Jonathan A. Salem

Glenn Research Center, Cleveland, Ohio

Slow Crack Growth and Fracture Toughness

of Sapphire for the International Space Station

Fluids and Combustion Facility

NASA/TM—2006-214023

January 2006
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Jonathan A. Salem

Glenn Research Center, Cleveland, Ohio

Slow Crack Growth and Fracture Toughness

of Sapphire for the International Space Station

Fluids and Combustion Facility

NASA/TM—2006-214023

January 2006

National Aeronautics and

Space Administration

Glenn Research Center
Acknowledgments

Thanks to Robert Polvani and David Black of National Institute of Standards and Technology (NIST),

Manufacturing Metrology Division, Gaithersburg, MD for the X-ray topography. Thanks to

Shelly Wiederhorn of NIST for many useful discussions. Thanks to William Scott of the

University of Washington for many useful discussions and for etching work. Thanks to

R.O. Ritchie of the University of California, Berkeley for supplying a fractured m-plane test

specimen. The Microgravity Sciences Fluids and Combustion Facility funded the work.

Available from

NASA Center for Aerospace Information

7121 Standard Drive

Hanover, MD 21076

National Technical Information Service

5285 Port Royal Road

Springfield, VA 22100

Available electronically at http://gltrs.grc.nasa.gov
NASA/TM—2006-214023 1 

Slow Crack Growth and Fracture Toughness of Sapphire for the 

International Space Station Fluids and Combustion Facility 

Jonathan A. Salem 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Summary 

The fracture toughness, inert flexural strength, and slow crack growth parameters of the r- and 

a-planes of sapphire grown by the Heat Exchange Method were measured to qualify the sapphire for 

structural use in the International Space Station. The fracture toughness in dry nitrogen, KIpb, was 

2.31±0.12 MPa√m and 2.47±0.15 MPa√m for the a- and r-planes, respectively. Fracture toughness 

measured in water via the operational procedure in ASTM C1421 was significantly lower, KIvb = 

1.95±0.03 MPa√m, 1.94±0.07 and 1.77±0.13 MPa√m for the a- , m- and r-planes, respectively. The mean 

inert flexural strength in dry nitrogen was 1085±127 MPa for the r-plane and 1255±547 MPa for the 

a-plane. The power law slow crack growth exponent for testing in water was n = 21±4 for the r-plane and 

n ≥ 31 for the a-plane. The power law slow crack growth coefficient was A = 2.81×10–14 m/s⋅(MPa√m)–n

for the r-plane and A ≅ 2.06×10–15 m/s⋅(MPa√m)–n for the a-plane. The r- and a-planes of sapphire are 

relatively susceptible to stress corrosion induced slow crack growth in water. However, failure occurs by 

competing modes of slow crack growth at long failure times and twinning for short failure time and inert 

environments. Slow crack growth testing needs to be performed at low failure stress levels and long 

failure times so that twinning does not affect the results. 

Some difficulty was encountered in measuring the slow crack growth parameters for the a-plane due 

to a “short” finish (i.e., insufficient material removal for elimination of the damage generated in the early 

grinding stages). A consistent preparation method that increases the Weibull modulus of sapphire test 

specimens and components is needed. This would impart higher component reliability, even if higher 

Weibull modulus is gained at the sacrifice of absolute strength of the component. The current 

specification frequently used for the preparation of sapphire test specimens and components (e.g., a 

“60/40” scratch-dig finish) is inadequate to avoid a “short” finish. 

1. Introduction 

The U.S. “Destiny” laboratory module resides within the International Space Station and contains six 

science facilities or “racks.” One of these facilities is the Fluids and Combustion Facility (FCF), which 

will be used to study combustion and fluids processes in microgravity. Within this facility is the 

Combustion Integrated Rack (CIR). This rack consists of a combustion chamber circumferentially 

surrounded by diagnostics, which image the interior of the chamber through eight viewing “windows” 

shown in Figure 1. Previously, such windows were made from quartz, however, due to its superior optical 

properties in certain infrared ranges, sapphire is being considered for the windows. 

Sapphire fails from a combination of slow crack growth and twinning, and the interaction between the 

mechanisms is unknown. The dominant mechanism is likely dependent on the temperature, environment, 

relative stress level and surface finish. Twinning can be induced with compression or tension, and thus all 

surfaces of test specimens and components need to be polished adequately. The chamber and windows are 

required to tolerate 90 percent RH (relative humidity) during storage and transportation, and 75 percent 

RH atmosphere in the International Space Station. The reported properties of sapphire vary substantially 
NASA/TM—2006-214023 2 

Figure 1.—(a) Cross section of Combustion Integrated Rack chamber assembly and (b) the aluminum window chamber. 

for the r-plane. Because the application requires highly polished windows to be exposed to low stresses in 

low temperature, high humidity air, the slow crack growth (SCG) parameters and fracture toughness of 

the candidate sapphire were measured in water on the a- and r-planes. 

2. Experimental Procedure 

2.1 Material 

Flexure specimens were ground from sapphire grown by the heat exchange method (HEM)1 and 

mechanically diamond polished2

 on the tensile face and sides to a 60/40 scratch dig specification (Ref. 1) 

by using a random circular motion in order to provide an optical quality surface. The compressive face of 

the test specimens was left with a ~320 ground finish. Several sets of specimens were made with a-, m-, 

or r-plane cross sections. For the test specimens with an r-plane cross section, the sides corresponded to 

the a-plane. For the specimens with a- and m-plane cross sections, the tensile surface corresponded to the 

c-plane. These orientations will be referred to as “a-plane,” “m-plane,” and “r-plane,” respectively, as 

shown in Figure 2. The specimens were oriented within ±2° of the plane specified. During test specimen 

preparation, the manufacturer requested that some of the test specimen bevels be increased to 0.250 mm 

from the 0.125 mm specified because difficulty was encountered in eliminating chips from the bevels. 

The request was approved and the test results have been corrected for the small change (~2.9 percent) in 

the stress state. 

2.2 Elastic Properties 

The elastic moduli of some of the flexural specimens were measured by using the impulse excitation 

technique (Ref. 2). Test specimens were supported at the nodal points with foam strips and the specimen 

impacted lightly. The resultant impact induced a flexural vibration that was detected with a microphone 

and automatically recorded with an electronic circuit. The elastic moduli were then calculated from the 

specimen geometry and measured resonant frequency assuming a Poisson’s ratio of 0.26, which 

corresponds to that of polycrystalline alumina. The elastic constants of Teft (Ref. 3) indicate that 

Poisson’s ratio for the c-plane in standard orientation is ν = -S12/S11 = 0.30. The elastic moduli in the a-, 

m- and r-directions were 430±5, 426±4 and 345±2 GPa respectively, in good agreement with the data of 

Teft (Ref. 3). The individual results are summarized in Table A1 of the appendix. 

 

1

Crystal Systems, Salem, MA. 

2

Insaco, Quakertown, PA.

(a) (b) 100 mm
NASA/TM—2006-214023 3 

Figure 2.—Test specimen orientation and sapphire crystal showing mineralogical and Miller index 

notation. The c-axis is a 3-fold symmetry axis, but sapphire is indexed as a hexagonal unit cell with 

c/a = 2.730. 

2.3 Hardness 

The hardness of the c-plane was measured by using Knoop and Vickers indentation. For Knoop 

hardness testing, a load of 0.1 kg was applied for 15s. Occasionally, cracking from the indentation 

corners occurred and such results were not used in calculations. The measured hardness was HK0.1/15 = 

19.6±1.3 GPa. Knoop testing was also attempted at 0.5 kg, however, large lateral cracks formed, making 

measurements difficult. For Vickers hardness testing, loads of 0.1 and 0.5 kg were applied. The resulting 

hardness was HV0.1/15 =20.6±1.2 GPa and HV0.5/15 =19.6±1.3 GPa, respectively. Little or no cracking 

was evident from the indentation corners for 0.1 kg. However for 0.5 kg, significant cracking occurred. 

In addition to testing on the c-plane, tests were conducted on the a- and m-planes. Vickers indentation 

on planes other than the c-plane resulted in severely cracked or spherically shaped indentations. The 

results are summarized in Table A2 of the appendix. 

2.4 Fracture Toughness 

The fracture toughness of the a-, m-, and r-planes was measured in accordance with standardized 

procedures (Ref. 4) by using the PB (single-edged-precracked-beam) and VB (chevron-notch beam) 

methods. Initially, the machining vendor had difficulty grinding chevron notches within the required 

tolerances due to the high hardness of sapphire. Thus, several sets of PB and VB tests were performed. 

Witness tests using α SiC were run before each series of tests and gave virtually identical results that 

compare very well with data from this and other laboratories for the same batch of material (Ref. 5). 

Water does not wet notches in sapphire well, and thus the test specimens and fixture were submerged in 

water for 10 to 20 min. prior to load initiation. 

r-plane 

4 mm 

a-plane 

m- or aplane 

c-plane 

4 mm 

, x3

3 mm 52 mm 3 mm 52 mm 

c

c (0001)

57.6°

32.4°

61.2°

n n n

n

n

n

r r

r

a a a

r {1012}

m

30°

30°

30°

a

a

m

m

{1120}

{1010}

r

r r

n

n n

n

n n

c

n

{2113}

View down c-axis

Optical axis (3-fold symmetry)

2-fold

symmetry axis
NASA/TM—2006-214023 4 

2.5 Non-Destructive Inspections 

All of the strength and slow crack growth test specimen bevels were checked for severe chips by 

carefully running a thumbnail along the bevel. This simple procedure identified five r-plane and seven 

a-plane specimens with severe chips. These specimens were used for fracture toughness testing rather 

than strength testing. The majority of r-plane specimen bevels were also inspected by optical microscopy, 

and the gage sections and bevels of six of the a-plane specimen were optically inspected in detail at NIST 

(Ref. 6). In addition to the “thumbnail” and optical inspections, three r-plane and 33 a-plane specimens 

were viewed using X-ray topography at Brookhaven National Laboratory (Ref. 6). 

2.6 Slow Crack Growth and Inert Strength Testing 

The SCG (slow crack growth) behavior of the sapphire was characterized by using the constant stress 

rate (“dynamic fatigue”) method in accordance with ASTM C 1368 (Ref. 7). Testing was conducted by 

using four-point flexure with 20 mm load and 40 mm support spans in distilled, deionized water at stress 

rates ranging from 0.002 to 2 MPa/s for the r-plane test specimens and 0.002 to 23 MPa/s for the a-plane 

test specimens. An electro-mechanical test system with load control capability was employed.3

 A sheet of 

0.125 mm thick graphite was placed between the rollers and the compressive face of the test specimens 

(Ref. 8). In addition, the compressive face and one-half of the sides of the test specimens were covered 

with cellophane tape to aid fractography. The stress rates were intentionally kept much smaller than 

typically used in strength (~35 MPa/s) and SCG testing in order to minimize averaging of the three 

regions of the SCG curve (Ref. 9). 

The inert strength was determined by testing specimens in high purity (99.98 percent) dry nitrogen 

after vacuum drying for more than 4 hours. The test fixtures and equipment were identical to that used for 

the SCG measurements. In addition to using the inert environment, the stress rate was increased to 

300 MPa/s in order to minimize any stable crack extension associated with region III of the SCG curve. 

3. Analyses 

3.1 Slow Crack Growth Parameter Analysis 

The stress rate as a function of load rate for four-point flexure is 

( )

2

o i

bw

P S S

2

3 − =  σ (1) 

where P is the load rate, b is the specimen width, w is the height and So and Si are the outer and inner 

spans. The use of different stress rates provides time for the flaws in the material to grow to different sizes 

via stress corrosion prior to specimen failure, thereby indicating the sensitivity of the material to the test 

environment. For ceramics and glasses, the slow crack growth rate under Mode I loading above the SCG 

limit is generally expressed by the power-law relations (Ref. 10)

n

I

n

IC

* I ] = AK

K

K = A [ dt

da

v= (2) 

where v, a and t are crack velocity, crack size and time, respectively. The material/environment dependent 

SCG parameters are A or A* and n. The values KI and KIC are, respectively, the applied Mode I stress 

 

3

Instron model 8562, Instron Corp., Canton MA. 
NASA/TM—2006-214023 5 

intensity factor and the critical stress intensity factor or fracture toughness of the material. For SCG 

testing that employs a constant load or stress rate, the corresponding SCG fracture strength, σf , can be 

expressed as a function of stress rate, σ , as follows (Ref. 10) 

[ ( ) ]

1 n 1 n 2

f i B n 1 + − σ = + σ σ (3) 

where σi is the inert strength, and B is a parameter associated with A or A*, n, fracture toughness, crack 

geometry and loading configuration. The parameter n is determined from the slope of log σf plotted as a 

function of log σ with Equation (3) written as 

log log D n 1

1 log f +

+ σ = σ (4) 

where 

[ ( ) ] n 2

i log B n 1 n 1

1 log D − +

+ = σ (5) 

Once the slope α and intercept β of Equation (4) are estimated via linear regression, the parameters n, 

D, B and A or A*, and their standard deviations, SDn, etc., can be estimated (Ref. 11) from 

1 1 n = − α (6) 

2 α

SDα SDn ≈ (7) 

β D = 10

 

( )( ) β β SDD ≈ 2.3026 SD 10 (8) 

1 3

i

10 B αβ α

σ

α

= (9) 

( )

2

2

Ic

2

3 1

i

2

Ic *

B( n 2 )Y

2K

10 (1 3 )Y

2K

A − = − =

−

α

σ

β α

α

 (10) 

( )

2

2 n Ic

2

3 1

i

1 3

Ic

B( n 2 )Y

2K

10 (1 3 )Y

2K

A − = − =

−

− −

α

σ

β α

α α

 (11) 

where Y is the geometry correction factor for the stress intensity formulation. The crack depth to halflength ratio, a/c, was taken as 0.1 and the corresponding value of the geometry correction factor as 

Y = 1.95 where KI = σY√a. 
NASA/TM—2006-214023 6 

3.2 Window Stresses 

During orbit, the windows are required to sustain an annual differential pressure spectrum of 

0.93 MPa (135 psid) applied for 0.5 hours, followed by 0.21 MPa (30 psid) for 14.5 hours. The desired 

life of the windows is 10 years with an additional safety factor of 4×, leading to a 40 cycle loading 

spectrum and 600 hours under variable pressure. 

