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Patent Searching and Data


Title:
MATERIALS JOINING
Document Type and Number:
WIPO Patent Application WO/1990/006208
Kind Code:
A1
Abstract:
A method for improving the reliability of a ceramic-metal joint by reducing the dynamic mismatch stresses and strains on the ceramic due to the dynamic temperature differential and mismatch in thermal expansions of the two materials. This is done by bonding with a metal layer the ceramic to metal to form a bonding interfacial region between the ceramic (11) and metal (12); and grading laterally (14, 15) or parallel to this interfacial region the thermal conductivity, thermal expansion coefficient, or softness of the metal layer. An article in the form of a laterally graded, metallic bonding composite disc for overcoming severe dynamic mismatch stresses and strains in, e.g., electronic device packages, is also disclosed.

Inventors:
LI, Chou, H.
Application Number:
PCT/US1989/005276
Publication Date:
June 14, 1990
Filing Date:
November 28, 1989
Export Citation:
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Assignee:
LI, Chou, H.
International Classes:
B22F1/02; B23K20/16; B23K35/00; C04B37/00; C04B37/02; C22C1/10; C22C47/08; C22C49/14; C23C26/02; H05K1/03; H05K3/38; (IPC1-7): B23K103/16; B23K103/18
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Claims:
Claims
1. A method for improving the reliability of a join between a first body of a first solid material and a second bod of a second solid material, said method comprising reducing o at least one of the solid materials the dynamic mismatc stresses due to dynamic temperature differentials and th resultant mismatches in thermal expansions between the two soli materials during transient uneven heating or cooling of th joint, comprising: providing a metal layer on a selected surface of the bod of the first solid material; bonding with the metal layer the second solid material bod to the selected surface of the first solid material body thereb forming the joint and a bonding interfacial region between th two bodies and including the metal layer; and grading from a central portion toward a peripheral portio of the joint at least one thermomechanical property of th material of the metal layer in the interfacial region to ensur that the maximum residual dynamic mismatch stresses will no exceed the local material strength in either solid material a any point and time, said one thermomechanical property bein selected from the group consisting of thermal conductivity thermal expansion coefficient, and softness or shockabsorbin ability.
2. A method as in claim 1 including heattreating th joint to modify by controlled elemental diffusion th composition profiling in the metal layer.
3. A method as in claim 1 wherein said grading ste includes grading ,at least two of the thermomechanica properties.
4. A method as in claim 3 wherein said grading ste consists of grading simultaneously all the three liste thermomechanical properties.
5. A method as in claim 1 including grading laterally fro the central portion toward the peripheral portion of th interfacial region the thermal expansion coefficient of th material of the metal layer; and providing at the peripheral portion a material of th metal layer having a thermal expansion coefficient higher tha that next to the periphery.
6. A method as in claim 1 wherein said grading ste comprises monotonically changing the thermal expansio coefficient of the metal layer material from the central portio toward the peripheral portion of the interfacial region withou generating additional stresses after the joint is made an cooled to room temperature.
7. A method as in claim 1 including reducing the transien or dynamic temperature differentials between the solid material at the central portion of the joint where dynamic mismatc stresses are maximum.
8. A method as in claim 1 wherein said first soli material is dissimilar from and relatively weaker than sai second solid material, said method comprising reducing th dynamic mismatch stresses due to the dynamic mismatches on th relatively weaker first solid material body, and said providing step comprises supplying a first metalli surface layer onto the selected surface of the first soli material body and furnishing a bonding metallic layer fo joining the first metallic surface layer to the second materia body.
9. A method as in claim 8 wherein said supplying furnishing, bonding and grading stpes are done in a singl processing operation.
10. A method as in claim 8 wherein the' processing of sa supplying, furnishing, bonding and grading steps alone, witho any additional heat treatment of the joint, is effective improving the reliability of the joint against the dynam mismatch stresses.
11. A method as in claim 8 wherein the grading ste comprises providing at least the bonding metallic' layer with t highest thermal expansion coefficient material in the peripher portion relative to the central portion of the joint.
12. A method as in claim 8 wherein the grading ste comprises grading the thermal expansion coefficient of at leas the material of the bonding metallic layer without generati additional stresses after the joint is made, and providing a least the bonding metallic layer with the highest therma conductivity material in the central portion relative to t peripheral portion of the joint.
13. A method as in claim 8 wherein the grading ste comprises grading the softness or shockabsorbing ability of t material of the bonding metallic layer, and providing at leas the bonding metallic layer with the softest material in t central portion relative to the peripheral portion of the joint.
14. A solid metallic bonding layer for bonding a cerami body to a second solid material body thereby forming a joi with a bonding interfacial region between the two materials, said ceramic body and said second material body havi different thermophysical properties thereby creating residua dynamic mismatch stresses in the bonding layer, comprising: a first solid metallic bonding material in one part of th bonding layer creating thereat a first actual or residua dynamic mismatch stress which is the theoretical local mismatc stress with a first portion thereof absorbed at least in t first bonding material; a second solid metallic bonding material in another part o the bonding layer creating thereat a second actual or residua dynamic mismatch stress which is the theoretical local mismatc stress with a second portion thereof absorbed at least in t second bonding material; said two solid bonding materials substantially differing i 22 at least one physical property thereby absorbing different portions of the respective actual mismatch stresses, said first theoretical local mismatch stress being highe than said second theoretical local mismatch stress but sai first portion being larger than said second portion so that bot actual or residual dynamic mismatch stresses after th absorptions are less than those that can damage the bond betwee the ceramic and the second solid material.
15. A bonding layer as in claim 14 wherein the firs bonding material is in the form of a central solid disc of th first metallic bonding material, while the second bondin material is in the form of an outer ring of the second metalli bonding material surrounding said central disc, said two bonding materials substantially differing in sai at least one physical property selected from the grou consisting of thermal conductivity, thermal expansio coefficient, and softness.
16. A bonding layer as in claim 14 wherein the cerami body is relatively weaker than the second solid material bod and including a soft, yieldable metal material at a centra portion of the bonding interfacial region to absorb within th yieldable metal material a large portion of the dynamic mismatc stresses so that the relatively weaker ceramic is no longe subjected to high dynamic mismatch stresses thereby preventin failure of the joint, the remainder of the interfacial regio having no such soft, yieldable metal material.
17. A method for reducing the dynamic mismatch stres failure of a joint between two solid material bodies joine together with a solid metallic bonding layer therebetwee thereby forming the joint and a bonding interfacial regio between the bodies, comprising: providing a plurality of physically integrated joi elements in parallel at a plurality of locations of the bondi interfacial region, the dynamic mismatch stress being differe at different locations; at least one physical property of the joint element at on location being substantially different from the same property o the joint element at another location, so that the maximum o critical transient or dynamic mismatch stresses never excee the local material strength at any point in the joint, at an time during the uneven heating or cooling of the joint i processing or service.
18. A method as in clai 17 wherein said one physica property is selected form the group consisting of therma conductivity, thermal expansion coefficient, and softness o shockabsorbing ability.
19. A method as in claim 17 wherein the maximum transien or dynamic mismatch stresses vary along the bonding interfacia region and are respectively the reduced theoretical mismatc stresses due to absorption in at least the metallic bondin layer and including the additional step of increasing th absorption of such mismatch stresses with the most absorption a the location where maximum theoretical dynamic stresses occur.
20. A method as in claim 19 including laterally tailor grading the physical property profile of the bonding interfacia region according to the maximum or critical transient o dynamic mismatch stresses so that these mismatch stresses neve exceed the local material strength at any point inside the joint, at any time during the heating or cooling of such joints in processing or service.
Description:
DESCRIPTION

