Volume 08 - Mechanical Testing and Evaluation Part 14 pdf

ASTM D 1062 specifies reporting the test results as force required, per unit width, to initiate failure in the specimen, while in fracture mechanics, the results are given as Gc with uni

All peel tests have the common characteristic that failure propagates from an initially debonded area They also generally involve large displacements/deformations For these and other reasons, linear elastic stress analysis is often not well suited to peel tests The stresses and strains in the peel configuration are complex and seldom well understood Test results are generally not given in terms of stress but rather as force per unit length required to peel the specimen It is, therefore, generally difficult to compare the results from a peel test with those from other testing methods

Because of the large deformations involved in peel tests, the analysis of such geometries is very difficult except under certain simplifying assumptions (Ref 3, 4, 6, 9, and 10) Some very interesting and informative observations can be made on the basis of simplifying assumptions and approximations Indeed, considerable useful work has been completed using peel tests The informative work of Gardon (Ref 10) and Kaelble (Ref 11) is noteworthy The polymer research group at The University of Akron, under the direction of Professor A Gent, has been particularly adroit in applying peel techniques and the concepts of fracture mechanics (see the section “Adhesive Fracture Mechanics Tests” in this article) to obtain critical information and insight into the behavior of adhesive joints (Ref 12, 13) The peel specimen is, in principle, a very versatile geometry for obtaining adhesive fracture energy because various combinations of mode I and mode II loadings can be applied by varying the peel angle (Ref 3) The stress analyses of Adams and Crocombe (Ref 14) have provided additional insight into the peeling mechanisms They examined the stress distributions in peel specimens using elastic large-displacement, finite-element analysis techniques

References cited in this section

3 G.P Anderson, S.J Bennett, and K.L DeVries, Analysis and Testing of Adhesive Bonds, Academic

Press, 1977

4 A.J Kinlock, Adhesion and Adhesives, Chapman and Hall, 1987

6 K.L Mittal, Adhesive Joints, Plenum Press, 1984

7 Adhesives, Annual Book of ASTM Standards, Vol 15.06, ASTM (updated annually)

9 G.P Anderson and K.L DeVries, Predicting Strength of Adhesive Joints from Test Results, Int J Fract., Vol 39, 1989, p191–200

10 J.L Gardon, Peel Adhesion, I Some Phenomenological Aspects of the Test, J Appl Polym Sci., Vol 7,

1963, p 654

11 D.H Kaelble, Theory and Analysis of Peel Adhesion: Mechanisms and Mechanics, Trans Soc Rheol.,

Vol 3, 1959, p 161

12 A.N Gent and G.R Hamed, Peel Mechanics, J Adhes., Vol 7, 1975, p 91

13 A.N Gent and G.R Hamed, J Appl Polym Sci., Vol 21, 1977, p 2817

14 R.D Adams and A Crocombe, J Adhes., Vol 12, 1981, p 127

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

Lap Shear Tests

The most popular test geometry for testing adhesive joints is the lap shear specimen Its appeal is probably based on the fact that it closely duplicates the geometry used in many practical joints These lap joints are popular for several reasons:

• They facilitate use of larger contact areas than, for example, a butt joint

• They are easier to make and align than butt joints

• The adhesive is not exposed to “direct” tensile stresses Direct tensile stresses are known to have deleterious effects on adhesives

Typical lap shear test specimens for which ASTM standards have been written are presented in Fig 3 (Ref 7) The specimens shown in this figure conform most closely to ASTM Standards D 1002, D 3163, D 3164, D

3165, and D 3528 for testing adhesives used to bond metals, plastics, and laminates These represent only a

small sampling of the more than two dozen standards in the Annual Book of ASTM Standards, Volume 15.06,

that relate to shear testing These other standards range from descriptions of block-type sample configurations for testing lumber and wood bonding in shear by compression loading, through descriptions of devices to simultaneously expose samples to lap shear stresses and extremes in temperature Still others describe apparatus for exposing lap joints to sustained loads (using springs) to measure long-term creep or time to failure

Fig 3 Typical lap shear geometries (a) ASTM D 1002, D 3163, and D 3164 (b) ASTM D 3165 (c) ASTM

D 3528 Source: Ref 7

The results from lap shear tests are generally reported as the force at failure divided by the bonded area (overlap area) Such values are listed in a number of reference books and manufacturers' literature for a wide variety of adhesives The reference book on types of adhesives (Ref 1) lists typical lap shear strength values for literally thousands of commercial adhesives Such tables of “shear strength” values are without doubt of considerable utility for comparison and other purposes However, their use also can lead to faulty expectations and

conceptions Otherwise knowledgeable designers might logically assume from the tabulations that these average stress values could, in a straightforward manner, be used to design an adhesive joint

For example, the tabulated shear stress value for a given adhesive from an ASTM D 1002 test might be given as

3000 psi It might be assumed that this adhesive is to be used to bond two 25 mm (1 in.) wide by 3 mm (0.12 in.) thick 7075-T6 aluminum pieces together to carry a tensile load of 3200 lb with a safety factor of two First, the designer must ascertain whether the aluminum pieces can carry such a load Typically, 7075-T6 aluminum has a yield strength slightly in excess of 65 ksi for an allowable stress of 32.5 ksi The pieces in question would have an allowable load of 4000 lb, which is more than the 3200 lb required in the design The “straightforward” method to design the joint would be to assume that the allowable shear strength for the adhesive used in the joint would be 3000/2 = 1500 psi, suggesting that an overlap of 3200/1500 = 2.13 in would be sufficient to support the load This is, in fact, the approach taught by a variety of otherwise very good texts on material science and mechanical design However, doubling the length of a lap joint almost never doubles its load-carrying capacity, and the increased joint strength is usually much less than doubled The length of overlap recommended in ASTM D 1002 is 12.7 mm (0.50 in.) Typically, quadrupling the amount of overlap does not increase the load at failure by anywhere near a factor of four For reasons given in the next few paragraphs, it is likely that it is not even the value of the maximum shear stress that determines the failure of the “lap shear joint.” As this article reveals, joint failure is more likely determined by the value of secondary induced cleavage stresses

The stresses along the bond line of lap specimens are not constant The bond stress distribution is highly dependent on the thickness of the adherends and the adhesive as well as the length of overlap As a consequence, the load to initiate failure also varies markedly with both the adherend(s) and adhesive-bond thicknesses The failure load increases very nearly linearly with width of the overlap but increases in a very nonlinear manner with length of the overlap As the load is increased in a lap shear test, the debonding generally initiates at or near one of the bond terminations Elastic stress analysis generally indicates that the stresses are singular at these termination points Debond initiation in lap shear specimens can perhaps, therefore, be best characterized in terms of fracture mechanics parameters, which are discussed in the section

“Adhesive Fracture Mechanics Tests” in this article In addition, it has been demonstrated that for debonds after initiation, crack propagation is dominated by crack- opening mode displacements (mode I) For this reason and

reasons given in the next couple of paragraphs, the word shear in the test titles and generally reported in test

results may, therefore, be a misnomer

It has been known for many years that the shear stresses in the bond line of lap specimens are accompanied by tensile stresses Many analyses have been completed for lap shear geometries, almost all of which have clearly demonstrated the presence of induced tensile stresses in so-called lap shear specimens under load In 1938, Volkersen (Ref 15) obtained expressions for the stresses in a lap shear joint by considering the differential displacements of the adherends and neglecting bending This study was followed in 1944 by the now classical treatment of Goland and Reissner (Ref 16) who used standard beam theory and strength of materials concepts

to obtain expressions for the joint stresses Plantema (Ref 17) combined the results of these two earlier investigations to include shear effects in the system

Because the stress state of the lap shear joint is so complex and does not lend itself to closed-form solutions, it

is only logical that as numerical methods became available, researchers would apply them to analyze adhesive joints Wooley and Carver (Ref 18), for example, used finite-element methods to calculate the joint stresses They compared their results with the results obtained by Goland and Reissner and reported very good agreement Adams and Peppiatt (Ref 19) used a two- dimensional finite-element code to analyze the stresses in

a standard lap shear joint and also reported good agreement with Goland and Reissner These authors also investigated the effect of a spew (triangular adhesive fillet) on the calculated stresses A nonlinear finite-element analysis of the single lap joint was completed by Cooper and Sawyer (Ref 20) in 1979

Anderson and DeVries conducted a linear elastic stress analysis of a typical single lap joint (Ref 21) making use

of plane-strain finite- element computer programs using elements as small as 0.00025 cm (0.0001 in.) They considered steel (modulus of elasticity, 207 GPa; Poisson's ratio, 0.30) adherends of various thicknesses bonded with a 0.25 mm (0.01 in.) thick epoxy (modulus of elasticity, 2.76 GPa; Poisson's ratio, 0.34) The overlap region was taken as 13 mm (0.5 in.) long The results of these analyses are shown in Fig 4 Note that both the shear and tensile stresses are distributed very nonlinearly over the length of the bond region Reference 21 reports stresses resulting from other adherend thicknesses As the bond termini is approached, both shear and normal stresses appear to become singular Careful analysis in this region suggests that the local mode I stresses

(tensile or crack opening) are significantly higher than mode II stresses (shear) Perhaps even more importantly, the mode I energy release rate is greater than that for mode II From these results, it might be concluded that lap shear specimens fail by mode I crack growth Therefore, the failure of lap shear specimens is usually governed

by tensile stress rather than shear stresses This is true for double lap joints as well as single lap joints (Ref 22, 23)

Fig 4 Bond line tensile and shear stresses in lap shear specimen (adherend thickness = 1.6 mm, or 0.06 in.)

As noted, the end(s) of the overlap on bond termini on lap shear specimens are points of stress concentration and of large induced tensile stresses While this closely simulates many practical situations, some have suggested that for determination of intrinsic adhesive properties, it would be useful if these termini could be eliminated ASTM E 229 “Standard Test Method for Shear Strength and Shear Modulus of Structural Adhesives” is a test designed specifically for this purpose In this test, the adhesive is applied in the form of a thin annulus ring bonded between two relatively rigid adherends in circular disc form Torsion shear forces are applied to the adhesive through this circular specimen, which produces a peripherally uniform stress distribution The maximum stress in the adhesive at failure is taken to represent the shear strength of the adhesive By measuring the angle of twist experienced by the adhesive and having knowledge of sample geometry, it is possible to calculate the strain A stress- strain curve can then be established from which the adhesive's effective shear modulus can be determined

References cited in this section

1 Adhesives, Edition 6, D.A.T.A Digest International Plastics Selector, 1991

7 Adhesives, Annual Book of ASTM Standards, Vol 15.06, ASTM (updated annually)

15 O Volkersen, Die Nietraftverteilung in Zugbeanspruchten Nietverblendugen mit Knastaten

Laschenquerschntlen, Luftfahrt forsch., Vol 15, 1938, p 41

16 M Goland and E Reissner, The Stresses in Cemented Joints, J Appl Mech., Vol 11, 1944, p 17

17 J.J Plantema, “De Schuifspanning in eme Limjnaad,” Rep M1181, Nat Luchtvaart-laboratorium, Amsterdam, 1949

18 G.R Wooley and D.R Carver, J Aircr., Vol 8 (No 19), 1971, p 817

19 R.D Adams and N.A Peppiatt, Stress Analysis of Adhesive-Bonded Lap Joints, J Strain Anal., Vol 9

22 J.K Strozier, K.J Ninow, K.L DeVries, and G.P Anderson, Adhes Sci Rev., Vol 1, 1987, p 121

23 G.P Anderson, D.H Brinton, K.J Ninow, and K.L DeVries, A Fracture Mechanics Approach to

Predicting Bond Strength, Advances in Adhesively Bonded Joints, Proceedings of a Conference at the Winter Annual Meeting of ASME, 27 Nov-2 Dec 1988 (Chicago), S Mall, K.M Liechti, and J.K

Vinson, Eds., ASME, 1988, p 98–101

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

Tensile Tests

Generally, the idea of mechanical failure produces a vision of an object being pulled apart by tensile force As noted previously, most practical adhesive joints are designed to avoid (or at least reduce) direct tensile forces across the bond line Examples of such joints are lap joints and scarf joints It was also pointed out that for many joints, where it appears that the primary loading is shear, failure might be initiated by the induced secondary tensile stresses There are, therefore, reasons why an adhesive's or adhesive joint's tensile strength might be of interest Accordingly, the third most common type of adhesive joint strength test is the tensile test ASTM has also formalized this type of test

