggle-lever deformation critical value when the joint tensile press (F = 63,000 N). All the others technological strength surpasses the base material ultimate strength. parameters of butt cold welding process of aluminium Thus, in case of 8 = 0.68, the formation of cold bars were determined at this up-setting (pressing) welding was noticed, but the joint had a poor strength force value. due to the small area of the weld (fig. 4). At tensile A special design clamping device ensured the tests, the samples failure occurred, without bars self-blocking for low values of elongation, due to material loos of elasticity (cold the radial direction. At higher deformation values, hardening caused by pressing), in the weld area. as the microscopic images present, the flow lines Moreover, a weld critical area was defined as the weld orientations on radial direction are observed. area when the joint strength is equal with the base Initiation of the typical forging subgrain structures metal strength. Once surpassed the weld critical area, (fig. 6b), with dimensions lower than 0.3 m, the joint failure initiates in base metal. allows for the fusion of the two lattices, thus achieving the cold welding. -

Fig.5. Ultimate strength of butt cold welded aluminium bars vs. bar deformation

Figure 5 presents the experimentally determined diagram of ultimate strength of butt cold welded aluminium bars as a function of deformation. Three b) domains corresponding to different deformation ranges provide information on the progress of cold Fig.6. Macro and microscopic images of CW joint welding process. It can be noticed that butt cold welding cannot be achieved at deformations inferior (99.5%Al), ( = 0.75); a) CW joint macrostructure; b) CW joint microstructure to = 0.68. Figure 5 also illustrates the moment min when the product of the weld area multiplied to the At microscopic level, the butt cold pressure correspondent stress is equal with the value of the welding process of the aluminium bars is produced ultimate strength of base metal, for the deformation due to the up-setting force when, in the contact area, = 0.73. Good quality joints are possible for bigger the value of the normal stresses couple allows the deformations, with respect to other technological initiation of the subgrain structures, with dimensions parameters, such as the preparation of the contact less than 0.3 m, capable to fusion and create a surfaces or the standoff value. common lattice. Figure 6 shows the macro and microscopic images corresponding to the cold weld formation; the material flowing lines are visible on the joint 3. FSW/FSP PROCESSES macrostructure. Macro and microscopic images featured the 3.1. FSW/FSP material flow and aluminium behaviour during CW, showing: temperature - The increase of the material flow in the up-setting force direction, on the longitudinal axis of the bars, Nowadays, new techniques as solid-state in accordance with the deformation value. The Friction Stir Welding – FSW are currently used for grains are compressed on the direction of the obtaining different aluminium alloys qualitative predominant stress developed in the longitudinal joints. Although the welding may produce high tensile axis of the bar, a typical forging structure being stresses (up to the yield stress) balanced by lower thus obtained. The initial grains form (typical for compressive residual stresses elsewhere in the the drawing manufacturing process) modifies by component, FSW results in a much lower distortion pressing. The increased values of the normal and residual stresses owing to the low heat input stresses couple at high deformation values lead to characteristic of the process . grains refinement and furthermore, to their Recently, a derivative from FSW, Friction Stir reorientation on radial direction. Processing – FSP namely, was proved as being useful - The material flowing outside the clamps and weld for inducing directed, localized, and controlled seam and burr formation, are presented in fig. 6a. materials properties in any arbitrary location of Due to internal stress values, grains slide mainly in components. Basically, the FSP/FSW process has three This result arises because of the large local variations stages: the penetration of the tool, when the local in the plastic flow and from the thermal history plasticity properties of the material quickly changes resulted from the material interaction with the tool. with temperature and the tool travel speed is The microstructure in the stir zone is characterised by acceleration from zero to the characterized by refined grains in a discrete series of working value (fig. 2b-a,b,c); the working stage, when bands and some precipitate mainly distributed at the the travel speed and the pressing force are constant, as grain boundaries. There is also still some debate well as the rotating speed and the tool angles (fig. 2b- concerning the origin of the annular rings observed d); the tool retracting phase, when the travel speed is within the nugget zone attributed to an abrupt decreased by zero value and the tool is removed from variation in the grain size and precipitate density the workpiece. The nugget zone grains suggest effective strains FSP/FSW process is characterised by some main together with a microstructural evolution that occurs technological parameters, namely: tool geometry, tool by a combination of hot working and a dynamic tilt and concordation angles (angles of the tool axis recovery or recrystallization. The temperature reached with the vertical direction in the longitudinal and in the nugget zone is known as being situated in the transverse plane, respectively), rotating and travel range of 450-500 °C for the 6061- Al alloy speeds and plunging/working force. Distinct precipitates and coarsened grains are The material flow during FSP/FSW is quite a observed at the deformed regions of TMAZ. HAZ complex deformation process of practical importance grains are severely coarsened by FSP (Fig. 8: 1, 4). for tools design and materials microstructure The characteristic annular-banded structure is transformations. Therefore, an overall pattern of the distinctly observed to be asymmetric and more material flow hasn't been reported yet. As example, obvious on the advancing side (A) of processed zone the paper approaches the processing of as-cast AA as shown in fig. 8, positions 1, 2, 4. A severe 6061. The processed layer macrostructure is presented deformation has also occurred along the top surface of in figure 7. the processed layer where the shoulder of the tool is in contact with the material. The flow lines from fig. 8 - positions 4, 2, 5 seem to represent plastic deformation increments that develop as the rotating tool moves through the processing line. Although, it is well known that the material is transported from A to the retreating side (R), Colligan showed that with a threaded pin tool, the material from the upper part of the processed zone is pressed down, whereas the material from the lower part processed zone is moved toward the top surface. The material may travel many Fig.7. FSP macrostructures of AA6061 as cast cycles around the tool before being redeposited. A aluminium alloy; 1 -6 microstructures positions little flow of material was observed near the bottom of the processed zone. The base metal (BM) microstructure of as-cast The effect of processing parameters on AA 6061 consists of Al solid solution dendrites along temperature was investigated by Arbegast and Hartley with coarse silicon and intermetallic phases. . They reported that for a given tool geometry and Shrinkage porosity is also prevalent. FSW/FSP closed depth of penetration, the maximum temperature the shrinkage porosity and homogenized the as-cast depends on the rotation rate, while the rate of heating microstructures by breaking up and evenly dispersing depends on the traverse speed. A higher temperature initial phases. Moreover, the resulting microstructures on the advancing side was noticed. do not have a uniform grain size distribution for any From different experimental investigations and one set of process parameters. Grain size varies from process modelling, several conclusions can be the top to the bottom as well as from the advancing to underlined about the FSP/FSW thermal profile: the retreating side. The differences in grain size likely - the maximum temperature developed within the stir are associated with differences in both peak zone is below the melting point of the materials; temperature and time of application of temperature. - tool shoulder dominates the heat generation during Figure 8 shows the typical features of all FSP; different zones in a single processed layer cross- - the maximum temperature increases with increasing section of as cast AA 6061 under processing tool rotation rate at a constant tool traverse speed condition of 1,120 rpm for the rotational speed and and decreases with increasing traverse speed at a 320 mm/min for the welding speed. The positions 1-6 constant tool rotation rate. Furthermore, maximum from fig. 8 are located in different micro structural temperature increases with increasing the ratio of zones. The micrographs show that the microstructure tool rotation rate/traverse speed. of the processed layer is complex and highly - the maximum temperature occurs at the top surface dependent on the position within the processed zone. of the stir zone. Fig.8. Typical features of all different zones in a friction stir processed single layer cross-section of as cast AA6061: 1- flow patterns in the appendage zone; 2 -nugget zone; 3 - the retreating side of TMAZ; 4 -the advancing side of TMAZ; 5 - nugget bottom side; 6 - processed layer bottom side; (200x)

