Industrial Materials for the Future 

Managed by 

UT-Battelle, LLC 

Final Technical Report 

Compressive Creep and Thermophysical 

Performance of Refractory Materials 

June 2006 

Principal Investigators:

Dr. M. K. Ferber 

Dr. A. A. Weresczak 

Dr. J. G. Hemrick 

Oak Ridge National Laboratory 

ORNL/TM-2005/134 
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FINAL TECHNICAL REPORT

Project Title: Compressive Creep and Thermophysical Performance of Refractory 

Materials 

Contract No.: DE-AC05-00OR22725 

Project Period: April 1997–June 2000 

PI(s): Dr. M. K. Ferber 

 (865) 576-0818 

 ferbermk@ornl.gov

 Dr. A. A. Wereszczak 

 (865) 576-1169 

 wereszczakaa@ornl.gov

 Dr. J. G. Hemrick 

 (865) 574-7601 

 hemrickjg@ornl.gov

Recipient: Oak Ridge National Laboratory (ORNL) 

 Bethel Valley Road 

 P. O. Box 2008 

 Oak Ridge, TN 37831 

Subcontractor University of Missouri–Rolla 

 Department of Materials Science Engineering 

 222 McNutt Circle 

 1870 Miner Circle 

 Rolla, MO 65409 

ORNL/TM-2005/134 

Compressive Creep and Thermophysical Performance 

of Refractory Materials 

Dr. M. K. Ferber, Dr. A. A. Wereszczak, Dr. J. G. Hemrick 

Oak Ridge National Laboratory

June 2006 

Prepared by 

OAK RIDGE NATIONAL LABORATORY 

P.O. Box 2008 

Oak Ridge, Tennessee 37831-6283 

managed by 

UT-Battelle, LLC 

for the 

U.S. DEPARTMENT OF ENERGY 

under contract DE-AC05-00OR22725 
ii 

Acknowledgments, Disclaimer, and 

Proprietary Data Notice 

Acknowledgments 

This report is based upon work supported by the U.S. Department of Energy, Office of Energy 

Efficiency and Renewable Energy, Industrial Technologies Program, Industrial Materials for the 

Future, under Contract No. DE-AC05-00OR22725. The authors are also indebted to the following 

individuals for supplying the mullite refractories: Paul Hutchinson (DSF Refractories & Minerals), 

Michael Nelson (Corhart Refractories), Larry Stover (Minteq International), D. J. Thomas, (VGTDYKO Industrial Ceramics), Jim Houp (National Refractories & Minerals), and Don McIntyre 

(NARCO); the guidance provided by the Glass Industry Advisory Committee, which comprises the 

following individuals: Larry Kotacska (Corning), Al Poolos (Owens-Brockway), Randy Sturtz 

(Owens-Corning), Warren Curtis (PPG), Jim Shell (Techneglas), Mike Nelson (Corhart), Steve 

Winder (UKSS/Monofrax), John Tracey/Dilip (Patel VRD-Americas), and Bob Moore (University of 

Missouri-Rolla); ORNL/HTML machinists L. O’Rourke, T. Jenkins, and R. Parten for the specimen 

preparation; and Peter Angelini, ORNL, for project guidance and review. 

Oak Ridge National Laboratory is operated by UT-Battelle, LLC, for the U.S. Department of Energy 

under contract DE-AC05-00OR22725. 

Disclaimer 

This report was prepared as an account of work sponsored by an agency of the United States 

Government. Neither the United States Government nor any agency thereof, nor any of their 

employees, makes any warranty, express or implied, or assumes any legal liability or responsibility 

for the accuracy, completeness, or usefulness of any information, apparatus, product, or process 

disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any 

specific commercial product, process, or service by trade name, trademark, manufacturer, or 

otherwise, does not necessarily constitute or imply its endorsement, recommendation, or favoring by 

the United States Government or any agency thereof. The views and opinions of authors expressed 

herein do not necessarily state or reflect those of the United States Government or any agency thereof. 

Proprietary Data Notice 

If there is any patentable material or protected data in the report, the recipient, consistent with the data 

protection provisions of the award, must mark the appropriate block in Section K of the DOE F 241.3, 

clearly specify it here, and identify them on appropriate pages of the report. Other than patentable 

material or protected data, reports must not contain any proprietary data (limited rights data), 

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subject to release. Protected data are specific technical data, first produced in the performance of the 

award, which is protected from public release for a period of time by the terms of the award 

agreement. Reports delivered without such notice may be deemed to have been furnished with 

unlimited rights, and the Government assumes no liability for the disclosure, reproduction, or use of 

such reports. 
Contents 

List of Figures .................................................................................................................................... v

List of Tables...................................................................................................................................... vii

Abbreviations and Acronyms............................................................................................................. ix

1. Executive Summary..................................................................................................................... 1

1.1 Research and Development................................................................................................ 1

1.2 Technology Transfer .......................................................................................................... 4

1.3 Commercialization ............................................................................................................. 4

1.4 Recommendations.............................................................................................................. 4

2. Introduction.................................................................................................................................. 7

3. Background.................................................................................................................................. 9

4. Materials and Experimental Procedures ...................................................................................... 11

4.1 Materials............................................................................................................................. 11

4.2 Creep Testing ..................................................................................................................... 11

4.3 Specimen Preparation......................................................................................................... 14

4.4 Creep Testing and Data Interpretation ............................................................................... 16

4.5 Microstructural Characterization ....................................................................................... 18

4.6 Supplemental Analysis....................................................................................................... 19

5. Results and Discussion ................................................................................................................ 23

5.1 Silica Refractories .............................................................................................................. 23

5.2 Mullite Refractories ........................................................................................................... 28

5.3 Fusion-Cast Alumina Refractories..................................................................................... 34

5.4 Fusion-Cast Spinel Refractories......................................................................................... 47

6. Accomplishments......................................................................................................................... 53

6.1 Research-Based Accomplishments .................................................................................... 53

6.2 Technology Transfer .......................................................................................................... 53

6.3 Publications and Patents..................................................................................................... 53

6.4 Commercialization ............................................................................................................. 54

7. Summary and Conclusions .......................................................................................................... 55

7.1 Silica................................................................................................................................... 55

7.2 Mullite................................................................................................................................ 56

7.3 Fusion-Cast Alumina ......................................................................................................... 58

7.4 Fused Spinel....................................................................................................................... 61

8. Recommendations........................................................................................................................ 63

9. References.................................................................................................................................... 65

 Appendix: Publications................................................................................................................ 67

iii 

List of Figures 

 1 Superstructure refractories in glass production furnaces....................................................... 7

 3 Primary components of creep frame...................................................................................... 12

 2 Schematic of creep frame and supporting instruments.......................................................... 12

 4 Schematic of creep frame extensometer................................................................................ 13

 5 Dimensions of core-drilled test specimens............................................................................ 14

 6 Schematic of how specimens were core-drilled from as-received bricks.............................. 14

 7 Refractory block core-drill locations for physical and chemical analysis............................. 15

 8 Schematic representation of the sectioning plan used to fabricate creep specimens 

from the bulk region of the refractory blocks........................................................................ 16

 9 Creep specimen assembly H-shell arrangement required for the creep testing of the 

fusion-cast alumina................................................................................................................ 16

10 Example of temperature and stress history used in the creep testing .................................... 18

11 Schematic describing how disk sections were machined from core-drilled specimens 

for (a) as-received and (b) post-test RL and CL imaging...................................................... 18

12 Schematic of transient plane source direct thermal conductivity method experimental 

set-up and picture of Kapton sensor ...................................................................................... 19

13 Schematic of dual-rod LVDT system used to measure thermal expansion........................... 20

14 Creep behavior of Vega as a function of temperature........................................................... 25

15 Change in volume of “aged” specimens as a function of time and temperature ................... 26

16 Schematic of the change in microstructure from the as-received material state to that 

after creep or high-temperature exposure.............................................................................. 27

17 Vega’s microstructural changes as a function of temperature as revealed by 

reflected-light microscopy..................................................................................................... 28

18 Vega’s microstructural changes as a function of temperature as revealed by 

cathodoluminescence imaging .............................................................................................. 29 

19 Amount of refractory recession when exposed to sodium carbonate at 1400°C for 24 h ..... 30

20 Amount of recession as a function of temperature and sodium carbonate addition 

for Gen-Sil silica ................................................................................................................... 31

21 Thermal conductivity of mullite refractories as measured by the “hot-disk” method........... 33

22 Creep rate measured at 0.6 MPa as a function of reciprocal temperature ............................. 33

23 Refractory recession when exposed to sodium carbonate at 1400oC for 24 h 

(according to ASTM C98)..................................................................................................... 35

24 x3-, x7-, and y-core density profiles for the (a) Corhart refractory and (b) Monofrax 

refractory ............................................................................................................................... 36

25 Optical microscopy of characteristic grain structures for (a) bulk structure, and 

(b) columnar structure ........................................................................................................... 37

26 Optical microscopy of columnar region (micron bar = 200 μm) .......................................... 39

27 Optical microscopy of nephelitic grain boundary phases...................................................... 39

28 Static Young’s modulus values for fusion-cast alumina sample types ................................. 40

29 Specimen elongation as a function of time for (a) the bulk Corhart and (b) the bulk 

Monofrax refractories............................................................................................................ 42

30 Monofrax thermal conductivity block section denoting test surfaces ................................... 43

31 Average laser flash thermal conductivity for as-cast bulk refractories: (a) Corhart and 

(b) Monofrax ......................................................................................................................... 44

32 Combined plots of expansion and creep effects in Corhart bulk samples............................. 45

33 Combined plots of expansion and creep effects in Monofrax bulk samples ......................... 46

34 Fused spinel accumulated strain curves ................................................................................ 48

v 
35 Bonded spinel accumulated strain curve at 1650°C.............................................................. 48