The geometries of several tensile tests for which there are specific ASTM test procedures are shown in Fig 5 (Ref 7) Some of these test geometries seem relatively simple; however, it has been demonstrated that the stresses along the bond line have a rather complex dependence on geometric factors and adhesive and adherent properties (adhesive thickness and its variation across the bonded surface, modulus, Poisson's ratio, and so on) (Ref 21)

Fig 5 Typical specimen geometries for testing the tensile strength of adhesive joints Source: Ref 7

It is almost always difficult to load tensile adhesion specimens in an axisymmetric manner, even if the sample itself is axisymmetric Nonaxisymmetric loads have been shown to reduce the bond failure load capability and

to cause large scatter in the resulting failure data Superficially, the geometry for standard tensile adhesion tests

is deceptively simple The result of the tensile adhesion test, as normally reported by experimentalists, is simply the failure load divided by the cross-sectional area of the adhesive (Ref 22) Such average stress at failure can

be very misleading Because of the differences in mechanical properties of the adhesive and adherend, the stresses may become singular at the bond edges when analyzed using linear elastic analysis (Ref 21, 23) Even

if the edge singularity is neglected, the stress field in the adhesive is very complex and nonuniform, with maximum values differing markedly from the average value (Ref 21, 23)

Some sense of the complex nature of the stresses can be obtained by visualizing a butt joint of a low modulus polymer (e.g., a rubber) between two steel cylinders As these are pulled apart, the rubber elongates much more readily than the steel Poisson's effect will cause a tendency for the rubber to contract laterally However, if it is tightly bound to the metal, it is restrained from contracting, and shear stresses are induced at the bond line Reference 9 provides the results of a finite element analysis that demonstrates how these stresses vary across the sample As noted, for an elastic analysis, both the shear and tensile stresses are singular (tending to infinity)

at the outer periphery

For the tensile specimen configurations considered to this point, the applied loading is intended to be axisymmetric There is another class of specimen in which the dominant stress is deliberately tensile but in which the loading is obviously “off center.” At least four ASTM standards describe so-called cleavage specimens and tests These tests are a logical preface to the next section in this article, “Adhesive Fracture Mechanics Tests” The reader familiar with cohesive fracture mechanics will see a similarity between the test specimen in ASTM D 1062 (Fig 6) and the compact tensile specimen commonly used in fracture mechanics testing ASTM D 1062 specifies reporting the test results as force required, per unit width, to initiate failure in

the specimen, while in fracture mechanics, the results are given as Gc with units of J/m2, which might be interpreted as the energy required to create a unit surface A knowledgeable and enterprising reader may want

to adapt the D 1062 specimen for obtaining fracture mechanics parameters ASTM D 3807, “Standard Test

Method for Strength Properties of Adhesives in Cleavage Peel by Tension Loading,” uses a different geometry

to measure the cleavage strength In this case, two 25.4 mm (1 in.) wide by 6.35 mm (0.25 in.) thick plastic strips 177 mm (7 in.) long are bonded over a length of 76 mm (3 in.) on one end, leaving the other ends free and separated by the thickness of the adhesive Approximately 25 mm (1 in.) from the end of each of these free segments, a “gripping wire” is attached as shown in Fig 7 During testing, these wires are clamped in the jaws

of a universal testing machine and the sample pulled to failure The results are reported as load per unit width (kg/m or lb/in.) Again, it would be possible to analyze this sample in terms of fracture mechanics, but it is unnecessary because, as the next section explains, this analysis is done in ASTM D 3433 for a very similar beam geometry

Fig 6 Specimen for testing the cleavage strength of metal-to-metal adhesive bonds (ASTM D 1062)

Fig 7 Specimen for testing cleavage peel (by tension loading) (ASTM D 3807)

ASTM D 5041 also makes use of a sample composed of two thin sheets bonded together over part of their length In this case, forcing a wedge (45° angle) between the unbonded portion of the sheets facilitates the separation The results are typically given as “failure initiation energy” or “failure propagation energy” (i.e., areas under the load deformation curve)

This latter test is similar to another test, formalized as ASTM D 3762, that has been found very useful for studying time-environmental effects on adhesive bonds This test is called by various names, but the authors prefer the name “Boeing Wedge Test” (Ref 24, 25) The test has been used by personnel at this and other aerospace companies to screen various adhesives, surface treatment, and so on for long-term loading at high temperatures and humidities For testing, two long, slender strips of candidate structural materials are first treated with the prescribed surface treatment(s) and bonded over part of their length with a candidate adhesive (Fig 8) As in the test described in the previous paragraph, the free ends are forced apart by a wedge The amount of separation by the wedge (determined by wedge thickness and depth of insertion) determines the value of the stresses in the adhesive These stresses can, of course, be adjusted and the values calculated from mechanics of material concepts When the wedge is in place, the sample is placed in an environmental chamber

At periodic time intervals, the length of the crack is measured, and a plot of crack length versus time is constructed The more satisfactory adhesives and/or surface treatments are those for which the crack is arrested

or grows very slowly While the environmental chamber typically contains hot, humid air, there is no reason why other environmental agents cannot be studied by the same method, including immersion in liquids

Fig 8 Boeing wedge test (ASTM D 3762) (a) Test specimen (b) Typical crack propagation behavior at 49

°C (120 °F) and 100% relative humidity a, distance from load point to initial crack tip; Δa, growth

during exposure Source: Ref 49

References cited in this section

7 Adhesives, Annual Book of ASTM Standards, Vol 15.06, ASTM (updated annually)

9 G.P Anderson and K.L DeVries, Predicting Strength of Adhesive Joints from Test Results, Int J Fract., Vol 39, 1989, p191–200

21 G.P Anderson and K.L DeVries, Analysis of Standard Bond-Strength Tests, Treatise on Adhesion and Adhesives, Vol 6, R.L Patrick, K.L DeVries, and G.P Andersen, Ed., Marcel Dekker, 1989

22 J.K Strozier, K.J Ninow, K.L DeVries, and G.P Anderson, Adhes Sci Rev., Vol 1, 1987, p 121

23 G.P Anderson, D.H Brinton, K.J Ninow, and K.L DeVries, A Fracture Mechanics Approach to

Predicting Bond Strength, Advances in Adhesively Bonded Joints, Proceedings of a Conference at the Winter Annual Meeting of ASME, 27 Nov-2 Dec 1988 (Chicago), S Mall, K.M Liechti, and J.K

Vinson, Eds., ASME, 1988, p 98–101

24 V.L Hein and F Erodogan, Stress Singularities in a Two-Material Wedge, Int J Fract.,Vol 7, 1971, p

317

25 J.A Marceau, Y Moji, and J.C McMillan, A Wedge Test for Evaluating Adhesive Bonded Surface

Durability, 21st SAMPE Symposium, Vol 21, 6–8 April 1976

49 J.C McMillan, Developments in Adhesives in Engineering, 2nd ed., Applied Science, London, 1981, p

243

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

Adhesive Fracture Mechanics Tests

Fracture mechanics originated with the pioneering efforts of A.A Griffith in the early 1920s The field remained relatively dormant until the late 1940s when it was developed into a very effective and valuable design tool to describe and predict “cohesive” crack growth Interested readers are referred to a number of

excellent texts on fracture mechanics (e.g., Ref 26 and Fatigue and Fracture, Volume 19 of the ASM Handbook)

In the 1960s and 1970s, researchers began exploring the use of the concepts of fracture mechanics in adhesive joint analysis as reviewed in Ref 3 These methods have the potential to use the results from a test joint to predict the strength of other joints with different geometries

In a common fracture mechanics approach (including Griffith's papers), the conditions for failure are calculated

by equating the energy lost from the strain field as a “crack” grows to the energy consumed in creating the new

crack surface This energy per unit area, Gc, determined from standard tests, is called by various names, including the Griffith fracture energy, the specific fracture energy, the fracture toughness, or the energy release rate

In 1975, ASTM Committee D-14 adopted a test configuration and testing method with fracture mechanics ramifications based on the pioneering efforts of Mostovy and Ripling (Ref 27, 28) The method is described in ASTM D 3433 “Standard Test Method for Fracture Strength in Cleavage of Adhesives in Bonded Joints.” Figure 9 shows the shape and dimensions for one specimen type recommended for use in this standard The specimen is composed of two “beams” adhesively bonded over much of their length as shown Testing is accomplished by pulling the specimen apart by means of pins passing through the holes shown near the sample's left end This adhesive sample configuration and loading to failure gives rise to the sample's nickname,

“split-cantilever beam.” Another recommended geometry in ASTM D 3433 is similar except the adherends are not tapered

Fig 9 Specimen for the contoured double-cantilever-beam test (ASTM D 3433)

It should now be clear that the stress distribution in adhesive joints is generally complex Furthermore, the details of this distribution are highly dependent on specific details of the joint system The maximum stresses in the bond almost always differ markedly from the average value, and elastic analyses often exhibit mathematical singularities at geometric or material discontinuities From these observations, it should be clear that the use of the conventionally reported results from most tests (i.e., values of the average stress at failure) would be of little use in designing joints that differ in any significant detail from the sample test configuration

For the resolution of this problem, the concepts of fracture mechanics have much to offer One of the more popular and graphically appealing approaches to fracture mechanics views the joint as a system in which failure (often considered as the growth of a crack) of a material (or joint) requires the stresses at the crack tip to be sufficient to break bonds and an energy balance It is hypothesized that even if the stresses are very large (often theoretically infinite), a crack can grow only if sufficient energy is released from the stress field to account for the energy required to create the new crack (or adhesive debond) surface as the fractured region enlarges The specific value of this energy (J/m2, or in · lbf/in.2, of crack area) for the adhesive bonding problem uses the

same basic titles as given previously but prefaced with the term adhesive Hence, adhesive fracture toughness

might be used to distinguish adhesive failure from tests of cohesive fracture The word adhesion is dropped from the comparable term when cohesive failure is being considered The cohesive and adhesive embodiments

of fracture mechanics both involve a stress-strain analysis and an energy balance

The analytical methods of fracture mechanics (both cohesive and adhesive) are described in Ref 3 and 25 These are not repeated here other than a few comments on the concepts and a brief outline of a numerical approach that can be applied where analytical solutions are tedious or impossible Inherent in fracture mechanics is the concept that natural cracks or other stress risers exist in materials and that final failure of an object often initiates at such points For a crack (or region of debond) situated in an adhesive layer, modern computation techniques are available (most notably, finite element methods) that facilitate the computation of stresses and strains throughout a body, even if analytical solutions may not be possible The stresses and strains are calculated throughout the entire adhesive system (adhesive and all adherends), including the effects of a

crack in the bond These can then be used to calculate the strain energy, U1, stored in the body for the particular

crack size, A1 Next, the hypothetical crack is allowed to grow to a slightly larger area, A2, and the preceding

process is repeated to determine the strain energy, U2 This approach to fracture mechanics assumes that at critical crack growth conditions, the energy loss from the stress-strain field goes into the formation of the new

fracture energy The quantity ΔU/ΔA is called the energy release rate, where ΔU = U2 - U1 and ΔA = A2 - A1

The so-called critical energy release rate (ΔU/ΔA)crit is that value of the energy release rate that will cause the crack to grow Loads that result in energy release rates lower than this critical value will not cause failure to proceed from the given crack, while loads that produce energy release rates greater than this value will cause it

to accelerate This critical energy release rate value is equivalent to the adhesive fracture energy, or work of adhesion, previously noted While the model just described is conceptually useful, computer engineers have devised other convenient ways of computing the energy required to “create” the new surface, such as the crack closure method (Ref 29, 30)

It is hoped that this simple model of fracture mechanics will help the reader who is unfamiliar with fracture mechanics to visualize the concepts of fracture mechanics The molecular mechanisms responsible for the fracture energy or fracture toughness are not completely understood They generally involve more than simply the energy required to rupture a plane of molecular bonds In fact, for most practical adhesives, the energy to rupture these bonds is a small but essential fraction of the total energy The total energy includes energy that is lost because of viscous, plastic, and other dissipation mechanisms at the tip of the crack As a result, linear elastic stress analyses are inexact

While fracture mechanics has found extensive use in cohesive failure considerations, its use for analyzing failure of adhesive systems is more recent There has, however, been a significant amount of research and development in the adhesive fracture mechanics area To review it all, even superficially, would take more space than is allocated for this article A small sampling of publications in this extensive and rich area of research is listed as Ref 12, 13, 26, 27, 28, and 31, 32, 33, 34, 35, 36, 37, 38, 39, 40, 41, 42, 43, 44, 45, 46, 47,