MATERIALS.JOINING

Technical Field

This invention relates to ceramic-metal joining, and mo particularly relates to ceramic-metal joining with unifo ceramic metallizing compositions and specially graded seals make reproducibly strong and thermomechanically shock-resista oints.

Background Art This is a continuation-in-part of my pending U. applications Serial Nos. 07/277,666 filed November 29, 1988 07/277,672 filed December 14, 1988.

A serious problem with present ceramic metallizing metho is the difficulty of achieving uniform metallized layers form on the ceramic. Take, for example, the commonly used hea metal processes, such as -yttria (W-Y2O3), W-Fe, or Mo-Mn. these and many similar methods, segregation of the mixed met or other powders takes place due to their differing specif gravities, shapes, sizes, porosities, and surface smoothnes These segregations occur at all times: during the mixing of t powders, storing of the powder suspensions, application of t suspensions, settling of the suspended powder particles aft the application, and drying of the applied layer. Furthe these segregations occur so fast as to be practical uncontrollable, as will be shown shortly.

In general, spherical, heavy, large, smooth, and den particles settle first and early in the binder or suspensi medium. Upon settling, these particles tend to roll or mo sidewise or downward toward the corners or boundaries faster a further than odd-shaped, light, small, rough, and poro particles of otherwise identical characteristics.

Take the W-Y2O3 mixed powders in an organic binder of specific gravities of 19.3, 4.5, and 0.98, respectively. Such a suspension, even if perfectly mixed up by shaking, stirring, roller-milling, or otherwise, will immediately segregates. The initial settling acceleration due to gravitational minus buoyancy forces of powders is 980.6 X (19.3-.98)/19.3 = 930.8 cm 2 /sec, while that of Ϊ2°3 P owders ~- s only 767.0 cm 2 /sec.