36 Bonded spinel accumulated strain curve at 1550°C.............................................................. 49

37 Bonded spinel accumulated strain curve at 1550°C.............................................................. 49

38 Thermal diffusivity of fused and bonded spinel ................................................................... 50

39 RL (left) and CL (right) images of fused spinel.................................................................... 51

vi 
List of Tables 

 1 Project and survey participants.............................................................................................. 10

 2 Ranking of refractory categories whose creep data are desired ............................................ 10

 3 Silica refractories tested in the present study and their manufacturers ................................. 11

 4 Mullite refractories tested in the present study and their manufacturers............................... 11

 5 Creep test matrix for the silica refractories (stress, MPa) ..................................................... 17

 6 Creep test matrix for the mullite refractories (stress, MPa) .................................................. 17

 7 Creep-test matrix for the fusion-cast alumina refractories (stress, MPa) .............................. 17

 8 Dimensional changes of crept specimens.............................................................................. 23

 9 Density changes of crept specimens...................................................................................... 24

10 Summary of microstructural observation for the mullite refractories before 

and after creep testing............................................................................................................ 31

11 Fusion-cast alumina nanoindentation results......................................................................... 40

12 Calculated strain rates ........................................................................................................... 48

vii 

Abbreviations and Acronyms 

AIM Advanced Industrial Materials 

ASTM American Society for Testing and Materials 

AZS aluminum-zirconium silicate 

CL cathodoluminescent imaging 

DOE U.S. Department of Energy 

EDS energy-dispersive X-ray spectroscopy 

LVDT linear variable differential transducer 

ITP Industrial Technology Program 

NBA Norton-Bailey Arrhenius 

OIT Office of Industrial Technologies 

ORNL Oak Ridge National Laboratory 

RL reflected light microscopy 

SEM scanning electron microscopy 

TPS transient plane source 

XRD X-ray diffraction 

XRF X-ray fluorescence 

ix 

1. Executive Summary 

1.1 Research and Development 

Over the last 10 years, the U.S. Department of Energy Office of Industrial Technology, through the 

Industrial Materials for the Future activity, has funded a number of projects at industry, university, 

and national laboratory sites to increase the energy efficiency of energy-consuming industries, 

including glass, steel, pulp and paper, and chemical. In the case of the glass industry, using valid 

engineering creep data to optimize furnace superstructure design is a means to achieve these goals. 

The objective of the ORNL project described in this report was to facilitate that endeavor by 

conducting creep testing and analysis on refractories of interest to glass manufacturers at 

representative service temperatures, enabling the availability of new and improved refractories by 

suppliers, and generating creep data on “equivalent” refractories that furnace designers can use for 

optimizing the design of their superstructures or for predicting their long-term structural integrity. 

The choice of refractories evaluated in the project was based on the advice of technical experts from 

the glass manufacturing industry, refractory suppliers, and academia. Although the results cited in this 

report are specific to refractories for glass applications, the techniques developed in this project are 

applicable to refractories used in all industrial applications. 

The creep and corrosion behaviors of several refractories, including six types of conventional silica, 

ten types of mullite, two types of fusion-cast alumina, and fused spinel were measured. The data 

generated for the silica refractories served as a baseline. The mullite refractories were examined 

because they are used in borosilicate glass furnace crowns and superstructures and in sidewall 

applications. Additionally, despite their high cost (≈$500/ft3

) they are cheaper than other refractories 

such as chrome alumina (≈ $3000/ft3

) or fusion-cast alumina (≈ $900/ft3), which are used as 

replacements for traditional silica refractories in harsh oxy-fuel environments. Fusion-cast alumina 

refractories, which are a popular choice for crowns and sidewalls in oxy-fueled glass furnaces, were 

selected for this study because of their homogeneity, low porosity, high refractoriness, low levels of 

glassy phase, and good spall resistance. Although more expensive, the fused spinel offers improved 

creep and corrosion resistances. 

Compressive creep testing was performed at 1300 to 1650o

C and at static stresses between 0.2 and 

0.6 MPa on a pneumatically controlled compressive creep frame in an electrically heated furnace. The 

phase content, microstructure, and secondary phase composition of the refractories were characterized 

before and after their creep testing in an attempt to correlate the creep responses with any observed 

changes. Corrosion resistance of each of the refractories was also examined. Lastly, the effect of 

temperature and time (in the absence of stress) on the dimensional stability of the materials was 

examined. 

In the case of the silica refractories, the compressive creep rates of all six brands could not be 

represented as a function of temperature and compressive stress by using the conventionally used 

Norton-Bailey Arrhenius (NBA) creep equation. Concurrently active mechanisms other than creep 

resulted in larger or oppositely anticipated dimensional changes than those produced by creep; this 

effect limited the ability to identify or interpret the less-active creep mechanism in them. The change 

in dimensions of the compressively crept specimens indicated that their size (both diameter and 

length) increased (by 0.35 to 1.0%) as a consequence of the creep test conditions. The cause of the 

volume expansion was not identified; however, because of the absence of quartz, it can be concluded 

that the expansion was not due to quartz (a relatively high-density phase) converting to tridymite or 

cristobalite (lower-density phases). The secondary phase constituents remained constant in all brands 

1 
when they were tested at 1550°C. A fraction of these phases visibly evolved in the specimens at 

1600°C. This phenomenon was quite severe at 1650°C. The density changes in the crept specimens 

(a net effect of the volume expansion and loss in secondary phase mass) ranged from a decrease of 

1.6 to 3.9% with a subtle trend of greater density decreases at higher test temperatures. The density 

decreases among the six brands were statistically equivalent within the data scatter. The corrosion 

resistances of the six silicas were statistically equivalent when they were exposed to sodium 

carbonate at 1400°C (2550°F) for 24 h. The amount of recession increased linearly with temperature 

between 1250 and 1450°C for the Gen-Sil silica. 

Creep rates for three brands of the mullite refractories were on the order of 10!11 s!1

 at the lower test 

temperature and 10!10 s!1

 at the higher test temperature. Creep rates for four of the brands ranged 

from 10!10 s!1

 at the lower test temperature and to 10!9 s!1 at the higher test temperature. The 

remaining three brands exhibited creep rates ranging from 10!9

 s!1 to 10!8 s!1 and exhibited low stress 

exponents (< 0.5) and changes in chemistry or in microstructure due to the creep-testing conditions. 

Several of the mullite refractories exhibited a creep stress exponent approximately equal to unity, 

suggesting that creep was controlled by diffusion. One of mullite refractories had a stress exponent of 

2.4, indicating that the rate-controlling mechanism may be grain boundary creep. The creep 

mechanism for the remaining mullite materials, which had stress exponents much less that unity, was 

attributed to a non-steady-state condition existing during the entire extent of testing. This condition is 

a consequence of contraction or time-hardening effects occurring during testing and of the samples 

never reaching a state of equilibrium. Those brands possessing high levels of matrix porosity all 

showed compaction of the microstructure during creep testing, which was ultimately responsible for 

their poorer creep resistance. One of the mullite refractories was found to be composed of andalusite 

grains containing free quartz, which was converted to cristobalite during firing. Further, the 

andalusite grains were fully converted to mullite during creep testing, and the amount of glassy phase 

increased. These events in combination may lead to the brand’s poor creep resistance. 

Activation energies for the various mullite refractory brands ranged from 50 to 223 kcal/mol. 

Although no direct correlation between the magnitude of the activation energy and the amount of 

creep exhibited by a particular brand could be drawn from the data, it is known that materials 

possessing a high activation energy will have greater sensitivity to temperature than those possessing 

a lower activation energy. Also, the values for many of the calculated activation energies correspond 

to the activation energy of vitreous silica. This finding indicates that the glassy phases in these brands 

affect the overall creep behavior of the material. 

The distinction of a brand being fused-grained or non-fused-grained mullite did not appear to have a 

bearing on the creep resistance of the individual brands. 

Eight of the ten mullite refractory brands showed less than a 0.2% change in mass and less than a 

0.6% change in volume after thermal aging in the absence of an applied load. The two exceptions 

exhibited extraneous amounts of change (≈8.5%) in both mass and volume due to thermal exposure. 

All samples showed less than a 0.5% change in bulk density. The corrosion rates of the ten mullite 

brands could not be determined through the use of ASTM C987, but it was determined visually that 

only minimal amounts of recession occurred because of exposure as defined by the ASTM lid test. 

Fused-cast alumina refractories from Corhart and Monofrax were examined. Both the density and 

porosity of the as-cast refractory blocks were highly consistent. The phase profiles for the α- to 

β-alumina ratios also showed similar trends to those seen in the density/porosity profiles. X-ray 

diffraction (XRD) showed that both brands consisted of only peaks identifiable as corundum (PDF# 

10-0173) or sodium aluminum oxide (PDF# 31-1263). Closer examination of the XRD patterns 

showed that there were some differences in ratios of α- to β-alumina between the two brands and 

2 
between the bulk and columnar microstructures. The microstructures of the two brands of refractory 

were also comparable with a columnar region existing in the outer 1 to 1 1⁄2 in. of the blocks and then 

transitioning to a more equiaxed structure of α-alumina grains embedded in a β-alumina matrix 

toward the center of the block. Average grain sizes were found to be on the order of 172 μm and 

82 μm for the Corhart bulk and columnar regions, respectively. Average grain sizes were found to be 

on the order of 231 μm and 79 μm for the Monofrax bulk and columnar regions, respectively. 