48, 49, 50, 51, 52, 53, 54, 55, 56, 57, 58 Not only is this listing incomplete, but also many of the researchers listed have scores of other publications It is hoped that the one or two listed for each investigator will provide the reader with a starting point from which more details can be found from reference cross listings, searching of citation indexes, abstracting services, and so on These investigators have treated such subjects as theory; mode dependence, effects of shape, thickness, and other geometric dependence; plasticity and other nonlinearities; numerical methods; testing techniques; different adhesive types; rate and temperature effects; fatigue; and failure of composites, as well as a wide variety of other factors and considerations in adhesion

Modern finite element or other numerical methods have no difficulty in treating nonlinear behavior Physical understanding of material behavior at such levels is lacking, and effective use of the capabilities of such computer codes depends, to a large extent, on the experimental determination of these properties For many problems, it has become conventional to lump all dissipative effects together into the fracture energy and not be overly concerned with separating this quantity into its individual energy-absorbing components Another

fracture mechanics approach, called the J-integral, has some advantages in treating nonlinear as well as elastic

behavior (Ref 51, 52, 59, and 60)

It was noted previously that most adhesive systems are not linearly elastic up to the failure point Nevertheless, researchers have shown that elastic analyses of many systems can be very informative and useful Several adhesive systems are sufficiently linear so that it is possible to lump the plastic deformation and other energy dissipative mechanisms at the crack tip into the adhesive fracture energy (critical energy release rate) term There has recently been some significant success in explaining many aspects of adhesive performance and

predicting the strength of a bond from tests on other, quite different, joints by using linear elastic fracture mechanics

As noted, in principle, fracture mechanics lends itself to using test results from one test in the design of other joints that have significantly different geometries A number of adhesive geometries have been proposed to measure fracture toughness in addition to the split-cantilever beam, but to date, it is the only one formalized by ASTM (Ref 3, 4, 6, 12, 36, 47, and 48) A recent paper by the authors (Ref 61) has demonstrated how such factors as end rotation (at the cantilever point assumed rigidly fixed in the original ASTM analysis) shear, and presence of the adhesive and its thickness (also neglected in the original analysis) affect the energy release rate

It is shown that inclusion of these effects can dramatically affect the results and greatly reduce test scatter Furthermore, this paper demonstrates how fracture mechanics may be used to predict the locus of adhesive crack growth To accomplish this, various crack paths were assumed, and using finite element methods, the energy release rate calculated for each path

References cited in this section

3 G.P Anderson, S.J Bennett, and K.L DeVries, Analysis and Testing of Adhesive Bonds, Academic

Press, 1977

4 A.J Kinlock, Adhesion and Adhesives, Chapman and Hall, 1987

6 K.L Mittal, Adhesive Joints, Plenum Press, 1984

12 A.N Gent and G.R Hamed, Peel Mechanics, J Adhes., Vol 7, 1975, p 91

13 A.N Gent and G.R Hamed, J Appl Polym Sci., Vol 21, 1977, p 2817

25 J.A Marceau, Y Moji, and J.C McMillan, A Wedge Test for Evaluating Adhesive Bonded Surface

Durability, 21st SAMPE Symposium, Vol 21, 6–8 April 1976

26 D Broek, The Practical Use of Fracture Mechanics, Kluver Acad Press, Dordrect, NL, 1989

27 S Mostovoy and E.J Ripling, J Appl Polym Sci., Vol 15, 1971, p 661

28 S Mostovoy and E.J Ripling, J Adhes Sci Technol., Vol 9B, 1975, p 513

29 E.F Rybicki and M.F Kanninen, Eng.Fract Mech., Vol 9, 1974, p 921

30 G.P Anderson and K.L DeVries, J Adhes., Vol 23, 1987, p 289

31 E.H Andrews, T.A Khan, and H.A Majid, J Mater Sci., Vol 20, 1985, p 3621

32 E.H Andrews, H.A Majid, and N.A Lockington, J Mater Sci., Vol 19, 1984, p 73

33 D.W Aubrey and M Sherriff, J Polym Sci Polym Chem Ed., Vol 18, 1980, p 2597

34 W.D Bascom and J Oroshnik, J Mater Sci., Vol 10, 1978, p 1411

35 W.D Bascom and D.L Hunston, Fracture of Epoxy and Elastomer-Modified Epoxy Polymers, Treatise

on Adhesion and Adhesives, Vol 6, R L Patrick, K.L DeVries, and G.P Anderson, Ed., Marcel

Dekker, 1988, p 123

36 H.F Brinson, J.P Wightman, and T.C Ward, Adhesives Science Review 1, VPI Press, 1987

37 J.D Burton, W.B Jones, and M.L Williams, Trans Soc Rheol, Vol 15, 1971, p 39

38 G Danneberg, J Appl Polym Sci., Vol 5, 1961, p 125

39 F Erdogan, Eng Fract Mech., Vol 4, 1972, p 811

40 T.R Guess, R.E Allred, and F.P Gerstle, J Test Eval., Vol 5 (No 2), 1977, p 84

41 G.R Hamed, Energy Conservation During Peel Testing, Treatise on Adhesion and Adhesives, Vol 6, R

L Patrick, K L DeVries, and G P Anderson, Eds., Marcel Dekker, 1988, p 233

42 R.W Hertzberg and J.A Manson, Fatigue of Engineering Plastics, Academy, 1980

43 G.R Irwin, Fracture Mechanics Applied to Adhesive Systems, Treatise on Adhesion and Adhesives, Vol

1, R L Patrick, Ed., Marcel Dekker, 1966, p 233

44 W.S Johnson, J Test Eval., Vol 15 (No 6), 1987, p 303

45 D.H Kaelble, Physical Chemistry of Adhesion, Wiley-Interscience, 1971

46 H.H Kaush, Polymer Fracture, Springer-Verlag, Berlin, 1978

47 W.G Knauss and K.M Liechti, Interfacial Crack Growth and Its Relation to Crack Front Profiles, ACS Organic Coatings and Applied Polymer Science Proceedings, Vol 47, American Chemical Society,

50 D.R Mulville, D.L Hunston, and P.W Mast, J Eng Mater Technol., Vol 100, 1978, p 25

51 J.R Rice and G.C Sih, J Appl Mech., Vol 32, 1965, p 418

52 E.F Rybicki and M.F Kanninen, Eng Fract Mech., Vol 9, 1974, p 921

53 G.B Sinclair, Int J Fract., Vol 16, 1980, p 111

54 J.D Venables, D.K McNamara, J.M Chen, T.S Sun, and R.L Hopping, Appl Surf Sci., Vol 3, 1979, p

88

55 S.S Wang, J.F Mandell, and F.J McGarry, Int J Fract., Vol 14, 1978, p 39

56 J.G Williams, Fracture Mechanics of Polymers, Ellis Horwood, Chichester, 1984

57 M.L Williams, J Adhes., Vol 5, 1973, p 81

58 R.J Young, in Structural Adhesives: Developments in Resins and Primers, A.J Kinlock, Ed., Applied

Science, London, 1986, p 163

59 J.R Rice, A Path Independent Integral and the Approximate Analysis of Strain Concentration by

Notches and Cracks, J Appl Mach., 1979, p 379–386h

60 J.W Hutchinson and P.C Paris, Stability of J-Controlled Crack Growth, STP 668, ASTM, 1979, p 37–

64

61 K.L DeVries and P.R Borgmeier, Fracture Mechanics Analyses of the Behavior of Adhesion Test

Specimens, Mittal Festschrift, W.J Van Ooij and H.R Anderson, Jr., Ed., 1998, p 615–640

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

Conclusions

The adhesive researcher or technologist has many standard test methods from which to choose These techniques are designed with various goals and objectives in mind Many of these methods are useful for the purposes of comparing different adhesives and substrates, investigating the effects of different loading, investigating chemical or physical attacks on adhesives, exploring aging phenomena, determining the effects of radiation and moisture combined with sustained loading on adhesive properties, and so on On the other hand, care should always be exercised not to use the test results for purposes for which they are not well suited Results from many of the adhesive strength tests are conventionally reported as the failure force divided by the bond area Such average stress at failure results cannot, in general, be consistently and reliably used to predict failure of other joints that differ even slightly from the test geometry Fracture mechanics approaches, on the other hand, show promise and have been used to predict the strength of joints that differ considerably from the reference joint ASTM D 3433 and Ref 27 and 28 describe a standard adhesive fracture mechanics joint in the form of a tapered double-cantilever beam The specimen dimensions are shown in Fig 9 It is important to note, however, that fracture mechanics is not limited to this or any other specific testing geometry In principle, any

geometry for which the described energy balance (or alternatively, calculation of the stress intensity factor,

J-integral, and so on) can be accomplished might be used as an adhesive test

Sometimes, circumstances dictate the use of a nonstandard test geometry For example, a few years ago, the authors were given the problem of measuring the quality of natural barnacle adhesive The barnacle dictated the exact form of the joint between the barnacle's shell and the plastic sheets that were placed in the ocean This form did not lend itself to tensile, lap shear, or split-cantilever testing It was, however, possible to predrill holes in the plate and to fill these holes with dental waxes that were solid and hard at the ocean temperatures near San Francisco, CA, where the barnacle growth experiments were conducted The wax was later easily removed at a moderately elevated temperature The base of the barnacle covering this hole was thereby exposed and could be tested by application of fluid pressure, thus forming a blister Measurement of the pressurization at failure allowed the determination of the adhesive fracture energy (Ref 3, 62)

Once the adhesive fracture energy is determined by testing, fracture mechanics points the way that it, along with a knowledge of the flaw size and a stress-strain analysis of the joint, can be used to predict the performance of other joints Modern computational techniques greatly facilitate the application of these methods

Finally, it is noted that the stresses, strains, fracture energy, and other such parameters used in the adhesive analysis depend on loading rate, mode of stress at the crack tip, temperature, environment, and other factors Development of means for incorporating these parameters into joint design has been, and continues to be, an area of active research Such concepts and methodology can be found in the references cited previously in this section

References cited in this section

3 G.P Anderson, S.J Bennett, and K.L DeVries, Analysis and Testing of Adhesive Bonds, Academic

Press, 1977

27 S Mostovoy and E.J Ripling, J Appl Polym Sci., Vol 15, 1971, p 661

28 S Mostovoy and E.J Ripling, J Adhes Sci Technol., Vol 9B, 1975, p 513

62 R.R Despain, R.D Luntz, K.L DeVries, and M.L Williams, J Dental Res., Vol 52, 1973, p 742

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

Acknowledgments

Major portions of the authors' research has been supported by The National Science Foundation, most recently under grant No CMS-9522743

Testing of Adhesive Joints

K.L DeVries and Paul Borgmeier, University of Utah

References

1 Adhesives, Edition 6, D.A.T.A Digest International Plastics Selector, 1991

2 R.L Patrick, Ed., Treatise on Adhesion and Adhesives, Vol 1–6, Marcel Dekker, 1966–1988

3 G.P Anderson, S.J Bennett, and K.L DeVries, Analysis and Testing of Adhesive Bonds, Academic

Press, 1977

4 A.J Kinlock, Adhesion and Adhesives, Chapman and Hall, 1987

5 A Pizzi and K.L Mittal, Ed., Handbook of Adhesive Technology, Marcel Dekker, 1994

6 K.L Mittal, Adhesive Joints, Plenum Press, 1984

7 Adhesives, Annual Book of ASTM Standards, Vol 15.06, ASTM (updated annually)

8 E.P Plueddemann, Silane Coupling Agents, Plenum Press, 1982

9 G.P Anderson and K.L DeVries, Predicting Strength of Adhesive Joints from Test Results, Int J Fract., Vol 39, 1989, p191–200

10 J.L Gardon, Peel Adhesion, I Some Phenomenological Aspects of the Test, J Appl Polym Sci., Vol 7,

1963, p 654

11 D.H Kaelble, Theory and Analysis of Peel Adhesion: Mechanisms and Mechanics, Trans Soc Rheol.,

Vol 3, 1959, p 161

12 A.N Gent and G.R Hamed, Peel Mechanics, J Adhes., Vol 7, 1975, p 91

13 A.N Gent and G.R Hamed, J Appl Polym Sci., Vol 21, 1977, p 2817

14 R.D Adams and A Crocombe, J Adhes., Vol 12, 1981, p 127

15 O Volkersen, Die Nietraftverteilung in Zugbeanspruchten Nietverblendugen mit Knastaten

Laschenquerschntlen, Luftfahrt forsch., Vol 15, 1938, p 41

16 M Goland and E Reissner, The Stresses in Cemented Joints, J Appl Mech., Vol 11, 1944, p 17

17 J.J Plantema, “De Schuifspanning in eme Limjnaad,” Rep M1181, Nat Luchtvaart-laboratorium, Amsterdam, 1949

18 G.R Wooley and D.R Carver, J Aircr., Vol 8 (No 19), 1971, p 817

19 R.D Adams and N.A Peppiatt, Stress Analysis of Adhesive-Bonded Lap Joints, J Strain Anal., Vol 9

22 J.K Strozier, K.J Ninow, K.L DeVries, and G.P Anderson, Adhes Sci Rev., Vol 1, 1987, p 121

23 G.P Anderson, D.H Brinton, K.J Ninow, and K.L DeVries, A Fracture Mechanics Approach to

Predicting Bond Strength, Advances in Adhesively Bonded Joints, Proceedings of a Conference at the Winter Annual Meeting of ASME, 27 Nov-2 Dec 1988 (Chicago), S Mall, K.M Liechti, and J.K