The window is contained and supported on the periphery with a flexible seal, resulting in neither a 

completely fixed nor completely free boundary condition. If the boundary condition is taken as simply 

supported at the edge of the window and the window face presumed to be the c-plane, which is elastically 

isotropic, the stresses can be described by (Ref. 12): 

( ) ( )( ) s 2

s

2

2

d

2

s

2

2

s rr R

r 2 1 v 3 v R

R 1 v 8t

3PR σ + σ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ = − + + − + (12) 

( ) ( )( ) s 2

s

2

2

d

2

s

2

2

s

R

r 2 1 v 1 3v R

R 1 v 8t

3PR σθθ + σ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ = − + + − + (13) 

( )

4( ) 1 v

P 3 v s −+ σ = (14) 

where P is the applied pressure, RS is the support radius, Rd is the window radius, t is the window 

thickness, ν is Poisson’s ratio and r is the radius of interest. The term σs is a small correction factor to the 

simple plate theory for the effects of the shearing stresses and lateral pressure on the plate deflection 

(Ref. 13). Based on Equations (12) to (14), a pressure of 0.93 MPa (135 psid) and nominal dimension of 

8 mm thickness, 115 mm support diameter, and 121 mm window diameter, the maximum stress is 60 MPa. 

For orientations other than the c-plane, Equations (12) to (14) do not accurately describe the elastic 

stress distribution. The stresses based on a general displacement solution for an anisotropic plate 

subjected to pressure (Refs. 14 to 16) are shown in Figure 3(a) for a window with an a- or m-plane face. 

The anisotropy causes a variation in the radial and tangential stresses with circumferential position, 

particularly at the edge of the window. For comparison, the stresses in a window with a c-plane face are 

shown in Figure 3(b). The radial and tangential stresses do not vary with circumferential position.

4. Results 

4.1 Fracture Toughness 

Chevron-Notch Testing 

Initial attempts to measure a- and r- fracture toughness by using the VB were complicated by poor 

test specimen machining. The poor machining resulted in the notch parameters α1 and/or α0 being slightly 

out of specification (see Fig. 4), and relatively wide notches (0.3 to 0.5 mm vs. 0.27 to 0.35 typically) for 

the r-plane specimens. ASTM C1421 (Ref. 4) does not require a minimum notch width; however, it 

recommends a width less than 0.300 mm in order to promote stability. To remedy the notch parameter 

deviations in estimation of the fracture toughness, specific stress intensity factor coefficients for the actual 

parameters were calculated, rather than using those calculated from the narrow range equations supplied 

in C 1421. These corrections were less than 1 percent and not significant for engineering purposes. The 

results for tests in air are summarized in Table 1. Some tests results are not included in Table 1 because 

they were either invalid or could not be accurately remedied. This was usually due to crack growth 

instability or specimen damage by the machinist. The instability was attributed the wide notches. 
NASA/TM—2006-214023 7 

Figure 3.—Polar plot of the stresses in a 115 mm diameter, 8 mm thick, simply supported sapphire window subjected to a 

0.93 MPa pressure: (a) window face is the a- or m-plane, and (b) window face is the c-plane.

Figure 4.—Crack configurations and configuration for single-edge-precracked-beam and 

chevron-notch test specimens. 

TABLE 1.—VB TEST RESULTS FOR 50 PERCENT RH AIR. RESULTS ARE FROM THE 

FIRST SET OF SPECIMENS MACHINED 

Specimen 

Number 

Test Condition Measured 

Toughness, 

MPa√m 

Noncompliant 

Values 

Other 

Complications 

Remedy 

R58 50 percent RH 1.98 α0 and α1 Wide notch Calculate Y

R60 50 percent RH 1.94 α0 and α1 Wide notch Calculate Y

A57 50 percent RH 1.96 α1 None Calculate Y

A58 50 percent RH 1.96 α1 None Calculate Y

A59 50 percent RH 1.88 α1 None Calculate Y

A60 50 percent RH 1.84 α1 None Calculate Y

SiC JAS-A12 50 percent RH 2.60 None α SiC Witness test 

W

B 

ao

a1 a2 a3

a1 

α1= a1 /W

αo= ao /W α =( a1+a2+a3)/3W 

SEPB CNB

Stress, MPa

0 10 20 30 40 50 60 70

0

10

20

30

40

50

60

70

<a>

<m>

Stress, MPa

0 10 20 30 40 50 60 70

0

10

20

30

40

50

60

70

<c>

<m>

or 

<a>

Tangential Stress, r/Rs = 0.2

Radial Stress, r/Rs = 0.2

Tangential Stress, r/Rs = 0.8

Radial Stress, r/Rs = 0.8

Stress, MPa

(a) (b) 

{a} or {m} {c} 
NASA/TM—2006-214023 8 

In order to obtain more data than shown in Table 1, two tracks were taken: (1) additional test 

specimens were machined; and (2) the remnants of the VB test specimens were tested using the PB 

method. The second set of VB specimens conformed to the geometric requirement and the notches were 

narrower, however, the notch planes that form the chevron were occasionally offset slightly more than 

allowed in C1421. Despite the narrower notches, it was still difficult to attain stable extension as required, 

presumably due to the offset. In order to promote stable initiation, the specimens were loaded in 

compression several times and then tested in the usual manner, to no avail. A second unsuccessful attempt 

to get stability was performed by loading the specimens to near the expected initiation load and holding 

for several minutes to several hours. Finally, a PB precracking fixture was used to generate short 

precracks like those shown in Figure 5. This consistently resulted in stable extension as required, 

however, some of the precracks were longer than the crack length corresponding to minimum stress 

intensity factor coefficient, Y*

min. This was remedied by calculating a specific Y* value and pairing it with 

the corresponding load at the onset of stable extension. This generally lead to a lower measured fracture 

toughness than using the maximum load and minimum stress intensity factor coefficient (as is usually 

done), because the decrease in load was usually larger than the increase in Y*

, and because the crack 

velocity is lower than that occurring in a test with large, stable extension followed by unstable extension. 

The useful results are summarized in Table 2. 

Figure 5.—Precracked VB test specimens showing the precrack as viewed through (a) the bottom, and (b) the side. 

TABLE 2.—VB TEST RESULTS FROM THE SECOND SET OF SPECIMENS MACHINED. TESTS WERE 

CONDUCTED IN DISTILLED, DEIONIZED WATER AFTER SOAKING FOR 10 MINUTES 

Specimen 

Number 

Test 

Condition 

Measured 

Toughness, 

MPa√m 

Noncompliant 

Values 

Other Complications Remedy 

A-II-4 Water 2.07 Notch plane offset Long precrack Calculate Y*

M-II-4 Water 1.98 Semi-stable 

R-II-1 Water 1.82 Notch plane offset Long precrack Calculate Y*

R-II-2 Water 1.80 Long precrack Calculate Y*

R-II-3 Water 1.87 Manual precracking 

R-II-4 Water 1.70 Long precrack Calculate Y*

R-II-5 Water 1.93 Notch plane offset Manual precracking 

SiC JAS-A16 Air 2.58 None α SiC Witness test 

(a) (b) 

Precrack 
NASA/TM—2006-214023 9 

As more test results than in Table 2 were needed, a third set of VB specimens was machined. These 

specimens meet all geometric tolerances, and the notch widths and offsets were excellent. Despite this, 

stable extension was difficult to attain. 

This was speculated to be from residual stresses at the notch root, or a different grinding rate or wheel 

that left fewer micro cracks at the notch root. Communications with the machine shop indicated that the 

same wheel was used; however, the feed rate may have been slowed to improve the notch tolerances. To 

test the theory that residual stresses were hampering stable crack initiation, two r-plane specimens were 

annealed and tested along with the as-received specimens. Again stable extension could not be attained. 

Thus precracking was employed. This was successful, except for the annealed test specimens. These 

could not be precracked at the usual load, implying that healing of machining damage in the notch root 

hampers initiation of crack extension in the VB. The useful results for the third set are summarized in 

Tables 3 to 5. 

During testing and data analysis, it was noted that well aligned, smooth precracks resulted in lower 

fracture toughness values, particularly for the r-plane, as shown in Figures 6 to 8 and noted in the 

comments in Tables A4 to A6 in the Appendix. Further fractography indicated that for the r-plane, cracks 

initiate on both sides of the chevron and join, resulting in a cleavage step. Depending on the steps, 

different values resulted (from 1.51 MPa√m for no steps to ~2.0 MPa √m for large or multiple steps), as 

shown in Figure 8. Specimens that fractured unstably from the chevron tip or during machining/handling 

exhibited a very planar surface with no cleavage steps visible to the unaided eye. The increase in 

measured fracture toughness for crack growth on planes even slightly misaligned with the r-plane is 

supported by recent dry nitrogen data (Ref. 17) and from early data in liquid nitrogen. The recent results 

indicated difficulty in keeping a crack on the r-plane and stress intensities greater than 2.8 MPa√m in 

dry nitrogen. The early results in liquid nitrogen indicated fracture energies of 6.2 to 8.7 J/m2

 (2.14 to 

2.53 MPa√m) for a rough surface (Ref. 18), and were thought to be too large to represent extension on the 

r-plane. 

All of the a-plane test specimens exhibited a smooth, conchoidal fracture surface without cleavage 

steps, as shown in Figure 6, implying that the fracture resistance on and near the a-plane is similar. Tilted 

precracks resulted in slight increase in the measured fracture toughness. The m-plane specimens 

occasionally exhibited poorly defined step-like features. When the precrack was tilted significantly, the 

measured fracture toughness was elevated 10 to 15 percent, as shown in Tables A4 to A6 in the 

Appendices. 

Figure 6.—Fracture surfaces of a-plane VB specimens tested in air or water. 

0.5 mm

KA-III-1 

KIvb = 2.37 MPa√m 

0.5 mm

Precrack

0.5 mm

KA-III-9 

KIvb = 1.97 MPa√m 

KA-III-13 

KIvb = 1.98 MPa√m 
NASA/TM—2006-214023 10 

TABLE 3.—VB a-PLANE TEST RESULTS FROM THE THIRD SET OF SPECIMENS MACHINED. THE SPECIMENS 

WERE TESTED IN AIR OR DISTILLED, DEIONIZED WATER AFTER SOAKING FOR 20 MINUTES 

Specimen 

Number 

Test Condition Measured 

Toughness, 

MPa√m 

Noncompliant 

Values 

Other Complications Remedy 

KA-III-1 50 percent RH 2.37 None 

KA-III-2 50 percent RH 2.09 None Long precrack Calculate Y*

KA-III-3 50 percent RH 2.33 None 

KA-III-7 Water 1.90 Long precrack Calculate Y*

KA-III-8 Water 1.95 None Manual precracking 

KA-III-9 Water 1.97 None Manual precracking 

KA-III-12 Water 1.96 None Manual precracking 

KA-III-13 Water 1.98 None Manual precracking 

KA-III-14 Water 1.96 None Manual precracking 

SiC-JAS-A17 Water 2.63 None α SiC Witness test 

Figure 7.—Fracture surfaces of m-plane VB specimens tested in water. 