In a mixing, storing, or carrying bottle 10 cm high and containing a perfectly mixed suspension of these metallizing powders, the time to completely settle out is only 147 ms (milliseconds) for W powders, if uniform acceleration is assumed. At the tip of a paint brush having a suspension drop 0.3 cm in diameter, the complete settling time of these same powders is merely 25.4 ms, while on a horizontally painted or sprayed layer 0.1 cm thick, the same settling time is only 14.7 ms. In all these cases, the complete settling time for the 2O3 powders is always the square root of 930.8/767.0, i.e., 1.21 times or 21% longer.

Note in particular that the powder segregations with uniform accelerations may be completed within 147 to 14.7 ms.

Such short times indicate that the - 2O3 powder segregations are beyond human control. The painted or sprayed mixed powder layers are thus always not uniform.

In metallizing onto a horizontal ceramic surface to be metallized, most of the powders immediately settles out. The first layers are therefore always very rich in W and very poo in v 2°3* These first layers are too refractory for the prese metallizing temperature (up to about 1550°C) so that the cerami surfaces are not sufficiently metallized, or not at all. Th last settling layers, on the other hand, are too rich in th fluxing Y2O3. Again, the ceramic surfaces are improperl metallized, with only a glassy layer being formed which is ver weak in strength and thermal shock resistances.

Thus, common metallizing results on ceramics are ofte erratic and uncontrollable. The metallized surface may contai

loose and- unmetallized spots with high content of heav refractory metal, and also non-wettable spots due to high flu content. The entire process is critical and involved, and ye nonunifor . The resultant ceramic-metal joints or cerami coatings on metals are weak, costly, nonreproducible, an usually not vacuum-tight, or temperature-resistant.

Painting or spraying onto vertical or inclined surface results in vertical and additional lateral segregations an gradations, and gives added poor uniformity, reproducibility and bonding strength.

While only the effect of gravitational density segregatio has been considered, the other segregation variables such a powder shape, size, porosity, and surface roughness are als important. A second important problem with common joining processes i the lack of control, or even understanding, of dynami mismatches of temperature, stress, and strain profiles in th joint region, and their variations with time. Another aspec of this invention is therefore to describe such dynami mismatch phenomena, and to specially tailor-grade th composition and/or physical property profiles of the join region so that the maximum or critical transient mismatc stresses never exceed the local material strength at an point inside the joint region, at any time during the heatin or cooling of such joints in processing or service.

Accordingly, an object of this invention is to provid improved ceramic-metal joints and joining methods;

A further object of this invention is to provide improve ceramic metallizing methods for these joints; A broad object of this invention is to minimiz gravitational segregations of the components in the metallizin methods prior to the joining;

Another broad object of the invention is to speciall tailor-grade the composition and/or property profiles in th joint regions to ensure that the maximum dynamic or transien

stresses do not exceed the local material strengths at any point and time.

Further objects and advantages of my invention will appears as the specification proceeds.

Disclosure of the Invention

The present invention provides a method for improving the reliability of a ceramic-metal or other joint between two similar or dissimilar solid materials against dynamic mismatch stresses due to temperature differentials and differing thermal expansions during nonuniforra dynamic heating or cooling of the joint. This method comprises grading the thermal conductivity, thermal expansion coefficient, and/or softness or shock- absorbing ability of the bonding metal layer. The grading is done laterally or in directions parallel to the bondin interfacial region, rather than axially or normally of th interfacial region as to the thermal expansion coefficient alon for commonly reducing static thermal mismatch stresses.

Brief Description of Drawing

In the single drawing: Fig. 1 is a top view showing the ceramic end of the joint and

Fig. 2 is a cross-section view of the joint taken along th cross-section line 2-2 of Fig. 1.

Best Mode for Carrying Out the Invention It will be understood that the specific embodiment described herein are merely illustrative of the genera principles of the invention and that various modifications ar feasible without departing from the spirit and scope of th invention. That is,- the invention is of general applicabilit 0for improving the quality of the ceramic-metal or other joint and joining methods, or coatings of ceramics on ceramics o metals. It is also evident that materials, structures, an

methods other than those especially described can be used t practice the invention.