Vinson, Eds., ASME, 1988, p 98–101

24 V.L Hein and F Erodogan, Stress Singularities in a Two-Material Wedge, Int J Fract.,Vol 7, 1971, p

317

25 J.A Marceau, Y Moji, and J.C McMillan, A Wedge Test for Evaluating Adhesive Bonded Surface

Durability, 21st SAMPE Symposium, Vol 21, 6–8 April 1976

26 D Broek, The Practical Use of Fracture Mechanics, Kluver Acad Press, Dordrect, NL, 1989

27 S Mostovoy and E.J Ripling, J Appl Polym Sci., Vol 15, 1971, p 661

28 S Mostovoy and E.J Ripling, J Adhes Sci Technol., Vol 9B, 1975, p 513

29 E.F Rybicki and M.F Kanninen, Eng.Fract Mech., Vol 9, 1974, p 921

30 G.P Anderson and K.L DeVries, J Adhes., Vol 23, 1987, p 289

31 E.H Andrews, T.A Khan, and H.A Majid, J Mater Sci., Vol 20, 1985, p 3621

32 E.H Andrews, H.A Majid, and N.A Lockington, J Mater Sci., Vol 19, 1984, p 73

33 D.W Aubrey and M Sherriff, J Polym Sci Polym Chem Ed., Vol 18, 1980, p 2597

34 W.D Bascom and J Oroshnik, J Mater Sci., Vol 10, 1978, p 1411

35 W.D Bascom and D.L Hunston, Fracture of Epoxy and Elastomer-Modified Epoxy Polymers, Treatise

on Adhesion and Adhesives, Vol 6, R L Patrick, K.L DeVries, and G.P Anderson, Ed., Marcel

Dekker, 1988, p 123

36 H.F Brinson, J.P Wightman, and T.C Ward, Adhesives Science Review 1, VPI Press, 1987

37 J.D Burton, W.B Jones, and M.L Williams, Trans Soc Rheol, Vol 15, 1971, p 39

38 G Danneberg, J Appl Polym Sci., Vol 5, 1961, p 125

39 F Erdogan, Eng Fract Mech., Vol 4, 1972, p 811

40 T.R Guess, R.E Allred, and F.P Gerstle, J Test Eval., Vol 5 (No 2), 1977, p 84

41 G.R Hamed, Energy Conservation During Peel Testing, Treatise on Adhesion and Adhesives, Vol 6, R

L Patrick, K L DeVries, and G P Anderson, Eds., Marcel Dekker, 1988, p 233

42 R.W Hertzberg and J.A Manson, Fatigue of Engineering Plastics, Academy, 1980

43 G.R Irwin, Fracture Mechanics Applied to Adhesive Systems, Treatise on Adhesion and Adhesives, Vol

1, R L Patrick, Ed., Marcel Dekker, 1966, p 233

44 W.S Johnson, J Test Eval., Vol 15 (No 6), 1987, p 303

45 D.H Kaelble, Physical Chemistry of Adhesion, Wiley-Interscience, 1971

46 H.H Kaush, Polymer Fracture, Springer-Verlag, Berlin, 1978

47 W.G Knauss and K.M Liechti, Interfacial Crack Growth and Its Relation to Crack Front Profiles, ACS Organic Coatings and Applied Polymer Science Proceedings, Vol 47, American Chemical Society,

50 D.R Mulville, D.L Hunston, and P.W Mast, J Eng Mater Technol., Vol 100, 1978, p 25

51 J.R Rice and G.C Sih, J Appl Mech., Vol 32, 1965, p 418

52 E.F Rybicki and M.F Kanninen, Eng Fract Mech., Vol 9, 1974, p 921

53 G.B Sinclair, Int J Fract., Vol 16, 1980, p 111

54 J.D Venables, D.K McNamara, J.M Chen, T.S Sun, and R.L Hopping, Appl Surf Sci., Vol 3, 1979, p

88

55 S.S Wang, J.F Mandell, and F.J McGarry, Int J Fract., Vol 14, 1978, p 39

56 J.G Williams, Fracture Mechanics of Polymers, Ellis Horwood, Chichester, 1984

57 M.L Williams, J Adhes., Vol 5, 1973, p 81

58 R.J Young, in Structural Adhesives: Developments in Resins and Primers, A.J Kinlock, Ed., Applied

Science, London, 1986, p 163

59 J.R Rice, A Path Independent Integral and the Approximate Analysis of Strain Concentration by

Notches and Cracks, J Appl Mach., 1979, p 379–386h

60 J.W Hutchinson and P.C Paris, Stability of J-Controlled Crack Growth, STP 668, ASTM, 1979, p 37–

64

61 K.L DeVries and P.R Borgmeier, Fracture Mechanics Analyses of the Behavior of Adhesion Test

Specimens, Mittal Festschrift, W.J Van Ooij and H.R Anderson, Jr., Ed., 1998, p 615–640

62 R.R Despain, R.D Luntz, K.L DeVries, and M.L Williams, J Dental Res., Vol 52, 1973, p 742

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Introduction

IN WELDED STRUCTURES, the welds typically have a mechanical purpose Loads must be carried across the weld joint Standard mechanical tests have been devised to demonstrate that not only the base metals but also the entire welded joint can fulfill this mechanical purpose (Ref 1, 2) This article primarily discusses standard test methods that can be applied to many types of welds These include tension, bending, impact, and toughness testing

Residual stress measurement techniques and weldability testing also are discussed Residual stress can be imposed by the welding itself, as well as by cutting and forming processes The presence of high-tension or compression residual stresses can affect the ability of the welded structure to carry the mechanical loading Cracking due to welding can also affect the load-carrying capacity of welded joints, and weldability testing techniques that combine welding and mechanical loading to test the resistance to cracking are available Many other testing techniques can be applied to weld joints and welded structures Fatigue and creep are both important areas where mechanical tests on welded joints have indicated properties different from those of the base metal Testing of welded structure properties can also be done on structures that more closely model the service structure than the standard specimens described below

References cited in this section

1 “Standard Welding Terms and Definitions,” AWS 3.0, American Welding Society

2 L.P Connor, Ed., Welding Handbook, 8th ed., Vol 1, American Welding Society, 1991

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Reasons to Measure Properties of Welds

The mechanical properties of welded joints, the properties related to stress and strain, are most often measured

to show that such a weld and other similar welds will serve their purpose under loading More rarely, several welds are compared to see which welding techniques, processes, or chemistries provide the best combination of mechanical properties

Four stages in the qualification process for the weld joint can use standard mechanical test methods The weld metal can be chosen based on the mechanical properties from standard tests The base metal, in some situations, may also need to be qualified to demonstrate that its mechanical properties will not be substantially degraded

by welding Once base metal and weld metal are chosen, the other weld process parameters, such as weld shape

or heat input, can be qualified by standard mechanical testing Finally, after the weld is made, it may require qualification by mechanical testing

Each of the four qualification stages requires different types of tests and different approaches These approaches are described in the next four sections

Weld Material Qualification

Weld metals are qualified by making welds that pass mechanical property tests on the weld metal itself Such tests can be used to qualify filler materials, such as welding wire or electrodes Mechanical tests for such qualification are described in the individual specifications of the filler metals, such as those in American Welding Society (AWS) Specification A5.1 (Ref 3) and others of the AWS A5 series

Mechanical property tests applied to weld qualification are designed to determine a small number of standard values to check whether the weld metal passes or fails This approach will not reveal the entire range of properties that the weld metal can achieve For instance, weld metal toughness in AWS A5.1 is measured by a Charpy test specimen taken from the weld centerline Other locations, which may have different toughnesses, are not checked

Base Material Qualification for Welded Service

The heat of welding will modify the structure and properties of the region of the base metal adjacent to the weld

in the heat-affected zone (HAZ) To prevent this modification in properties from causing service failures, some standards require that a sample of the base material be tested after undergoing a representative heat treatment American Petroleum Institute standard (API) RP2Z is an example of a standard that requires base-material qualification for welded service (Ref 4) Multipass welding provides the heat treatment The welding parameters are chosen to represent the most severe heat treatment of the base material that may occur during fabrication The parameter of interest is the fracture toughness, commonly measured by crack tip opening displacement (CTOD) Particular regions of the HAZ are the most likely to show low toughness (the local brittle zones, or LBZs), so there is also a requirement that the crack sample the required portion of that kind of microstructure

Weld Procedure Qualification

Weldment properties are dependent not only on the materials used to make the joint but also on the other parameters of the welding process Weld metal properties may be modified by the admixture of base material melted by the heat of welding The region where the base metal was only partially melted, at the fusion line, may have local mechanical properties differing from those of the neighboring weld metal and HAZ regions In addition, the welding processes and procedures may induce specific imperfections, such as slag inclusions, blowholes, or cracks

Because the issues described cannot be resolved by either weld metal or base material tests alone, test procedures that use specimens containing weld metal, base metal, and HAZ are used to determine mechanical properties of welded joints in weldment procedure qualifications

Weld procedure qualification tests may be less quantitative than weld metal qualification tests and often provide only a “yes-or-no” answer They are often capable of being completed in a shop floor environment rather than a testing laboratory with calibrated equipment

While many of the weld procedure qualification tests in wide use are discussed subsequently in this article, several are not, because they are not properly mechanical tests; that is, neither a loading parameter, such as stress, or a displacement parameter, such as strain, is measured This group includes visual examination for surface flaws and the breaking of fillet welds to examine the weld root

Weld Service Assessment

Assessment of existing welds to determine if they meet the needs of future service may require that material properties be obtained from representative weldments In some cases, new weldments can be made with the same materials and process parameters, so that sections from these weldments can be tested to find representative properties for the existing welds Generally, neither the original materials nor full information is available to allow replication of existing weldments Instead, a sample must be taken from the existing weldments

Taking a sample requires trading the advantage of obtaining mechanical property data for the disadvantage of damaging the existing structure Choices often are made that limit the damage to the existing structure by limiting the amount of material to be tested Tests such as macrohardness or microhardness, which damage a small surface volume, may be appropriate Smaller-scale test specimens may also be used, for instance, subsize Charpy specimens Alternatively, specimens can be taken from regions where subsequent repair is easiest or from an area that is being removed as part of a modification

References cited in this section

3 “Specification for Carbon Steel Electrodes for Shielded Metal Arc Welding,” ANSI/AWS A5.1-91, American National Standards Institute/American Welding Society, Miami, 1991

4 “Recommended Practice for Preproduction Qualification for Steel Plates for Offshore Structures,” API 2Z, 2nd ed (includes November 1998 addenda), American Petroleum Institute, Washington DC, 1998

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Mechanical Testing for Weldment Properties

Tensile Strength and Ductility of Weldments

Testing for mechanical properties of strength and ductility for welded joints is somewhat more complicated than it is for base metal, because these properties vary across the weld metal, the adjacent HAZ, and the base metal Several different tests may be used or combined to assess the strength of the overall welded joints Tensile testing is widely used to measure the strength and ductility of the weld metal alone Tensile testing of welds in place, with weld metal, HAZ, and base metal, allows an overall strength to be determined but usually cannot provide the strengths of the individual parts of the weldment

Tensile tests of welds can also measure elastic modulus However, except in rare cases of dissimilar metal joining, the elastic modulus is not sensitive to the differences between weld, HAZ, and base metal So,

measurement during weld tensile tests is not usually required Also, most tensile testing procedures for weld joints cannot be relied upon to provide accurate values of elastic modulus The specific procedures for testing of elastic modulus distributed by ASTM should be used if required (Ref 5)

Testing of Weld Material Deposited weld metal can be tested for the mechanical properties of strength and ductility using the same test methods used for base metals (Ref 6, 7) However, a sufficient volume of deposited weld metal is required to remove a test specimen made entirely of weld metal Often, arc welds are long only in one direction (the longitudinal direction), while the through-thickness and cross-weld directions are much smaller This encourages all-weld-metal tensile test specimens to be removed with the long direction of the specimen corresponding to the longitudinal direction of the weld Such longitudinal tensile test specimens are standard for all-weld-metal tests

