June 2011

NASA/TM—2011–215876

Failure Assessment Diagram for Brazed 304 Stainless 

Steel Joints

Yury Flom

Goddard Space Flight Center, Greenbelt, MD
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*- Brush and Wellman trade name for 62%Be 38%Al metal matrix composite 

FAILURE ASSESSMENT DIAGRAM FOR BRAZED 304 STAINLESS STEEL JOINTS 

ABSTRACT 

Interaction equations were proposed earlier to predict failure in Albemet 162* brazed 

joints. Based on the test results of this study, it was determined that the same 

interaction equations can be used for lower bound estimate of the failure criterion in 304 

stainless steel joints brazed with silver-based filler metals as well as for construction of 

the Failure Assessment Diagrams (FAD). 

1.0 INTRODUCTION 

Prediction of failure in brazed joints subjected to complex loading conditions continues 

to challenge designers and structural analysts attempting to estimate margins of safety 

in critical assemblies fabricated by brazing. Despite the fact that brazed components 

and structures are extensively used in the aerospace industry, literature is lacking 

simple engineering procedures or guidelines for failure assessment of brazed joints. 

Earlier work [1] demonstrated that interaction equations could be used for failure 

assessment of Albemet 162 joints brazed with AWS BAlSi-4 (88%Al,12%Si) filler metal. 

In the current effort, different base / filler metal combinations consisting of 304 SS 

brazed with pure silver and AWS BAg8 (78%Ag, 28%Cu eutectic) filler metals were 

evaluated. 

This memorandum provides information on specimen design, fabrication and testing 

performed by the Materials Engineering Branch to verify that interaction equations used 

earlier could also be applied for failure assessment of other brazed systems. 

2.0 APPROACH 

This effort consisted of two parts. In first part design values of tensile σo and shear o

strength (allowables) were determined. These allowables were used in interaction 

equation (1). The purpose of the second part was to verify that this equation can be 

used as a lower bound FAD for 304SS-silver brazed joints. A semi-empirical interaction 

equation proposed in [1] is shown below: 

ߪ

௢ߪ

൅ 

߬

߬௢

 ൑ 1 (1) 
2 

In this equation σ and  are the maximum normal tensile and shear stresses acting on 

the filler metal layer in the brazed joint. σo was determined experimentally. Since the 

shear strength of stainless steel lap joints brazed with silver-based filler metals is a well 

established quantity [2-4], it was more cost effective to omit fabrication and testing of 

the lap shear specimens. Instead, it was decided to accept o = 15 ksi (103.5 mpa). 

This value is listed as the shear strength allowable in [4]. 

Equation (1) was validated by testing specially designed brazed specimens. Tensile and 

shear stresses acting on the brazed joints were calculated using engineering mechanics 

of materials, as described in Appendix. 

3.0 EXPERIMENTAL PROCEDURE 

3.1 Test Specimens 

All test specimens were fabricated from cold rolled 304SS bars purchased from 

McMaster Carr. In addition to brazed specimens, several blank tensile specimens were 

fabricated and tested to compare the properties of the base metal with the properties of 

the brazed joints. There were two types of specimens designed to subject the brazed 

joints to a combined action of shear and normal stresses: double scarf and V-type. A 

total of five different configurations of specimens used in this effort are shown in Figure 

1. Fabrication sequence of a double scarf specimen is shown in Fig.2. Figures 3-5 show 

photographs of various test specimens at different phases of fabrication process. 

The double scarf geometry reduces the tendency of specimen to rotate during tensile 

test. Also, since the double scarf test specimen has two geometrically identical brazed 

joints and the failure occurs in only one joint, each tested specimen has one brazed joint 

still intact. This allows for a metallographic examination of the brazed joint that 

experienced the condition of imminent failure. The V-type specimen geometry [5] 

eliminates rotation, provides fully axisymmetric loading conditions while subjecting the 

brazed joint to a combined tension and shear load. Test specimen identifications are 

explained in Table 1. 