Stokes in 1851 first considered the resistance R which fluid medium of density d m and viscosity n offers to th

5 movement of a spherical particle of diameter D and density

(dp>d m ) and settling with velocity v in the medium. He arrive at the equation: R = 37C Dvn *

The small sphere settling in the fluid or suspension mediu is acted on by the downward force of gravity with gravitationa - - constant g, 7 " D 3 d_ g/6; and by the upward buoyant force o the fluid, 7T D - a 9^- / giyen by Archimedes' principle. Th resultant net gravitational force G is T D3 ( d p" d m)9 y ' 6 actin downward, producing a downward acceleration, a.

When the resistance R exactly equals this net gravitationa 15 force G, the acceleration reduces to zero. The final velocity V£, becomes constant. There then results: 3 7T D n v f « 7c D 3 (dp-d j g/6 Hence, the final velocity is: ^ = (d p -d ) g D2 /18 n, th equation of Stokes'law which has been shown to be widely valid. 20 For a given fluid density (c^) at a specific temperatur (viscosity n) and a given sphere (of density d_ and mass M), t Stokes' equation gives a velocity constant: Also, the velocity at any time starting from rest, t, is: 25 v = (1 - exp(-R t /M)) x v f ;

The settling distance at time t is: s t = (t - (l-exp(-Rt/M)) x M/R) x G/R The velocity equation shows that the exact v^ is no reached until after infinitely long time when the exponentia 30term in the equation becomes zero and the velocity reduces t v=V£, as it should.

With the Stokes' law, one can calculate the velocit constants, v c in 1/cm-sec, for settling in water at 20 C (d =1. and n = 0.010), of various metal or oxide powders with densitie

35in g/cc in parentheses, as follows: W (19.35) 100,000, Y

(5.01) 21,-900, Fe (7.87) 37,400, Mo (10.2) 50,100, Mn (7.2) 33,800, W0 3 (7.16) 33,600, Fe 2 0 3 (5.24) 23,100, M0O3 (4.692) 20,100, and Mn0 2 (5.026) 21,900.

Thus, in the -Y2O metallizing process, because the powders are 3.9 (19.35/5.01) times heavier than Y2O3, th velocity constants c 's of the two components differ by a facto of 100,000/21,900 = 4.6 times. That is, for a given powder siz D, the final constant settling velocity ~ r- of W spheres is 4. times greater than that of Y2°3 spheres. As discussed above, 0 this wide difference in settling velocities results in sever gravitational segregation and early depletion of W particles i the settling mixtures and, therefore, poor metallizing results.

It can also be seen that the powders in the mixed oxid processes, e.g., C>3-Fe2C> , are much more uniform, or. les 5varying, in densities, d p , than mixed particles of the sam metals, e.g., W-Fe. Thus, the Wθ3~Fe2θ3 process shows densit and velocity constant ratios of 1.366 and 1.455, vs 2.459 an 2.674, respectively, for the W-Fe process.

Similarly, in the Mo-Mn process, replacing the meta 0powders by their respective oxides reduces the differences i the ratios of velocity constants, v c , and final velocities, Vf from 48.2 % to only 9.0 % and 19.2 % to 4.2 %, respectively. I addition, the metal particles, i.e., W, Fe, Mo, and Mn whe reduced by hydrogen during metallizing from their respectiv 5oxides are smaller than the initial oxide powders. Thes smaller sizes further promote uniform metallizing results.

Hence, if we select and mix the Fe2C>3 and WO3 spherica powders in sizes (diameters D) according to the square root o the inverse ratio of their velocity constants 33,600/23,100 301.455, i.e., 1»206, the final settling velocities of both thes size-ratioed powders will be exactly the same. That is, b simply making the .Fe 2 θ3 powders 20.6% la'rger than the WO powders, the mixed particles will finally settle in water 20°C at exactly the same velocity. This condition leads 35improved metallizing composition and uniformity.

The final settling velocities of the two mixed powders

Vf's, however, come only after some settling time, t s , when specific amount, Q, of the mixed powders has already settled ou at differing velocities. For the specific combination o

5 component powders, the settled amount Q and material us efficiency at this settling time, t s , can be computed from th materials remaining after t s . The materials already settle before t s is the presettled distances, s t , multiplied by th initial material densities. The already settled materials are

10however, not lost, since they can be recirculated and reused i subsequent metallizing runs.

By repeated iteration or computer simulation, the bes mixed-powder metallizing process for combined metallizin uniformity and material use efficiency can be determined. Base l5on these principles, method and equipment can be developed fo controlling the turn-on time for starting to deposit the mixe powder at a nearly equal final settling velocity, v^, int metallizing layers with the size-ratioed powders.

In practice, we specify that the two settling velocities o

20the mixed particles are within a certain prespecifie percentage, e.g., 20 or 10%, of each other. Still gravitational segregations are minimized. Naturally, th smaller the percentage of velocity or useful powder siz differences, ^v and ^ s r respectively, the lower the materia

25use efficiency on a particular, mixed-powder combination. A engineering compromise must, therefore, be struck..