TABLE 4.—VB m-PLANE TEST RESULTS FROM THE THIRD SET OF SPECIMENS MACHINED. THE 

SPECIMENS WERE TESTED IN DISTILLED, DEIONIZED WATER AFTER SOAKING FOR 20 MINUTES 

Specimen 

Number 

Test 

Condition 

Measured 

Toughness, 

MPa√m 

Noncompliant 

Values 

Other Complications Remedy 

KM-III-7 Water 1.82 None Long precrack Calculate Y*

KM-III-9 Water 1.91 None Long precrack Calculate Y*

KM-III-10 Water 1.97 None Manual precracking 

KM-III-12 Water 1.92 None Long precrack Calculate Y*

KM-III-13 Water 1.95 None Manual precracking 

KM-III-15 Water 1.90 None Manual precracking 

KM-III-16 Water 2.04 None Manual precracking 

KM-III-17 Water 2.01 None Manual precracking 

Figure 8.—Fracture surfaces of r-plane VB specimens tested in water. 

KR-III-3 

KIvb = 1.99 MPa√m

KR-III-6 

KIvb = 1.51 MPa√m 

KR-III-17 

KIvb = 1.74 MPa√m 

0.5 mm 0.5 mm 0.5 mm

KM-III-9 

KIvb = 1.99 MPa√m 

0.5 mm 0.5 mm 0.5 mm

KM-III-16 

KIvb = 2.04 MPa√m 

KM-III-15 

KIvb = 1.90 MPa√m 
NASA/TM—2006-214023 11 

TABLE 5.—VB r-PLANE TEST RESULTS FROM THE THIRD SET OF SPECIMENS MACHINED. THE 

SPECIMENS WERE TESTED IN DISTILLED, DEIONIZED WATER AFTER SOAKING FOR 20 MINUTES 

Specimen 

Number 

Test 

Condition 

Measured 

Toughness, 

MPa√m 

Noncompliant 

Values 

Other Complications Remedy 

KR-III-3 Water 1.99 None Manual Precracking 

KR-III-6 Water 1.51 None None 

KR-III-12 Water 1.73 None Long precrack Calculate Y*

KR-III-13 Water 1.70 Long precrack Calculate Y*

KR-III-16 Water 1.71 None Manual Precracking 

KR-III-17 Water 1.74 None Manual Precracking 

KR-III-18 Water 1.85 None Manual Precracking 

Single-Edge-Precracked-Beam Testing 

Precracking of the test specimens was performed in ambient laboratory air. Testing of the specimens 

was performed in high purity dry nitrogen or silicone oil after subjecting the precracked specimens to 

vacuum drying for ~20 hours. This was done to obtain a fracture toughness value in the absence of SCG, 

and thereby estimate a maximum value. Such an “inert” fracture toughness is needed to combine with the 

materials inert strength for estimation of the initial flaw size in the calculation of the SCG parameter A, 

and for estimation of the window proof test pressure. The crack extension and stability of the tests 

were monitored by a strain gage placed on the back-face of the test specimens (Ref. 19). The measured 

fracture toughness, KIpb, of the r-plane was 2.47±0.15 MPa√m for six valid tests. The measured fracture 

toughness, KIpb, of the a-plane was 2.31±0.12 MPa√m for ten valid tests. In total, 13 valid and invalid

r-plane tests were run for a mean of 2.47±0.20 MPa√m, and 14 valid and invalid a-plane tests were tested 

for a mean of 2.32±0.16 MPa√m. 

The precrack and fast fracture regions in the r-plane test specimens tended to be very planar, except 

when the precrack hooked or turned slight out of plane at the precrack front, as shown in Figures 9 and 

10. This was likely due to frictional effects in the precracking fixture rather than to any tendency of the 

crack to transition to other planes because the same effect is observed in polycrystalline test specimens 

(Ref. 19). Very small cleavage steps emanated from the indentations used to initiate the precracks. The 

crack hook and steps were quite beneficial in that the crack length would otherwise be difficult to 

measure because penetrant did not wet the cracks well. The precracks in the a-plane test specimens 

tended to exhibit a conchodial appearance. Observations of fractured specimens with low magnification 

polarized light indicated an absence of twinning. In order to check for small scale twinning activity and 

dislocations, surfaces of two a- and two r-plane test specimens were etched by using sodium tetraborate, 

which is quite specific to a-planes, at 800 °C for ~20 to 60s. Only a few etch pits could be observed on 

the a-plane sides of the r-plane specimen, implying a very low dislocation density (~1000/cm2

) for this 

material. Observations on both sides of the precrack front on the fracture surface of the a-plane specimens 

indicated no increase in the dislocation density and thus little plastic deformation (Ref. 20). Also, no 

evidence of twins could be observed on the r-plane fracture surfaces, except at the indentations used to 

initiate the precracks. Thus the slight increase in measured fracture toughness of the r-plane specimens as 

compared to the a-plane may be due to the small cleavage steps or a small difference in cleavage energy 

in the absence of water. 
NASA/TM—2006-214023 12 

Figure 9.—Top and side views of the fracture surfaces of a-plane PB test specimens. 

Figure 10.—Top and side views of the fracture surfaces of r-plane PB test specimens. 

R58 R60 RU5

1 mm

Cleavage 

steps 

1 mm

A59 AU3 AU5

Precrack

Indentations 
NASA/TM—2006-214023 13 

Comparison to Published Values 

As mentioned previously, Wiederhorn and Krause (Ref. 17) reported difficulty in obtaining extension 

on the r-plane when using DT (double torsion) test specimens. However, double torsion measurements 

were successful on the m-plane and indicated values of 1.95 to 2.1 MPa√m in dry nitrogen at a velocity of 

~10–4 m/s (Ref. 17). Although in this investigation, no such difficulty was encountered in producing very 

planar cracks in PB specimens, the PB and DT test method are quite different and in this study 

precracking was done in air. Evidently the susceptibility of the r-plane to cracking is very dependant on 

humidity. Other factors influencing fracture measurements may be impurities and residual stresses in the 

particular sapphire tested. 

It should be noted that a fracture toughness value measured under stress corrosion conditions 

represents a point on the stress corrosion induced, crack growth curve and thus depends on the stress 

intensity rate and the ability of the corrosive medium to get to the crack tip, which depends on the 

concentration and crack size, etc. Thus the measured fracture toughness is an operational value based on a 

stated procedure. It is likely that the only stress intensities corresponding to fracture properties are the 

threshold value and that measured under inert conditions. 

Table 6 summarizes data from published literature and this work. Most of the work indicates that the 

fracture toughness in dry nitrogen is greater than 2 MPa√m for all of the typical cleavage planes. The 

introduction of water to the environment drops the fracture toughness to just under 2 MPa√m for the 

a- and m-planes, however, for the r-plane the measured fracture toughness varies between 1.40 and 

1.77 MPa√m depending on the test method and procedures (i.e. crack velocity) employed. 

The large difference between the values on the r-plane may be due to the low n value and 

contaminant strong dependence on crack velocity, in addition to the flatness of the crack employed. As 

compared to polycrystalline alumina, sapphire exhibits substantially lower fracture toughness: For oil, air, 

and water a 96 percent polycrystalline alumina exhibited fracture toughness of 3.37±0.05, 3.19±0.07, and 

2.75±0.01 MPa√m (Ref. 5), respectively. 

For the m-plane, the measured fracture toughness in water is significantly larger than the threshold 

value in 85 percent RH air (1.94 vs. 1.65 MPa√m (Ref. 21)). The value reported by Reference 22 

(2.33 MPa√m) was measured by using a relatively blunt saw notch (tip radius = 0.1 mm) and thus likely 

overestimates the fracture toughness of the m-plane in air, despite being lower than the value 

(3.14 MPa√m) measured by Iwasa and Bradt (Ref. 23) in air by using surface cracks. Many cleavage 

steps were observed on the fracture plane, implying propagation on multiple planes rather than on the mplane only. For the r-plane, Michalske (Ref. 24) measured the crack velocity on the r-plane in water, 

ammonia, hydrazine, and acetonitrile. If the data is extrapolated to 10–4 m/s, a stress intensity of 

~1.40 MPa√m is estimated for water, while a stress intensity >2.0 MPa√m is estimated for acetonitrile. 

One troubling aspect of the reported data is the low values in Figure 23 of Reference 25. A fracture 

toughness of 1.62 MPa√m for the r-plane in dry nitrogen at 10–4 m/s is indicated in Figure 23; however, 

Table 4 of the same publication indicates much higher corresponding fracture energy of 0.6 J/m2

, which 

coverts to 2.10 MPa√m assuming isotropy and 2.18 MPa√m accounting for anisotropy. This difference 

nominally equates to √2 in the conversion. Private communication (Ref. 26) indicated the difference to be 

due to scatter. The m-plane fracture energy was reported as 7.3 J/m2

 (Refs. 18 and 25), which converts to 

2.59 MPa√m assuming isotropy and E = 430 GPa, in agreement with a later publication reporting a 

fracture toughness of 2.5 MPa√m (Ref. 27). 

It is critical that relatively planar cracks be generated, particularly for the r-plane, or elevated values 

will be measured. Further, environment and crack velocity play a strong role in the measured result, 

implying the need for a specific operational procedure or the use of threshold values. 
NASA/TM—2006-214023 14 

TABLE 6.—REPORTED FRACTURE TOUGHNESS AND FRACTURE ENERGY OF SAPPHIRE. VALUES 

WITH REFERENCE NUMBER WERE THOSE REPORTED; THE CORRESPONDING FRACTURE 

TOUGHNESS OR FRACTURE ENERGY VALUES WERE CALCULATED BY ASSUMING ISOTROPY 

AND THE ELASTIC MODULUS OF THE ASSOCIATED DIRECTION AND A POISSON’S RATIO OF 0.26 

Crystal plane Fracture toughness, 

MPa√m 

Cleavage energy, 

J/m2

Elastic modulus, 

GPa 

Watera

a {1210} 1.95±0.03 

m {1100} 1.94±0.07 

r {1012} 1.77±0.13 

r – v = 10–4 b1.40 (Ref. 24) 

Aira

a - 50 percent RH 2.06±0.21 

r - 50 percent RH 1.96 

m - 85 percent RH; KIth c1.64 (Ref. 21) 

m - 85 percent; v = 10–4 d2.22 (Ref. 21) 

c { } 0001 4.54±0.32 (Ref. 23) 21.54 465 

m 3.14±0.30 (Ref. 23) 11.43 430 

a 2.43±0.26 (Ref. 23) 430 

r 2.38±0.14 (Ref. 23) 7.68 e440

m 2.33±0.07 (Ref. 22) 

Dry Nitrogena

a 2.31±0.12 

r 2.47±0.15 

m – v = 10–4 f2.14 (Ref. 17) 

c >40 (Ref. 18) 

m g

2.59±0.29 7.3±1.1 (Refs. 18 and 

25)

430 

r g

2.10±0.17 6.0±0.6 (Refs. 18 and 

25)

344 

r – v = 10–4 i1.62 (Ref. 25) 

r –196 °C lN2 24.0 (Ref. 28) 

{1 126} 24.4 (Ref. 28) 

{1 126}–196 °C lN2 32.2 (Ref. 28) 

{1 123}–196 °C lN2 16.4 (Ref. 28) 

a

Room temperature unless stated otherwise. 

b

Value taken from Figure 2 of Reference 24 corresponding to a velocity of 10–4 m/s in water. c

Threshold stress intensity in 85 percent RH. 

d

Value taken from Figure 4 of Reference 21 corresponding to a velocity of 10–4 m/s in water. e

A value of E = 440 GPa was reported by Iwasa (Ref. 23) for the r-plane. E is closer to 344 GPa and 

thereby gives 7.68 J/m2

. f

Calculated from Equation (5) of Reference 17. 

g

Calculated from the formula KIc = √(2γE/(1 – ν

2

)). Nineteen tests were on conducted on two anneal 

boules. Testing of unannealed boules resulted in 5.3±1.3 J/m2

. h

If anisotropy is considered, the r-plane value becomes 2.18 MPa√m. i

Value taken from Figure 23 of Reference 25 corresponding to a velocity of 10–4 m/s in dry N2. 
NASA/TM—2006-214023 15 

4.2 Non-Destructive Inspection 

The optical inspections of the strength test specimens indicated a relatively well-polished gage section 

on both the a- and r-plane test specimens. However, the bevels of the a-plane test specimens were 

generally marked more strongly with grinding/polishing scratches than the r-plane test specimens, as 

shown in Figure 11. X-ray topographs indicated not only stronger grinding marks on the a-plane 

specimens, but also a “short” finish (i.e., insufficient material removal for elimination of the damage 

generated in the early grinding stages), as shown in Figures 12 and 13. Besides the remnant grinding 

damage, both the a- and r-plane test specimens exhibited asymmetric polishing of the tensile face bevels 

as shown in Figure 13. In discussions, the vendor2

 indicated that preparation of the a-plane specimens is 

more difficult due the hardness anisotropy of sapphire. Further, the vendor indicated that the polishing 

procedures used were developed for plates and lenses instead of narrow beams. This is collaborated by 

X-ray topographs of three disk test specimens that are shown in Figure 14. The disks were polished by 

the vendor at a similar time to the same specifications (60/40 scratch/dig) as the beam test specimens; 

however, they exhibit a much better finish. 