Q[rpmJ 2000 (Sample: AA6065-T4, 1=3.9mm) v[mrn/minj ~ 500 ~ It Q | tHAZ FSW [iv [4 TMAZ Classification _ .. \lQ 1 lHAZ • Coldo . =} . tv tTMAZ

Q [rpm] 1600 v [mm/minj 800

Fig.9. FSW/FSP typical material flow patterns

The characteristic annular-banded structure is . They reported that for a given tool geometry and distinctly observed to be asymmetric and more depth of penetration, the maximum temperature obvious on the advancing side (A) of processed zone depends on the rotation rate, while the rate of heating as shown in fig. 8, positions 1, 2, 4. A severe depends on the traverse speed. A higher temperature deformation has also occurred along the top surface of on the advancing side was noticed. the processed layer where the shoulder of the tool is From different experimental investigations and in contact with the material. The flow lines from fig. 8 process modelling, several conclusions can be - positions 4, 2, 5 seem to represent plastic underlined about the FSP/FSW thermal profile: deformation increments that develop as the rotating - the maximum temperature developed within the stir tool moves through the processing line. Although, it is zone is below the melting point of the materials; well known that the material is transported from A to - tool shoulder dominates the heat generation during the retreating side (R), Colligan [15] showed that with FSP; a threaded pin tool, the material from the upper part of - the maximum temperature increases with increasing the processed zone is pressed down, whereas the tool rotation rate at a constant tool traverse speed material from the lower part processed zone is moved and decreases with increasing traverse speed at a toward the top surface. The material may travel many constant tool rotation rate. Furthermore, maximum cycles around the tool before being redeposited. A temperature increases with increasing the ratio of little flow of material was observed near the bottom of tool rotation rate/traverse speed. the processed zone. - the maximum temperature occurs at the top surface The effect of processing parameters on of the stir zone. temperature was investigated by Arbegast and Hartley 3.2. FSW/FSPparameters and their Figure 10b present the hardness profile results in influence on the processed material case of processing a heat treatable aluminium alloy. Information about the typical location of the global minimum value of the harness field can be found The main result of the research regards the here, located in the interface between the heat affected influence of the friction stir main parameters (the tool zone and the thermo-mechanically heat affected zone. rotational and advancing speed) on the material flow Along the heat affected zone there is, typically, a local pattern around the tool (fig. 9) . In the case of minimum hardness value due to processed material hot conditions, the visco-plastic material flow is more over-ageing. concentrated around the pin and the heat affected zone is wider resulting in a basin shape nugget. In the The processed materials hardness profile opposite, under cold conditions, the thermo- enables a reliable assessment of its static mechanical mechanically heat affected zone is wider and the heat resistance. affected zone is smaller. 4. CONCLUSIONS