To completely eliminate gravitational segregations solution metallizing is ideal. One difficulty of metallizin MACOR, Corning Glass's machinable glass ceramic, by the solutio

30method is the relatively low, allowable metallizing temperatur of about 950°C. The solubilities of the metallizing compound are also restricting- factors. Still, many potential metallizin compounds are soluble or at least partly soluble. Zinc chlorid and sodium molybdate, for example, are soluble up to 432 and 6

35grams, respectively, per 100 cc of cold water. Such a mixe

solution may be used for MACOR or other ceramics.

Another important consideration in making joints between two similar or dissimilar materials relates to thermal mismatch stresses and strains. In any ceramic-metal joints, or for that matter, any joining of two materials, the match or mismatch of their. thermomechanical characteristics in general, and thermal expansion coefficients in particular, is extremely important.

From this mismatch of their thermal expansions, thermal stresses are generated. Mismatches in other thermomechanical characteristics also result in other thermomechanical mismatch stresses and strains.

The magnitude of these mismatch stresses and strains determines the failure probability of the joint.

Generally, the thermal expansion mismatch differentials of within 100 ppm (parts per million) are considered as allowable, according to Hagy and Ritland's paper on "Viscosity Flow in

Glass-to-Metal Seals," J. Amer. Ceram. Soc, Vol. 40, pp. 58-62,

1957. Such thermal expansion coefficients and differentials relate only to the static or equilibrium case, and may not trul represent dynamic or transient conditions when the joint is unevenly being heated up or cooled down. Such transien conditions often exist during the processing and services of the joint.

Unlike the commonly used static thermal expansion mismatch, the dynamic mismatch in thermal expansion coefficients is no constant, but varies with the bonded material shapes and sizes, physical and surface properties, and heating or coolin conditions and times.

As will be shown, the dynamic expansion strain mismatch ma exceed the yield point of the ceramic materials, while th dynamic mismatch stress may exceed the " flexure or eve comprehensive strengths of these same materials. What fail most ceramic-metal joints is the dynamic, rather than th static, thermal expansion mismatch.

Using- this dynamic mismatch technique, we can determine th location, magnitude, and occurrence time of the maximum o critical mismatch stresses, and take measures to reduce th dynamic mismatch stresses on the relatively weaker ceramic s that the ceramic is no longer failing from the high stresses.

In a joint of two unevenly heated or cooled soli materials, particularly of dissimilar materials such as cerami and metal, dynamic mismatches exist. Dynamic mismatches i ceramic-metal joints result partly from the fact that metals an ceramics have widely different thermal conductivities. Th conductivities for metals range from 0.014 cal/sq. cm/cm/degre C/sec for tellurium to 1.0 for silver (same unit), while thos for ceramics are from 0.0018 for glass to 1.8 for beryllia.

During heating of a ceramic-metal joint, the cerami temperature lags behind that of the metal, often markedly so Under cooling, the opposite is true. This produces differen temperature profiles in the metal and ceramic at a particula time instant on either heating or cooling. Dynamic mismatche in temperatures, strains (i.e., expansions on heating o shrinkages on cooling), and stresses (strains multiplied b Young's moduli) then result.

Take the special example of the case of a long cylinde joined end-to-end to a similarly sized metal cylinder. Th ceramic may be, for example, Corning Glass's machinable glas ceramic (MACOR), while the metal may be SAE 1010 carbon steel The joint is brazed at 950°C and is, for the worst-cas condition, suddenly air quenched in a room-temperature (20°C ambient.

The Fourier equation for independent radial heat conductio in long ceramic and metal cylinders is well known. The solutio of the cylindrical heat conduction problem consists of a infinite series. Each term of this series' is a product of Bessel's function and an exponential function, as given i textbooks on heat conduction. Data tables and master charts fo cylindrical heat diffusion have been compiled. See, e.g., 196

Gebhart's '-'Heat Transfer," McGraw-Hill, New York). With these equations, data tables, and master charts, one can determine the temperature profiles at different locations (i.e., radial positions, r, in a cylindrical end-to-end joint) at various time instants. From these temperature profiles, one can compute the associated, maximum transient mismatch stresses and strains.

The cooling down of a MACOR-metal joint from the brazing to room temperatures represents one of the most severe thermal changes, because of the wide temperature range involved. The step-by-step temperature changes, i.e., u^ and u s for the temperatures of MACOR and steel, respectively, at cooling time t in seconds, at the center, (r = 0) of the interfacial regions of a 5.08-cm diameter, rod-type MACOR-steel joint are given i Table 1. Other tables have also been prepared for cylindrical rods of different diameters.