Brazing was performed in a vacuum furnace. All specimens were electrolitically Ni 

plated prior to brazing. One set of specimens was brazed using AWS BAg-0 (pure 

silver) and another set was brazed with BAg-8 (silver-copper eutectic) filler metals. Filler 

metal foils were preplaced between the faying surfaces. For simplicity, AWS BAg-0 filler 

metal is referred to in the rest of the text as Ag. Typical time-temperature records of the 

brazing cycles are shown in Figure 6. After brazing, all specimens were machined into 

the standard round tensile test coupons, as shown in Fig.5 
3 

Table 1. Test Specimen Identification 

Spec. ID What it means 

ButtAg-1 Specimen # 1 butt brazed using Ag filler metal 

ButtBAg8-1 Specimen # 1 butt brazed using AWS BAg8 filler metal 

D60Ag-2 Specimen # 2, double 60 ̊ scarf joint, brazed with Ag filler metal 

D45BAg8-3 Specimen # 3, double 45 ̊ scarf joint, brazed with AWS BAg8 filler metal 

V60Ag-1 Specimen # 1, 60 ̊ V-shape joint, brazed with Ag filler metal. 

304SS-Ag Base metal tensile coupon exposed to Ag brazing cycle 

Fig.1 Geometry of the test specimens used in this effort. Base metal blanks (a) were 

tested to establish the property baseline. Butt brazed (b), V60 (c) and double scarf 

60 ̊(d) and 45 ̊(e) were tested to determine the failure loads used to calculate normal 

and shear stresses at failure acting on the braze layer. 
4 

Fig. 2 Showing schematic of scarf specimen fabrication sequence. 

3.2 Mechanical testing 

All specimens, including the blank ones, were tested on Instron 4115 test frame using a 

crosshead speed of 0.05 in/min. A 1” gage length extensometer was used to record the 

elongation of the test specimens. In case of the double scarf specimens, the 

extensometer was recording the elongation across only one joint. When elongation 

reached approximately 1%, extensometer was removed to avoid possible exposure to 

the shock event during brazed joint failure, and the Bluehill 2 testing software continued 

to acquire the load / displacement record. This feature allows for an uninterrupted plot of 

the entire test up to the failure point. Figures 7-8 show typical stress-strain curves for all 

specimens tested in this effort. Prior to tensile testing, the stainless steel blanks were 

exposed to the same brazing cycle time/temperature conditions as the brazed 

specimens. Typical appearances of the fractured specimens are shown are Figure 9. 

3.3 Metallographic examination 

The remaining brazed joints in each 45 ̊ and 60 ̊ double scarf specimens tested at the 

highest strength were cross sectioned and metallographically polished, as schematically 

shown in Figure 10. Metallographic cross sections were examined under optical 

microscope to observe the condition of the braze filler metal interlayer immediately prior 
5 

to fracture. Various microstructural features observed on the metallographic cross 

sections and discussed in the following section are shown in Figures 11 through 14. 

Fig.3. Top view shows the alignment features on the interfaces of male and female 

halves forming the butt-brazed specimens. These features are similar to the ones 

suggested in European standard EN 12797 entitled “Brazing-Destructive tests of brazed 

joints”. The filler metal foil is placed between the interfaces prior to brazing. Bottom 

view shows as brazed butt joint prior to machining. 
6 

Fig. 4 A close-up of the V60 joints in as-brazed condition. 

Fig.5. Top view shows one of the D45 type specimens after machining. A close-up view 

of the brazed joints in the same specimen is shown at the bottom. 
7 

4. RESULTS AND DISCUSSION 

4.1 Mechanical Testing 

As one can see, all specimens brazed with AWS BAg8 filler metal failed at higher loads 

than their counterparts brazed with pure silver (Fig.7, 8). Two factors are responsible for 

this difference. One factor is related to the difference in properties between 304SS base 

metal exposed to two different braze cycles. 304SS is considerably weaker and more 

ductile after Ag braze cycle than after the BAg8 one. The annealing temperature for 

304SS is somewhere between 1010 ̊ and 1121 ̊C [6]. Consequently, the Ag brazing 

cycle (see Fig. 6) brings the 304SS base metal very close to fully annealed condition. 

 

Fig. 6 Brazing cycles for Ag-brazed (left) and BAg8-brazed test specimens. 