The main conclusions emerged after B4 experiments related to butt cold welding of aluminium i 82 - '^>"*T y bars are: V, BO w S, - Small clamping force is needed at process r1- .^•J .\i."V>" ,. »--* beginning, ensuring only the initial bar self- ..• — •v- blocking in clamping dies. The actuation of the squeezing devices must be designed to provide this small value force, whilst the clamping device itself should be able to carry out bigger 16 -14-13 -10-8 -8 -4 -1 0 2 4 6 B 10 12 14 IB loads generated by the pressed material Distance from the nugget centre [mm) reaction forces. - Material flow due up-setting reflects its a) simultaneous displacement outside and inside the clamps; 100 • - The structural refinement microscopically 95 observed has only mechanical origin. 90 . - At microscopic level, the butt cold pressure B5 • *-, welding process of the aluminium bars is BO -

Study establishes a rational basis for atom-to-atom bonding that supports neither the film theory nor the dif fusion principle

BY H. A. MOHAMED AND J. WASHBURN

ABSTRACT. The mechanism of pres welded area was in agreement with brought into intimate contact a weld sure welding in polycrystalline alumi measured fracture strength. It was will be created. The theory attributes num, copper, silver and gold was in concluded that the strength attained the different weldability of different vestigated. The role of the oxide film after a given deformation is deter metals to the relative hardness of the was studied and it was found that no mined by the fractional welded area at bulk metal and oxide. Initiation of metal to oxide bonding contributes to that deformation. welding is controlled by the degree of the strength of the weld. From scan fragmentation of the oxide film. ning electron microscope observa Introduction In the present work, scanning elec tions a two-stage model has been tron microscope photographs show suggested which could explain the Pressure welding is the establish ment of an atom-to-atom bond be that oxide films on aluminum, copper different behavior of the metals and silver crack at deformations studied. tween the two pieces to be joined through intimate contact between much less than the minimum welding The first stage of welding involves oxide-free areas achieved under deformation (see Fig. 1); the way this the formation of overlapped oxide- pressure and without the formation of deformation was determined will be free metallic areas; this is controlled liquid phase. discussed. by: (a) difference on a microscale of In order to develop this bond, sur This suggests that it is not only the the local plastic strain occurring on face films have to be removed or at fragmentation of the oxide film which matching opposite faces of the weld least reduced in amount. Surface controls the initiation of welding. interface, (b) relative hardness of the films fall into two categories: metal and its oxide film, and (c) Energy Barrier Theory mechanical properties of the oxide. Oxide Film The second stage involves: (a) The energy barrier theory (Refs. plastic flow of the metal to the over All metals except gold possess an 6, 7) suggests that even if clean sur faces are brought into contact no weld lapped areas; the stress at which this oxide film at room temperature. In will result. The theory states that an can take place is influenced by the most metals the oxide film reaches a energy barrier exists that must be stacking fault energy of the metal, and limiting thickness in the range 20-100 overcome before welding can take (b) some relative shear displacement angstroms at room temperature place. Parks (Ref. 7) thought that the at the points where metal cleaned (Ref. 1). of oxide comes into contact; this is barrier is recrystallization while Erd- mann-Jesnitzer (Ref. 6) thought that influenced by surface roughness. Contaminant Film it is diffusion. Semenov (Ref. 6) sug The different weldability of differ gested that the energy barrier comes ent metals is attributed to differences This film consists of a thin layer of from the misorientation of the crys in stacking fault energy, hardness moisture and greases. The best tech tals at the contact surface, since he ratio and plastic properties of the ox nique which has proved (Refs. 2-4) to could weld aluminum, copper and ide. The weld strength calculated be successful in reducing these films silver at the temperature of liquid theoretically on the basis of measured is a combination of chemical and nitrogen. It is impossible to assume mechanical cleaning. that welding at this temperature oc Two theories have been proposed curred due to diffusion or recrystal The authors are associated with the Inor so far to explain the mechanism of lization. pressure welding: ganic Materials Research Division. McEwen and Milner (Ref. 8) have Lawrence Berkeley Laboratory and the shown that immiscible metal pairs Department of Materials Science and The Film Theory Engineering, College of Engineering, re can be joined satisfactorily. spectively, at the University of California, This theory (Refs. 2, 5) proposes In the present work the rate of ap Berkeley, California 94720. that if two clean metal surfaces are plying the pressure was found to have