Figure 11.—Optical macrographs of (a) a-plane, and (b) r-plane test specimen bevels. The 

observations were made at 200×. 

Figure 12.—X-ray topographs of (a) a-plane, and (b) r-plane test specimens. 

 1.22 mm 

(a) (b)

(a) (b)

0.20 mm 
NASA/TM—2006-214023 16 

Figure 13.—X-ray topographs of the tensile face of (a) an r-plane test specimen and (b) an 

a-plane test specimen. Note the asymmetric bevel finish on both test specimens. 

Figure 14.—X-ray topographs of the faces of three 50.4 mm disk test specimens showing a near dislocation level finish. 

The face of the disks is the c-plane. 

4.3 Slow Crack Growth and Inert Strength Testing: r-plane Data 

Weibull distributions of fracture strength in dry nitrogen and water at the slowest and fastest stress 

rates are plotted in Figure 15, and the fracture strength as a function of stress rate is plotted in Figure 16. 

A significant loss in strength is exhibited as the stress rate is reduced, implying stress corrosion induced 

crack growth. The SCG parameters corresponding to the data in Figures 15 and 16 are summarized in 

Tables 7 and 8, and the data is given in detail in Table A8 of the appendix. The r-plane of HEM sapphire 

exhibits a relatively low value of the fatigue exponent n, which is comparable to that of glass (Ref. 29). 

The inert strength and SCG data exhibit four low out-lying data points that tend to skew the slope of 

the SCG curve because they occur so infrequently. One approach to minimize these points is the median 

deviation technique (Ref. 10), in which regression is performed by using the median fracture strength at 

each stressing rate. Unfortunately, ASTM C 1368 does not currently provide for the application of this 

technique. It is likely that the outlying values are associated with the bevels of the flexure specimens: a 

geometric feature that is less likely to influence the failure of a window because the stresses at the beveled 

edge of a window are lower than the stresses at the window center. The effect of censoring the outlying 

data points is shown in Figures 17 and 18 to illustrate what better polishing, handling, and inspection 

might achieve. If the outlying values were associated with bevel damage, then the censored fatigue 

parameters would be a better estimate of the SCG behavior. However, fractographic justification is 

needed for censoring. Without supporting fractography, which is left to future work, the uncensored SCG 

parameters should be considered the best estimates. 

(b) 4.1 mm 

(a)

{c} 

0.5 mm
NASA/TM—2006-214023 17 

Figure 15.—Uncensored r-plane Weibull strength distributions for dry nitrogen 

and the slowest and fastest stress rates in water. 

Stress Rate, σ, MPa/s

0.001 0.01 0.1 1 10 100 1000

Strength,

σf , MPa

300

350

400

450

500

600

700

800

900

1000

Water

Regression Line

95% Confidence Interval

95% Prediction Interval

Inert (Dry N2)

Figure 16.—Experimentally measured failure stress as a function of stress rate for 

the r-plane of sapphire in water. 

 300 400 500 600 800 1000 1700 

Probabilit

y of Failure, % 

Fracture Strength, MPa 

Stress Rate, σ , MPa/s

(Dry N2) 

σθ m
NASA/TM—2006-214023 18 

Figure 17.—Weibull strength distributions for the r-plane in dry nitrogen and for 

the slowest and fastest stress rates in water. Four outlying data points were 

censored. 

 

Figure 18.—Experimentally measured failure stress as a function 

of stress rate for the r-plane of sapphire in water. Four outlying 

specimens were censored. 

 300 400 500 600 800 1000 1700 

Probabilit

y of Failure, % 

Fracture Strength, MPa 

(Dry N2) 

σθ m
NASA/TM—2006-214023 19 

TABLE 7.—MEASURED STRENGTH STATISTICS AND 90 PERCENT CONFIDENCE 

BANDS FOR THE r-PLANE 

Stress Rate 

MPa/s 

(Number) 

Mean 

Strength.a

MPa 

Characteristic 

Strength, σθ

MPa 

90 percent 

Confidence 

Bands on σθ

Weibull 

Modulus 

m 

90 percent 

Confidence 

Bands on m

Uncensored Data: 

300 – Dry N2 (15) 1085±127 1135 1093; 1179 12 8.5; 16 

2 (18) 806±150 857 817; 899 8.5 6.3; 11 

0.002 (16) 582±61 607 587; 627 13 9.4; 18 

Censored Data: 

300 – Dry N2 (14) 1108±93 1149 1113; 1186 15 10; 21 

2 (16) 851±69 881 857; 906 15 11; 21 

0.002 (15) 592±46 613 595; 630 16 11; 22 a

± one standard deviation. 

TABLE 8.—SLOW CRACK GROWTH PARAMETERS AND STANDARD 

DEVIATIONS FOR THE r-PLANE 

4.4 Slow Crack Growth and Inert Strength Testing: a-plane Data 

The Weibull distributions and SCG data for the a-plane test specimens are shown in Figures 19 and 

20. The data is extremely scattered, and the corresponding confidence in the SCG parameters is thus low. 

The source of scatter observed at any stress rate is likely the “short” finish and the asymmetric bevel 

polishing. Such scatter is common is sapphire, and although treatments such as “super polishing” and 

“annealing” significantly improved the mean strength of sapphire, such treatments did not decrease the 

scatter and increase Weibull modulus in one case reported in the literature (Ref. 30). 

During testing, two test specimens (A23 and A32) failed while preloading and were thus subjected to 

very little slow crack growth. This was an unexpected occurrence because the r-plane, which was tested 

first, was reasonably consistent and strong, and the a-plane is generally considered tougher than the 

r-plane. The treatment of these data points has a strong influence on the SCG results but relatively little 

on the Weibull statistics for a specific stress rate. Three scenarios can be envisioned for processing of the 

data: (1) Censor the two test results because they are not representative of the slow crack growth process; 

(2) Use the data at the stress rate for which the tests were to proceed (i.e. treat them as upper bound 

Parameter n D B 

MPa2s 

A* 

m/s 

A 

m/s⋅(MPa√m)-n

Least Squares, 

All Data 21.0±4.4 762±28 24,454 7.12 x 10–6 2.81 x 10–14 

Least Squares, 

Censored 17.7±1.3 819±12 251,204 8.38 x 10–7 6.29 x 10–14 

Median Deviation 17.6±2.4 267±7 209,276 1.01 x 10–6 9.43 x 10–14
NASA/TM—2006-214023 20 

estimates of the SCG strength at the assigned stress rate); (3) Assign the data a stress rate based on the 

actual preloading rate. Scenario (1) seems the most plausible because the two tests are not representative 

of SCG due to the very short failure times. However, from a statistical sampling standpoint, case (2) 

seems appropriate because a slightly lower preload would have allowed measurement of a SCG strength. 

Case (3) seems implausible because the failure time is so short (< 2s) that the results likely approach inert 

strength measurements. 

Tables 9 and 10 summarize the Weibull and SCG statistics associated with a-plane for all three 

scenarios. The detailed data is given in detail in Table A9 of the appendix. The parameter n varies from 

31 for scenario (2) to 83 for scenario (3). Considering the relatively low values for SCG parameters for 

the r-plane (n = 21), it is suggested that n be taken as 31. Again, better test specimen preparation would 

eliminate the problems encountered in estimation of the SCG parameters for the a-plane. 

TABLE 9.—MEASURED STRENGTH STATISTICS AND 90 PERCENT CONFIDENCE 

BANDS FOR THE a-PLANE. 

Stress Rate 

MPa/s 

(Number) 

Mean 

Strengtha

MPa 

Characteristic 

Strength, σθ

MPa 

90 percent 

Confidence 

Bands on σθ

Weibull 

Modulus 

m 

90 percent 

Confidence 

Bands on m

300 – Dry N2 (15) 1255±547 1416 1195; 1677 2.64 1.89; 3.67 

23 (19) 1078±456 1214 1047;1408 2.67 1.99; 3.58 

0.002 (21) 785±334 888 769; 1024 2.63 1.99; 3.49 a

± one standard deviation. 

TABLE 10.—SLOW CRACK GROWTH PARAMETERS AND STANDARD 

DEVIATIONS FOR THE a-PLANE 

Scenario n D B 

MPa2s 

A* 

m/s 

A 

m/s⋅(MPa√m)-n

(1) 37±23 887±66 93.5 8.26 x 10–4 5.35 x 10–17 

(2) 31±16 868±66 499 1.89 x 10–4 2.06 x 10–15 

(3) 83±112 832±63 2.45 x 10–8 1.37 x 10+6 4.62 x 10–24 

Median 

Deviation 25±26 324±56 27,686 4.26 x 10–6 5.55 x 10–15
NASA/TM—2006-214023 21 

Figure 19.—Uncensored a-plane Weibull strength distributions for dry nitrogen and 

for the slowest and fastest stress rates in water. 

Stress Rate, σ, MPa/s

0.001 0.01 0.1 1 10 100 1000

Strength,

σf , MPa

150

200

300

400

500

600

800

1500

2000

3000

100

1000

Water Regression Line

95% Confidence Interval

95% Prediction Interval Inert (Dry N2)

Figure 20.—Experimentally measured failure stress as a function of stress rate for 

the a-plane of sapphire in water. 

Stress Rate, σ , MPa/s

 100 200 500 1000 2000 4000 

Probability of Failure, % 

Fracture Strength, MPa 

(Dry N2) 

σθ m
NASA/TM—2006-214023 22 

Comparison to Published Literature 

Table 11 and Figure 21 summarize data from this work and the literature. Some references did not 

report the SCG parameters and the original data was not available (Ref. 31), so the data was estimated by 

scanning and digitizing original reprints with commercial software4

. In some cases, the digitized data was 

plotted on the same scale and size plot as the original figure, and overlaid on the original to assure that the 

process was accurate. The crack growth parameters A and n were determined by linear regression of log v 

as a function of log KI. 

The a-plane and m-plane data exhibit similar trends regardless of the humidity, with peak stress 

intensity of just over 2 MPa√m. The r-plane data varies widely. In general, older vintage data exhibits 

slow crack growth at lower stress intensities and much scatter as a function of RH (relative humidity). If a 

√2 adjustment is applied to the early NIST data, the agreement is better, but a significant difference still 

exists between recent NIST data (Ref. 17) and early NIST/Sandia data (Refs. 24 and 25). Dynamic fatigue 

testing resulted in lower n values than macro-crack techniques, and thus indicates more sensitivity to 

water vapor. The estimation of parameter A from the dynamic data requires an estimate of the initial flaw 

size distribution. This is estimated via the strength and fracture toughness measured in an inert 

environment (e.g., dry nitrogen or silicone oil). Figure 23 shows the sensitivity of the SCG curve on the 

inert fracture toughness. Even reducing the inert fracture toughness to 1.75 MPa√m does not make the 

r-plane data agree. 

Based on the fractography of Reference 17, the r-plane of sapphire does not cleave easily in the 

absence of water vapor. Both large and small scale cracks (i.e., those from DT and dynamic fatigue test, 

respectively) exhibited extension at small angles (5 to 10°) to the r-plane on which the resolved mode I

stress intensity is lower. In macro-crack DT test specimens, stable extension off the r-plane at very large 

stress intensities was followed by crack jumps onto the r-plane and a drop in the stress intensity to 

2.8 MPa√m. In the dynamic fatigue specimens, a smooth mirror region followed by a rough surface was 

exhibited, implying stable extension and unstable extension off the r-plane prior to fast fracture. Two 

possible explanations for the crack not readily propagating on the r-plane are misalignment of the DT 

fixture and bend-twist coupling. Discussions indicated that misalignment was likely not the case 

(Ref. 32). Bend-twist coupling occurs in anisotropic materials, like sapphire, when insufficient symmetry 

occurs, as might be the case when the secondary orientation of the test specimen is not controlled. Other 

possibilities include residual stresses, which were noted to decrease fracture energy from 7.3±0.9 J/m2

 to 

5.6±0.6 J/m2

 (23 percent) (Ref. 18), and cation dopant level, which increased the fracture toughness for 

0.049 percent Ti3+ or Ti4+ on the a- and m-planes (Ref. 33). Results of Cr3+ doping were less clear.