Both columnar orientations show much less accumulated strain than the bulk samples under similar 

stresses because the elongated needle-like grains, which are oriented perpendicular to the direction of 

the applied compressive load, inhibit creep in the transverse columnar samples. In the longitudinal 

columnar case, it is expected that the unrecoverable expansion effect plays a role in the reduced strain 

accumulation. The columnar samples at 1650°C showed less accumulated strain than the bulk sample 

under similar stress, but the behavior between the bulk and columnar samples was more similar. Both 

microstructures showed traditional creep behavior under 1.0 MPa of stress and much larger amounts 

of accumulated strain than the bulk samples at 0.6 and 0.8 MPa. 

The three bulk 1650°C/1.0 MPa samples all had similar creep rates, with the Corhart material on the 

order of 5 x 10!6

 h!1 and the Monofrax material on the order of 6 x 10!6 h!1. The columnar structures 

at 1650°C/1.0 MPa had a lower creep rate, on the order of 4.25 to 4.50 x 10!6

 h!1. Creep rates for the 

other samples were not obtainable because nontraditional creep behavior occurred. Stress exponents 

and activation energies could not be calculated from the data because steady-state creep rates were 

only obtained at one stress and temperature. 

Unrecoverable expansion was seen in all materials and was thought to be connected to the 

nontraditional creep behavior seen in the compressive creep testing at temperatures and loads up to 

1650°C and 0.8 MPa. As the temperature increased, the amount of unrecoverable expansion occurring 

over a given time interval also increased. The occurrence of the unrecoverable expansion was thought 

to be caused by thermal expansion hysteresis due to microstresses or microcracking. Because of that 

behavior, the generated creep curves could not be represented by the empirical NBA creep equation. 

Instead, an empirical fit utilizing an exponential expansion series with multiple time constants was 

used to represent the behavior exhibited by the fusion-cast alumina refractories. The same exponential 

fit used to model the raw creep curves was also utilized to describe this composite behavior. 

Several samples were analyzed after being crept, and results were compared with those previously 

obtained from virgin samples. Analysis included thermal conductivity, optical microscopy, and 

scanning electron microscopy/energy-dispersive X-ray spectroscopy (SEM/EDS) chemical analysis. 

The thermal conductivity values for the Monofrax post-creep cores were similar to those obtained 

earlier from the virgin cores, which had thermal conductivities in the range of 5.9 W/mK. The value 

obtained from the Corhart post-creep core was significantly lower than the average value of 

6.61 W/mK found earlier for the Corhart bulk material. Initial optical microscopy and SEM/EDS 

analyses were performed on Corhart and Monofrax bulk samples exposed to 1550°C and 1.0 MPa. 

This work was conducted to determine whether any changes in the microstructure or chemistry could 

be detected at that intermediate condition. No discernible deformation and only slight chemical 

changes were seen in the samples. Subsequent optical microscopy and SEM/EDS analyses were 

performed on Corhart and Monofrax bulk samples exposed to the extreme conditions (1650°C and 

1.0 MPa). The structure of the Corhart sample was found to resemble the virgin structure; the 

Monofrax crept structure displayed a smaller α-alumina grain size and a tighter α-alumina grain 

structure with more equiaxed angular grains than in the virgin structure. More severe chemical 

changes were also seen between the virgin and crept samples. 

3 
The fused spinel materials had negligible creep rates up to 1650°C at 8.5MPa of stress. The bonded 

spinel, surprisingly, crept substantially compared to the fused material. However, this must be 

compared with other common crown materials; the total creep of the bonded material is less than 

0.5% over 100 h at 1550°C and 0.3 MPa which are fairly standard temperature and pressure for 

testing refractory materials The creep of silica materials is approximately 0.4% over 100 h at 1450°C 

and 0.3 MPa. The spinel materials did not react with soda in the exposure tests. There was some 

limited reaction with silica and calcia from the gasifier atmosphere. Thermal conductivity was 

determined by laser flash and was found to decrease as temperature increases from 14 (100°C) to 

8 (700°C) W/mK for the fused spinel and 9.5 (100°C) to 4 (1300°C) for the bonded spinel. A model 

of the glass tank crown was developed that takes into account nonlinear changes in the refractory 

material due to time, temperature, and corrosion. 

1.2 Technology Transfer 

The results of this project have established hardware and a test methodology that can be used by the 

refractories community to evaluate the creep and aging behaviors of most refractory materials. Of 

particular merit is the development of specialized extensometry, which offers extremely high 

resolution required for the measurement of small deformations. The project staff made presentations 

at national meetings and published results in industry technical journals. 

1.3 Commercialization 

The data generated in this project were provided to participating refractory manufacturers and are 

available from those manufacturers to any users of these types of refractories. It has been confirmed 

with industrial partners involved in this project that creep data and other thermomechanical property 

information generated as part of this project was used in analysis of glass furnaces currently in service 

and for the planning and design of refractory systems for future furnace constructions. The data has 

not only been used by glass companies in their analysis and planning of furnace systems, but by 

refractory companies and industrial consultants as well. The information generated was said to be 

“invaluable” by one industrial participant and more data of similar scope is felt to be of general 

interest to the industry as a whole. 

1.4 Recommendations 

Technical efforts should be undertaken to measure and model the composite effects of unrecoverable 

expansion (aging behavior) and creep response. Given that these processes may be highly nonlinear, 

the traditional equations, which assume that steady-state deformation rates are dominant, are not 

appropriate. New models must be derived from basic theoretical principles that can describe the 

nonlinear, transient, nonsteady equilibrium behavior observed in this work. Creep testing needs to 

include simultaneous measurement of specimen elongation/contraction occurring in the absence of an 

applied stress. Such testing could be conducted using a system consisting of furnace of sufficient size 

to accommodate multiple mechanical load trains. During testing one load train would be used to 

monitor elongation/contraction in the absence of any appreciable stress. The stress range employed 

with the other load stations must be large enough to promote significant and measurable changes in 

the deformation response. It is recommended that the stress range cover at least one order of 

magnitude (e.g., 5 to 50 MPa). Testing times must also be much longer (> 1000 h) to ensure that 

representative deformation data are generated. A second approach for distinguishing creep from aging 

effects is to use the stress relaxation technique for measuring creep response. Because this method is 

able to generate an entire creep-rate vs stress curve in a matter of a few hours, one could use it to 

4 
measure creep behavior at periodic intervals during a standard aging test. Thus a single sample could 

be used to assess both aging and creep behavior. 

To facilitate technology transfer and commercialization, software tools should be developed which 

allow users to readily link the key properties to finite element codes used in furnace construction. 

This software would allow one to quickly assess the effects of specific materials upon long-term 

thermal and structural reliability. It should also include a routine that allows the user to optimize 

performance of a given material by changing the material properties; these altered properties would 

serve as new targets for those involved in materials development. 

Any follow-on effort must also address new materials. 

5 

2. Introduction 

Furnace designers and refractory engineers recognize that optimized furnace superstructure design 

and refractory selection are needed as glass production furnaces are continually striving toward 

greater output and efficiencies. Harsher operating conditions test refractories to the limit, while 

changing production technology (such as the conversion to oxy-fuel from traditional air-fuel firing) 

can alter the way the materials perform [1–3]. Refractories for both oxy- and air-fuel fired furnace 

superstructures (see Fig. 1) are subjected to high temperatures that may cause them to creep 

excessively or subside during service if the refractory material is not creep resistant, or if it is 

subjected to high stress, or both. Furnace designers can ensure that superstructure structural integrity 

is maintained if the creep behavior of the refractory material is well understood and well represented 

by appropriate engineering creep models. 

Fig. 1. Superstructure refractories in glass production furnaces. 

Several issues limit the abilities of furnace designers to (1) choose the optimum refractory for their 

applications, (2) optimize the engineering design, or (3) predict the service mechanical integrity of 

their furnace superstructures. Published engineering creep data are essentially nonexistent for almost 

all commercially available refractories used for glass furnace superstructures. The limited data that do 

exist are supplied by the various refractory suppliers. Unfortunately, the suppliers generally have 

different ways of conducting their mechanical testing, and they interpret and report their data 

differently. This inconsistency makes it hard for furnace designers to draw fair comparisons between 

competing grades of candidate refractories. Furthermore, the refractory suppliers’ data are often not 

available in a form that can be readily used for furnace design or for the prediction and design of 

long-term structural integrity of furnace superstructures. 

As a consequence, the U.S. Department of Energy (DOE) Industrial Technology Program (ITP) Glass 

Industry of the Future sponsored research and development at industry, university, and national 

laboratory sites with the intent to help domestic glass manufacturers improve their energy and 

operating efficiencies. The optimization of furnace superstructure design using valid engineering 

creep data is a means to achieving these ITP goals. The present project at Oak Ridge National 

Laboratory (ORNL) aided in this endeavor by conducting creep testing and analysis on refractories of 

interest to glass manufacturers at representative service temperatures, enabling the availability of new 

and improved refractories by refractories suppliers and by generating creep data on equivalent 

refractories that furnace designers could use for optimizing the design of their superstructures or for 

predicting their long-term structural integrity. Similar refractory creep-testing projects have been 

7 
conducted at ORNL [4–6], so many of the unique experimental nuances and difficulties associated 

with the high-temperature creep testing of refractories have been encountered and overcome. 

8 
3. Background 

The ITP program has recognized that a reduction in overall domestic energy consumption will occur 

if the primary energy-consuming industries improve their own energy efficiencies [7]. Recognizing 

this need, the glass industry is currently converting older, conventional air-fuel-fired furnaces to oxyfuel firing, or in the case of new construction, is building new oxy-fuel-fired furnaces instead. This 

has caused oxy-fuel technology to become one of the fastest growing technologies in the glass industry because it promises pollution abatement, increased glass-pull effectiveness, capital cost savings, 

and increased energy efficiency. For example, a recent study [8] has shown that approximately 

$202M in energy savings per year in 2005 and a $445M per year savings by 2020 could be expected 

with the conversion of air/fuel to oxy-fuel–fired glass manufacturing furnaces. These results, which 

reflect energy savings of 2.8 and 14.2 TBtu/year, respectively, are based on the projection that 61% 

and 100% furnace conversions will occur by the years 2005 and 2020, respectively. 