Another factor is related to the difference in strength between Cu-Ag alloy and pure Ag 

due to solid solution hardening effect. It is expected that the silver-copper alloy (AWS 

BAg8) would have higher strength than unalloyed silver. For example, tensile strength 

of annealed silver-copper eutectic is reported somewhere between 40 and 44 ksi (276 – 

304 Mpa), whereas typical tensile strength of pure silver is only around 18.2 ksi (130 

Mpa) [7,8]. The difference in strength between the joints brazed with AWS BAg8 and Ag 

filler metals is evident from the engineering tensile stress vs. strain plots shown in Fig.7. 

For clarity, only the data from the specimens that showed the highest strength are 

present on the plots. All specimens, regardless of the joint geometry, demonstrated 

much higher strengths than the strengths of their respective filler metals. This is a well 

known property of butt brazed joints [9,10]. As one can see from the plots, scarf and Vtype brazed joints also behave in a similar manner. As a general observation, Agbrazed joints displayed higher ductility than their counterparts brazed with silver-copper 

eutectic with the exception of the 45 ̊ double scarf joints, which showed the same 

ductility 
8 

a) 

b) 
9 

c) 

d) 
10 

e) 

Fig.7. Stress-strain curves comparing BAg8- and Ag-brazed test specimens. All 

specimens regardless of configuration and brazed with BAg8 filler metal demonstrated 

higher strength than their pure silver brazed counterparts. 

regardless of the filler metal. Fig. 8 shows the same plots grouped in accordance with 

the filler metal used. Ag-brazed joints essentially follow the deformation behavior of the 

304SS base metal. It appears that D60 Ag scarf joints required higher stresses to 

sustain their plastic deformation compared to the rest of the joints including the base 

metal. BAg8 – brazed specimens also follow the base metal stress strain curve. In this 

case, however, the scarf joints yielded earlier than butt- and V-brazed joints as well as 

the base metal itself. It appears that behavior of the scarf-brazed joints was not 

consistent in terms of their yield onset. Yielding could occur either below (Fig.8a) or 

above (Fig.8b) their respective base metals. This observation is most likely due to an 

experimental artifact caused by a complex interaction between the slip along the braze 

interfaces and extensometer readout. 

Results of all mechanical tests are presented in Table 2. Selection of the brazed joint 

tensile strength allowable σo was based on test results of butt Ag specimens. 

Comparing the test results of the butt joints, BAg8-brazed specimens showed 

approximately 50% higher strength than their Ag-brazed counterparts. Consequently, 
11 

a) 

b) 

Fig. 8. The same plots as in Fig.7 only grouped according to the AWS BAg8 (a) and Ag 

(b) filler metals. All specimens demonstrated higher strength than their respective filler 

metals. 
12 

Table 2. Test Results 

* Determined by dividing maximum load by the initial cross sectional area 

** Rσ = ߪಳ

ߪ ௢ൗ ;  R = ߬ಳ

௢ ൗ where  σo = 38.4 ksi (265 Mpa); o = 15 ksi (103.5 Mpa) 