302-s I SEPTEMBER 1975 no effect on either the welding defor Experimental Procedure eluded high purity aluminum, silver, mation or the weld strength. It will be Materials copper, gold (purity, 99.999%) and shown that the misorientation factor commercial purity aluminum (purity, has an important effect in the ini Materials for Lap Welding. The 99%). The dimensions of the strips tiation of welding. materials used for this study in- were 75 mm length, 18.75 mm width and the thickness ranged from 0.8 to 1.2 mm. The overlapping distance was 25 mm.

Materials for Butt Welding. These materials included high purity alumi num (purity. 99.995%) and 6061T6 aluminum alloy (main alloying ele ments are silicon and magnesium). The dimensions of the rods were 60 mm length and 9 mm diameter.

Procedure Welding Dies. Two welding dies were designed, one for lap welding and the other for butt welding. Figure 2 is a schematic illustration of these dies. Surface Preparation. In order to reduce surface films, all the spec imens were first degreased in ace tone and then wire brushed using a motor driven wire brush. Figure 3 shows a scanning electron micro scope photograph for an aluminum surface. It is seen that the surface consists of a series of hills and valleys. Welding. After surface prepara tion, the specimens were imme diately set in the welding die and then the pressure was applied. At the beginning of the experiments, the pressure was applied at very slow rate and then at a much higher rate. It was

lOOmm -* — 50mm l-T-rH i

Specimen 20mm I30rr -" / %i 45 mm r—L-J—I t _i_ - 35mm - 45 mm - -. -

(o) Butt Welding Die

- 40mm iTt^Xtir'

• '

3000X 4 mm >-25mm

Fig. 1 — Scanning photographs for the surfaces of aluminum, copper and silver of defor (b) Lop Welding Die mations below the minimum lap welding deformation. High purity aluminum: (A) 3% defor mation, (B) 6% deformation. High purity copper: (C) 5% deformation, (D) 23% defor Fig. 2 — Schematic illustration of the mation. High purity silver: (E) 25% deformation. All reduced 9% welding dies

WELDING RESEARCH SUPPLEMENT! 303-s the minimum welding deformation and at deformations higher than the minimum, i.e., investigating the frac ture surface of the weld.

Experimental Results and Discussion Role of the Oxide Film To investigate the role of the oxide film the following experiments were carried out: 1. In the case of aluminum butt welding, a hard sleeve was set around the rods, as shown in Fig. 5, so as to prevent the metal from deformation at the interface. If the oxide film does contribute to welding, it would be ex pected that welding can occur in the presence of the sleeve. It was ob served that whatever the pressure applied in the presence of this sleeve no welding did occur. Figure 6 is a scanning photograph of two aluminum specimens pressed at the same pressure, one in absence of the sleeve (A) and the other in the 11000X presence of it (B). By comparing the two photographs it is seen that in the Fig. 3 — Scanning photograph for aluminum surface after wire brushing. Reduced 41% presence of the sleeve the oxide layer did not break. found that the rate of applying the pressure does not have a marked In lap welding, gold could be weld effect on either the welding deforma ed at zero macroscopic deformation. tion or the weld strength. The weld 2. By welding silver at 200 C, it ing time (time of applying and re was found that the minimum lap leasing the pressure) was then set to welding deformation was reduced be 1 min in all the experiments. The from about 75% at room temperature experiments were carried out at room to negligible value at 200 C. This temperature. behavior cannot be attributed only to Deformation Measurements. For a reduction in the flow stress of the lap welding the deformation was metal but is associated with the dis measured as a percentage reduction sociation of silver oxide (it is well in the total thickness of the two strips. known that silver oxide dissociates For butt welding it was measured as a completely at 190 C at atmospheric percentage increase in the cross-sec pressure). tional area. 3. A thin (about 100 angstroms) Measurement of Minimum Welding layer of aluminum was deposited on a Deformation. Figure 4 charts pressure scratch brushed gold surface by and strength vs strain for aluminum. It evaporation in vacuum. The gold is seen that there is a stage of easy specimens were then exposed to the plastic flow; the onset of this stage atmosphere so that the deposited IO 20 30 40 50 layer consisted mostly of aluminum DEFORMATION, % corresponds to a sudden increase in the weld strength. oxide. It was found that no welding did Fig. 4 — Pressure and strength vs strain occur up to about 70% deformation. diagram for aluminum The deformation at which this oc curs was taken to be the minimum These results lead to the following welding deformation. The minimum conclusions: welding deformation for the other 1. The oxide film does not con metals was determined in the same tribute to welding; in order to initiate way. welding this film has to be broken. Mechanical Testing. Lap joints 2. Plastic flow of the metal is a pre were tested in tensile shear. Due to requisite for oxide breakage. the relatively high welding deforma Mechanism of Oxide Film Fracture tion of copper and silver, most Sleeve Specimen failures occurred outside the welding It was concluded in the last section at the weld metal junction. Data for that plastic flow of the metal is a pre lap welding were obtained for alu requisite for oxide film breakage. minum and gold only. Butt joints were Consider a spot on the interface first machined and then tested in ten where the mating surfaces come into sion. contact. Upon applying a deforming Scanning Electron Microscope In pressure the dislocations most favor vestigations. The specimens were ably oriented with respect to the applied stress start to move on their Fig. 5 — Schematic illustration of sleeve viewed in a direction normal to the used to prevent plastic flow at the Inter weld surface. This investigation was slip planes. These dislocations con face in butt welding carried out at deformations less than tinue their motion until the outermost