Unfortunately, the new internal furnace environment produced by oxy-fuel firing creates many problems with refractory life not previously encountered with conventional firing. These problems include 

increased potential for corrosion and “rat holing” along with deterioration of traditionally used conventional silica crown refractories. Additionally, due to a significant increase in the concentration of 

volatiles in the furnace, the refractories can experience a more intense alkali attack as water vapor 

reacts with alkali to form hydroxide vapors. Therefore, alternatives to silica must be considered for 

refractory superstructure applications, and data must be obtained to predict the performance of the 

refractories in service. Furthermore, the cost of the alternate refractories must be balanced with their 

effects on furnace life and glass quality while also considering the comparative performance of the 

other refractories in the furnace and overall furnace life.

Given the need for data on alternative refractories, the first objective of this project was to identify 

superstructure refractory categories of greatest interest to the glass manufacturing industry via a survey. To achieve this, the advice of project participants and technical experts from the glass manufacturing industry, refractory suppliers, and academia was first sought to organize a list of distinct 

refractory categories to include in the survey (Table 1). 

The mechanical deformation of thirteen different candidate refractory categories was determined to 

be of possible interest to the flat, fiber, container, and specialty glass manufacturers (Table 2). A list 

of the categories was then forwarded to nine refractory suppliers (Table 1), and company representatives were asked to indicate if any of their marketed refractory brands belonged to those categories. 

From their input, anywhere between four to ten competing or “equivalent” brands were identified per 

category. 

A survey consisting of the thirteen categories (and subcategory brands) was then forwarded to a furnace or refractory specialist at more than thirty domestic glass manufacturers (Table 1). The person 

taking the survey was asked to rank each refractory category as “high,” “middle,” or “low” priority in 

terms of their desire to have access to engineering creep data for that particular refractory category. 

To quantify the results of the survey, a score of ten, five, and zero points was respectively given to a 

“high,” “middle,” and “low” priority. The total points were summed and then used to rank and compare the interest among the thirteen categories. 

Eight categories received relatively high total scoring (> 100 points) while the remaining five 

received little priority (< 100 points) (Table 2). Creep data for fusion-cast alumina was of greatest 

interest to the most respondents followed in descending order by creep data for conventional silica, 

9 
Table 1. Project and survey participants 

Subject matter experts 

P. Antimarino, Harbison-Walker M. Nelson, Corhart S. Winder, UKSS 

R. Antram, NARCO J. Petreanu, Owens-Corning J. Wosinski, Corning 

W. Curtis, PPG P. Ross, Glass Industry Consulting 

R. Moore, University of Missouri-Rolla J. Shell, Techneglas 

Refractory suppliers 

A. P. Green Harbison-Walker NARCO 

Corhart/SEPR Monofrax VGT-DYKO 

DSF National Zedmark 

Survey recipients 

AFG Industries Glenshaw Philips Lighting 

Anchor Glass Guardian Industries Pilkington 

Ball-Foster Glass Holophane Pittsburgh Corning 

Cardinal FG Kimble Glass PPG 

Carr-Lowrey Knauf Fiber Glass PQ Corporation 

CertainTeed Lancaster Glass Schuller International 

Corning Leone St. George Crystal 

Corning Glass - RBMG Libbey Techneglas 

Ford Osram Sylvania Thomson Consumer Electronics 

Gallo Glass Owens-Brockway Vetrotex 

General Electric Lighting Owens-Corning Wheaton 

fusion-cast AZS, fused-grain mullite, bonded AZS, andalusite, fused silica, and bonded zircon. The 

relatively high interest in creep data for fusion-cast alumina and AZS is indicative of the glass 

industry’s desire to convert to oxy-fuel firing and of the belief that these refractory categories may be 

attractive alternatives to traditionally used (but perhaps unsuitable) conventional silica refractories. 

The availability of creep data for mullite, spinel, or-containing, and 5–15% chromium-containing 

refractories received little interest from the respondents. 

Creep testing was initiated in 1998. In addition to creep 

testing, the brand in each category was characterized by 

analyzing its (1) grain size and grain size distribution, (2) 

before- and after-testing microstructure using 

catholuminescence, (3) before- and after-testing phase 

content using X-ray diffraction (XRD), (4) before- and aftertesting density, and (5) corrosion resistance [9]. The results 

from efforts (1)–(4) were used to explain any observed 

differences among the measured creep responses of the 

refractories, while the results from effort (5) were used to 

assess their corrosion responses. When the program was 

completed in 2000, the creep and corrosion behaviors of six 

types of conventional silica, ten types of mullite, and two 

types of fusion-cast alumina had been examined. This report 

summarizes the key findings obtained from the studies and 

provides recommendations for follow-on research. Detailed 

results are provided elsewhere [10–13]. 

Table 2. Ranking of refractory 

categories whose creep data are 

desired. The top eight ranked 

categories were creep tested at ORNL. 

Ranking Refractory category 

1 Fusion-cast alumina 

2 Conventional silica 

3 Fusion-cast AZS 

4 Fused-grain mullite 

5 Bonded AZS 

6 Andalusite 

7 Fused silica 

8 Bonded zircon 

9 MgO 

10 Spinel 

11 50–80% Chrome 

12 ≈ 30% Chrome 

13 5–15% Chrome 

10 
4. Materials and Experimental Procedures 

4.1 Materials 

Six conventional silica refractories were analyzed in the present study (see Table 3). All brands were 

“Type A” [14] conventional silicas except SI96AU (a Type B silica refractory); namely, their flux 

factors (alumina content plus twice the alkali content) were less than 0.50%. The chosen refractories 

provided a comparison between similar Type A silicas at well as comparison between Type A and 

Type B silicas. According to the published Teltech Report Crown Refractories for Glass 

Manufacturing with Oxy-Fuel Combustion [1], Type B conventional silica is not recommended for 

crowns; however, its mechanical performance at representative furnace crown temperatures and 

stresses were characterized for the sake of comparison. The major secondary-phase constituent in all 

six brands was calcia (CaO), with additional reported constituents consisting of alumina (Al2O3), iron 

oxide (Fe2O3), magnesia (MgO), titania (TiO2), sodium oxide (Na2O), and potassium oxide (K2O). 

The amounts of these secondary phase constituents were semiquantitatively measured by the authors 

using SEM/EDS and are presented and described in later sections. 

Ten mullite refractories were analyzed in the 

present study (see Table 4). Seven of the 

materials were fused-grained and three were 

not (although one non-fused-grain material 

did contain a small amount of fine fused 

mullite grains). Many of the manufacturers 

listed have merged or changed names since 

program initiation. 

Table 3. Silica refractories tested in the present study 

and their manufacturers 

Refractory Brand Name Manufacturer 

Gen-Sil A. P. Green 

SI96AU VGT-DYKO 

Stella GGS NARCO 

Vega Harbison-Walker 

Vega H Harbison-Walker 

Unnamed developmental brand Harbison-Walker 

Two brands of M-type fusion-cast 

alumina refractories (Corhart 

Refractories of Louisville, 

Kentucky, and Monofrax Inc. of 

Falconer, New York) were 

examined. All specimens were 

fabricated from regular fusion-cast 

alumina blocks of approximate 

dimensions 0.60 H 0.50 H 0.30 m 

(24 H 20 H 12 in.). Key findings are 

summarized below. 

Table 4. Mullite refractories tested in the present study and 

their manufacturers 

Brand Name Type Manufacturer 

BP mullite fused-grained Corhart 

DURITAL S75 non-fused-grained NARCO 

Frimul F fused-grained DSF 

GEM fused-grained NARCO 

HF 17 fused-grained DSF 

MU75AF fused-grained VGT-DYKO 

SL75AD non-fused-grained VGT-DYKO 

UFALA UCR non-fused-grained Harbison-Walker 

ZED FMC fused-grained Minteq 

ZED FM fused-grained Minteq 

Two types of spinel were examined: fusion cast and reaction bonded. In the case of the fusion cast 

material, specimens having both columnar and equiaxed grains were examined. 

4.2 Creep Testing 

The testing frame used for the compressive creep testing consisted of a large clamshell furnace that 

used resistance heating, a contacting extensometer that was used to measure axial dimension changes, 

and silicon carbide push rods that transferred the applied compressive load from an actuator to the test 

specimen (see Figs. 2 and 3). 

11 
Fig. 2. Schematic of creep frame and supporting 

instruments. 

Fig. 3. Primary components of creep frame. 

Creep tests were accomplished by using either digitally controlled pneumatic power (by LabVIEWTM

software, National Instruments, Austin, Texas) or hydraulic power controlled by an analog controller. 

The latter feature provides both static (creep) and dynamic loading. The test specimen was heated 

with a high-temperature furnace capable of 1800o

C. Within ASTM guidelines, temperature 

fluctuations were approximately ± 2°C and load fluctuations were less than 1% of the test load. The 

specimen was compressed by two long push rods (each 0.3 m long) with the unheated ends 

connecting to water-cooled anvils. The upper compression anvil was fixed to an upper platen by a 

universal joint and was leveled by four turnbuckles. The lower compression anvil rests on a selfaligning coupler that plays an important role in maintaining load-train stability and uniform 

compression to the specimen. Details of the compressive creep testing system are given elsewhere 

[10]. 

The generated creep strain was measured by a scissors-type mechanical extensometer, a schematic of 

which is shown in Fig. 4. The extensometer employs a capacitance sensor whose gap was 

conditioned, calibrated, and measured as a function of time. The lever-arm assembly was made from 

a continuous rod of silicon carbide. The probes monitored creep deformation of the specimen 

continuously in real time with a capacitive transducer attached to the cool end of the extension rod. 