 

4

Un-Scan-It 5.0, Silk Scientific, P.O. Box 533, Orem, UT 84059 
NASA/TM—2006-214023 23 

Stress Intensity, KI, MPa- m

0.8 0.9 1.0 1.2 1.5 2.0 2.5 3.0 1.8

Velocity, v, m/s

1e-9

1e-8

1e-7

1e-6

1e-5

1e-4

Water NIST/Sandia 1986 [24]

85% RH UCB 2000 [21]

Dry N2, NIST 2002 [17]

Dry N2, NIST 2002 [17]

Water, NASA 2001 case (2)

Water, NASA 2001

Dry N2, NIST 1968 [25]

50% RH, NIST 1968 [25]

r-plane

m-plane

a-plane

1968 1986

2002

2001

Figure 21.—Crack velocity as a function stress intensity factor for sapphire. 

√
NASA/TM—2006-214023 24 

Stress Intensity, KI, MPa- m

0.8 0.9 1.5 2.0 2.5 3.0 1.0

Velocity, v, m/s

1e-9

1e-8

1e-7

1e-6

1e-5

1e-4

50% RH, 1968 NIST [25] (correction)

85% RH UCB 2000 [21]

Dry N2, 2002 NIST [17]

Dry N2, NIST 2002 [17]

Water, NASA 2001 case (2)

r-plane

m-plane

a-plane

Water, Sandia/NIST 1986 [24]

Water, NASA 2001

Dry N2, NIST 1968 [25] (correction)

Figure 22.—Crack velocity as a function stress intensity factor for sapphire. 

A √2 factor was applied to the NIST 1968 data sets. 

(√2 adjustment) 

(√2 adjustment) 

√
NASA/TM—2006-214023 25 

TABLE 11.—CRACK GROWTH PARAMETERS FOR SAPPHIRE IN AIR, WATER AND DRY NITROGEN 

Data Source 

(number of data points) Test Method1 A 

m/s (MPa√m)–n n ± SDn 

Nominal Vintage 

& Growth Method1 

a-plane

This work 4 PB, Water 2.06 x 10–15 31±16 2001, HEM7 

m-plane

(Ref. 21), As reported in text CT, 85% RH 1.8 x 10–20 45.6±2.5 2000, EFG8 

(Ref. 17), Fig. 4, Data points2 DT, Dry N2 1.44 x 10–21 52.0±4.3 2002, EFG9 

(Ref. 17), Eq. 5, 20 °C DT, Dry N2 1.11 x 10–21 51.5 2002, EFG9

r-plane

(Ref. 25, private) Fig. 23 (8)3 DCB, 50% RH 2.22 x 10–5 49.7±4.1 1968, Verneuil10 

(Ref. 25, private) Fig. 23 (11)4 DCB, Dry N2 3.98 x 10–17 59.6±26 1968, Verneuil10 

(Ref. 24), Fig. 2, Data points (17)5 DCB, Water 3.18 x 10–14 64.4±3.5 198611 

This work 4 PB, Water 2.81 x 10–14 21.0±4.4 2001, HEM7 

(Ref. 17), Fig. 7, Data points (8)6 DT, Dry N2 3.88 x 10–32 60.6±6.8 2002, EFG9 

 1. CT =compact tension; DT = double torsion; DCB = double cantilever beam; 4PB = four-point bend, 

HEM = heat exchange method; EFG = edge-defined film-fed growth. 

 2. The value of A was not reported, thus the data was digitized and fit. The reported average n for four 

measurements on three test specimens at 20 °C was 54.5±11.9. 

 3. The estimated standard deviation of log(A), SDlog(A) = 0.17. Propagation of errors gives 

SDA = 2.3026 A (SDlog(A)) = 8.68 x 10–6. 

 4. The estimated standard deviation of log(A), SDlog(A) = 4.81. Using propagation of errors gives 

SDA = 4.41 x 10–16. 

 5. The estimated standard deviation of log(A), SDlog(A) = 0.39. Using propagation of errors gives 

SDA = 2.83 x 10–14. 

 6. The estimated standard deviation of log(A), SDlog(A) = 1.57. Using propagation of errors gives 

SDA = 2.58 x 10–31. 

 7. Manufactured by Crystal Systems, Salem, MA with the HEM (Heat-Exchange Method). 

 8. Manufactured by Kyocera, with the EFG (Edge-defined, Film-fed Growth) method. 

 9. Manufactured by Saphikon Inc./St. Gobain, Milford NH, with EFG. 

10. Manufactured by Linde Corp by the Verneuil method. 

11. Manufactured by Atomergic Chemetals Corp., Plainview NY, probably by EFG. 
NASA/TM—2006-214023 26 

Stress Intensity, KI, MPa- m

0.8 0.9 1.0 1.2 1.5 2.0 2.5 3.0 1.8

Velocity, v, m/s

1e-9

1e-8

1e-7

1e-6

1e-5

1e-4

Water NIST/Sandia 1986 [24]

Dry N2, NIST 2002 [17]

Water, NASA 2001

Dry N2, NIST 1968 [25]

50% RH, NIST 1968 [25]

r-plane

1.75 2.00

2.50

KIc = 

1968 1986

2002

Figure 23.—Crack velocity as a function stress intensity factor for the r-plane 

of sapphire. The effect of the value of inert fracture toughness used to 

estimate the SCG curve is shown for constant stress rate data from r-plane 

test specimens. 

4.5 Fractographic Observations 

Casual observation of the test specimens during testing indicated failure from origins located on both 

the bevels and within the tensile face gage section. However, during the strength testing, occasional pops 

were heard and jumps in the load–time recordings were noted. In an initial set of inert strength tests 

(10 r-plane specimens and 10 a-plane specimens) none of the load drops were sufficient (10 percent 

load drop within 10 milliseconds) to trigger the test systems “break detect” function and unload the test 

specimen. However, in a second set of tests (5 r-plane specimens and 5 a-plane specimens) one inert, 

r-plane test that triggered the detector prior to failure displayed cracks running between the compressive 

face to the neutral axis, as shown in Figure 24. The cracks developed from twins that run between the 

compressive face and the neutral axis (Ref. 34). As the material to the left of the twins tried to move 

upward relative to that to the right, the cracks formed at the neutral axis and propagated upward as the 

faulted material is squeezed out. The faults were relatively uniformly distributed within and just outside 

the inner load span. No damage on the tensile surface was apparent to the unaided eye and the specimen 

remained whole. Apparently, failure under inert conditions occurs by two competing effects: (1) tensile 

√
NASA/TM—2006-214023 27 

face fracture from large flaws prior to significant twining; and (2) tensile face fracture after severe twin 

faulting has diminished the ability of the compressive section to bear load, thereby shifting the test 

specimen neutral axis and elevating the tensile face stress. Thus, tensile face failure origins are observed 

despite the formation of twins. The audible pops and load-drops are probably the result of the faults 

forming and the material popping-out of the compressive side of the gage section. The result of twinning 

is likely a lower measured strength. The use of graphite foil, which has been reported to suppress 

compression face twins and increase strength (Ref. 8), was not sufficient in this case. Polishing of the 

compressive side of the test specimens should minimize the formation of the compression-side twins 

during inert strength testing. 

Testing in water at lower stress rates (2 MPa/s for 15 r-plane specimen and at 23 MPa/s for 

15 a-plane specimens) did not result in load drops sufficient to trigger the break detect for an initial set of 

test specimens. As the stress level at failure is lowered and the environment made more corrosive, SCG 

becomes the dominant failure mechanism. However, during testing of three r-plane specimens 7 months 

later and four a-plane specimens 1 month later, one r-plane test specimen (R61) and one a-plane test 

specimen (A49) triggered the break detect. Additional fractography of r-plane specimens indicated the 

presence of many twins in the compression region of all but two specimens (R3 and R18). The frequency 

and delineation of twinning was dependant on the stress level, with the number of twins ranging from 

~one dozens at ~500 MPa to nearly one hundred at 1 GPa. The fracture mirror and trace along the tensile 

face appeared parallel to the r-plane; however, the surrounding planes containing river marks were 

consistently at a ~26° to the r-plane (Figure 25(a)). The line of intersection between the r-planes and the 

a-plane sides is normal to the c-plane twins, implying that the r-plane mirror transits onto the m-plane. 

This is consistent with the lower inert fracture toughness for the m-plane as compared to the r-plane. No 

distinct bifurcation occurred, however, at distances well away from the origin, bifurcation-like cracks 

were apparent (Fig. 25(b)). 

Figure 24.—Side view of an r-plane test specimen (R67). The test was interrupted at 1198 MPa during 

loading in dry nitrogen at 300 MPa/s. Faults leading from the compressive side to just beyond the initial 

neutral axis of the beam are apparent (arrows marked “F”). Basel plane twins are apparent as the dark 

shadows running 32° to the neutral axis. Note that the twins align cross the cracks, implying that they 

occurred first. 

Neutral 

Axis 

Tension 

3.1 

Compression 

F F Tape 

< r >

{a} 

Twins 

Basel Twins 
NASA/TM—2006-214023 28 

 

Figure 25.—(a) Side view of an r-plane test specimen (R3) showing typical m-plane crack extension from the mirror. 

Failure was from a long surface scratch. (b) Cleaves occurring away from the origin in an r-plane test specimen 

(R10) that failed at 733 MPa. River marks running along the m-plane are also apparent by looking through the 

tensile surface. Bifurcation did not occur along the tensile face. 

 

Figure 26.—View of an a-plane test specimen (A54) that failed at 1288 MPa: (a) a-plane mirror followed by r-plane 

branches, and (b) secondary r-plane cleaves that increase in frequency and size with distance from the origin. 

The a-plane specimens exhibited mirrors on or near the a-plane that branched to the r-plane. The 

r-plane branches were surrounded by a symmetric series of r-plane cracks of increasing frequency and 

size with increasing failure stress and distance from the origin (Fig. 26). Test specimens failing at 

~500 MPa or less usually exhibited no secondary r-plane cleaves (though the branches were on the

r-plane) while specimens failing over ~1 GPa exhibited dozens. 

Thus, a specific procedure describing the stages of removal, the feed-rate, grit-sizes and amounts to 

be removed in each stage is needed for sapphire to avoid the complications described in this section and 

section 4.2. For purposes of producing a specimen with a ground finish (e.g., 400 grit) the methods 

described ASTM C 1161 (Ref. 35) might be adequate for test specimens or components with a uniaxial 

stress state. However, the procedures in ASTM C 1161 impart more damage in the direction parallel to 

grinding than in the direction perpendicular to grinding. Thus components or test specimens with a 

multiaxial stress state will fail preferentially along the direction parallel to grinding (Ref. 36). Thus, any 

polishing needs to remove all of the grinding damage imparted in prior stages. 

Tensile 

face 

Cleaves Origin 

< r >

< r >

{a}

m-plane 

m-plane 

river marks 

Bevel 1 mm

1 mm

1 mm

a-plane 

mirror 

r-plane 

branches 

r-plane 

cleaves 

(a) (b)

(a) (b)
NASA/TM—2006-214023 29 

4.6 Window Strength 

Because the windows are larger than the test specimens used to generate the strength and SCG data, 

the window strength is expected to be lower. The strength of a window at a probability of failure can be 

estimated from the Weibull scaling equation for a given probability of failure: 

( )

1 / m σ 1 σ 2 = Ae2 Ae1 (15) 

where σ1 and σ2 are the strengths corresponding to “effective area” Ae1 and Ae2. Because the specimens in 

this study probably failed from surface connected flaws, the use of an effective area is appropriate. The 

“effective area” of a simply supported circular plate subjected to a uniform pressure can be calculated 

from Reference 37 

( ) () ()

( )( ) ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

+ +

+ + −

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

+

+

≅

ν ν

π ν ν ν

3 1 3

2R 1 R 1

R

R

1 m

4 1 Ae

2

s 2d

2

d

s (16)

within 1 percent for m > 5, ν > 0.17 and Rs /Rd < 0.9. For uniaxial flexure specimens the effective area is 

( )

( ) [ ] o 2

i o b w S

b w

mb 1 m 1

1 m S / S Ae + ⎥

⎦

⎤ ⎢

⎣

⎡

+

+

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

+

+ = (17) 

where b is the specimen width, w is the height, and Si and So are the load and support spans respectively. 

The expected strength of a window for 50 percent probability of failure is summarized in Table 12 for the 

data sets. It should be noted that the predicted strength based on the a-plane data is low due to the poor 

polishing of the test specimens. 

TABLE 12.—ESTIMATED FAST FRACTURE STRENGTH OF A 121 MM DIAMETER, 

SIMPLY SUPPORTED WINDOW SUBJECTED TO A UNIFORM PRESSURE 

Data Set Test Specimen 

Characteristic 

Strength, 

MPa 

Test Specimen

Effective 

Area, 

mm2

Window 

Effective 

Area, 

mm2

Estimated Window 

Strength for 

50 percent Failure, 

MPa 

 r-plane, All data, Dry N2 1135 93.75 2200 843 

 r-plane, Censored, Dry N2 1149 90.33 1805 913 

 r-plane, All data, Water 895a 93.75 2335 671 

 a-plane, Case (2) 1416 109.4 7762 245 a

Estimated by extrapolation of the slow crack growth curve to the ASTM C1161 rate of 35 MPa/s. 