304-s I SEPTEMBER 1975 loops approach the surface where the the oxide, the oxide ultimately failing dislocations will also multiply in the oxide film may act as a barrier to their in a brittle tensile manner. Evidence oxide. In that case the metal and ox emergence. The stress concentra for the piling up of dislocations ide undergo plastic deformation tion associated with the pile-ups of against surface oxide films has been together. This behavior is possible dislocations can be relieved by either provided by Barrett (Ref. 9). when the oxide film is relatively duc opening a crack in the oxide or mov Figures 1A and 1B are scanning tile. In this case, the oxide may fail in a ing pre-existing dislocations, and/or photographs for high purity alumi shear manner. It is seen in Figs. 1C generating new ones in the oxide, de num surfaces at two deformations and 1D, which are scanning photo pending on the relative hardness of below the minimum welding defor graphs for a copper surface at two the metal and its oxide film and the mation. It is seen that the surface has deformations below the minimum mechanical properties of the oxide. If discontinuities which may corre welding deformation, that the dis the surface is covered with an oxide spond to cracks in the oxide film. continuities are quite different from film harder than the metal, the dislo The direction of propagation of the those on aluminum surfaces. They are cations experience an image force crack is seen to be normal to the in the form of steps or striae which which is a repulsion reflecting the applied stress which is a character may suggest that the oxide film on strain energy of the elastically harder istic feature of brittle tensile failures. copper fails in a shear manner. Figure material. The stress concentration If, on the other hand, the surface is 1E shows the surface of silver, the dis can be relieved by opening a crack in covered with a deformable oxide film, continuities are similar to those on copper surfaces; therefore, silver ox ide also fails in shear. Table 1 gives the relative hardness of the metals investigated and their oxides. If the oxide film is completely brittle at the welding temperature, then the proportions of oxide-free metallic area formed at a certain deformation is equal to the surface extension at that deformation. This is not the case if the oxide is de formable. It may be concluded that the proportion of oxide-free metallic area revealed is dependent upon the relative hardness of the metal and its oxide film, and the mechanical prop erties of the oxide.