The data used in the computations for Table 1 are: ro diameter D = 5.08 cm, rod radius r = 2.54 cm, surface hea transfer coefficient = 0.1 per inch (0.039 per cm) for bot steel and MACOR, thermal diffusivities =0.108 cπr/sec for stee and 0.0054 for MACOR, initial temperature of both MACOR an steel = 950°C, and final or room temperature = 20°C.

The computed data in Table 1 show, for the particular cas treated, the maximum temperature differential between MACOR an steel at the axial center point, (or r = O), i.e., . u = u_ u s , at different cooling times t in seconds. Thus, immediatel upon cooling after brazing (t = 0), this differential is zer because both the MACOR and steel are at the same brazin temperature of 950°C. Subsequently, faster cooling of the stee rod relative to MACOR causes this differential to increase wit time t, until both rods are significantly cooled when th temperature differential decreases. After 29,900 seconds (8. hours), for example , both rods are within a* few degrees of th room temperature at 20°C. The maximum temperature differentia reaches 775°C at about 1,000 seconds after the air cooling giving rise to the maximum or critical dynamic mismatch stres

Table 1: Nonsteady Heat Transfer Computations For a 5.08-cm MACOR-Steel Joint Cooling from 950°C to 20°C

u s -^s

950 0

947 3

935 14

901 48

867 82 835 113

804 144

731 217

665 282

456 478

316 703

220 681

155 729

112 756

82 769

62 773

39 765

23 708

22 643

22 528 21 436

21 358

21 295

21 199

21 134

21 91

20 42

20 19

20 7

20 3

20 l also shows that the temperature differential ^u = - u B reaches 113, 144,.217, 282, 478, and

703°C at 47.8, 59.8, 89.6, 119, 239, and 358 seconds, respectively, after the cooling from 950°C.

By comparison, the maximum temperature differential of 727°C at the axial center point of a 2.54-cm (or r = 1.27 cm) diameter MACOR-steel joint is reached sooner, at about 440 seconds, after cooling. The linear thermal expansion coefficients, f, are define as the thermal expansion per unit length per unit degre Centigrade. As given in the literature, they refer only to th static case. For a given material, these coefficients ar constants in given temperature ranges. Withi these ranges, they do not depend on specimen geometries, sizes, diffusivities surface characteristics, heating or cooling conditions, an initial and final temperatures.

The static thermal shrinkage (or negative expansion strain, e, for a given material is, by definition, the stati thermal expansion coefficient, f, multiplied by the temperatur range of cooling, A n, i.e., e = f x A u. Thus, for the stee rod, this strain is: e g = f s x _4u s . For the MACOR rod, it is e m = f x ^"ro¬ unde equilibrium conditions, the materials of the joint i.e., MACOR and steel, are supposed to be in constant therma equilibrium. That is, u m = u s . Both materials are thus at th same brazing temperature of U Q at the beginning of cooling (t 0) . Also, at any time t during the cooling, the coolin temperature range for MACOR and steel are always the same in th static case. Thus:

^ = u 0 " = u 0 " u s β ^ u s β u ' and the static expansion mismatch strain between steel and MACO is:

A e = e s - e m = (f s - f m j x A u = constant x A n. On the other hand, dynamic thermal expansion coefficients f*, and the resultant dynamic mismatch ' strains, e*, an stresses, s , strongly depend on the " joint materials geometries, sizes, physical and surface properties, and heati or cooling conditions.

Starting with zero strain on cooling from the brazin temperature of 950°C, the dynamic strain in the steel rod is e* s = f s x Δn s where j n s = 950 - u s , while that in the MACO rod is: e* m = f m x Δ ^ where ^x^ = 950 - . The difference in dynamic mismatch strain is: Δ e* = f s x ^u s - f m x ^u m . This strain reaches a maximum of about 0.0123 at t = 1,00 seconds for the 5.08-cm MACOR-steel joint. Such high strain exceed even the yield point of steel. The dynamic thermal expansion coefficient mismatch. A f can be computed by dividing the dynamic mismatch strain, e s e* m , by the average cooling temperature range, i.e.,^1^ = 950 (u s + U JJJ J/2. This dynamic coefficient mismatch, for the 5.08-c MACOR-steel rod joint cooling from 950°C to 20°C, still depend greatly on the cooling time t. It reaches a maximum rate o about 29.6 ppm/degree C at a cooling time of about 90 seconds but continuously drops down to less than 5.6 ppm/degree C at t 1,000 seconds, as can be computed from the data in Table 1. Th total dynamic coefficient mismatch over the temperature range o 930°C far exceeds the maximum of 100 ppm considered allowable b Hagy and Ritland. According to their criterion, cooling only few degrees would cause failures.