4.7 Choice of Window Orientation 

Sapphire windows with c-plane and a-plane faces have been made, and the choice of window face has 

usually been based on the size of the window and the size of the sapphire boules available rather than the 

crystal properties (Ref. 38). However, for high precision applications the c-plane is preferred over the 

a- or m-planes because it is the zero birefringence orientation. For c-orientations, the a- and m- planes are 

subjected the principal stress, whereas the r-plane is subjected to a normal stress of 71 percent of the 

principal (cos2

 32.4°), thereby mitigating the effect of the lower r-plane fracture toughness in high 

humidity. 
NASA/TM—2006-214023 30 

Recent testing of twenty ring-on-ring biaxial strength specimens at room temperature indicated that 

c-face disks are stronger (1053±421 MPa) than a-face disks (607±182 MPa) (Ref. 39). The strengths and 

standard deviations correspond to Weibull moduli of ~3 and ~4 respectively, implying an inconsistent 

surface finish on disks of both orientations. 

4.8 Recommended Values 

Because the reported properties of sapphire vary widely, it is recommended that more recent values or 

values determined for the specific type of sapphire (i.e., crystal systems HEM grade) be used. Based on 

the preceding sections, the values in Table 13 are recommended. The m-plane data is based on reference 

(Ref. 21) and is similar to the a-plane data generated in this work. The r-plane values are based on the 

measurements of HEM sapphire in water. Because design requirements typically use mean strength, the 

strength corresponding to 50 percent probability of failure for the widow is given. 

TABLE 13.—RECOMMENDED PROPERTIES 

Plane A 

m/s⋅(MPa√m)-n n f σ MPa 

(Water) 

KIc 

MPa√m 

(Water)

KIc 

MPa√m 

(Dry N2)

KI threshold 

MPa√m 

(85 percent RH)

a- or m-plane 1.8 x 10–20 46 ± 2.5 ---- 1.94 ± 0.07 2.14 1.65 

r-plane 2.81 x 10–14 21 ± 4 671 ± 71 1.77 ± 0.13 2.47 ± 0.15 ---- 

5. Summary and Conclusions 

The fracture toughness, inert flexural strength, and slow crack growth parameters of the r- and 

a-planes of HEM sapphire were measured. The fracture toughness in dry nitrogen, KIpb, was 

2.47 ± 0.15 MPa√m and 2.31 ± 0.12 MPa√m for the r- and a-planes. Fracture toughness measured in 

water via the operational procedure in ASTM C1421 was significantly lower, 1.95 ± 0.03 MPa√m, 

1.94 ± 0.07 and 1.77 ± 0.13 MPa√m for the a- , m- and r-planes. The mean uniaxial inert strength in dry 

nitrogen was 1085 ± 127 MPa for the r-plane and 1255 ± 547 MPa for the a-plane. The measured 

strengths are a result of the polishing of the material rather than an inherent material property. The power 

law slow crack growth exponent was n = 21 ± 4 for the r-plane and n ≥ 31 for the a-plane. The r-plane of 

HEM sapphire is relatively susceptible to stress corrosion induced slow crack growth in water. Data for 

HEM a-plane sapphire tested in water agreed with recent m-plane data sets for 85 percent RH and dry 

nitrogen. The published literature and the data reported herein indicate that the fracture toughness and 

SCG resistance of the r-plane of sapphire have improved since the 1960s, with the stress intensity levels 

nearly doubling. 

Interferences to testing due to test specimen preparation included the frequent occurrence of a “short” 

finish, occasional and severe chips on bevels, and an asymmetric bevel polishing in which one bevel 

retained severe grinding damage. A consistent, well defined method for preparation of sapphire is needed, 

even if consistency is gained at the sacrifice if absolute strength. This would impart higher component 

reliability despite the sacrifice in absolute strength of the component. 

The strength difference between the two planes tested is due to the difference in specimen preparation 

rather than the difference in the stress corrosion properties associated with the specific crystallographic 

planes. If the strength scatter (i.e., the low Weibull modulus) associated with the a-plane material can be 

eliminated either by more consistent polishing or proof testing, a better estimate of the a-plane slow crack 

growth parameters could be made from strength data. 
NASA/TM—2006-214023 31 

In order to minimize twinning in the failure process and insure that SCG is the dominant mechanism 

in dynamic fatigue testing, the stress rates should be relatively low, and water should be used. Testing 

circular disks via pressure loading might improve data quality. Manufacturing of disk is simpler and the 

issues associated with bevels and contact stresses are reduced. Further, pressurized disks provide a better 

analogy to an actual window and extrapolation of the test specimen characteristic strength to that of the 

component is lessened. Unfortunately, an ASTM standard for such testing does not exist. However, a 

standard for concentric ring testing, which produces a similar stress state, has been developed (Refs. 40 

and 41). 

Recommendations 

Determination of the threshold stress intensity for the a-, m-, and r-planes for HEM sapphire in air 

and water is recommended. 

NASA/TM—2006-214023 33 

Appendix—Summary Reports on Testing 

 
NASA/TM—2006-214023 34 TABLE A1.—ELASTIC MODULUS OF SAPPHIRE TEST SPECIMENS Length mm Height mm Width mm Mass g Flexural Frequency (Hz) Density g/cc Poisson’s Ratio v E GPa Corrected E GPa (No Bevel on KA-, KM-, or KR- specimens) KA-21 51.10 3.023 4.066 2.5010 12137 3.982 0.26 424.4 424.4 KA-22 51.07 3.028 4.060 2.4968 12169 3.977 0.26 423.8 423.8 KA-23 51.10 3.024 4.080 2.5121 12179 3.985 0.26 427.4 427.4 KA-24 51.06 3.013 4.077 2.4933 12097 3.975 0.26 422.3 422.3 KA-25 51.07 3.025 4.079 2.5094 12173 3.982 0.26 425.4 425.4 A56 51.34 3.100 4.125 2.5946 12256 3.952 0.26 416.6 433.1 A57 51.34 3.099 4.114 2.5887 12250 3.956 0.26 416.9 433.4 A58 51.33 3.100 4.125 2.5972 12252 3.957 0.26 416.5 433.0 A59 51.34 3.101 4.110 2.5897 12250 3.958 0.26 416.4 432.9 A60 51.35 3.100 4.121 2.5964 12247 3.958 0.26 416.9 433.4 A-46 51.35 3.100 4.112 2.5901 12246 3.957 0.26 416.8 433.2 A-55 51.34 3.099 4.116 2.5865 12258 3.950 0.26 416.8 433.2 

a-plane average = 430 

± 5

KM-21 51.07 3.052 4.053 2.5088 12246 3.971 0.26 422.0 422.0 KM-22 51.04 3.042 4.040 2.4960 12228 3.979 0.26 423.3 423.3 KM-23 51.03 3.052 4.042 2.5022 12242 3.975 0.26 420.8 420.8 KM-24 51.05 3.047 4.055 2.5069 12230 3.974 0.26 421.9 421.9 KM-25 51.05 3.047 4.053 2.5050 12241 3.973 0.26 422.6 422.6 

MU-1 51.58 3.083 4.088 2.565 12036 3.946 0.26 413.0 429.3 MU-2 51.49 3.083 4.086 2.5655 12046 3.955 0.26 411.9 428.1 MU-3 51.33 3.095 4.086 2.5654 12161 3.952 0.26 411.2 427.4 MU-4 51.53 3.082 4.075 2.5612 12058 3.958 0.26 414.4 430.8 MU-5 51.54 3.085 4.08 2.5576 12110 3.943 0.26 416.0 432.4 MU-6 51.55 3.087 4.086 2.5639 12075 3.943 0.26 413.4 429.8 

m-plane average = 426 

± 4
NASA/TM—2006-214023 35 TABLE A1 (CONTINUED).—ELASTIC MODULUS OF SAPPHIRE TEST SPECIMENS Length mm Height mm Width mm Mass g Flexural Frequency (Hz) Density g/cc Poisson’s Ratio v E GPa Corrected E GPa (No Bevel on KA-, KM-, or KR- specimens) KR-21 51.00 3.035 4.074 2.5160 10973 3.990 0.26 342.3 342.3 KR-22 51.07 3.039 4.063 2.5112 10973 3.982 0.26 342.6 342.6 KR-23 51.01 3.034 4.062 2.5099 10973 3.993 0.26 343.0 343.0 KR-24 51.02 3.035 4.063 2.5108 10983 3.991 0.26 343.5 343.5 KR-25 51.09 3.039 4.065 2.5138 10979 3.983 0.26 343.6 343.6 RU-1 53.21 3.065 4.099 2.6352 10092 3.942 0.26 331.8 344.9 

RU-2 53.39 3.072 4.097 2.6471 10027 3.939 0.26 330.3 343.3 RU-3 53.32 3.069 4.095 2.6474 10073 3.951 0.26 333.2 346.4 RU-4 53.19 3.073 4.099 2.6366 10106 3.935 0.26 330.0 343.1 RU-5 53.34 3.075 4.099 2.641 10062 3.928 0.26 329.8 342.8 

R56 51.32 3.058 4.091 2.5384 10856 3.954 0.26 335.3 348.5 R57 51.28 3.057 4.093 2.5383 10847 3.956 0.26 334.2 347.4 R58 51.3 3.066 4.094 2.5477 10883 3.956 0.26 335.0 348.3 R59 51.27 3.070 4.096 2.5502 10895 3.956 0.26 334.0 347.2 R60 51.31 3.058 4.094 2.5394 10845 3.953 0.26 334.3 347.5 

r-plane average = 345 

± 2
NASA/TM—2006-214023 36 TABLE A2.—HARDNESS OF SAPPHIRE TEST SPECIMENS. TEN TESTS WERE RUN PER CONDITION. Orientation Knoop HK0.1/15 Vickers HV0.1/15 Vickers HV0.5/15 

c-plane 

diagonal parallel to the a-axis 19.6 

± 1.3 20.6 

± 1.2 19.6 

± 1.3 

a-plane 

diagonal parallel to the 

m-axis 19.0 

± 1.0 ** 

a-plane 

diagonal parallel to the r-axis 18.0 

± 1.1 ** 

m-plane 

diagonal parallel to the a-axis 21.7 

± 0.9 ** 

Tensile face of r-plane specimen 

Diagonal parallel to r-axis 21.0 

± 1.3 ** 

** For 0.1 kg load, spherically shaped indentations formed. For 0.1 kg, heavy cracking occurred.
NASA/TM—2006-214023 37 TABLE A3.—SUMMARY OF a-PLANE VB FRACTURE TOUGHNESS RESULTS FOR TESTING IN AIR OR WATER. GRAYED VALUES EXCEED THE SPECIFICATIONS OF ASTM C1421. STRUCK-THROUGH VALUES ARE INVALID. Tare = 1.631 N So = 19.951 mm Stroke Si = 9.977 Rate = 0.20 mm/min Bevel Correction = 1.029 Notch Plane Normalized Notch Parameters: Load Test Spec B W Width Offset 

α0 α11 α12 α1 P Stability, 

KIvb

Number mm mm mm N Comments MPa

√m 

Air: 

A56 3.100 4.125 0.224 0.000 0.199 0.952 0.948 0.950 23.05 Overload 2.505

A57 3.099 4.114 0.256 0.044 0.190 0.942 0.949 0.946 18.16 Yes, corrected Y* 1.96 

A58 3.100 4.125 0.242 0.025 0.198 0.953 0.945 0.949 17.91 Yes, corrected Y* 1.96 

A59 3.101 4.110 0.296 0.070 0.195 0.944 0.948 0.946 17.13 Yes, corrected Y* 1.88 A60 3.100 4.121 0.293 0.094 0.189 0.949 0.937 0.943 17.08 Yes, corrected Y* 1.84 

KA-III-1 3.019 4.081 0.292 0.005 0.190 0.953 0.960 0.957 21.64 Yes 2.37 KA-III-2 3.032 4.062 0.293 0.007 0.198 0.970 0.958 0.964 15.79 Yes, corrected Y* 2.09 KA-III-3 3.025 4.086 0.288 0 0.186 0.949 0.964 0.956 21.58 Yes 2.33 

Average = 2.06 

± 0.21 

Water: 

A-II-2 3.052 3.995 .321 0.096 0.206 0.987 0.988 0.987 17.68 No 2.19

A-II-4 3.048 3.995 .327 0.102 0.207 0.981 0.983 0.982 17.25 Yes 2.07 

KA-III-7 3.014 4.085 0.289 0.005 0.179 0.943 0.940 0.942 17.55 Yes, long precrack 1.90 