All-weld-metal tests are most commonly done on specimens with round cross section The diameter of the specimen may need to be reduced from that used for base metal so that the specimen can be taken entirely from weld metal Rectangular cross-section specimens also are used occasionally

Ultimate tensile strength, yield strength (usually based either on yield point or a specified offset), elongation, and reduction of area are all commonly recorded

While the specimen surface should be smooth, without deep machining marks, imperfections within the gage length due to welding should not be removed This requirement may increase the variability of results within a group of similar specimens

If the data required are for a class of weld material such as an electrode lot, the material can be taken from specimens that reduce the possibility of dilution of base metal into the weld, such as a built-up weld pad If the data required are for a particular weldment, the geometry as well as the welding process and procedure should model those of the weldment as closely as possible Some modifications of the weldment may be allowed, such

as increasing the root opening by 6 mm ( in.) or buttering the groove faces with the weld metal to be tested The surface of the tested section, in the gage length, is recommended to be 3 mm ( in.) or more from the fusion line

Testing of Welds in Place When the weld metal extends over only part of the tested gage length, tensile tests can be performed similar to those performed on the round and rectangular specimen tests of weld metal The nonuniformity of deformation or stresses of the weld, HAZ, and base metal combination limits the information normally recorded

For transverse tests, ultimate strength and the location of fracture are the only commonly recorded parameters, because strength, elongation, and reduction in area will all be affected by the constraint of the adjacent differing materials If the weld is undermatched, the yield strength tends to be higher than it is for an all-weld-metal specimen, while the elongation over the gage length and reduction in area are smaller If the weld yield strength exceeds that of the base material, that is, it is overmatched, the failure tends to occur not in the adjacent HAZ, but in the base material closer to the end of the gage length, because of the constraint provided by the high-strength weld metal

Local strain measurements, such as those made by strain gages, can add useful information to the results of transverse testing The local strain information can be correlated to the load and displacement information to allow local strengths to be determined

For longitudinal tests, the strain will be nearly uniform across the weld metal, HAZ, and base metal Differences in response to the applied strain may result in stresses varying across the cross section Only ultimate strength is commonly measured

Testing standards may need to be varied for some specific geometries For instance, girth welded tubes of less than 75 mm (3 in.) diameter are commonly tested in the form of tubes with central plugs at the grips The weld

is placed at the center of the gage length between the grips The additional constraint induced by the hoop direction continuity tends to increase the measured strengths and decrease the measured ductilities for tube welds tested in this manner compared to a similar joint between flat sheets

Shear Testing of Fillet Welds

Shear strength tests for fillet welds are described in AWS B4.0 for two orientations of fillet welds (transverse to the tension loading and longitudinal to the tension loading) (Ref 8, 9) The transverse specimen is a double lap specimen with loaded fillet welds, as shown in Fig 1 The longitudinal specimen is a combination of two

lapped shear plates that are tack welded back to back, as shown in Fig 2 These geometries are chosen to avoid rotation during loading The longitudinal specimen requires machining of grooves after the fillet welds are made The base plate is cut under the center of the lapped plate The lapped plate is cut near each end so that each length of weld connecting the base plate to the lapped plate across the gap in the base plate is 38 mm (1 in.)

Fig 1 Transverse fillet weld shear test specimen Source: Ref 9

Fig 2 Longitudinal fillet weld shear test specimen Source: Ref 8

Fillet-weld strength tests are sensitive to surface contour of the welds and to the condition of the weld root Excessive gaps between the lapped plates should be avoided, because these tend to magnify stresses at the weld root The specimens are also sensitive to underbead cracking and undercut

Fillet size is most accurately measured after failure in the test The stress is calculated based on assuming uniform stress across the entire weld throat

Bending Strength and Ductility

Bend tests are commonly used to evaluate the acceptability of weld procedures for providing sound welds (Ref 10) They allow rapid determination of strength and ductility on a specimen substantially simpler than the standard tensile specimen Bend tests tend to provide vivid demonstrations of difference between welds with surface or near-surface flaws and welds without flaws adjacent to the convex surface of the bend Bend tests are further described in the article “Bend Testing” in this Volume

Bending ductility can be calculated by determining the radius of the outer surface of the bend specimen at the completion of the test This ductility is usually smaller than that measured in a uniaxial tensile test The bend ductility is localized at the outer surface of the specimen, and the constraint is more severe because of the shear stresses generated through the thickness of the bend specimen

The thickness of the specimens and the size of the plunger or mandrel determine the outer surface ductility requirement Table 1 provides a summary of the radii of plungers or mandrels and the maximum test specimen thicknesses for several groups of materials as required by the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code (Ref 11)

Table 1 Bend test geometry for testing based on material thickness and material type

in With >20% elongation

(3 + )t in Alloy steels with <20% elongation

High strength Al alloys

(8 + )t in 4000 series Al alloy

Al alloy welded with 4000 series electrodes

Cu base alloys with Al and <20% elongation

Root, Side, and Face Bends The primary geometries for bend test specimens place the butt weld so that the bending stress is transverse to the weld axis Different areas of the welds reach the highest bending stress in transverse root bends, transverse face bends, and transverse side bends Root bends put the weld-root side of the tested butt weld on the convex side of the bend specimen Face bends put the weld-cap side of the tested butt weld on the convex side of the bend specimen Side bends put a cross section of the weld on the convex side of the bend specimen

Longitudinal bend tests may sometimes be used to replace transverse tests, particularly when the strengths of the regions within the specimen differ greatly However, longitudinal side bends are not possible since the weld cross section does not include the longitudinal direction Longitudinal root bends, with the convex side of the bent specimen on the weld-root side, and longitudinal face bends, with the convex side of the bent specimen on the weld-cap side, can both be made and tested

Longitudinal tests are less likely than transverse tests to fail from flaws that are long in the same direction as the weld

Wrap-Around Bend Testing While the plunger-type bend fixtures are by far the most widely used for guided bend testing, some circumstances require a fixture that creates a different distribution of strain The most common is a wrap-around testing fixture Both the plunger type and the wrap-around type force the material into a specified radius However, the wrap-around fixture moves the points of bending load application around

a mandrel rather than using a fixed location for the central force, as shown in Fig 3

Fig 3 Wrap-around bend testing Source: Ref 9

The wrap-around fixture is most commonly used for welds that have significant mismatch between base metal strengths, between base metal and weld metal strengths, or where the HAZ strength differs greatly from the weld or base metal If such welds are tested in a plunger-type fixture, the strain can be concentrated into the lower-strength material, leaving a sharper bend in that material and much less bending in the higher-strength material The wrap-around fixture forces a more uniform strain into the materials because the loading point in the center of the specimen moves across the weld

Wrap-around test fixtures are commonly used for aluminum alloys where the strength of the weld metal and HAZ may be chosen to be lower than the base metal For instance, 6061-T6 aluminum base metal that is welded with 4043 electrodes has a minimum yield strength of 240 MPa (35 ksi) in the base metal but only 100 MPa (15 ksi) for cross-weld tensile specimens (Ref 12) Wrap-around testing can limit the localization of the strain at the weld

Hardness

Hardness testing of welded joints is widely used as a rapid measurement of mechanical properties across the varying microstructures of the welded region It allows local regions and individual microstructures to be compared for strength, because strength is correlated to hardness A further discussion of hardness testing can

be found in the Section “Hardness Testing” in this Volume

Hardness has been primarily related to the tensile strength rather than to the yield strength or the ductility Standard conversion charts are available for conversion of one hardness measurement to another and from hardness to tensile strength measurement Such converted information should be used with caution, because the

variation of weld microstructure may cause the average hardness to correspond to values that cannot be obtained in larger scale specimens

Macrohardness testing of welds requires preparation of a small region of the surface The major techniques are Brinell testing, which uses a spherical indenter, and Rockwell testing, which uses a diamond penetrator or a sphere The Brinell indentation is typically 2 to 6 mm in diameter while the Rockwell indentation is much smaller but still is visible, unaided Rockwell methods use several different loads for different hardness scales,

so it is possible for a weld to require different hardness scales for different regions

Macrohardness testing results can be limited by the microstructural gradients around the welds A result of 240

HB may represent a hardness for one uniform microstructure or an average over the regions deformed by the indenter Welds and HAZs often have gradients of microstructure and chemistry that can cause variations in hardness across the indentation Interpretation of the hardness from the impression may be made more difficult

if there is a large gradient in the hardness of the material under the indenter This can result in noncircular Brinell impressions and Rockwell tests with the deepest point not under the deepest point of the indenter Microindentation hardness testing using an indenter requires an even smaller region of the surface to be used than macrohardness testing, but the surface preparation requirements are more stringent Thus the Knoop and Vickers microindentation hardness tests are primarily applied to ground and polished cross sections or to ground, polished, and etched cross sections Microindentation hardness traverses are often used to determine the variation of hardness within the weld, across the fusion line, and across the HAZ

Impact Toughness

Several methods are available for measuring the material resistance to starting and growing cracks that can be applied to welded joints This section discusses test methods that cause a crack to grow from a notch under the rapid load of an impact Methods that use sharp crack tips and thus can apply the loading more slowly are discussed in the next section on fracture toughness

Charpy The Charpy V-notch impact test is the most common measurement method for fracture toughness of welded joints Specifications for the test are given in ASTM E 23 (Ref 13) and AWS B4.0 The test uses a pendulum hammer to rapidly fracture a notched bar with dimensions of 55 mm by 10 mm by 10 mm (2.165 in

by 0.394 in by 0.394 in.)

Several measures of toughness can be obtained from a Charpy test Absorbed energy, measured in ft · lbf or joules, is the most commonly reported, but the percent shear fracture and the lateral expansion in inches or millimeters are also sometimes reported Greater toughness material will have higher values of each of these three parameters Occasionally, percent fibrous fracture, which is 100% minus the percent shear fracture, is reported

Many metals, including carbon and alloy steels, have toughnesses that vary strongly with temperature So tests

on welded joints are often conducted at several temperatures, and the absorbed energy or other parameter is plotted as a function of temperature Material specifications and weld qualifications that include Charpy V-notch testing normally require a minimum absorbed energy at a particular temperature In this case, testing is routinely conducted only at the temperature of interest

The choice of minimum absorbed energy and test temperature are often varied between standards or within a standard, based on service conditions For instance, welded joints on bridges to be used in cold climates are qualified to lower temperatures than those used in warm climates

The absorbed energy in a Charpy V-notch test includes both the energy to start the crack from the 0.25 mm (0.010 in.) radius notch and the energy to propagate the crack across the Charpy specimen For many cases, including constructional steels, these two parts are of comparable magnitude In fact, the popularity of the Charpy V-notch test was originally based on its ability to predict both crack initiation and crack arrest in ship steel plates This means that both the metal microstructure at the notch tip and through the specimen thickness contribute to the reported toughness For welded joints with heterogeneous microstructures, the position of the notch tip will be important in determining the measured absorbed energy The absorbed energy, however, will also depend on the microstructures through which the fracture passes

The dependence of Charpy impact test results on microstructure for many metals causes weld joints with heterogeneous microstructures to have a range of Charpy values depending on specimen orientation in the weld and notch position Often weld centerline values are reported or compared with standards Sometimes the HAZ

is tested at a particular location, such as 1 mm from the fusion line These tests cannot determine a toughness

appropriate to all microstructures in the weld or HAZ Additional tests of a greater variety of specimens may reveal zones of lower toughness, such as unrefined columnar weld metal or coarse-grained HAZ, or zones or higher toughness, such as reheated weld metal or fine-grained HAZ

Subsize Charpy specimens are sometimes taken from thin material or areas where the geometry prevents a size specimen Only one dimension is reduced, the distance from the notched face to the unnotched surface opposite Reductions of this dimension can be from 10 mm (0.394 in.) to 7.5 mm (0.296 in.), called three-quarter size; to 5 mm (0.197 in.), called half-size; and to 2.5 mm (0.099 in.), called quarter-size These are the most common reductions Reduced thickness Charpy tests can be used to test the toughness of the root or cap regions of fillet welds

full-Charpy toughness test specimens can be taken from welded joints in several orientations These orientations can

be given two-letter designations to show the orientation The first letter is the direction normal to the crack plane (the long direction of the Charpy specimen), while the second letter is the direction in which the crack will propagate The letter designations are L, longitudinal direction; T, long transverse direction (the weld width direction); and S, short transverse direction (the through thickness direction) The letter designations are shown for compact tension specimens in Fig 4 Care should be taken that the orientation letters describe the weld area, because different combinations of these letters may apply to the same orientation of specimen in base metal For instance, in a girth weld in a pipe, the long direction of the weld is the hoop direction of the pipe, not the longitudinal or axial direction of the pipe