The First Requirement for Welding It was found from scanning electron microscope photographs that the ox ide films on aluminum, copper and Fig. 6 — Scanning photographs of aluminum surface (A) in absence and (B) in presence of silver crack at deformations much the sleeve. Reduced 48% less than the minimum welding defor mation (see Fig. 1). This shows that initiation of welding is not deter mined only by the fragmentation of Table 1 — Hardness of Metals and Their the oxide film as was suggested in the Oxides at Room Temperature simple film theory. Table 2 shows the Hardness Hardness minimum lap welding deformations Metal HV Oxide HV for the metals investigated at room temperature. Table 3 shows this Al 15 Al203 1800 deformation for butt welding. Cu 0 Cu 40 2 160 In the mechanism of oxide film Ag 26 Ag20 135 Au 20 (a) Original Interface fracture presented in the last section — — it would not be expected that disloca tion pile-ups at both sides of the inter face would have the same distribu Table 2 — Minimum Lap Welding Defor mation at Room Temperature tion and would always line up. There fore, it seems likely that the oxide film Defor- ^^ Metal mation, % on the two surfaces should break (b) Fracture of Brittle Oxide Film independently not as one layer as was High purity Al 10 suggested by Vaidyanath and Milner Commercial Purity Al 30 (Ref. 10). Although a metallic area is High purity Cu 64 revealed on one surface, yet, the cor High purity Ag 75 responding area on the other surface High purity Au 0 (c) First Requirement for Welding I Formotion of may still be covered by an oxide. Overlopped Oxide-free Metallic Areas Welding cannot commence unless freshly revealed areas overlap one Table 3 — Minimum Butt Welding Defor mation at Room Temperature VVVV above the other as shown sche LVXCVZ matically in Fig. 7. Defor (d) Second Requirement for Welding: Extrusion of This may explain why there is a Metal the Metol through the Gops Creoted in the Oxide mation, % ond some Relative Shear Displacement at the considerable additional deformation Points of Contact of Oxide-free Metal High purity Al 130 required between the first initiation of 6061T6AI alloy 227 Fig. 7 — Schematic illustration of the weld cracks in the oxide and the initiation ing process of welding. Therefore, the first weld-

WELDING RESEARCH SUPPLEMENT! 305-s ing may occur after the cracks begin represents an internal crack. During come into contact. This was con to overlap. mechanical testing these voids grow firmed by butt welding gold with dif until the material between two voids ferent methods of surface prepara Fracture Surface of the Weld thins down and separates by rup tion; gold was selected due to the turing. Figure (14) is a macrophoto- complete absence of the oxide film. It Figures 8 to 13 are scanning photo graph for the surface of a butt welded was found that, at the pressures graphs of the fracture surface of aluminum specimen just before fail where welding did occur in wire welds. It is seen that the fracture sur ure; deformation bands can be seen. brushed specimens, highly polished faces of all the metals investigated specimens degreased in acetone did have the same feature. They are Role of Surface Roughness not weld. In lap welding of gold at zero made of concave depressions which macroscopic deformation there was a is the prominent feature of fracture The fact that initially rough sur critical pressure for welding although created by coalescence of voids (dim faces are required for welding sug the macroscopic deformation is ex ple rupture), (Refs. 11, 12). The gests that bringing oxide free metals cluded. This behavior should be asso into contact does not result in weld dimples are seen to be equiaxed ciated with the local deformation of ing unless there is also some shear which is typical of normal rupture. high spots since there should be a displacement as the two surfaces Each of the voids of the interface minimum applied pressure neces sary to deform these spots. Agers and Singer (Ref. 13) suggested that in lap welding the local deformation at the interface is more important than the macroscopic deformation. It seems that the importance of the shear displacement, besides in creasing the contact area, is that it de stroys the continuity of any adsorbed oxygen layer which may contaminate the oxide-free area due to trapped air at the interface. Second Requirement for Welding The second requirement for weld ing is to force the metal to flow through the gaps created in the oxide. This flow will result in welding when the metal flowing from one side of the interface comes into contact with the metal flowing from the other side and 1000X then some relative shear displace Fig. 8 — Fracture surface of lap welded high purity aluminum. (A) 29% deformation, (B) ment occurs to destroy the continuity 46% deformation. Reduced 50% of the adsorbed oxygen monolayers

F0£&l

Jam l30"" TB^F^ ,p7ooor Fig. 9 — Fracture surface of lap welded commercial purity alu Fig. 10 — Fracture surface of butt welded high purity aluminum. minum. (A) 45% deformation, (B) 58% deformation. Reduced53% (A) 130% deformation, (B) 360% deformation. Reduced 55%

306-s I SEPTEMBER 1975 AA which probably contaminate the sur N N AZ AA, (2) face as a result of the trapped air at R = 2- and ~A~ R + 1 the interface. 1 At 1 Z i High stacking fault energy would be which is in the same form as that for AA expected to facilitate extrusion of the R butt welding, Eq. (1). Now, if welding metal through cracks because of easy Ac i is assumed to occur whenever clean cross-slip of glide dislocations. metal surfaces come into contact, Except for gold, aluminum, which and the true fraction al metallic area then the fractional welded area, f, has high stacking fault energy (~110 revealed is glive n by: should be given by: ergs/cm2), smallest hardness ratio (Hmetai/Hoxide) and possesses a com pletely brittle oxide film, was the easiest metal to weld. Although gold has low stacking fault energy, it is easier to weld than aluminum be cause the first requirement of weld ing is satisfied everywhere due to the complete absence of the oxide. It is only necessary for deformation to in crease the area of contact.