KA-III-8 3.016 4.082 0.293 0.013 0.185 0.950 0.956 0.953 17.74 Yes 1.95 KA-III-9 3.025 4.058 0.287 0 0.193 0.965 0.957 0.961 17.44 Yes 1.97 KA-III-10 3.032 4.081 0.288 0 0.195 0.966 0.959 0.963 -------- Broken KA-III-11 3.038 4.088 0.292 0 0.191 0.957 0.963 0.960 19.060 Yes, tilted precrack 2.09

KA-III-12 3.018 4.073 0.298 0.02 0.185 0.952 0.951 0.951 17.890 Yes 1.96 KA-III-13 3.031 4.062 0.289 0.04 0.186 0.959 0.962 0.960 17.940 Yes 1.98 KA-III-14 3.032 4.080 0.284 0.035 0.192 0.963 0.957 0.960 17.600 Yes 1.96 KA-III-15 3.024 4.075 0.288 0 0.188 0.959 0.965 0.962 16.430 Yes, curved front 1.93

Average = 1.95 

± 0.03
NASA/TM—2006-214023 38 TABLE A4.—SUMMARY OF m-PLANE VB FRACTURE TOUGHNESS RESULTS FOR TESTING IN AIR OR WATER. GRAYED VALUES EXCEEDED THE SPECIFICATION IN ASTM C1421. STRUCK-THROUGH VALUES ARE INVALID. Tare = 1.631 N So = 19.951 mm Stroke Si = 9.977 Rate = 0.20 mm/min Bevel Correction = 1.029 Notch Plane Normalized Notch Parameters: Load Specimen B W Width Offset α0 α11 α12 α1 P Test KIvb Number mm mm mm N Stability, MPa√m Comments Water: M-II-4 3.086 4.083 0.309 0.078 0.205 0.964 0.954 0.959 16.65 Semi-stable 1.98 KM-III-5 3.047 4.058 0.280 0.007 0.185 0.959 0.969 0.964 21.68 Yes, tilted 2.365 KM-III-6 3.045 4.069 0.285 0.007 0.192 0.965 0.968 0.966 ---- Broken KM-III-7 3.052 4.038 0.288 0.013 0.191 0.970 0.966 0.968 15.23 Yes, long precrack 1.82 KM-III-8 3.039 4.070 0.291 0.009 0.189 0.962 0.967 0.964 18.51 Yes, tilted 2.05 KM-III-9 3.046 4.051 0.291 0 0.185 0.966 0.965 0.966 17.22 Yes, long precrack 1.91 

KM-III-10 3.049 4.074 0.288 0.028 0.196 0.969 0.975 0.972 17.49 Yes 1.97 

KM-III-11 3.044 4.046 0.281 0.005 0.191 0.972 0.966 0.969 19.70 Yes, tilted 2.20

KM-III-12 3.046 4.043 0.285 0 0.190 0.975 0.964 0.969 17.02 Yes, long precrack 1.92 

KM-III-13 3.041 4.070 0.289 0 0.195 0.967 0.962 0.965 17.32 Yes 1.95 KM-III-14 3.052 4.080 0.282 0 0.196 0.966 0.967 0.966 20.65 Yes, tilted 2.29

KM-III-15 3.045 4.067 0.286 0 0.193 0.968 0.961 0.965 16.92 Yes 1.90 KM-III-16 3.048 4.055 0.282 0 0.188 0.968 0.964 0.966 18.36 Yes 2.04 

KM-III-17 3.046 4.076 0.280 0 0.192 0.965 0.967 0.966 18.11 Yes 2.01 

Average = 1.94 ± 0.07

5Strikethrough values were not used in estimation of averages. 
NASA/TM—2006-214023 39 TABLE A5.—SUMMARY OF r-PLANE VB FRACTURE TOUGHNESS RESULTS FOR TESTING IN AIR OR WATER. GRAYED VALUES EXCEEDED THE SPECIFICATION IN ASTM C1421. STRUCK-THROUGH VALUES ARE INVALID. Tare = 1.631 N So = 19.951 mm Stroke Bevel Correction = 1.029 Si = 9.977 Rate = 0.20 mm/min Notch Plane Normalized Notch Parameters: Load Test Specimen B W Width Offset 

α0 α11 α12 α1 P Stability, 

KIvb

Number mm mm mm N Comments MPa

√m 

Air: 

R58 3.066 4.094 0.498 0.157 0.167 0.939 0.938 0.938 18.97 Yes, corrected Y* 1.98 

R59 3.070 4.096 0.475 0.172 0.150 0.938 0.947 0.942 21.9 No 2.196

R60 3.058 4.094 0.516 0.206 0.148 0.942 0.946 0.944 19.12 Yes, corrected Y* 1.94 

Water: 

R-II-1 2.936 4.012 0.314 0.114 0.183 0.954 0.947 0.951 9.14 Yes 1.82 

R-II-2 2.940 4.010 0.322 0.092 0.192 0.948 0.950 0.949 14.20 Yes 1.80 R-II-3 2.939 4.011 0.350 0.093 0.182 0.953 0.950 0.952 16.02 Yes 1.87 R-II-4 2.938 4.009 0.326 0.088 0.181 0.952 0.950 0.951 14.50 Yes 1.70 R-II-5 2.936 4.013 0.321 0.121 0.195 0.952 0.956 0.954 16.08 Yes 1.93 

KR-III-3 3.040 4.065 0.293 0 0.194 0.959 0.955 0.957 17.77 Yes, large step 1.99 

KR-III-6 3.043 4.069 0.285 0 0.181 0.958 0.974 0.966 13.47 Yes 1.51 

KR-III-12 3.041 4.060 0.294 0.009 0.194 0.971 0.962 0.966 8.40 Yes, long precrack 1.73 

KR-III-13 3.033 4.059 0.280 0 0.188 0.967 0.970 0.968 14.86 Yes, long precrack 1.70 

KR-III-14 3.034 4.056 0.290 0 0.191 0.964 0.968 0.966 ------- Broken 

KR-III-15 3.032 4.063 0.292 0 0.190 0.967 0.966 0.966 17.82 Yes, tilted precrack 1.99

KR-III-16 3.036 4.055 0.285 0 0.190 0.969 0.959 0.964 15.07 Yes 1.71 

KR-III-17 3.035 4.057 0.289 0 0.194 0.970 0.965 0.968 15.17 Yes 1.74 KR-III-18 3.042 4.069 0.287 0 0.190 0.960 0.964 0.962 16.54 Yes 1.85 KR-III-19 3.043 4.047 0.291 0 0.192 0.964 0.968 0.966 12.23 Yes, slow loading 1.42

KR-III-20 3.043 4.063 0.290 0.006 0.190 0.966 0.960 0.963 18.15 Yes, tilted precrack 2.02

Average = 1.77 

± 0.13

6Strikethrough values were not used in estimation of averages. 
NASA/TM—2006-214023 40 TABLE A6.—SUMMARY OF r-PLANE PB FRACTURE TOUGHNESS RESULTS FOR TESTING IN DRY NITROGEN AND SILICONE OIL. GRAYED VALUES EXCEEDED THE SPECIFICATION IN ASTM C1421. STRUCK-THROUGH VALUES ARE INVALID. Tare = 1.631 N So = 19.951 mm Stroke Si = 9.977 Rate = 0.20 mm/min Bevel Correction = 1.029 Crack Length: Load Specimen Indent a1 a2 a3 Δa1 Δa2 Δa3 a/W B W P Test KIpb Number Load mm mm Mm % % % mm mm N Stability MPa√m Dry N2: R56 5 x 3 kg 2.502 2.55 2.355 1.3 3.3 -4.6 0.60 3.058 4.091 46.72 Yes 2.48 R57 5 x 3 kg 2.47 2.37 2.193 5.4 1.1 -6.5 0.57 3.057 4.093 51.85 Yes 2.45 R58 5 x 3 kg 1.932 2.044 2.063 -4.0 1.5 2.5 0.49 3.066 4.094 65.53 Yes 2.35 R59 5 x 4 kg 1.743 1.642 1.512 6.8 0.6 -7.4 0.40 3.070 4.096 97.3 Yes 2.67 R60 5 x 4 kg 2.196 2.135 2.109 2.3 -0.5 -1.8 0.52 3.058 4.094 64.89 Yes 2.58 RU4 5 x 5 kg 2.612 2.587 2.501 1.8 0.8 -2.6 0.63 3.073 4.099 41.35 Yes 2.397

RU5 5 x 5 kg 2.076 2.024 1.959 2.8 0.2 -3.0 0.49 3.075 4.099 63.20 Yes 2.27 

Oil:

R57-2 5 x 4 kg 2.581 2.592 2.532 0.5 0.9 -1.4 0.63 3.057 4.093 47.28 Yes 2.76

R58-2 5 x 4 kg 2.857 2.866 2.819 0.3 0.7 -1.0 0.70 3.066 4.094 32.58 Yes 2.58R59-2 5 x 4 kg 3.552 3.532 3.512 0.6 0.0 -0.6 0.86 3.070 4.096 12.28 Yes 2.47R60-2 5 x 4 kg 2.843 2.872 2.932 -1.4 -0.4 1.7 0.70 3.058 4.094 31.99 Yes 2.65

RU1 5 x 5 kg 2.820 2.881 2.911 -1.8 0.4 1.4 0.70 3.075 4.099 30.97 Yes 2.51RU2 5 x 5 kg 2.981 3.091 3.105 -2.5 1.0 1.5 0.75 3.075 4.099 22.45 Yes 2.31

7Strikethrough values were not used in estimation of averages. 
NASA/TM—2006-214023 41 TABLE A7.—SUMMARY OF a-PLANE SEPB FRACTURE TOUGHNESS RESULTS FOR TESTING IN DRY NITROGEN AND SILICONE OIL. GRAYED VALUES EXCEEDED THE SPECIFICATION IN ASTM C1421. STRUCK OUT VALUES ARE INVALID. Tare = 1.631 N So = 19.951 mm Stroke Si = 9.977 Rate = 0.20 mm/min Bevel Correction = 1.029 Crack Length: Load Specimen Indent a1 a2 a3 Δa1 Δa2 Δa3 a/W B W P Test KIpb Number Load mm mm mm % % % mm mm N Stability MPa√m Dry N2: A56 5 x 3 kg 2.849 2.758 2.755 2.2 -1.1 -1.2 0.68 3.100 4.125 36.87 Yes 2.608 A57 5 x 5 kg 2.307 2.001 1.676 15.7 0.3 -16.0 0.48 3.099 4.114 58.07 Yes 2.01 A58 5 x 5 kg 2.046 2.345 2.571 -11.8 1.0 10.8 0.56 3.100 4.125 49.69 Yes 2.21 A59 5 x 5 kg 2.277 2.227 2.2 1.9 -0.3 -1.6 0.54 3.100 4.125 53.27 Yes 2.20 A60 5 x 5 kg 2.207 2.256 2.288 -1.9 0.3 1.7 0.55 3.100 4.121 56.79 Yes 2.38 AU1 5 x 5 kg 2.302 2.334 2.302 -0.5 0.9 -0.5 0.56 3.104 4.122 54.91 Yes 2.42 AU2 5 x 5 kg 2.148 2.348 2.458 -7.3 1.3 6.0 0.56 3.103 4.126 45.89 Yes 2.17 

AU3 5 x 5 kg 2.009 2.063 2.077 -2.0 0.7 1.3 0.50 3.111 4.113 65.12 Yes 2.33 AU4 5 x 5 kg 2.108 2.053 1.927 3.9 1.2 -5.0 0.49 3.107 4.111 60.78 Yes 2.16 AU5 5 x 5 kg 2.082 1.979 1.887 5.0 -0.2 -4.8 0.48 3.106 4.115 64.19 Yes 2.19 

Oil:

A56-2 5 x 4 kg 2.516 2.536 2.521 -0.3 0.5 -0.1 0.61 3.100 4.125 45.40 2.43 A58-2 5 x 4 kg 2.232 2.332 2.313 -2.6 1.7 0.9 0.56 3.100 4.125 57.18 2.47 

A59-2 5 x 4 kg 2.55 2.582 2.577 -0.8 0.5 0.3 0.63 3.101 4.110 43.7 2.48

A60-2 5 x 4 kg 2.06 2.125 2.161 -2.6 0.5 2.2 0.51 3.100 4.121 63.96 2.40 

8Strikethrough values were not used in estimation of averages. 
NASA/TM—2006-214023 42 

TABLE A8.—SUMMARY OF r-PLANE FRACTURE STRENGTH 

FOR TESTING IN WATER AND DRY NITROGEN 

Sapphire Load Span: 20.176 mm 

r-plane Support Span: 40.292 mm 

 Tare: 2.26 N 

Instron 8562, 5000N Environment: Distilled, Deionized Water or 99.98% Dry Nitrogen 