It can also be shown that the maximum dynamic expansio coefficient mismatch, - f* the 5.08-c MACOR-steel rod joint cooling two to fiv times greater than the corresponding mismatches for the stati or equilibrium case.

Statically, MACOR only marginally "matches" with a few low expansion metals such as the 42 Ni-Fe alloy. MACOR and stee joints now become totally "mismatched" dynamically for th above specimen configuration, size, and brazing conditions.

To approximately compute the dynamic mismatch stresses, on may further neglect the presence of the braze and the metallize layers, and use a Timoshenko approach as follows. Consider portion of the steel specimen of unit length and cross-sectiona

area, brazed together with a MACOR specimen of equal length an cross-sectional area. At time t = t after cooling from the brazing temperature of 950°C, the temperature of the steel is u g and A --- - 950 - u s , while the temperature of MACOR is u m an A ~ __ ~ 950 - U JJ .. The steel specimen has thus shrunk from uni length to 1 - f g xj g , while the MACOR to 1 - f ffi x A u^. Th steel has shrunk more than MACOR, since both f s and A u g ar greater than ^ and respectively. To maintain join integrity at the ends, the originally stress-free but overshrun steel must be stretched with dynamic tensile stress s s by th adjoining MACOR, to length y from length 1 - f s x ^ u g Simultaneously, the undershrunk MACOR must be compressed wit dynamic compressive stress s m by the steel, to the same lengt y from length of 1 - ^ x u^. 5 Hence, the tensile stress in the steel, s s , is s s* = E s x (Y " λ + f s x ^ n s / ( 1 * f s x -^ u s) where e s is the Young's modulus of steel, i.e., 30,000,000 ps

(2,109,000 kg/mm 2 ); while the compressive stress in MACOR, s m * is 0 s m* - Em ( -" f m x ^ " D'v 1 - f m x ■* > where E m is the Young's modulus of MACOR, i.e., 5,000,000 ps

(3,515,000 kg/mm 2 ).

Apparently, s s = s m . Hence: y=((l - f m x -4 )^ + (1 -f s ^ n s ) ) / (E s +E m ) 5 From these equations, one can compute and show that th common maximum dynamic mismatch stress in MACOR and steel, s * s g , reaches over 52,800 psi (3,712 kg/mm 2 ), well above MACOR' flexual strength of 15,000 psi (1,055 kg/mm 2 ) or even it comprehensive strength of 50,000 psi (3,515 kg/mm 2 ). Similarly, . dynamic or transient differences i temperatures, thermal expansion coefficients, thermal expansio strains, and thermal mismatch stresses have been computed f differently sized cylindrical MACOR-steel joints, at vario radial locations and cooling time instants. The dynam

- > -> mismatch stresses and strains are all unexp.ectedly hig

Similarly, . unevenly heated or cooled joints of even simila materials can also give high dynamic stresses and strains

Measures must therefore be taken to reduce the dynamic mismatch stresses on the relatively weak ceramic or one of th similar materials so that the ceramic or this one material is n longer subjected to the high stresses. This reduction can b achieved by, e.g., reducing the temperature differential o absorbing a major portion of the dynamic mismatch stresse normally present in the ceramic through the use of a soft yieldable metallic braze. These measures prevent the braze joint failures particularly from these dynamic mismatc stresses. This is because residual or actual mismatch stres between the two joined materials is the theoretical mismatc stress with a portion thereof absorbed in the metallized o brazed layer.

The following methods, used singly or in combination, wil minimize or neutralize these high mismatch stresses and strains

1) Using a soft, yieldable metal layer to braze th metallized ceramic to the metal, and to absorb within the braz layer a large or major portion of these mismatch stresses s that the relatively weak MACOR or other ceramic is no longe subjected to high stresses thereby preventing fractures;

2) Grading radially of (i.e., parallelly to rather tha perpendicular to) the bonding interfacial region as to th thermal conductivity, expansion coefficient, and softness o tensile strength of the braze metal, to ensure that the maximu residual mismatch stress, after absorption in the braze, wil not exceed the local material strength in the joint at any poin and time; and 3) In combination with the radial grading, grading axiall or normally of the bonding interfacial region, from one soli material to the other, the thermal expansion' coefficient of-th braze layer to minimize direct mechanical interaction betwee the two materials. This axial grading, practiced alone, is old