Fig 4 Orientations of toughness specimens in relation to welds L, longitudinal direction; T, long transverse direction (weld width direction); S, short transverse direction (weld thickness direction) In the two-letter code for specimen designation, the first letter designates the direction normal to the crack plane, and the second letter designates the expected direction of the crack plane Source: Ref 9

Nil-Ductility Temperature Drop weight tests use a notched weld bead as the starting point for a crack The test determines the ability of the base metal to arrest the crack running from an overlay of brittle weld metal The test results do not describe the properties of the overlay weld metal, so the overlay weld metal is standardized

A brittle hard-facing alloy with good surface adhesion is used as the crack starter

ASTM E 208 describes the test procedure (Ref 14) A standard weight is dropped onto test specimens at different temperatures The lowest temperature without full-section fracture is determined as the nil-ductility temperature (NDT)

Weld metal can be tested for crack arrest by placing the notched weld overlay across a machined butt weld of the weld metal of interest The notch is typically placed so its long direction is above the longitudinal direction

of the butt weld

Fracture Toughness

Fracture toughness testing of welded joints introduces several complications to standard fracture toughness measurement as described in the Section “Impact Toughness Testing and Fracture Mechanics” in this Volume The weld and adjacent HAZ will have heterogeneous microstructures that can have widely varying strength and toughness In addition, welding residual stresses may be retained

Fracture initiation testing, using slow loading and a crack tip sharpened by precracking, allows determination of only the crack initiation portion of the fracture toughness Charpy testing determines a combination of crack initiation and arrest properties Fracture initiation testing of welded joints thus is even more sensitive to the local microstructure around the tip of the precrack than Charpy tests are to the microstructure at the notch Weld heterogeneity also causes welds to be particularly sensitive to the rules for data interpretation for fracture initiation tests Fatigue precracks are less likely to be straight in a heterogeneous material Tests may have multiple events of crack initiation and arrest in local areas (“pop-ins”) Varying strengths may cause validity criteria based on yield strength and specimen size to give ambiguous results Each of these issues must be accounted for in a test protocol appropriate to welds

Fracture toughness specimens can be taken from welded joints in several orientations These orientations can be given two-letter designations to show the orientation The first letter is the direction normal to the crack plane, while the second letter is the direction in which the crack will propagate The letter choices are L, longitudinal direction; T, long transverse direction (the weld width direction); and S, short transverse direction (the through thickness direction) The letter designations are shown for compact tension specimens in Fig 4, but the same designations can be used for other shapes of test specimen Care should be taken that the orientation letters describe the weld area, because different combinations of these letters may apply to the same orientation of specimen in base metal For instance, in a girth weld in a pipe, the long direction of the weld is the hoop direction of the pipe, not the longitudinal or axial direction of the pipe

Fracture toughness testing to measure the fracture resistance of weld HAZs presents particular problems, because several different microstructures can cluster within the HAZ, based on the different histories of heating from the welding in different locations A fracture toughness measured from the HAZ is likely to be affected both by the properties of the several HAZ microstructures that the crack tip passes through and by the properties of the adjacent weld metal and base material

Fracture initiation toughness is measured most commonly using the stress intensity factor, K; the J-integral, J;

or the crack tip opening displacement (CTOD) All of these are regularly applied to welded joints CTOD measurements are somewhat more commonly specified in welded regions than in base metal, because the test

was originally developed for welded joints Conversions between K, J, and CTOD are routinely performed but

should be noted, because the correlations have limited precision

Test Modifications for Welds Fracture toughness testing of welds may require precise positioning of the notch

to test the microstructure of interest Testing of welds may also require modification of the test specimen Two methods of modification, local compression and gull-winging (Ref 15), are described in the following section Local compression counteracts the effects of welding residual stress Gull-winging allows a curved piece with a weld to be tested as a full-thickness specimen

Local Compression The sharp crack tip needed for a fracture toughness test to determine initiation toughness is commonly provided by fatigue precracking at low levels of stress In welds, fatigue precracking may produce a crack front that is not straight across the specimen This is particularly likely when the specimen is thick and was removed from a weld with as-welded residual stresses The crack tip of the fatigue precrack can deviate from the average straight line so much that the fracture toughness test results are invalidated because the crack depth is poorly described by a single average value

Determining valid or invalid precracks is part of the standard to which the testing is done, such as ASTM E

1290 for CTOD testing (Ref 16) or ASTM E 1737 for J-integral testing (Ref 17)

A local compression treatment can be applied before precracking to avoid excessive deviation of the precrack from a straight line Compression is applied in circular regions around the points where the notch tip reaches the surface, as shown in Fig 5 Local compression can be applied to the most common fracture test specimen shapes, including compact tension specimens and single-edge notched bend bars

Fig 5 Local compression of fracture specimens before fatigue precracking, P, load Source: Ref 15

Gull-Winging Some geometries of weld may be difficult to test for fracture toughness because the shape of the base metal around the weld limits the thickness or area of any standard geometry specimen, such as a compact tension specimen or a three-point bend bar Welds in curved shapes, such as spheres and cylinders, can be particularly difficult

As shown in Fig 6(a), a flat specimen taken for fracture toughness testing from a curved part may be limited in dimension by both the inside surface and the outside surface of the curved part This could force the use of small specimens preferentially determining the toughness near the inside or concave surface of the weld Gull-winging is a mechanical bending process that allows the full thickness of the curved part to be used for fracture testing

Fig 6 Gull-winging of single-edge notched-bar weld fracture toughness specimen (a) Small scale of specimen that can be obtained from a longitudinal weld in a cylinder compared to full thickness (b) Gull-winged specimen at full cylinder thickness ready to be loaded Source: Ref 15

Full-thickness transverse tests using a modified three-point bend bar geometry may be used f or geometries where the weld is straight but the curvature is transverse to the weld by gull-winging The gull wings are introduced by plastically bending the base material away from the weld These bends must allow the three locations of loading to be the same as if the specimen were flat, as shown in Fig 6(b) During gull-winging, the area adjacent to the weld must be supported to prevent plastic deformation of the weld or the area directly adjacent to it

Gull-winged specimens should have limited deviation from the flat centerline of the equivalent flat specimen Greater deviation may allow plastic deformation during testing at the point of greatest deviation

References cited in this section

5 “Standard Test Method for Young's Modulus, Tangent Modulus, and Chord Modulus,” E 111, Annual Book of ASTM Standards, Vol 3.01, ASTM, 1999

6 “Standard Methods of Tension Testing of Metallic Materials,” E 8, Annual Book of ASTM Standards,

Vol 3.01, ASTM, 1999

7 “Standard Methods of Tension Testing Wrought and Cast Aluminum, and Magnesium Alloy Products,”

B 557, Annual Book of ASTM Standards, Vol 3.01, ASTM, 1999

8 B4 Committee on Mechanical Testing of Welds, Mechanical Testing of Welds, Part 1: Summary of

Tension Testing of Welds, Weld J., Jan 1981, p 33–37

9 “Standard Methods for Mechanical Testing of Welds,” ANSI/AWS B4.0-98, American National Standards Institute/American Welding Society, Miami, 1998

10 B4 Committee on Mechanical Testing of Welds, Mechanical Testing of Welds, Part 2: Bend Testing of

Welds A Summary, Weld J., Feb 1981, p 34–37

11 ASME Boiler and Pressure Vessel Code, Section IX, American Society of Mechanical Engineers

12 Aluminum Design Manual, The Aluminum Association, 1994

13 “Standard Test Methods for Notched Bar Impact Testing of Metallic Materials,” E 23, Annual Book of ASTM Standards, Vol 3.01, ASTM

14 “Standard Method for Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature

of Ferritic Steels,” E 208, Annual Book of ASTM Standards, Vol 3.01, ASTM

15 M.G Dawes, H.G Pisarski, and S J Squirrell, Fracture Mechanics Tests on Welded Joints, Nonlinear Fracture Mechanics Vol II: Elastic-Plastic Fracture, STP 995, J.D Landes, A Saxena, and J.G

Merkle, Ed., ASTM, 1989, p 191–213

16 “Standard Test Method for Crack-Tip Opening Displacement (CTOD) Fracture Toughness

Measurement,” E 1290, Annual Book of ASTM Standards, Vol 3.01, ASTM

17 “Standard Test Method for J-Integral Characterization of Fracture Toughness,” E 1737, Annual Book of ASTM Standards, Vol 3.01, ASTM

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Residual Stress Measurement

Residual stress measurement on welded joints must take account of the characteristics of the welded joint Welds usually have gradients of residual stress through the weld area and have these residual stress gradients in the same regions as gradients in microstructure These features may limit the amount of information that can be obtained from some techniques Some residual stress measurement techniques, for instance, work by comparing the distances between atoms in the crystal structure from an unstressed area to the stressed area of interest If no

unstressed area is available with the same crystal structure as the weld, the zero stress level for that weld region would be uncertain

Two general types of measurements for residual stress in welds are most common: locally destructive techniques and nondestructive techniques The locally destructive techniques include hole drilling, chip machining, groove machining, and block sectioning These measure changes in strain as a new surface is created and determine residual stresses most sensitively around the area where the machining took place The nondestructive techniques measure the local strain of the material by inputting a physical change, receiving a signal based upon that change and the residual stress, and then decoding the part of the signal induced by the residual stress These techniques include x-ray diffraction, neutron diffraction, Barkhausen noise analysis, and ultrasonic propagation analysis

Both groups of techniques do not measure stress directly Both measure strain Both groups have limitations for measuring rapidly varying residual stress fields because they need to sample a volume of material to get sufficient signal to determine the residual stresses with precision

Sectioning Methods

One extreme way of imagining a residual stress measurement of a welded region is that many tiny pieces of the weldment could all be separated simultaneously from neighboring pieces and allowed to change shape to reach zero residual stress The shape change in each tiny piece could be measured and that change correlated using the elastic properties of the material to the residual stress originally in each piece of the weldment

Sectioning methods can come close to this extreme case, but each cut cannot happen simultaneously So analysis techniques for determining residual stress by multiple cuts need to include calculations of the effects of previous cuts on the residual stress to find the original stress rather than the stress determined for an individual cut

Block sectioning and block layering and sectioning are significantly more destructive than hole drilling The part for which the residual stresses will be determined is cut into sections and layers while surface strain gages are monitored These techniques can provide distributions of residual stress in multiple dimensions

Hole Drilling and Similar Local Measurements

Hole drilling techniques, often called center-hole drilling or blind-hole drilling, measure the change in strain on the adjacent surface as the residual stress field is disturbed by the machining of a hole from that surface (Ref 18) The hole depth is generally between 1 and 2 mm (0.04 and 0.08 in.), although several organizations make residual stress measurements as a function of hole depth as the hole is drilled

The drilling techniques should cause as little as possible additional stress around the hole, so techniques are commonly used that avoid contact of the walls of the hole as the bottom is drilled Air turbine and air abrasion systems can provide holes with close dimensional tolerance and little machining-induced surface stress on the hole Specialized strain gage rosettes are available for measurement, both with the hole placed in the center of the rosette and with the rosette on one side of the hole The rosette on one side of the hole is used for cases where the surface is obstructed on one side of the measurement position, as may happen adjacent to a weld toe The measured strain must be converted to a local residual stress in the area where the hole was drilled This conversion can be done most simply by assuming a uniform residual stress distribution Only components of the residual stress that are parallel to the surface are measured by this technique, giving two directions of axial stress and one of shear stress

Chip machining has been used similarly to hole drilling to determine near-surface residual stresses In this case,

a small region of the surface is removed with the strain gage rosette on the surface of the chip rather than attached to the base material The advantage is that the small chip can be assumed to reach essentially zero residual stress, so the change of strain measured on the chip can be directly correlated to the stress in the chip region However, the chip must be larger than the most common size of drilled hole to carry the strain gage The chip machining method will average the strains over a larger volume of material

Other shapes of local machining can be used, including deep holes and partial thickness slits (Ref 19) These methods can measure residual strains either with surface strain gages or with measurements of post-cut displacement As with hole drilling, only part of the residual stress tensor is obtained by these methods, but the direction of cutting can be oriented to find the residual stresses of most interest