Spec.ID

Ultimate

Tensile

Strength*

TUS

Normal stress

braze plane

σB

Shear stress

braze plane

B

Normal

stress

ratio**

Rσ

Shear

stress

ratio**

R

Ksi Mpa ksi Mpa ksi Mpa 

ButtAg-1 53 363 53 366 0 0 1.4 0.0 

ButtAg-2 48 332 48 331 0 0 1.3 0.0 

ButtAg-3 46 316 46 317 0 0 1.2 0.0 

ButtBAg8-1 75 517 75 518 0 0 2.0 0.0 

ButtBAg8-2 71 490 71 490 0 0 1.8 0.0 

ButtBAg8-3 75 515 75 518 0 0 2.0 0.0 

D60BAg8-1 61 420 35 238 20 137 0.9 1.3 

D60BAg8-2 61 424 37 254 21 146 1.0 1.4 

D60BAg8-3 57 395 34 233 19 134 0.9 1.3 

D60Ag-1 46 318 46 316 26 182 1.2 1.8 

D60Ag-2 49 335 46 316 26 182 1.2 1.8 

D60Ag-3 45 313 43 295 25 170 1.1 1.6 

D45Ag-1 41 280 21 141 21 141 0.5 1.4 

D45Ag-2 40 274 20 138 20 138 0.5 1.3 

D45Ag-3 38 265 19 131 19 131 0.5 1.3 

D45BAg8-1 55 377 28 190 28 190 0.7 1.8 

D45BAg8-2 59 410 30 204 30 204 0.8 2.0 

D45BAg8-3 56 387 28 193 28 193 0.7 1.9 

V60Ag-1 48 333 36 248 21 143 0.9 1.4 

V60Ag-2 46 319 35 238 20 137 0.9 1.3 

V60Ag-3 52 361 39 269 23 155 1.0 1.5 

V60BAg8-1 67 462 50 347 29 200 1.3 1.9 

V60BAg8-2 64 441 48 331 28 191 1.3 1.8 

V60BAg8-3 54 375 41 279 23 161 1.1 1.6 

304SS-Ag 90 620 

304SS-BAg8 100 690 
13 

pooling the data would not be statistically justified. Since we are interested in 

determining the lower bound failure criteria, the logical choice would be to use Agbrazed butt specimen showing the lowest strength. A simple, 3 sigma statistical analysis 

was performed to estimate σo : 

 σo  = AVG‐3×SIGMA  = ((48+53+46)/3)‐3×STDEV= 38.4 ksi (265 Mpa),

as indicated in the foot note of Table 2. A rational for selecting shear allowable o = 15 

ksi (103.5 Mpa) is provided earlier in Section 2. 

 a) b) c) 

 d) e) f) 

Fig. 9. Shows fracture surfaces of butt (a), scarf (b, c) and v (d) joints after pull test. 

Some scarf joints, like the ones shown in e) and f), had discontinuity areas that most 

likely contributed to a reduction in their strength and/or ductility, although no effort was 

made in this study to correlate a lack of braze with reduction in properties. 
14 

a) 

b) 

Fig. 10. The view (a) is showing metallographic preparation sequence of the scarf joint 

specimens after mechanical tests. The brazed joint which is still intact (left) is cut from 

the specimen, cross sectioned and polished. The unetched polished surfaces were 

examined under optical microscope. Typical polished cross section of 45 ̊ scarf joint is 

shown in view (b). 

4.2 Metallographic Examination 

Examination of the metallographic cross sections revealed a number of interesting 

features, as noted below: 

 High integrity of the braze layer. 

Since the cross sectioned brazed joints were exposed to the failure loads, it was 

expected to find severe voiding and cracking in the braze layer indicating imminent 

fracture. However, the microstructures in both Ag- and BAg8-brazed scarf joints appear 

to be fundamentally sound, showing no evidence of extensive damage such as massive 

micro-voiding and void linking typically observed in ductile metals immediately prior to 

failure. The shape and location of the voids found in the braze layers point to their pre- 
15 

 

a) b) 

 

c) d) 

Fig. 11. Images of various locations in the cross section of the D60Ag scarf joint after 

exposure to failure load. Image (a) is typical. No voids were observed except at the very 

edge of the joint (b). The shape of the void points to the shear deformation-related 

origin. Images (c) and (d) show the same edge. Amount of shear strain can be 

estimated as 72/44 = 164%! All dimensions are in microns. 

existing nature. It appears that shear plastic deformation of the braze layer resulted in 

expansion and stretching of pre-existing brazing flaws. 

 Uniform shear deformation 

Both types of Ag- and BAg8-brazed scarf joints displayed a fairly uniform plastic shear 

strain along the braze plane. It was very surprising to see typical shear strains in either 

45 ̊ or 60 ̊ scarf joints in excess of 100%! This indicates tremendous resilience of the 

304SS/Ag and BAg8-brazed joints and their ability to undergo large plastic deformation 
16 

 

a) b) 

 

c) d) 

Fig. 12 Cross section of the D45Ag scarf joint. Most of the braze is free from voids and 

cracks (a). This joint has also experienced very large shear strains (b) that can be 

estimated as 65/40 = 163%. A number of voids could be seen along the braze (c, d). 