Concurrently, a laser-based scanner was used to measure the movement of two fiducial flags attached 

to the ends of the extension rods for verification. The extensometer provided an accuracy and 

resolution of about 5 μm, which is equivalent to the sensitivity yielded by conventional strain gages. 

Strain measurements with the transducer and the laser system were in good agreement. 

12 
Fig. 4. Schematic of creep frame extensometer. 

Extensometry methods other than the capacitance contacting extensometers used in this study are 

often used for the measurement of refractory creep; however, the capacitance contacting extensometer 

circumvents problems that other techniques inherently possess. The deformation of refractories is 

frequently measured continuously during compressive creep testing using two (sometimes more) 

linear variable differential transducers (LVDTs) whose mutual displacements coincide with the 

continuous position of the specimen ends. The creep results generated from this technique are 

accurate only as long as the accumulated measured deformation coincides with the actual specimen 

heights measured before and after testing. Deformation and/or translation of the load train during the 

creep testing of the specimen, reaction of the specimen ends with the fixturing, and “bedding-down” 

of the specimen all have been shown to cause a lack of correlation between the measured deformation 

during testing and the change in pre- and post-test specimen height. If any of these events occurs then 

the measured contraction during testing is not solely due to creep. Consequently, caution must be 

exercised when interpreting deformation data generated by the LVDT technique because it will only 

be representative creep data if the experimenter verified and correlated pre- and post-test specimen 

height with the accumulated measured deformation. The advantage of the contacting extensometers 

used in the present study is that their accurate operation is independent of any rigid body motion or 

deformation of the test hardware, specimen (e.g., “end-crushing”), and push rods. 

13 
4.3 Specimen Preparation 

In the case of the silica and mullite refractories, 

cylindrical creep test specimens having 

dimensions shown in Fig. 5 were core-drilled 

from the supplied conventional bricks in the 

manner illustrated in Fig. 6. The primary axis of 

the machined cylinder-shaped specimens was 

oriented parallel to the pressing direction. 

Although not verified in the present study, the 

measured creep deformation was assumed to be 

isotropic, so specimens were not machined from 

the bricks in different orientations. After core 

drilling, the specimen ends were ground parallel 

to within 0.013 mm (0.001 in.). All specimens 

were dried in an oven at 300°C (570°F) for 4 h 

after they were core drilled and before they were 

creep tested. 

Using an appropriate aspect ratio of compression 

test specimens is an important consideration. A 

specimen aspect ratio of 2.6 or greater is 

recommended for compressive strength tests so 

that the friction effects between the specimen ends and loading ram are small compared to the axially 

applied compressive stresses [15]. However, in the present study (aspect ratio = 2.0), the applied axial 

compressive stresses during creep testing were already low (i.e., much less than the refractories’ 

compressive strengths). Consequently, the stresses and strains due to such friction were likely 

negligible, so adherence to the height/diameter ratio of 2.6 or greater was deemed unnecessary. 

3.00 

± 0.02

1.50 Dia. 

± 0.01

0.001

Dimensions 

in inches

Fig. 5. Dimensions of core-drilled test specimens. 

3"

Fig. 6. Schematic of how specimens were core-drilled from as-received bricks. 

In the case of the fusion-cast alumina, large blocks of approximate dimensions 0.60 H 0.50 H 0.30 m 

(24 H 20 H 12 in.) were prepared by both Corhart Refractories and Monofrax Inc., using their normal 

processing conditions. These blocks were sectioned in half with one half being shipped to ORNL and 

one half being retained by the preparing company. Cores of 19 mm (0.75 in.) diameter were drilled 

from whole regular cast blocks at three locations prior to being sectioned. Location one was in the 

center of the bottom of the block in the positive y-direction (designated y-core). Locations two and 

14 
Sectioned face

Fig. 7. Refractory block core-drill locations for physical and chemical analysis. 

three were in the side of the block at 76 mm (3 in.) and 178 mm (7 in.) from the base in the 

y-direction with the cores taken in the x-direction (designated x3-core and x7-core, respectively). Core 

locations and orientations are shown in Fig. 7. Individual cores were then sectioned into 25-mm 

(1-in.) pieces for density/porosity and X-ray fluorescence (XRF) analysis. The coordinate system 

used in Fig. 7, which is a nontraditional left-handed coordinate system, was chosen to comply with 

sample orientations utilized by Monofrax Inc. for its physical and chemical analysis techniques. 

The regular cast blocks were further sectioned as shown in Fig. 8 to obtain creep specimens that were 

representative of the interior bulk region of each cast block. The individual bulk creep specimens 

were 38 mm (1.5 in.) in diameter and 76 mm (3 in.) in length. A specially cast block of M-type 

fusion-cast alumina was also prepared at Monofrax Inc. by incorporating an increased cooling rate. 

The block possessed a columnar zone of ≈ 50 mm (2-in.) thickness (as opposed to a columnar zone of 

≈ 25 mm (1 in.) thickness on the regular cast blocks). A 0.30 H 0.20 H 0.05 m (12 H 8 H 2 in.) slab 

was cut from the exterior of the block and was shipped to ORNL. Monofrax retained the remainder of 

the block. The entire 50-mm cross section contained the characteristic columnar zone microstructure. 

Creep specimens were sectioned such that the orientation of the columnar structure was oriented in 

both a longitudinal direction [dimensions: 38 mm (1.5 in.) diameter, 50 mm (2 in.) length] and 

transverse direction [(38 mm diameter (1.5 in.), 70 mm (2.75 in.) length] with respect to the length of 

the cylinder. 

Cylindrical specimens [dimensions: 13 mm (0.5 in.) diameter, 38 mm (1.5 in.) length] for thermal 

expansion and elastic modulus determination were also taken from the sectioned regular cast 

Monofrax and Corhart blocks at locations between the creep specimens. Equivalent specimens were 

taken from the specially cast Monofrax block possessing the columnar structure. Similar to that of the 

creep specimens, the columnar structure in the specimens was oriented in both the longitudinal and 

transverse directions with respect to the length of the cylinders. 

The cylindrical specimens were encased in a fine-grained H-type (100% β-alumina) fusion-cast 

alumina shell (shown in Fig. 9) to inhibit soda evolution from the test specimens into the furnace 

atmosphere during testing. Attempts were made to ensure that a gap existed between the sample and 

H-shell during testing, but that was not always the case.

15 
 

Fig. 8. Schematic representation of the sectioning plan used to fabricate creep specimens from the bulk 

region of the refractory blocks. 

4.4 Creep Testing and Data 

Interpretation 

Tables 5, 6, and 7 describe the creep test matrices 

for the silica, mullite, and fusion-cast alumina 

refractories, respectively. A typical stresstemperature-time history is provided in Fig. 10. All 

specimens were preloaded in compression to 

approximately 0.04 MPa (≈6 psi) during furnace 

heat-up to keep all the load train components and 

the specimen in continuous contact. The specimens 

were then heated from room temperature up to the 

test temperature in 4 h. Each of the specimens was 

soaked at temperature for 25 h prior to the 

application of the first stress. 

Most creep analyses for long-term applications 

(i.e., where steady-state creep accumulation 

dominates the accumulation of strain, not the 

amount of primary creep) involve the determination 

of a steady-state or minimum creep rate and its 

examination as a function of applied stress and 

temperature. The steady-state or minimum 

compressive creep rate (dε/dtmin) can be related to 

the applied compressive stress and temperature by using an empirical Arrhenius power law or the 

familiar Norton-Bailey creep equation [16] 

FusionCast 

Alumina 

Specimen

Dendrytic 

Alumina 

Spacer

Monofrax M 

Push Rod

Monofrax M 

Push Rod

Sapphire 

Extensometer 

Knife

Sapphire 

Extensometer 

Knife

High Purity 

Alumina Disk

High Purity 

Alumina Hemisphere

H Fusion-Cast 

Alumina Disk

H Fusion-Cast 

Alumina Shell

Fig. 9. Creep specimen assembly H-shell 

arrangement required for the creep testing of 

the fusion-cast alumina. 

16 
 dε/dtmin = A σ n exp(-Q/RT), (1) 

where A is a constant, σ is the applied stress, n is the stress exponent, Q is the activation energy, R is 

the universal gas constant, and T is the absolute temperature. Multilinear regression may be 

performed to determine the constants A, n, and Q for each material. By performing the analysis in this 

manner, it is implied that the same dominant (or rate-controlling) creep mechanism is active at all 

temperatures and stresses. The validity of this assumption is assessed by the goodness-of-fit of this 

equation to the experimental data and the reasonableness of the obtained values for n and Q. 

Table 5. Creep test matrix for the silica refractories (stress, MPa)a

Refractory 1550°C (2820°F) 1600°C (2910°F) 1650°C (3000°F) 

Gen-Sil 0.2, 0.4, 0.6 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

SI96AU 0.2, 0.4, 0.6 0.2, 0.4, 0.6 b 

Stella GGS 0.2, 0.4, 0.6 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

Vega 0.2, 0.4, 0.6 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

Vega H 0.2, 0.4, 0.6 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

Unnamed development brand 0.2, 0.4, 0.6 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

a

0.2 MPa = 29.0 psi, 0.4 MPa = 58.0 psi, 0.6 MPa = 87 psi. 

b

Material exhibited excessive creep deformation at 1600°C, so its testing at 1650°C was omitted. 