Theoretical Calculation of the Weld Strength Assume that the oxide film is com pletely brittle, so the metallic area re vealed, which could be welded, is equal to the amount of surface exten sion that occurs after the surfaces are in intimate enough contact to ex clude oxygen. In butt welding the experimentally measured extension R is given by

A - Ao R = Ao where A0 is the original cross-sec tional area and A is the instanta neous area at extension R. The true fractional metallic area revealed at a certain extention R is then: AA R A R + 1 (D I In case of lap welding, if we divide the surface area A0 into small areas A, (i = 1, 2, 3, . . .N) then A0 = NAi and upon applying a deforming stress, the area A; extends to be A, + AAj where 1000X Ai (x, +x,)(y, + y,)-x,y, Fig. 11 Fracture surface of lap welded high purity copper, 64% deformation. Reduced 33% A, XIVI Ax, and Ay, are the extensions in the x and y directions respectively and therefore:

AA; Ax ^Vi +

From the first law of plasticity:

AXi -Wt Az, + =0

Therefore:

AAi Azi

(minus sign means that the strain is compressive)

The total measured reduction Fig. 12 — Fracture surface of lap welded high purity silver, 75% deformation. Reduced 50%

WELDING RESEARCH SUPPLEMENT! 307-s (3) —) + 1 ' where C is a proportionality constant. If there is a complete matching be tween the metallic areas over the in terface, f has its maximum value given by Eq. (1) or (2), i.e.,

R + 1

but if there is no correlation, f should be given by Eq. (3) if the strength of the weld is Sw and the strength of the metal is Sm, then Jems "1000X SW **S*el^* »l = fn •f—y F/g. 13 — Fracture surface of lap welded high purity gold. (A) 3% deformation, (B) 10% Sm V R + 1/ (4) deformation. Reduced 50% where 1 < n < 2. It would be expected that f cannot attain this maximum value because flow of the metal will be restricted at the edges of the oxide gaps.

Experimental values of (f) were ob- ained from tensile tests, where

true ultimate tensile strength of the weld fexp — true ultimate tensile strength of the metal

The average fractional metallic areas revealed were estimated from the fractographs. The true welded area cannot be estimated directly be cause of the deformation of the weld ed region which occurs during me chanical testing. The fractional weld ed area

R R + 1

was calculated from the estimated Fig. 14 — Macrophotograph tor the surface ot butt welded aluminum specimen just before fractional revealed metallic area failure in tension. Reduced 42%

0.6 I I I I A A o_ i r A r O—-— A^. CT~ 0.4 — ^^o° - -

0.2 - - y - — Calculated o High Purity Aluminum Calculated A 6061 T6 Aluminum Alloy O High Purity Aluminum A 606I T6 Alluminum Alloy

^/ I I I I Deformation (R) I 2 3 Fig. 15 — Fractional revealed metallic Deformotion (R) areas estimated from fractographs vs deformation in butt welding of aluminum Fig. 16 Weld strength vs deformation of butt welded aluminum

308-s j SbHIbMBEH 1975 0.4 1 1 1 1 1 0.15 III 1 1 — Calculated O High Purity Aluminum | 0.3 - - A Commercial Purity Aluminum A 0.10 A

S 0.2- /o 5 0.05 - /o — Calculated •2 0.1 - O High Purity Aluminum A Commercial Purity Aluminum 0 -\ 1 1 1 1 1 1 1 1 1 0 0.1 0,2 0.3 0.4 0.5 06 0.1 0.2 0.3 0.4 0.5 06 Oeformation (R) Deformation (R) Fig. 18 — Weld strength vs deformation of lap welded aluminum Fig. 17 — Fractional revealed metallic areas estimated from frac tographs vs deformation in lap welding of aluminum