These voids appear to form during brazing and became distorted, tilted and stretched 

during the pull test. All dimensions are in microns. 

prior to failure. A uniform shear strain within the braze layer may also be indicative of 

the end-to-end uniformity of von Mises stress in the scarf brazed joints tested in this 

effort. This can be explained by the small aspect ratios (length of the brazed interface 

divided by the specimen diameter) of 60 ̊ and 45 ̊ scarf joints. These ratios are 1/cos30 ̊

and 1/cos45 ̊ or 1.15 and 1.41 respectively. These values are very close to the aspect 

ratio of 1 for the lap shear joints having overlap length equal to the thickness of the base 

metal T. Such joints are commonly referred to as having 1T overlap. Stress analysis of 

stainless steel lap joints brazed with silver filler metal showed that von Mises stress 

distribution within 1T joints was also quite uniform [3]. Such similarity between scarf 

joints and lap shear joints indicates that behavior of the 45 ̊ and 60 ̊ scarf joints is 

dominated by shear. This is consistent with earlier observations of the behavior of the 

scarf joints for different base metal / filler metal combinations [5]. 
17 

 

 a) b) 

 

 c) d) 

Fig. 13. Cross section of D60BAg8 scarf joint showing more voids. Judging from the 

shape of these voids, it appears that majority are pre-existing braze defects (a-c) that 

became deformed during tensile test. Some of the voids, however, were most likely 

caused by fracture of the filler metal (d). Note that the copper rich phase is tilted during 

shear deformation within the filler metal. The amount of this tilt or shear strain is fairly 

uniform from end-to-end of the brazed joint. All dimensions are in microns. 

 Benefit of multiphase system 

A comparison between the BAg8 and Ag metallographic cross sections in 45 ̊ and 60 ̊

scarf joints illustrates that a two phase microstructure in BAg8-brazed joints is a much 

better indicator of the plastic flow than a relatively featureless structure of the silver 

interlayer. It is evident how the copper-rich phase aligns itself with the shear flow 

pattern, see Figs. 13 -14. These uniform tilt patterns suggest end-to-end uniformity of 

shear strain and shear stresses in the scarf joints. Consequently, a ductile multiphase 

structure is very useful in studying plastic deformation in brazed joints. 

4.3 Failure Assessment Diagram (FAD) 

Stress ratios presented in Table 2 are plotted in Fig.15. All Ag-brazed joints are denoted 

by solid symbols and all BAg8-brazed joints are shown with open symbols. A line 

connecting points with coordinates (1, 0) and (0, 1) represents interaction equation 

equation (1) or 
18 

Rσ + R = 1, where Rσ = ߪಳ

ߪ ௢ൗ and R = ߬ಳ

௢ ൗ are tensile and shear stress ratios. 

It is quite clear that this line is very conservative and quite adequate to be used as lower 

bound FAD even for 304SS brazed with pure silver. All BAg8-brazed joints tested higher 

than the Ag-brazed ones and, therefore, are located further away from the FAD line. 

Since it is fairly safe to say that pure silver has the lowest strength among the rest of the 

high temperature silver-based filler metals, it can be concluded that interaction equation 

Rσ + R = 1 

can be used as lower bound FAD for 304SS brazed with high temperature silver-based 

filler metals. 

 

 a) b) 

 c) d) 

Fig.14 Shows typical (a) optical image of the brazed joint region in tested D45BAg8 

specimen as well as the brazed joint edge (b). Extent of plastic deformation can be 

ascertained from the shear strain estimated as 30/25 = 120%. A copper-rich second 

phase is tilted by the shear deformation in the filler metal. The angle of this tilt is the 

same from end-to-end which indicates shear strain uniformity within the brazed joint. 

Most of the voids seem to be related to pre-existing brazing flaws (c, d). All dimensions 

are in microns. 
19 

Fig.15 Plot of the stress ratios at failure for all tested specimens. Lap shear test results 

from the previous investigations [2,3] are marked with “x” and denoted with asterisk in 

the legend. As one can see all experimental results are located noticeably far away from 

the lower bound FAD. Consequently, the region inside the FAD line can be considered 

a “safe” zone. 