Table 6. Creep test matrix for the mullite refractories (stress, MPa)a

Refractory 1300oC (2370oF) 1375oC (2500oF) 1450oC (2640oF) 

BP mullite 0.4, 0.6 0.4, 0.6 

DURITAL S75 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

Frimul F 0.4, 0.6 0.4 & 0.6 

GEM 0.2, 0.4, 0.6 0.2, 0.4, 0.6 

HF 17 0.4, 0.6 0.4, 0.6 

MU75AF 0.4, 0.6 0.4, 0.6 

SL75AD 0.4, 0.6 0.4, 0.6 

UFALA UCR 0.4, 0.6 0.4, 0.6 

ZED FMC 0.4, 0.6 0.4, 0.6 

ZED FM 0.4, 0.6 0.4, 0.6 

a

0.2 MPa = 29.0 psi, 0.4 MPa = 58.0 psi, 0.6 MPa = 87 psi. 

Table 7. Creep-test matrix for the fusion-cast alumina refractories (stress, MPa)a

Refractory 1450°C (2640°F) 1550°C (2820°F) 1650°C (3000°F) 

Corhart Jargal M cast block, 

regular bulk 

0.6, 0.8, 1.0 0.6, 0.8, 1.0 0.6, 0.8, 1.0 

Monofrax M cast block, 

regular bulk 

0.6, 0.8, 1.0 0.6, 0.8, 1.0 0.6, 0.8, 1.0 

Monofrax M cast block, 

special longitudinal columnar 

0.6 1.0 

Monofrax M cast block, 

special transverse columnar 

0.6 1.0 

a

0.6 MPa = 87.0 psi, 0.8 MPa = 116 psi, 1.0 MPa = 145 psi. 

17 
0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0

500

1000

1500

2000

0 50 100 150 200 250 300

Compressive Stress (MPa)

Temperature (°C)

Time (h)

100

75

50

25

0

3500

3000

2500

2000

1500

1000

500

0

Temperature (°F)

Stress

Temperature

4h to

temp

Note: shown temperature is

1600°C, but same history was

used for 1550 and 1650°C tests

Fig. 10. Example of temperature and stress history used in the creep testing. 

4.5 Microstructural Characterization 

The microstructures of the various refractories were examined using reflected light (RL) imaging, 

cathodoluminescent (CL) imaging, and SEM. In the case of RL and CL, disks for examination were 

sectioned from both as-received and crept specimens Disks were cut from the top of each as-received 

specimen as shown in Fig. 11(a); disks were sectioned from crept specimens in their centers as shown 

in Fig. 11(b). The disk sections were cleaned and dried and were then impregnated with a lowviscosity resin to fill the pores. The impregnated specimens were mounted on Co-Cast mounting 

media. Cured sections were ground with diamond wheels down to 600 grit and then were finally 

polished with diamond paste and lapping oil. RL and CL microscopy were utilized for the 

characterization. A cold-cathode CL microscopy system (Model Mk4, Cambridge Imaging 

0.2"

Metallographically 

Polished 

Surface

Core Drilled 

Specimen

Used for 

Creep 

Test

"As-Received" 

Microstructure

0.2"

Metallographically 

Polished 

Surface

Crept 

Specimen

Sectioning 

of Specimen

 (a) (b) 

Fig. 11. Schematic describing how disk sections were machined from core-drilled specimens for 

(a) as-received and (b) post-test RL and CL imaging. 

18 
Labophot-Pol with a Nikon UFX-DX photomicrographic system) was used for these microstructural 

analyses. The CL microscopy characterization technique has been described in detail by Karakus and 

Moore [17] and is briefly described here. The CL imaging system utilizes an energetic electron beam 

that is produced from a cold cathode ray tube. The electron beam is trained on the surface of 

uncovered polished or unpolished specimens in a low vacuum environment. As a result of the 

electron beam-solid specimen interaction, minerals or phases in the specimen produce characteristic 

colored light known as “cathodoluminescent” emission. The CL technique is unique and provides 

microstructural information that often cannot be obtained by any other technique. The CL microscopy 

system can provide an immediate assessment of the phases in refractory materials through the 

identification of characteristic CL color and crystal habits that minerals exhibit. The CL microscopy 

system can also be used in conjunction with RL and SEM for elemental analysis of individual phases. 

Certain minerals produce a characteristic CL color. For example, corundum (α-Al2O3) is characterized by a characteristic bright red CL color while spinel (MgAl2O4) produces a characteristic green 

color. These CL emissions are due to “activator” elements present in trace amounts in these mineral 

structures; Cr3+ in corundum produces a characteristic CL emission at approximately 694 nm, and 

tetrahedrally coordinated Mn2+ in spinel produces a CL emission at approximately 520 nm. Activator 

elements related to CL emission centers in these minerals often require low voltages (≈ 8 kV) for 

excitation. 

Important for the materials analyzed in this study are the CL behaviors of mullite, alumina, and silica 

grains along with that of the glassy matrix phase. Regular mullite grains do not produce bright CL but 

rather exhibit a dull response. Contrary to this, fused mullite grains produce intense dark blue or 

violet CL. Secondary mullite bonds between primary fused mullite grains can produce intense violet 

reddish CL, distinguishing them from primary mullite. Fused alumina grains exhibit very bright violet 

blue CL, while tabular alumina grains appear as an intense violet. Andalusite grains give off dark 

reddish CL and zirconia, a blue-white CL. 

4.6 Supplemental Analysis 

The bulk densities of various core-drilled specimens were measured prior to creep testing. Masses 

were measured on a balance that had a resolution of ± 0.0005 g. The masses were then divided by the 

respective specimens’ volumes (diameters and lengths measured with calipers having a resolution of 

± 0.005 mm) to calculate density. 

Specimen

Specimen

Specimen

Styrofoam

Standard Measurement

Single-side Measurement

Sensor

(side view)

Sensor

(side view)

 

Kapton Sensor

(for low temps)

Fig. 12. Schematic of transient plane source 

direct thermal conductivity method 

experimental set-up and picture of Kapton 

sensor. 

Room-temperature thermal conductivity of the “asreceived” bricks was measured by a transient plane 

source (TPS) technique in the “hot-disk” arrangement [18]. In this arrangement, a Kapton TPS element was used as both a heat source and a temperature sensor to measure the time required to conduct a 

known amount of heat across a prescribed distance 

of sample. Measurements were taken in the “standard method” by sandwiching the Kapton element 

between two identical cylindrical creep specimens 

and passing 1 W of power through the sample for 

20 s. Five measurements were taken on each 

specimen type with 15 min. or longer between 

measurements. The specimen configuration along 

with the sensor itself is shown in Fig. 12. 

19 
The phase content of the materials in their as-received state was characterized and was compared with 

that after creep testing at 1450°C. Specimen preparation involved grinding the refractory to 

approximately 325 mesh in an agate mortar and pestle. XRD was performed with all ten mullite 

materials in a Scintag 2000 diffractometer. The samples were scanned at a rate of 1°/min with Cu-Ka

radiation (l = 1.54059 Å), and a step size of 0.02° over a scan range of 10 to 120° two-theta. 

In the case of the fusion-cast alumina, thermal diffusivity was also measured from room temperature 

up to 1400°C (2552°F) by the laser flash method [19]. In the laser flash method, a cylinder of the bulk 

refractory [13-mm (0.5-in.) diam., 32 mm (1.25-in.) length] was sectioned into small disks [2 mm 

(0.08 in.) thick], and the disks were subjected to a high-intensity, short-duration thermal pulse 

through their thicknesses. The temperature resulting at the rear face of each specimen, due to the 

pulse, was recorded and was used to calculate the time required to raise the rear face temperature a 

certain percentage of its maximum value. Thermal conductivity values were then calculated: 

λ = ρ α Cp, (2) 

where λ is the thermal conductivity, ρ is the density, and Cp is the heat capacity. The latter parameter 

was measured by calorimetery at ORNL and Monofrax, Inc., in New York. 

To examine dimensional stability due entirely to only exposure at temperature, specimens from each 

refractory were subjected to aging at times and temperatures that were equivalent to those applied to 

the creep specimens. Two methods were used for measuring these aging characteristics. The first 

involved heating the specimen at temperature for a prescribed number of hours and then cooling the 

furnace to room temperature. The specimen was then weighed, and its dimensions were measured 

before it was put back in the furnace for subsequent reheating. The mass, dimensions (volume), and 

density of the specimens were thus determined as a function of cumulative time and temperature. 

The second method involved measuring the real-time specimen displacement at temperature using a 

dual-rod LVDT system (Fig. 13). The specimen was supported by a high-purity alumina disk resting 

on a fixed SiC push rod while the test assembly was supported by the body of the furnace. The test 

assembly for monitoring the linear dimensional changes consisted of two sapphire rods in contact 

with individual LVDTs. One sapphire rod and LVDT pair was used to monitor changes in the test 

Fig. 13. Schematic of dual-rod LVDT system used to measure thermal expansion. 

20 
sample (designated LVDT-1) while the other rod and LVDT were placed in contact with the highpurity alumina disk supporting the test specimen (designated LVDT-2). The difference between these 

readings was used to calculate the dimensional changes of the specimen [13]. 

 

The ASTM lid test was used to evaluate the corrosion of the various refractories [9]. A refractory 

specimen (2.5 H 2.5 H 1 cm) sectioned from the as-received bricks was placed over a small alumina 

crucible containing 12.5 g of sodium carbonate. The crucible and refractory specimen "lid" were then 

placed in a secondary alumina container and isothermally exposed at 1400°C for 24 h. After 

exposure, samples were removed and were examined for recession. 

21 

5. Results and Discussion 

5.1 Silica Refractories 

A major finding from the compressive creep tests of the silica refractories was that concurrently 

active mechanisms, other than creep, resulted in positive dimensional changes. Specifically, as shown 

in Fig. 14, some specimens exhibited axial expansion during creep testing even though an axial 

compressive load was applied. This peculiarity is a consequence of the bulk expansion that these 

refractories exhibited. This effect limited the ability to identify or interpret the lesser-active creep 

mechanism in these silica refractories. Consequently, the compressive creep rates of all six brands 

were not able to be represented as a function of temperature and compressive stress using the 

conventionally used NBA creep equation (i.e., the Arrhenius power-law creep model). 