R + 1

Figures 15 to 18 show that: 0.30 (a) The experimental values of the fractional metallic areas revealed are in agreement with those calculated theoretically from R/(R+1). (b) The measured strength of the weld is in agreement with that calcu lated theoretically from Eq. (3) with 0.25 — n = 2 using the measured welded area. (c) The proportionality constant C in Eq. (3) is almost constant and inde pendent of deformation. It has a value of the order of 0.7-0.8. (d) The experimental values of the 0.20 weld strength are comparable to 0.2 0.3 05 those calculated on the basis of un- Deformation (R) correlated cracking of the oxide layers. Fig. 19 Weld strength vs deformation of lap welded gold It may be concluded from these results that cracks in the oxide layers are uncorrelated and that the weld tact at which the stress is highly con two oxide layers. strength, as a fraction of the strength centrated may be considered as 3. Factors that affect the relative of the metal, attained after a given nuclei of the weld which grow with difficulty of pressure welding are: (a) deformation is determined by the deformation, and as the number of stacking fault energy of the metal, (b) fractional welded area at that defor these nuclei increases, welding relative hardness of the metal and its mation according to the relation: becomes easier. It was also shown oxide film and the mechanical proper that C is almost independent of the ties of the oxide, and (c) surface deformation. roughness prior to welding. f = C —Y It is proposed that C depends on 4. The weld strength attained after R + 1/ the method of preparation of the sur a given deformation is determined by With reference to gold, the formula face, highly rough surface gives high the fractional area of clean metal sur which gives the revealed fractional value of C and vice versa. Therefore, face that has been brought into con metallic area while assuming the one of the principal functions of wire tact at that deformation. presence of a brittle oxide film, is not brushing is to increase the param 5. Some relative shear displace useful. For gold, the area of clean eter C. ment at the points where clean metal metal in contact will depend only on surfaces come into contact is neces the initial roughness of the surface sary for welding. and on the deformation of the high Conclusions spots as the pressure increases. 1. Metal to oxide bonding does not Acknowledgment Figure 19 shows the variation of fexp contribute significantly to weld obtained from mechanical tests with strength. This work was done under the auspices deformation. 2. Initiation of welding of a given of the U.S. Atomic Energy Commission. The Parameter C. It was shown that metal is determined by cracking of some relative shear displacement at the oxide films which permits contact References the points of contact of oxide-free to gradually develop between clean 1. Tylecote, R. F., The Solid Phase metal is required for welding and that metal surfaces. Localization of slip Welding of Metals, St. Martin's Press, an initial rough surface facilitates this into heavy slip bands may be the 1968, pp. 38, 57, 201. displacement. These points of con- cause of uncorrelated cracking of the 2. Tylecote, R. F., British Welding Jour-

WELDING RESEARCH SUPPLEMENT! 309-s nal, Vol. 1, 117, 1954. 32, 5, May, 1953, Res. Suppl., pp. 209-s to Study of Micromechanisms of Fracturing 3. Vaidyanath, L. R. and Milner, D. R., 222-s. Processes," in Fracture Toughness British Welding Journal, Vol. 7, 1, 1960. 8. McEwan, K. M. B. and Milner, D. R., Testing and Its Applications, STP 381, 4. Donelan, J. A., British Welding Jour British Welding Journal, Vol. 9, 406, 1962. ASTM, Philadelphia, p. 210. nal, Vol. 6, 5, 1959. 9. Barrett, C. S., Acta Met. 1, 2, 1953. 12. Telelman, A. S. and McEvily, A. J. 5. Bowden, F. P. and Tabor, D., Struc 10. Vaidyanath, L. R., Nicholas, M. G. Jr., "Fracture of Metals," in Fracture, An ture and Properties of Solid Surfaces, and Milner, D. R., British Welding Journal, Advanced Treatise, Liebowitz, H., ed., Gomer, R. and Smith, C. S., eds. Vol. 6, 13, 1959. Academic Press, 1969, p. 156. 6. Semenov, A. P., Wear 4, 1, 1961. 11. Beachem, C. D. and Pelloux, R. M. 13. Agers, B. M. and Singer, A. R., Brit 7. Parks, J. M., Welding Journal, Vol. N., "Electron Fractography, A Tool for the ish Welding Journal, Vol. 11, 313, 1964.

WRC Bulletin No. 187 Sept. 1 973

"High-Temperature Brazing"

by H. E. Pattee

This paper, prepared for the Interpretive Reports Committee of the Welding Re search Council, is a comprehensive state-of-the-art review. Details are presented on protective atmospheres, heating methods and equipment, and brazing proce dures and filler metals for the high-temperature brazing of stainless steels, nickel base alloys, superalloys, and reactive and refractory metals. Also included are an extensive list of references and a bibliography. The price of WRC Bulletin 187 is $5.00 per copy. Orders should be sent to the Welding Research Council, 345 East 47th Street, New York, N.Y. 10017.

WRC Bulletin No. 197 August 1974

"A Review of Underclad Cracking in Pressure-Vessel Components"

by A. G. Vinckier and A. W. Pense

This report is a summary of data obtained by the PVRC Task Group on Under clad Cracking from the open technical literature and privately sponsored re search programs on the topic of underclad cracking, that is, cracking underneath weld cladding in pressure-vessel components. The purpose of the review was to determine what factors contribute to this condition, and to outline means by which it could be either alleviated or eliminated. In the course of the review, a substantial data bank was created on the manufacture, heat treatment, and clad ding of heavy-section pressure-vessel steels for nuclear service. Publication of this report was sponsored by the Pressure-Vessel Research Committee of the Welding Research Council. The price of WRC Bulletin 197 is $5.50. Orders should be sent to the Welding Research Council, 345 E: 47th St., New York, N.Y. 10017.

310-s I SEPTEMBER 1975