Acknowledgments 

The author would like to thank D. Thomas of D&L Engineering for machining the test 

specimens as well as D. Puckett, M. Powell and D. Kolos, NASA GSFC, for their 

assistance in performing mechanical testing, metallographic examination and technical 

editing. Special thanks go to Mr. Thomas Deming, ATK, for his technical comments. 

The author also would like to thank Dr. G. Alcorn, NASA GSFC, for sponsoring this 

effort. 

0.0

0.5

1.0

1.5

2.0

2.5

0.0 0.5 1.0 1.5 2.0 2.5

R

σ

R

FAILURE ASSESSMENT DIAGRAM

D45BAg8

ButtAg

Lap Shear*

V60Ag

D45Ag

Rσ + R = 1

Safe zone 
20 

Appendix 

Stresses acting on the plane of the brazed joints were hand calculated using 

engineering mechanics. Figure 16 shows forces acting on the portion of the tensile test 

specimens adjacent to the brazed joint. 

 

Fig. 16 Force equilibrium diagram used to estimate normal and shear stresses acting on 

the braze plane. 

From equilibrium conditions we can write: 

1) fn • cos - fs • cosα = 0 along X axis and 

2) F - fs • sinα - fn • sin = 0 along Y axis, α +  = 90 ̊

Solving these equations for fn and fs will result in: 

fn = F•cosα = F•sin and fs = F•sinα = F•cos

Average normal and shear stresses acting on brazed joint can be estimated as: 

σb = fn / Braze area and b = fs / Braze area. Braze area is an ellipse. Its area equals 

to: Braze area ܣ ൌ ሺߨ • ݀ • ݈ሻ/4 , where l can be calculated as d/sin ; For example, for 

D60Ag-1 joint that failed at F=4432 lbs (19700 N);  = 60 ̊ we get: 

σb = (4432 •sin60 ̊)/((3.14 • 0.350 • (d/sin60 ̊)/4) = 35 ksi (242 Mpa). 
21 

References 

1. Flom, Y., Wang, L., Powell, M. M., Soffa, M. A., and Rommel, M. L., 2009. 

Evaluating Margins of Safety in Brazed Joints, Welding Journal, 88(10): 31-37. 

2. Bredzs, N, and Miller, F., 1968. Use of the AWS standard shear test method for 

evaluating brazing parameters. Welding Journal, 47(11): 481-s to 496-s. 

3. Flom, Y., Wang, L., 2004, Flaw Tolerance in Lap Shear Brazed Joints-Part 1. 

Welding Journal, 83(1): 32-s to 38-s. 

4. Metallic Materials Properties Development and Standardization (MMPDS), April 

2005 

5. Peaslee, R.L., The Brazing Test Specimen – Which One? Welding Journal 

55(10): 850-858 

6. Carpenter Stainless Steels – Selecting Alloy Data Fabrication, Carpenter 

Technology Corporation 2/87/15M, 1987 

7. Metals Handbook, 9th edition, Vol.2, ASM, 1979 

8. H.E. Cline and D. Lee, “Precious Metals: Science and Technology”, The 

International Precious Metals Institute, 1991, p.645 

9. Handy & Harman, “Easy-Flo and Sil-Fos” Bulletin 20, 1950, p.8 

10. Brazing Handbook, 5th edition, AWS, 2007 
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09-06-2011 Technical Memorandum

Failure Assessment Diagram for Brazed 304 Stainless

Steel Joints

Yury Flom

Goddard Space Flight Center

Greenbelt, MD 20771

National Aeronautics and Space Administration

Washington, DC 20546-0001

NASA

TM–2011-215876

Unclassified-Unlimited, Subject Category: 31

Report available from NASA Centerfor Aerospace Information, 7115 Standard Drive, Hanover, MD 21076. (443) 757-5802

Interaction equations were proposed earlier to predict failure in Albemet 162 brazed joints. Present study demonstrates that the same

interaction equations can be used for lower bound estimate of the failure criterion in 304 stainless steel joints brazed with silver-based

filler metals as well as for construction of the Failure Assessment Diagrams (FAD).

Brazing, Structural Analysis, Interaction Equations, Failure Assessment Diagrams

Unclassified Unclassified Unclassified Unclassified 21

Yury Flom

(301) 286-3274