The change in dimensions of the compressively crept specimens indicated that their size (both 

diameter and length) had actually increased as a consequence of the employed creep test conditions 

(see Table 8). This expansion phenomenon greatly complicates creep analysis in which Eq. (1) is used 

and likely invalidates its use. All six brands showed this expansion effect. The increases in diameter 

and length of the creep specimens were between 0.35 and 1.0%. 

Table 8. Dimensional changes of crept specimens. A negative value constitutes contraction and 

a positive value represents expansion. 

Length change 

(%) Gen-Sil SI96AU 

Stella 

GGS Vega Vega H 

H-W Dev. 

Brand 

1550°C 

Diameter 1.0 0.65 0.45 !0.08 0.49 0.38 

Length 0.55 0.55 0.36 0.94 0.42 0.41 

1600°C 

Diameter 1.0 9.12 0.42 0.24 0.46 0.38 

Length 0.52 !24.0 0.70 0.35 0.47 0.49 

1650°C 

Diameter 0.50 0.71 0.50 0.53 0.50 

Length 0.47 0.71 0.59 0.46 

Consistent with bulk changes exhibited by the creep specimens, the unstressed or “aged” silica 

refractory specimens from all six materials also exhibited loss of mass, dimensional increases, and 

density decreases between 1550 and 1650°C (see Table 9). The mass losses of the six brands at 1550 

and 1600°C were equivalent; however, some silica brands lost more mass than others at 1650°C: 

Harbison-Walker’s developmental brand lost approximately 0.1% of its mass; SI96AU lost ≈ 0.2%, 

Vega H lost ≈ 0.3%, Gen-Sil lost ≈ 0.5%, Vega lost ≈ 1.0%; and Stella GGS lost ≈ 1.5%. Most of the 

mass loss occurred in less than 25 h for all six brands. 

The dimensional increases of the silica specimens were a function of temperature, as illustrated in 

Fig. 15. Additionally, some brands expanded more than others. Each of the specimens expanded 

isotropically in each of three directions (within statistical confidence). The ranges of expansion were 

≈ 1!2.5% at 1550°C, ≈ 2.25!3.75% at 1600°C, and ≈ 3.5!6% at 1650°C. Harbison-Walker’s 

developmental brand expanded the least amount at all three temperatures, followed by Gen-Sil. The 

ascending order of expansion for the other four brands varied with temperature. Stella GGS expanded 

23 
Table 9. Density changes of crept specimens. A negative value constitutes a decrease in density 

and a positive value represents an increase. 

Gen-Sil SI96AU 

Stella 

GGS Vega Vega H 

H-W Dev. 

Brand 

After 1550°C creep testing 

Density (g/cc) 1.75 1.80 1.81 1.73 1.83 1.90 

Change (%) !1.8 !2.6 !1.9 !3.2 !2.8 !1.6 

After 1600°C creep testing 

Density (g/cc) 1.73 1.79 1.76 1.82 1.90 

Change (%) !2.7 !2.0 !1.8 !1.9 !2.2 

After 1650°C creep testing 

Density (g/cc) 1.72 1.79 1.74 1.80 1.88 

Change (%) !3.0 !2.2 !3.1 !3.9 !3.0 

the most at 1550°C (≈ 2.5%), SI96AU expanded the most at 1600°C (≈ 3.75%), and Vega expanded 

the most at 1650°C (≈ 6%). Dimensional increases in these aged specimens were greater than axial 

expansion of the crept silica specimens at the same temperatures, indicating that there may have been 

a specimen-size effect. The change in the as-received assemblage of tridymite, amorphous silicate, 

Ca-silicate, and cristobalite to the crept microstructure consisting entirely of cristobalite likely caused 

the observed volume expansion. The same phenomenon appeared to have occurred in the aged 

specimens. Temperature is the activator of this phenomenon; the application of stress is not a 

prerequisite for it. 

Based upon the data given in Fig. 15 it is reasonable to assume that the deformation rate measured 

during the application of a compressive stress was composed of two parts: (1) specimen contraction 

arising from creep-related processes (a function of time, stress, and temperature) and (2) specimen 

elongation arising from the change in the phase assemblage (a function of time and temperature only). 

One can therefore estimate the contribution arising from creep alone by subtracting the rate of 

elongation occurring at zero stress from the rates measured under a fixed stress. To this end, the 

curves in Fig. 15 were subsequently used to calculate the linear expansion rates occurring at 250 h for 

the 1600 and 1650°C aging tests. The calculated rates were then subtracted from the rates measured 

under an applied stress at comparable times. 

For the 1600 and 1650°C aging tests the stress of 0.6 MPa was used because the total time at 

temperature was comparable to that used to calculate the zero-stress expansion rate: 

• Gen-Sil: dε/dtmin (1/s) = 3.6 H 1020 σ exp[-25 (kcal/mole)/RT] 

• SI96AU: dε/dtmin (1/s) = 4.2 H 1044

σ exp[-44 (kcal/mole)/RT] 

• Stella GGS: dε/dtmin (1/s) = 5.5 H 1014

σ exp[-20 (kcal/mole)/RT] 

• Vega: dε/dtmin (1/s) = 1.9 H 1012

σ exp[-18 (kcal/mole)/RT] 

• Vega H: dε/dtmin (1/s) = 1.4 H 1013 σ exp[-18 (kcal/mole)/RT] 

XRD results showed that all six as-received silica refractories consisted of cristobalite (PDF# 

39-1425) and tridymite (PDF# 18-1170). Even though a calcium silicate was present in the grain 

boundaries (as detected by SEM/EDS), its concentration was sufficiently small that XRD did not 

24 
-0.5

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

0.5

-1.0

-0.5

0.0

0.5

1.0

1.5

2.0

0 50 100 150 200 250 300

Creep Strain (%)

Stress (MPa)

Time (h)

Strain

Stress

Vega

1600°C

4.5 x 10-9 s-1

3.5 x 10-9 s-1 -6.3 x 10-10 s-1

-0.50

-0.40

-0.30

-0.20

-0.10

0.00

0.10

0.20

0.30

0.40

0.50

-1.0

-0.5

0.0

0.5

1.0

1.5

2.0

0 50 100 150 200 250 300

Creep Strain (%)

Stress (MPa)

Time (h)

Strain

Stress

Vega

1550°C

9.9 x 10-11 s-1 -6.2 x 10-11 s-1 3.6 x 10-11 s-1

-0.5

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

0.5

-1.0

-0.5

0.0

0.5

1.0

1.5

2.0

0 50 100 150 200 250 300

Creep Strain (%)

Stress (MPa)

Time (h)

Strain

Stress

Vega

1650°C

1.9 x 10-9 s-1

6.1 x 10-9 s-1

3.8 x 10-9 s-1

Fig. 14. Creep behavior of Vega as a function of temperature. 

25 
identify it. It was qualitatively 

determined that Gen-Sil had the 

highest tridymite content and that 

SI96AU had the highest 

cristobalite content. Very small 

amounts of quartz were detected 

only in SI96AU and Vega. 

0

2

4

6

8

10

0 50 100 150 200 250 300

Vega

Vega H

Gen-Sil

H-W Dev. Brand

Stella GGS

SI96AU

Change in Volume (%)

Time (h)

1550°C

(2820°F)

0

2

4

6

8

10

0 50 100 150 200 250 300

Vega

Vega H

Gen-Sil

H-W Dev. Brand

Stella GGS

SI96AU

Change in Volume (%)

Time (h)

1600°C

(2910°F)

Based on the density differences in 

the silica minerals (quartz = 

2.64 g/cc, tridymite = 2.26 g/cc, 

cristobalite = 2.32 g/cc), it takes 

the conversion of approximately 

6.5% of residual quartz to 

tridymite or cristobalite to cause a 

bulk expansion of 1%. Because no 

quartz was detected in four of the 

conventional silicas, and because 

its concentration was minute in 

SI96AU and Vega, the phasechange phenomenon could not 

have caused the volume expansion 

that these materials exhibited as a 

consequence of the creep testing. 

Subsequent creep testing changed 

the phase content in the six silica 

refractories in that all the tridymite 

in the silica refractories was 

converted to cristobalite. The 

stable silica phase above 1470°C 

(2680°F) is cristobalite, so the 

XRD results are not surprising. 

The microstructures of the six 

silica refractories changed as a 

consequence of the applied test 

conditions. The changes in asreceived to crept microstructures in 

the six materials had commonality 

and are shown schematically in 

Fig. 16. The as-received silica 

microstructures of the six materials 

typically consisted of cristobalite 

aggregates or “islands” in a matrix 

or “sea” or of prismatic tridymite 

(approximately 10 to 20 μm in size 

near the cristobalite islands, 20 to 

80 μm farther away from them), 

calcium silicate crystals, and an amorphous silicate. The structure of the cristobalite in the islands 

often resembled fish scales. The individual cristobalite grains comprising the cristobalite islands had a 

0

2

4

6

8

10

0 50 100 150 200 250 300

Vega

Vega H

Gen-Sil

H-W Dev. Brand

Stella GGS

SI96AU

Change in Volume (%)

Time (h)

1650°C

(3000°F)

Fig. 15. Change in volume of “aged” specimens as a function of 

time and temperature. 

26 
rather uniform size distribution, 5 to 10 μm 

across; the size of the cristobalite islands ranged 

from approximately 40 μm to more than 1 mm. 

During creep testing at 1550, 1600, and 1650°C, 

the tridymite converted to cristobalite, and a new 

amorphous silicate (formed between the new and 

growing cristobalite islands). Additionally, many 

of the crept cristobalite microstructures 

contained pores that resembled tunnels. They 

appeared to be avenues for vaporization, but their 

structure was not examined, so their cause is not 

known. 