Nondestructive Techniques

Nondestructive techniques for measuring residual strains in welded joints use a variety of inputs, including rays, neutrons, magnetic signals, and ultrasonic waves Each of these inputs can induce different outputs, depending on the residual strains in the welded joint area, but also depending on other parameters, such as the crystal structure or grain size Nondestructive techniques thus are best applied where a standard for comparison without residual strains but with the same microstructure is available The comparison of test material with an unstrained standard may be easiest in base metal outside the HAZ and much more difficult for HAZ or weld metal (Ref 20)

x-X-ray diffraction determines the residual strains by measuring the average distances between atoms x-X-rays, being limited in surface penetration in metallic materials, can be used for detection of residual strains only within approximately 0.5 mm (0.02 in.) of the surface on which they impinge The surfaces of welds may be more difficult locations than smooth surfaces of base material for measurement of residual strains The rough surface of the weld cap and notches at weld toes may cause some of the area of interest for residual strain measurement to be hidden from residual stress detection by x-rays

Neutron diffraction uses an input that is much more difficult and costly to generate than x-rays but has the advantage of penetrating much further into metallic materials Neutrons can be used for detection of residual strains in steel thicknesses beyond 25 mm (1 in.) Average values of the residual strains are determined in a volume approximately 1 mm3 (6 × 10-4 in.3) The power of the source, the efficiency of the detector, and the time of exposure all influence the volume required for residual stress detection Neutron diffraction is easiest when the grain size is significantly smaller than the detection volume and the grains are oriented randomly Large grains and highly oriented microstructures can eliminate the diffracted neuron signal

Barkhausen noise analysis uses an external varying magnetic field as input and monitors the magnetic response

of the area of interest The response comes from the jumps in magnetization as magnetic domain walls move within the metal This response is a surface response, so the surface condition of the part is important

Ultrasonic propagation analysis inputs ultrasonic waves and then monitors the response of the time taken for the waves to travel through the metal and reach the detector Since the residual stress distribution can change the speed of propagation all along the path of the ultrasound, the resulting effect is summed over the entire path Multiple path analysis can provide local results

Barkhausen noise and ultrasonic propagation analysis both are limited by the microstructural variations around welds The combination of welding induced changes cannot be easily deconvoluted, because both techniques are comparing unstressed areas to stressed areas and also are sensitive to microstructural variation

References cited in this section

18 N.J Rendler and I Vigness, Hole-Drilling Strain-Gage Method of Measuring Residual Stresses,

Experimental Mechanics, Vol 6, 1966, p 577–586

19 W Cheng and I Finnie, A Method for Measurement of Axisymmetric Axial Residual Stresses in

Circumferentially Welded Thin-Walled Cylinders, J Eng Mater Technol (Trans ASME), Vol 107,

July 1985, p 181–185

20 D.S Kim and J.D Smith, Residual Stress Measurements of Tension Leg Platform Tendon Welds, Proc

of the Offshore Mechanics and Arctic Engineering Conference 1994, Vol 3, American Society of

Mechanical Engineers, 1994

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Weldability Testing

Weldability, while sometimes defined in the general sense as the measure of the compatibility of base metal and any added filler metal with the heating cycles used for welding, is more commonly defined as the measure of the resistance of the materials to the formation of cracks during welding Methods are available for assessing both hot cracking, at or near the solidification temperature, and cold cracking, at or near room temperature Test methods do not check for all types of cracking at once, but, instead, each test checks for the susceptibility to a certain type of crack

Weldability tests have been designed to allow small-scale specimens to mimic the cracking behavior of large, rigid welded structures These tests either use specimen geometry to force the shrinkage of the weldment as it cools to be counteracted by plastic extension of the weld and HAZ, or they use additional loading to achieve plastic extension of weld and HAZ in addition to that caused by weld shrinkage The bulk of this section will discuss the types of tests where additional loadings are required, because these are properly mechanical tests A limited discussion of some test methods without additional loading is also provided for comparison

The techniques that include welding and loading to measure weldability have been designed primarily for hot cracking at temperatures at or just below that for the last solidification of weld metal The loading augments the shrinkage strains Cracking when the weld has cooled, such as cracks due to hydrogen, is usually tested using welding without additional loading

Varestraint Testing

Varestraint testing uses a cantilever beam specimen that is bent downward by a rapid application of force while

a weld is being made on the top surface of the beam (Ref 21, 22) The weld is made parallel to the direction of tension once the force is applied with the welding torch moving toward the support point, as shown in Fig 7 A die block 51 mm (4 in.) long is placed beneath the specimen to force it to a limiting value of augmented strain The radius of the die block can be varied to examine the dependence of cracking on augmented strain, or a single radius can be used to examine the effect of other welding variables such as base metal chemistry or weld heat input Common values of the augmented strain are between 0 and 4% Standard varestraint testing methods are discussed in AWS B4.0

Fig 7 Varestraint test fixture and specimen Source: Ref 9

The specimen size is usually 305 mm (12 in.) long and 51 mm (2 in.) wide with a thickness of either 6 or 12.5

mm ( or in.), although minivarestraint specimens are sometimes tested, which reduce all dimensions by a factor of 2 To prevent local kinking of the specimen, auxiliary plates are bent along with the test plate The auxiliary plates, rolled steel 305 mm by 51 mm by 12.5 or 6 mm (12 in by 2 in by or in.), are clamped to the edges of the test plate A minimum of three specimens are tested for any experimental condition The weld on the top surface is produced by gas tungsten arc welding (GTAW)

Measurements are based on crack length of cracks found in the HAZ or fusion zone Hot cracks typically form radially along the trailing edge of the weld pool and in the HAZ Either the length of the longest crack or the total combined crack length of all cracks is the measured parameter Cracks are counted and sized on the

surface under a low power microscope (40 to 80×) One parameter of interest is the maximum augmented strain without cracks However, because only one augmented strain is measured per specimen, the value is most closely estimated by extrapolating to zero the crack length measured at higher augmented strains

Transvarestraint testing also uses welding on the top surface on a piece that is bent over a die block (Ref 23) However, the force is applied so that tension is rapidly applied across the direction of welding rather than along

it This method detects cracking sensitivity in weld metal more effectively than the original varestraint method Cracks in weld metal are more likely to be subsurface, so weld cross sections or nondestructive methods, such

as radiography, are required to observe the subsurface cracks

One problem with transvarestraint tests is that cracks formed during the rapid loading may continue to propagate after the specimen has been bent

Spot Varestraint Testing

The spot varestraint test, or TIG-A-MA-JIG test, is a modification of the varestraint test This test uses a stationary welding torch (Ref 24) The torch is shut off at the instant of bending of the specimen or a specified short time before

The size of the weld pool should remain relatively constant between tests that are to be compared This may require changes of welding current or arc time when materials of differing thicknesses are compared

The die block for the spot varestraint test is curved in only one direction, that is, it is shaped like the surface of

a cylinder Cracking, thus, will be primarily found transverse to the tension above the highest part of the surface

of the die block Radii for the die block have ranged from 44.5 to 1270 mm (1.75 to 50 in.)

Like the varestraint test, cracks are measured visually on the surface using a low-power microscope Because this test is primarily used to test for HAZ liquation cracks, crater cracks in the center of the weld metal are ignored

A measure of the relative susceptibility to HAZ liquation cracks that is not available in the varestraint test is the shortest time from torch shut-off to bending that produces no cracks

Weldability Testing without Augmented Strain

Weldability testing without augmented strain requires that the weld be made in a highly restrained specimen Highly restrained geometries prevent weld shrinkage in more than one direction Cracking is typically from the root of the weld, either at a fillet weld or a partial penetration butt weld Specimen shapes sometimes surround the weld metal with base metal, as in the Lehigh restraint test Other test geometries, such as the oblique Y-groove test, join two plates with welds at the ends and then put the test weld in the center

References cited in this section

9 “Standard Methods for Mechanical Testing of Welds,” ANSI/AWS B4.0-98, American National Standards Institute/American Welding Society, Miami, 1998

21 W.F Savage and C.D Lundin, The Varestraint Test, Weld J., October 1965, p 433s–442s

22 C.D Lundin, A.C Lingenfelter, G.E Grotke, G.G Lessmann, and S.J Matthews, The Varestraint Test,

Bulletin 280, Welding Research Council, August 1982

23 B.F Dixon, R.H Phillips, and J.C Ritter, Cracking in the Transvarestraint Test, Part 1: A New

Procedure for Assessment of Cracking, Met Constr., Feb 1984 and Aust Weld J., Vol 28 (No 4),

summer 1983

24 W Lin, J.C Lippold, and W.A Baeslack III, An Evaluation of Heat-Affected Zone Liquation Cracking

Susceptibility, Part I: Development of a Method for Quantification, Weld J., Vol 72 (No 4), 1993, p

135s-153s

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

Weldability Testing

Weldability, while sometimes defined in the general sense as the measure of the compatibility of base metal and any added filler metal with the heating cycles used for welding, is more commonly defined as the measure of the resistance of the materials to the formation of cracks during welding Methods are available for assessing both hot cracking, at or near the solidification temperature, and cold cracking, at or near room temperature Test methods do not check for all types of cracking at once, but, instead, each test checks for the susceptibility to a certain type of crack

Weldability tests have been designed to allow small-scale specimens to mimic the cracking behavior of large, rigid welded structures These tests either use specimen geometry to force the shrinkage of the weldment as it cools to be counteracted by plastic extension of the weld and HAZ, or they use additional loading to achieve plastic extension of weld and HAZ in addition to that caused by weld shrinkage The bulk of this section will discuss the types of tests where additional loadings are required, because these are properly mechanical tests A limited discussion of some test methods without additional loading is also provided for comparison

The techniques that include welding and loading to measure weldability have been designed primarily for hot cracking at temperatures at or just below that for the last solidification of weld metal The loading augments the shrinkage strains Cracking when the weld has cooled, such as cracks due to hydrogen, is usually tested using welding without additional loading

Varestraint Testing

Varestraint testing uses a cantilever beam specimen that is bent downward by a rapid application of force while

a weld is being made on the top surface of the beam (Ref 21, 22) The weld is made parallel to the direction of tension once the force is applied with the welding torch moving toward the support point, as shown in Fig 7 A die block 51 mm (4 in.) long is placed beneath the specimen to force it to a limiting value of augmented strain The radius of the die block can be varied to examine the dependence of cracking on augmented strain, or a single radius can be used to examine the effect of other welding variables such as base metal chemistry or weld heat input Common values of the augmented strain are between 0 and 4% Standard varestraint testing methods are discussed in AWS B4.0

Fig 7 Varestraint test fixture and specimen Source: Ref 9

The specimen size is usually 305 mm (12 in.) long and 51 mm (2 in.) wide with a thickness of either 6 or 12.5

mm (¼ or ½in.), although minivarestraint specimens are sometimes tested, which reduce all dimensions by a factor of 2 To prevent local kinking of the specimen, auxiliary plates are bent along with the test plate The auxiliary plates, rolled steel 305 mm by 51 mm by 12.5 or 6 mm (12 in by 2 in by ½or ¼in.), are clamped to the edges of the test plate A minimum of three specimens are tested for any experimental condition The weld

on the top surface is produced by gas tungsten arc welding (GTAW)

Measurements are based on crack length of cracks found in the HAZ or fusion zone Hot cracks typically form radially along the trailing edge of the weld pool and in the HAZ Either the length of the longest crack or the total combined crack length of all cracks is the measured parameter Cracks are counted and sized on the surface under a low power microscope (40 to 80×) One parameter of interest is the maximum augmented strain

without cracks However, because only one augmented strain is measured per specimen, the value is most closely estimated by extrapolating to zero the crack length measured at higher augmented strains

Transvarestraint testing also uses welding on the top surface on a piece that is bent over a die block (Ref 23) However, the force is applied so that tension is rapidly applied across the direction of welding rather than along

it This method detects cracking sensitivity in weld metal more effectively than the original varestraint method Cracks in weld metal are more likely to be subsurface, so weld cross sections or nondestructive methods, such

as radiography, are required to observe the subsurface cracks

One problem with transvarestraint tests is that cracks formed during the rapid loading may continue to propagate after the specimen has been bent

Spot Varestraint Testing

The spot varestraint test, or TIG-A-MA-JIG test, is a modification of the varestraint test This test uses a stationary welding torch (Ref 24) The torch is shut off at the instant of bending of the specimen or a specified short time before

The size of the weld pool should remain relatively constant between tests that are to be compared This may require changes of welding current or arc time when materials of differing thicknesses are compared

The die block for the spot varestraint test is curved in only one direction, that is, it is shaped like the surface of

a cylinder Cracking, thus, will be primarily found transverse to the tension above the highest part of the surface

of the die block Radii for the die block have ranged from 44.5 to 1270 mm (1.75 to 50 in.)