 Bevel Correction: 1.029 % 

 Preload: 300 N 

 

Specimen Loading Stressing Width Height Failure Fracture 

ID Rate Rate b w Load Strength 

N/s MPa/s mm mm N MPa 

Distilled Water 

R1 2.4 1.884785 4.098 3.062 919.5 744.9 

R2 2.45 1.926872 4.092 3.062 923.1 748.9 

R3 1.923550 4.096 3.063 423.3 343.8 

R4 1.932943 4.089 3.058 1054.0 857.5 

R5 1.930534 4.084 3.062 1162.0 944.0 

R6 1.936374 4.096 3.053 1091.0 889.1 

R7 1.918775 4.086 3.071 1130.0 912.5 

R8 1.914016 4.096 3.071 999.3 805.1 

R9 1.925939 4.091 3.063 1170.0 948.2 

R10 1.924541 4.087 3.066 903.9 732.5 

R11 1.917383 4.092 3.070 1121.0 904.6 

R12 1.914808 4.091 3.072 1026.0 826.9 

R13 1.943805 4.077 3.054 1097.0 897.4 

R14 1.930139 4.092 3.059 984.1 799.6 

R15 1.921556 4.093 3.066 1063.0 859.7 

R45 1.916142 4.098 3.069 1115 899.1 

R61 1.913102 4.092 3.073 679.3 547.6 

R63 1.914036 4.094 3.072 1051 846.7 

R16 0.245 0.191857 4.089 3.070 929.0 750.4 

R17 0.191758 4.088 3.071 913.9 737.9 

R18 0.192694 4.092 3.062 485.8 395.0 

R19 0.192434 4.091 3.065 921.7 746.8 

R20 0.193194 4.088 3.059 891.4 725.1 

R21 0.0245 0.019248 4.091 3.064 754.3 611.6 

R22 0.019186 4.091 3.069 779.2 629.7 

R23 0.019244 4.088 3.066 867.4 702.9 

R24 0.019459 4.088 3.049 736.2 603.5 

R25 0.019269 4.092 3.062 858.4 696.5 
NASA/TM—2006-214023 43 

TABLE A8 (CONTINUED).—SUMMARY OF r-PLANE FRACTURE STRENGTH 

FOR TESTING IN WATER AND DRY NITROGEN 

Sapphire Load Span: 20.176 mm 

Rhombohedral Support Span: 40.292 mm 

Tare: 2.26 N 

Instron 8562, 5000N Environment: Distilled, Deionized Water or 99.98% Dry Nitrogen 

 Bevel Correction: 1.029 % 

 Preload 300 N 

 

Specimen Loading Stressing Width Height Failure Fracture 

ID Rate Rate b w Load Strength 

N/s MPa/s mm mm N MPa 

Distilled Water

R26 0.00245 0.001922 4.092 3.066 526.4 426.7 

R27 0.001920 4.094 3.067 661.2 534.9 

R28 0.001934 4.093 3.056 754.3 614.5 

R29 0.001937 4.094 3.053 745.8 608.7 

R30 0.001940 4.092 3.052 770.3 629.3 

R31 0.001934 4.095 3.055 780.3 635.8 

R32 0.001926 4.089 3.064 793.4 643.5 

R33 0.001923 4.089 3.066 760.4 616.1 

R34 0.001924 4.088 3.066 825.4 668.8 

R35 0.001921 4.094 3.066 743.7 601.8 

R36 0.001928 4.087 3.063 636.8 517.4 

R37 0.001935 4.088 3.057 720.6 587.5 

R38 0.001932 4.091 3.058 718.8 585.2 

R39 0.001939 4.093 3.052 633.6 517.9 

R40 0.001930 4.088 3.061 680.3 553.4 

R69 0.001930 4.093 3.060 703.1 571.7 

R44 30 23.694968 4.096 3.054 1246 1014.5 

 

Dry Nitrogen 

R46 385 301.0 4.096 3.070 1453 1170.83 

R47 303.4 4.094 3.058 1231 999.96 

R48 304.5 4.090 3.055 1301 1060.54 

R49 304.4 4.091 3.055 934.2 761.87 

R50 301.3 4.092 3.070 1529 1232.93 

R51 303.9 4.086 3.059 1342 1091.98 

R52 303.4 4.092 3.059 1386 1125.70 

R53 304.4 4.093 3.054 1448 1179.93 

R54 301.4 4.093 3.069 1451 1170.84 

R55 301.4 4.095 3.068 1288 1039.35 

R64 301.9 4.097 3.065 1373 1109.81 

R65 301.4 4.097 3.068 1449 1168.99 

R66 300.3 4.097 3.073 1179 947.99 

R67 303.0 4.087 3.063 1198 971.91 

R68 303.1 4.089 3.062 1531 1241.91 
NASA/TM—2006-214023 44 

TABLE A9.—SUMMARY OF a-PLANE FRACTURE STRENGTH FOR 

TESTING IN WATER AND DRY NITROGEN 

Sapphire Load Span: 20.176 mm 

a-plane Support Span: 40.292 mm 

Tare: 2.26 N 

Instron 8562, 5000N Environment: Distilled, Deionized Water or 99.98% Dry Nitrogen 

 Bevel Correction: 1.029 % 

 Preload 300 to 500 N 

 

Specimen Loading Stressing Width Height Failure Fracture 

ID Rate Rate b w Load Strength 

N/s MPa/s mm mm N MPa 

Distilled Water 

A1 30 22.74 4.128 3.106 478.4 374.9 

A2 22.76 4.130 3.104 2303.0 1799.6 

A3 22.74 4.130 3.105 1930.0 1507.4 

A4 22.86 4.112 3.104 1084.0 851.6 

A5 22.79 4.121 3.105 1391.0 1089.3 

A7 22.83 4.118 3.103 1351.0 1059.7 

A8 22.70 4.129 3.108 1754.0 1367.4 

A9 22.72 4.122 3.110 728.1 569.0 

A10 22.81 4.118 3.105 921.3 722.6 

A12 22.83 4.116 3.104 1661.0 1302.7 

A13 22.83 4.121 3.102 2077.0 1628.1 

A14 22.77 4.128 3.104 1521.0 1189.7 

A17 22.75 4.128 3.105 369.2 289.8 

A18 22.82 4.119 3.103 1764.0 1382.8 

A19 22.74 4.126 3.106 954.6 746.4 

A47 22.73 4.133 3.105 1477.0 1153.0 

A48 22.83 4.120 3.103 2316.0 1815.3 

A49 22.79 4.129 3.102 679.6 533.1 

A50 22.88 4.112 3.102 1393.0 1095.0 

A15 0.3 0.228442 4.110 3.105 597.3 469.8 

A16 0.229263 4.108 3.100 1242.0 978.4 

A21 0.03 0.022844 4.120 3.101 1383.0 1085.4 

A22 0.022910 4.115 3.099 1469.0 1156.1 

A23 196.33413 4.125 3.101 499.5 392.7 

A24 0.022905 4.112 3.100 1613.0 1269.0 

A25 0.022917 4.114 3.099 671.0 529.2 

A26 0.00245 0.001870 4.116 3.099 1657.0 1303.3 

 
NASA/TM—2006-214023 45 

TABLE A9 (CONTINUED).—SUMMARY OF a-PLANE FRACTURE STRENGTH 

FOR TESTING IN WATER AND DRY NITROGEN 

Sapphire Load Span: 20.176 mm 

a-plane Support Span: 40.292 mm 

Tare: 2.26 N 

Instron 8562, 5000N Environment: Distilled, Deionized Water or 99.98% Dry Nitrogen 

 Bevel Correction: 1.029 % 

 Preload 300 to 500 N 

 

Specimen Loading Stressing Width Height Failure Fracture 

ID Rate Rate b w Load Strength 

N/s MPa/s mm mm N MPa 

Distilled Water

A27 0.003 0.002291 4.121 3.096 732.8 577.7 

A28 0.002289 4.119 3.099 720.4 567.3 

A29 0.002294 4.110 3.099 629.3 496.9 

A30 0.002280 4.124 3.103 1575 1233.7 

A31 0.002285 4.116 3.103 1587 1245.4 

A32 176.999 4.121 3.100 449.3 354.0 

A33 0.002285 4.115 3.103 1709 1341.4 

A34 0.002294 4.110 3.099 713.5 563.1 

A35 0.002281 4.126 3.101 1291 1011.8 

A36 0.002286 4.111 3.104 638 501.9 

A37 0.002287 4.116 3.101 697 548.5 

A38 0.002283 4.115 3.104 1561 1224.4 

A39 0.002283 4.115 3.104 1121 879.8 

A40 0.002291 4.115 3.099 879.8 693.1 

A51 0.002282 4.118 3.104 489.5 384.9 

A66 0.002269 4.128 3.109 1064 830.0 

A68 0.002284 4.110 3.105 640.8 503.9 

A69 0.002276 4.123 3.106 1213 948.8 

A70 0.002278 4.126 3.104 491.8 386.1 

A71 0.002281 4.117 3.105 1136 890.7 

Dry Nitrogen

A41 385 292.3570 4.126 3.104 1923 1504.4 

A42 292.8184 4.119 3.104 1793 1405.0 

A43 292.1920 4.127 3.104 2792 2182.2 

A44 292.3683 4.128 3.103 2285 1787.3 

A45 291.9573 4.125 3.106 1882 1470.3 

A61 293.1241 4.124 3.100 2224 1744.1 

A62 293.5635 4.117 3.101 1711 1344.2 

A63 292.7931 4.126 3.101 1846 1446.4 

A64 292.9472 4.123 3.102 2257 1768.9 

A65 292.6776 4.117 3.105 928.9 728.4 

A52 291.9126 4.119 3.109 758.3 593.4 

A53 292.5465 4.122 3.104 676.6 530.8 

A54 292.7939 4.122 3.103 1643 1287.5 

A72 292.5355 4.119 3.105 861.9 675.7 

A73 292.6653 4.119 3.105 463.4 364.2 

NASA/TM—2006-214023 47 

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41. J.A. Salem and L.M. Powers, “Guidelines for the Testing of Plates,” pp. 357–364 in 

Proceedings of the 27th International Cocoa Beach Conference on Advanced Ceramics and 

Composites: B, Ceramic Engineering and Science Proceedings, Vol. 24, No. 4, Waltraud M. 

Kriven and H.T. Lin, editors (January, 2003). 
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52

Slow Crack Growth and Fracture Toughness of Sapphire for the International

Space Station Fluids and Combustion Facility

Jonathan A. Salem

Sapphire; Flexural strength; Fracture toughness; Crack propagation; Stress corrosion;

Fatigue life; Weibull; Structural reliability; Polishing

Unclassified -Unlimited

Subject Category: 27

Responsible person, Jonathan A. Salem, organization code RXL, 216–433–3313.

The fracture toughness, inert flexural strength, and slow crack growth parameters of the r- and a-planes of sapphire grown by the Heat

Exchange Method were measured to qualify sapphire for structural use in the International Space Station. The fracture toughness

in dry nitrogen, KIpb, was 2.31 ± 0.12 MPa√m and 2.47 ± 0.15 MPa√m for the a- and r-planes, respectively. Fracture toughness

measured in water via the operational procedure in ASTM C1421 was significantly lower, KIvb = 1.95 ± 0.03 MPa√m, 1.94 ± 0.07 and

1.77 ± 0.13 MPa√m for the a- , m- and r-planes, respectively. The mean inert flexural strength in dry nitrogen was 1085 ± 127 MPa for

the r-plane and 1255 ± 547 MPa for the a-plane. The power law slow crack growth exponent for testing in water was n = 21 ± 4 for the

r-plane and n ≥ 31 for the a-plane. The power law slow crack growth coefficient was A = 2.81 x 10–14 m/s·(MPa√m)–n for the r-plane

and A 2.06 x 10–15 m/s·(MPa√m)–n for the a-plane. The r- and a-planes of sapphire are relatively susceptible to stress corrosion

induced slow crack growth in water. However, failure occurs by competing modes of slow crack growth at long failure times and

twinning for short failure time and inert environments. Slow crack growth testing needs to be performed at low failure stress levels and

long failure times so that twinning does not affect the results. Some difficulty was encountered in measuring the slow crack growth

parameters for the a-plane due to a “short” finish (i.e., insufficient material removal for elimination of the damage generated in the

early grinding stages). A consistent preparation method that increases the Weibull modulus of sapphire test specimens and components

is needed. This would impart higher component reliability, even if higher Weibull modulus is gained at the sacrifice of absolute

strength of the component. The current specification frequently used for the preparation of sapphire test specimens and components

(e.g., a “60/40” scratch-dig finish) is inadequate to avoid a “short” finish.