The first two methods are achieved by providing a nove composite metallic braze disc used for joining the metallize ceramic cylinder 11 to the metal cylinder 12 to form the join 10. The ceramic cylinder has a metallized layer 13 at its lowe end. The composite braze disc has a soft, pure copper centra core 14 within the opening of an outer harder copper allo (e.g., 70:30 Zn Cartridge brass) ring or washer 15. The entir metallized layer 13 and the composite braze disc 14 and washe 15 form the joint 10 or bonding interfacial region. The linea thermal expansion coefficient of pure copper is 16.5 ppm/°C while that of 70 Cu:30 Zn Cartridge brass is 19.9 ppm/°C. Also the tensile strength of the brazing-annealed, soft pure coppe is only 15,000 psi (1,055 kg/mm 2 ), while that of the 70:3 Cartridge brass is over 40,000 psi (2,812 kg/mm 2 ), or .abou three times greater.

These combinations of linear thermal expansion and tensil properties achieve the required results. In a ceramic-stee joint, the maximum or critical transient mismatch temperatures thermal strains and stresses occur in the axial centers of th interfacial regions. We therefore have dead soft, brazing annealed, pure copper at the core regions which are highly an easily yieldable to absorb most of the dynamic mismatch therma strains and, therefore, stresses. Pure copper also ha relatively low thermal expansion to reduce these mismatc effects in the first place. In addition, the pure copper is good thermal conductor, equalizing the temperature between th centers, as well as their outer and regions, of the joint t further minimizes mismatch strains and stresses.

On the other hand, the outer peripheral regions of t braze disc is made of relatively highly expansive but the l thermal-conducting brass. At these peripheral regions, t mismatch temperature differentials are relative small. T higher tensile strength is even desirable at the peripher regions to enhance the joint strength.

This composite braze disc design will thus provide t required radially tailor-graded profiles of braze compositio thermal expansion coefficient, braze softness, and therm conductivity needed to overcome the critical dynamic mismat stresses in, e.g., the preforms for electronic device packages. The composite braze discs can be made by, for exampl metallurgically cladding; or mechanical press-forming a sphe and a washer, at least two layers or two tubes, or oth combinations together into a single layer. Elemental interdif usion during the braze manufacture brazing operation, or special pre- or post-brazing heat treatments can modify or provide any reasonable compositi profiling in the braze discs for even improved results.

It is possible to prepare the specially graded joint 10 Figs. 1-2 in two separate processing steps, i.e., metallizing the ceramic 11 and brazing with the special braze disc copper 14 a d brass ring 15. It is also possible to combine, in a sing processing step, the ceramic metallizing and brazing. The sa specially graded brazing disc and ring is useful. I metallizing with 10-20 w/o of Fe in W or Mn in Mo, or the oxides in a reducing (e.g., hydrogen or 10-45 % hydrogen formi gas) atmosphere at, e.g., 800-1500°C, the solubilities of bra copper, brass, iron, manganese, or nickel alloys in the W/M based metallized layers are limited, even at the hi metallizing temperatures. The above-described special gradi effect therefore still exist. The boundaries between t metallized layer 13, braze disc 14, and braze ring 15 ma however, be less distinct or even practically disappear.

Since M0O3 melts at 801°C, special metallizing mixtur of this compound can be used together with low-zinc (or ti copper alloy disc and high-zinc (or tin) copper alloy ring achieve better te pearture matches and phase compatibiliti between these materials. The single-step metallizing-brazi step can be done at temperatures as low as about 800°C.

If all* these measures still do not prevent dynamic thermal mismatch failures, the common axial elemental grading or sudden composition changes may be added. One method consists of providing a disc of low-expansive metals such as Sylvania #4, Dumet, 50% nickel alloy, chrome-iron stainless, platinum, Sealmet, and titanium placed intermediately between the steel and the copper braze. In this way, the ceramic MACOR is mechanically isolated from the highly expansive steel. The desired axial elemental profiling can also be achieved through controlled diffusion.

Skilled persons can, of course, select other soft metals such as gold, silver, tin, lead, indium, zinc, or even iron or nickel, or other materials to replace copper, and select other chemical elements to replace the copper-strengthening zinc (or tin). The resultant new alloys will, of course, be different i compositions, strengths, softnesses, or shock-absorbin abilities, diffusivities, thermal conductivities, melting o softening points, and other properties.

The invention, as described above, is not to be construe as limited to the particular forms disclosed herein, since thes are to be regarded as illustrative rather than restrictive. Various combinations, equivalent substitutions, or othe modifications of the preferred embodiments described herein ar obviously possible in light of the description, withou departing from the spirit of the invention. In. particular other ceramics including alumina, zirconia, silicon carbide silicon nitride, carbon, ..., may be used instead of MARCO with the same or modified metallizing compositions.