Like the varestraint test, cracks are measured visually on the surface using a low-power microscope Because this test is primarily used to test for HAZ liquation cracks, crater cracks in the center of the weld metal are ignored

A measure of the relative susceptibility to HAZ liquation cracks that is not available in the varestraint test is the shortest time from torch shut-off to bending that produces no cracks

Weldability Testing without Augmented Strain

Weldability testing without augmented strain requires that the weld be made in a highly restrained specimen Highly restrained geometries prevent weld shrinkage in more than one direction Cracking is typically from the root of the weld, either at a fillet weld or a partial penetration butt weld Specimen shapes sometimes surround the weld metal with base metal, as in the Lehigh restraint test Other test geometries, such as the oblique Y-groove test, join two plates with welds at the ends and then put the test weld in the center

References cited in this section

9 “Standard Methods for Mechanical Testing of Welds,” ANSI/AWS B4.0-98, American National Standards Institute/American Welding Society, Miami, 1998

21 W.F Savage and C.D Lundin, The Varestraint Test, Weld J., October 1965, p 433s–442s

22 C.D Lundin, A.C Lingenfelter, G.E Grotke, G.G Lessmann, and S.J Matthews, The Varestraint Test,

Bulletin 280, Welding Research Council, August 1982

23 B.F Dixon, R.H Phillips, and J.C Ritter, Cracking in the Transvarestraint Test, Part 1: A New

Procedure for Assessment of Cracking, Met Constr., Feb 1984 and Aust Weld J., Vol 28 (No 4),

summer 1983

24 W Lin, J.C Lippold, and W.A Baeslack III, An Evaluation of Heat-Affected Zone Liquation Cracking

Susceptibility, Part I: Development of a Method for Quantification, Weld J., Vol 72 (No 4), 1993, p

135s-153s

Mechanical Testing of Welded Joints

William Mohr, Edison Welding Institute

References

1 “Standard Welding Terms and Definitions,” AWS 3.0, American Welding Society

2 L.P Connor, Ed., Welding Handbook, 8th ed., Vol 1, American Welding Society, 1991

3 “Specification for Carbon Steel Electrodes for Shielded Metal Arc Welding,” ANSI/AWS A5.1-91, American National Standards Institute/American Welding Society, Miami, 1991

4 “Recommended Practice for Preproduction Qualification for Steel Plates for Offshore Structures,” API 2Z, 2nd ed (includes November 1998 addenda), American Petroleum Institute, Washington DC, 1998

5 “Standard Test Method for Young's Modulus, Tangent Modulus, and Chord Modulus,” E 111, Annual Book of ASTM Standards, Vol 3.01, ASTM, 1999

6 “Standard Methods of Tension Testing of Metallic Materials,” E 8, Annual Book of ASTM Standards,

Vol 3.01, ASTM, 1999

7 “Standard Methods of Tension Testing Wrought and Cast Aluminum, and Magnesium Alloy Products,”

B 557, Annual Book of ASTM Standards, Vol 3.01, ASTM, 1999

8 B4 Committee on Mechanical Testing of Welds, Mechanical Testing of Welds, Part 1: Summary of

Tension Testing of Welds, Weld J., Jan 1981, p 33–37

9 “Standard Methods for Mechanical Testing of Welds,” ANSI/AWS B4.0-98, American National Standards Institute/American Welding Society, Miami, 1998

10 B4 Committee on Mechanical Testing of Welds, Mechanical Testing of Welds, Part 2: Bend Testing of

Welds A Summary, Weld J., Feb 1981, p 34–37

11 ASME Boiler and Pressure Vessel Code, Section IX, American Society of Mechanical Engineers

12 Aluminum Design Manual, The Aluminum Association, 1994

13 “Standard Test Methods for Notched Bar Impact Testing of Metallic Materials,” E 23, Annual Book of ASTM Standards, Vol 3.01, ASTM

14 “Standard Method for Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature

of Ferritic Steels,” E 208, Annual Book of ASTM Standards, Vol 3.01, ASTM

15 M.G Dawes, H.G Pisarski, and S J Squirrell, Fracture Mechanics Tests on Welded Joints, Nonlinear Fracture Mechanics Vol II: Elastic-Plastic Fracture, STP 995, J.D Landes, A Saxena, and J.G

Merkle, Ed., ASTM, 1989, p 191–213

16 “Standard Test Method for Crack-Tip Opening Displacement (CTOD) Fracture Toughness

Measurement,” E 1290, Annual Book of ASTM Standards, Vol 3.01, ASTM

17 “Standard Test Method for J-Integral Characterization of Fracture Toughness,” E 1737, Annual Book of ASTM Standards, Vol 3.01, ASTM

18 N.J Rendler and I Vigness, Hole-Drilling Strain-Gage Method of Measuring Residual Stresses,

Experimental Mechanics, Vol 6, 1966, p 577–586

19 W Cheng and I Finnie, A Method for Measurement of Axisymmetric Axial Residual Stresses in

Circumferentially Welded Thin-Walled Cylinders, J Eng Mater Technol (Trans ASME), Vol 107,

July 1985, p 181–185

20 D.S Kim and J.D Smith, Residual Stress Measurements of Tension Leg Platform Tendon Welds, Proc

of the Offshore Mechanics and Arctic Engineering Conference 1994, Vol 3, American Society of

Mechanical Engineers, 1994

21 W.F Savage and C.D Lundin, The Varestraint Test, Weld J., October 1965, p 433s–442s

22 C.D Lundin, A.C Lingenfelter, G.E Grotke, G.G Lessmann, and S.J Matthews, The Varestraint Test,

Bulletin 280, Welding Research Council, August 1982

23 B.F Dixon, R.H Phillips, and J.C Ritter, Cracking in the Transvarestraint Test, Part 1: A New

Procedure for Assessment of Cracking, Met Constr., Feb 1984 and Aust Weld J., Vol 28 (No 4),

summer 1983

24 W Lin, J.C Lippold, and W.A Baeslack III, An Evaluation of Heat-Affected Zone Liquation Cracking

Susceptibility, Part I: Development of a Method for Quantification, Weld J., Vol 72 (No 4), 1993, p

Rolling bearings include radial, thrust, and angular contact designs A review of the many versions of these

bearings can be found in the article“Friction and Wear of Rolling-Element Bearings” in Friction, Lubrication, and Wear Technology, Volume 18 of the ASM Handbook (Ref 1) The primary requirement of rolling bearings

is proper and adequate lubricant that provides separation of the moving surfaces under all conditions, maintains appropriate temperature, and provides an operating environment so that bearings will achieve their expected lives

Sliding bearings include sleeve and thrust bearings of various designs Based on designs and materials selection, plain bearings operate under dry or boundary lubrication conditions, partial film or mixed lubricant film conditions, or a full film, which means the “rubbing” surfaces are essentially separated More details of

sliding bearings can be found in the article “Friction and Wear of Sliding Bearings” in Friction, Lubrication, and Wear Technology, Volume 18 of the ASM Handbook (Ref 2) and in Ref 3

References cited in this section

1 T.A Harris, Friction and Wear of Rolling-Element Bearings, Friction, Lubrication, and Wear Technology, Vol 18, ASM Handbook, ASM International, 1992, p 499

2 R Pike and J.M Conway-Jones, Friction and Wear of Sliding Bearings, Friction, Lubrication, and Wear Technology, Vol 18, ASM Handbook, ASM International, 1992, p 515–521

3 M.M Khonsari and J.Y Jang, Chapters 61, 62, and 63, Tribology Data Handbook, E.R Booser, Ed.,

CRC Press, LLC, 1997, p 669–707

Testing of Bearings

Charles A Moyer, The Timken Company (Retired)

Rolling Element Bearings

Rolling bearings date back to the Neolithic period, or the new Stone Age (Ref 4) When heavy materials needed

to be moved, or when primitive vehicles had wheels, someone devised bronze or wood “bearings,” with or without oils or fats Archaeologists have found bronze balls, cylindrical rollers, and wooden tapered roller bearings dating back to 300 to 500 B.C

Considering the different materials used, it is clear that finding the proper material, shapes, and sizes was then, and continues to be, important and is the core of testing to determine what performance can be expected from rolling element bearings Advancements continued over the years, but bearings that resemble those in use today did not appear until the late 19th century Modern rolling bearings developed gradually from 1850 to 1925 (Ref 4) One driving force was the need for bearings in bicycles

While working in Berlin, Germany, Professor Richard Stribeck undertook early bearing tests, the results of which were published in 1901 and 1902 (Ref 5, 6) His goal was to determine safe ball loads statically and in complete bearings over different speed ranges He made use of Hertz's work covering elastic bodies in contact (Ref 7) and started with a press arrangement with three hardened steel balls in contact, two steel balls with a steel plate between them, and finally, a single ball between two short cylinders set on end with cup shapes to fit

the ball He found the relation of load P on the ball diameter d to be:

K was a constant based on steel type and contact geometry Based on the materials he used, Stribeck established

that Eq 1 held to the elastic limit and somewhat beyond He also ran tests on complete bearings and established the way balls shared loads under radial load conditions Using bearings with 10 to 20 balls and no clearance

between balls and rings, he determined that the most heavily loaded ball (P0) was related to the total load P by:

where Z is the number of balls in the bearing To be conservative, Stribeck changed 4.37 to 5.0

In 1912, Professor John Goodman published in England studies on rolling element bearings (Ref 8), covering life, friction, and wear He knew of Stribeck's work and determined reductions in bearing load capacity based

on bearing speed (N) Goodman's equation for this reduction was:

where D is the ball-race path diameter; A is a constant; and P,K, and d are the same as in Eq 1

Studies of bearings and bearing materials (primarily steel) were followed by the emergence of the first manufacturing companies to patent and produce rolling element bearings Bearing lubrication at this time was greatly influenced by the Reynold's equation (Ref 9) and the successes experienced in the hydrodynamic films generated in conformal bearings It was assumed that the nonconformal contacts in “antifriction” bearings generated much thinner and, thus, less protective films for the contact region The actual lubricating means for rolling bearings was thus a puzzle

However, in 1949, Grubin (Ref 10) considered the elastic behavior of the contact materials and the pressure characteristics of the viscosity under the high stresses in the contact region in an equation that gave lubricant films one or two orders thicker than with the usual rigid cylinders assumption used for determining lubricant films as in hydrodynamic bearings For rolling bearings, elastohydrodynamic lubrication came to be recognized

as an important factor in determining life expectancy under a wide range of operating conditions Tallian (Ref 11) described the stages of rolling element bearings as an empirical state until late 1920 A classical period then emerged that saw the engineering inclusion of Hertz's elasticity theory, the Weibull function, and statistics leading to worldwide bearing standards Finally, the modern period began in the 1950s The advancements in bearing technology allowed the new era to develop meaningful testing for rolling bearings

References cited in this section

4 D Dowson, History of Tribology, 2nd ed., Professional Eng Pub., Ltd., London, Edmunds, U.K., 1998

5 R Stribeck, Kugellager für beliebige Belastungen, Z Ver dt Ing., Vol 45 (No 3), 1901, p 73–125

6 R Stribeck, Die Wesentlichen Eigen-schaften der Gleit und Rollenlager, Z Ver dt Ing., Vol 46 (No

38), 1902, p 1341–8, 1432–8; (No 39), 1463–70

7 H Hertz, On the Contact of Elastic Solids, J reine und angew, Math., Vol 92, 1881, p 156–71

8 J Goodman, (1) Roller and Ball Bearings, (2) The Testing of Antifriction Bearing Materials, Proc Inst Civ Engrs., clxxxix, Session 1911-2, Pt 111, 1912, p 4–88

9 O Reynolds, On the Theory of Lubrication and Its Application to Mr Beauchamp Tower's

Experiments, Including an Experimental Determination of the Viscosity of Olive Oil, Philos Trans R Soc., 177, 1886, p 157–234

10 A.N Grubin and I.E Vinogradova, Investigation of the Contact of Machine Components, Book No 30

(DSIR Trans No 337), Kh F Ketova, Ed., Central Scientific Research Institute for Technology and Mechanical Engineering, Moscow, 1949

11 T.E Tallian, “Progress in Rolling Contact Technology,” Report AL 690007, (SKF) Industries, King of Prussia, PA, 1969

Others, such as Robert Thurston (Ref 13), who studied journal bearing friction and lubrication; Nikolai Petrov, who explained the hydrodynamic characteristics of journal bearing friction (Ref 14); and Beauchamp Tower (Ref 15), who discovered the pressure that developed within an operating journal bearing and that a bearing with sufficient lubrication floated on a film of oil, followed Hirn's studies Although the studies first