RL and CL imaging supplemented one another 

well to explore these microstructural changes. 

Examples of the RL images of the as-received 

and crept microstructures are shown for the Vega 

refractory in Fig. 17. Pores, tridymite, and Casilicate were consistently observed in the asreceived microstructures of each refractory. 

Changes in the microstructure arising from creep 

testing included the formation of larger pores, 

larger cristobalite islands, the new amorphous 

silicate, and the creation of gaseous pores. The 

bulk expansion of the crept specimens may be 

explained by the destruction of original tridymite 

bonds and the enlargement of pores and of the 

cristobalite islands. 

CL imaging (Fig. 18) provided an interesting and 

different perspective into how the microstructures changed as a consequence of the creep 

testing. The CL imaging revealed the tridymite in 

the as-received state better than the RL imaging, 

and it provided new insights into how the 

cristobalite islands changed or grew. 

Specifically, the cores of most of the cristobalite 

aggregates in these specimens exhibited a light 

blue color while their outer zones or rims were dark blue. The light blue core represents the original 

fish-scale cristobalite island. The core or center of the islands after creep testing was relatively free of 

impurity phases, resulting in the light blue CL color. During the creep testing, the tridymite converted 

to cristobalite or added to the size of existing cristobalite islands. Additionally, the original calcium 

and amorphous silicates (containing transition metals and alkalis) formed a liquid phase during creep 

testing that either diffused into the enlarging cristobalite islands or was present during their growth. 

This outer zone or rim was characterized by dark blue CL. The Ca-silicate phase luminesced bluishwhite and locally bright yellow, with some of its crystals being as small as 15 μm, in the as-received 

microstructures. However, during creep testing, iron impurities in the original amorphous silicate 

combined with the Ca-silicate phase to form a new amorphous silicate that because of the presence of 

the iron, did not produce CL color. 

Boundary Pore Cristobalite

Ca-silicate crystals + Tridymite

amorphous silicate

Typical "As-received" Silica Microstructure

Boundary Pore Cristobalite

New amorphous silicate

Typical Crept Silica Microstructure

"Gas" or "Tunnel" 

Pore

Fig. 16. Schematic of the change in microstructure 

from the as-received material state to that after 

creep or high-temperature exposure. 

27 
In the case of the corrosion 

testing, a portion of the silica 

refractory lid was dissolved by 

the sodium carbonate vapor, 

forming a liquid phase at 

temperature. After the test, the 

maximum amount of recession 

was measured and summarized 

(Fig. 19). In general, there was 

only a slight variation in the 

amount of attack among the six 

silicas, and it may be concluded 

that their corrosion resistances 

were statistically equivalent. 

The amount of recession 

showed an almost linear 

relationship with temperature, 

as shown in Fig. 20. Because a 

vapor attack is expected, this 

effect seems reasonable. 

Varying the amount of sodium 

carbonate by 50% more 

(18.75 g) and by 50% less 

(6.25 g) from the standard 

12.5 g had very little effect. 

Apparently, the sodium 

carbonate is not fully consumed 

during the test, and the lowest 

amount does not limit the extent 

of attack. 

Fig. 17. Vega’s microstructural changes as a function of 

temperature as revealed by reflected-light microscopy. 

5.2 Mullite Refractories 

The crystalline phases present in each of the ten mullite refractories varied from material to material, 

but some trends or similarities existed among the brands (see Table 10). The ten brands could be 

broadly categorized as either fused-grained or non-fused-grained (see Table 4). The non-fusedgrained materials all contained mullite grains derived from andalusite or kyanite while the fusedgrained mullite was made by fusion (melting) of alumina and silica. All the refractories were 

composed of variously sized mullite grains floating in a matrix of secondary mullite with various 

levels of glass and porosity present in the matrix. The DURITAL S75, HF 17, and ZED FMC brands 

all contained very little glassy phase. All the others possessed significant quantities of glassy phase. 

The HF 17, UFALA UCR, and ZED FMC brands were all found to possess high levels of matrix 

porosity when examined by optical microscopy and SEM; the rest showed little or no matrix porosity. 

Room-temperature thermal conductivity testing resulted in the average values shown in Fig. 21. 

Frimul F showed the highest thermal conductivity at just under 3.5 W/mK. It was followed by Durital 

S75, GEM, MU75AF, and ZED FMC, which all had a thermal conductivity of approximately 

3.0 W/mK. BP mullite and HF 17 both had similar thermal conductivity values of 2.7 to 2.8 W/mK, 

followed by ZED FM with a value of just over 2.5 W/mK. The two brands with the lowest thermal 

conductivity values were SL75AF and UFALA UCR, which had values of approximately 2.5 W/mK. 

28 
These values were all higher 

than those given in the 

manufacturers’ literature, which 

ranged from 2.0 to 2.4 W/mK. 

The results from creep testing at 

an applied stress of 0.5 MPa are 

shown in Fig. 22. The amount 

of compressive creep for th

various refractories fell into 

three levels. BP mullite, HF 17, 

and Frimul F all showed low 

amounts of creep at 

temperatures between 1300 and 

1450

e 

 at 

test 

o

C (2370 to 2640oF). 

Although the creep rate for the 

MU75AF at 1450°C was also 

quite low, it exhibited 

anomalous behavior in that the 

creep rate increased as the 

temperature was lowered to 

1375°C. Slightly greater 

amounts of creep were seen in 

the GEM, DURITAL S75, and 

ZED FM brands at temperatures 

between 1300 and 1450o

C 

(2370 to 2640o

F). Creep rates 

for these refractories were found 

to be on the order of 10!10 s!1

the lower test temperature and 

10-9 s!1

 at the higher 

temperature. The SL75AD, 

UFALA UCR, and ZED FMC 

brands all showed significant 

amounts of creep at 

temperatures between 1375 and 

1450o

C (2500 to 2640oF). Their creep rates were on the order of 10!9 s!1 at the lower test temperature 

and 10!8

 s!1 at the higher test temperature. Because only one specimen per condition was tested, the 

authors were unable to statistically conclude that any one of the refractories in each group had 

superior creep resistance to the others. If differences exist within a grouping, they are believed to be 

insignificant. 

Fig. 18. Vega’s microstructural changes as a function of 

temperature as revealed by cathodoluminescence imaging. 

The activation energy in the Norton-Bailey creep equation was found for each brand by equating the 

respective slope in Fig. 22 to the quantity –Q/R. Although the values of the stress exponents could 

have also been determined from the experimental data, that procedure was not implemented because 

the uncertainties were high in the resulting estimates. Specifically, because n is typically very close 1, 

an increase in stress from 0.2 to 0.6 would only increase the creep rate by a factor of 3. Replicate 

creep tests of other structural ceramics have shown that the minimum creep rate value can vary by an 

order of magnitude or more. Consequently, the stress range used in the present investigation was 

probably not sufficient to generate a significant increase in creep rate with increasing stress. 

Therefore, the value of A in Eq. (1) was calculated by assuming n = 1. The effect of specimen 

29 
elongation or contraction in the absence of a stress was ignored in the calculations. The resulting 

expressions were then derived for the creep rate of the mullite refractories: 

• BP mullite dε/dtmin (1/s) = 4.2 H 10-1

σ exp[!73 (kcal/mole)/RT] 

• DURITAL S75 dε/dtmin (1/s) = 1.5 H 105 σ exp[!110 (kcal/mole)/RT] 

• Frimul F dε/dtmin (1/s) = 3.3 H 107 σ exp[!130 (kcal/mole)/RT] 

• GEM dε/dtmin (1/s) = 7.4 H 101

σ exp[!80 (kcal/mole)/RT] 

• HF 17 dε/dtmin (1/s) = 6.6 H 1016 σ exp[!200 (kcal/mole)/RT] 

• MU75AF [Could not be determined] 

• SL75AD dε/dtmin (1/s) = 6.0 H 10-2

σ exp[!50 (kcal/mole)/RT] 

• SL75AD dε/dtmin (1/s) = 2.4 H 104

σ exp[!90 (kcal/mole)/RT] 

• ZED FMC dε/dtmin (1/s) = 1.1 H 1012

σ exp[!150 (kcal/mole)/RT] 

• ZED FM dε/dtmin (1/s) = 5.1 H 108 σ exp[!130 (kcal/mole)/RT] 

0

1

2

3

4

5

Gen-Sil #1Gen-Sil #2

Stella GGS

Vega H

H-W Dev. Brand

Vega

SI96AU

Maximum Recession (mm)

Fig. 19. Amount of refractory recession when exposed to sodium carbonate at 

1400°C for 24 h. 

30 
1

2

3

4

5

6

7

Recession (mm)

1200 1250 1300 1350 1400 1450 1500

Temperature (°C)

18.75g

6.25 g

12.5 g

Fig. 20. Amount of recession as a function of temperature and 

sodium carbonate addition for Gen-Sil silica. 

Table 10. Summary of microstructural observation for the mullite refractories before and after creep testing 

Brand name Starting microstructure Microstructure after creep 

BP mullite RL&CL: large and small fused mullite grains dense matrix of 

secondary mullite and glass 

XRD: mullite and alumina 

SEM/EDS: mullite is stoichiometric while the glassy phase contains 

K2O, Na2O, and CaO 

No change 

DURITAL S75 RL&CL: non-fused mullite grains derived from and alusite along 

with tabular and fused alumina grains floating in a matrix of 

secondary mullite, little detectable glass 

XRD: mullite and alumina 

SEM/EDS: mullite is stoichiometric 

No change 

Frimul F RL&CL: fused mullite aggregates floating in a very glassy matrix of 

secondary mullite, small mullite grains present in the glassy matrix 

XRD: mullite and traces of alumina 

SEM/EDS: mullite is stoichiometric while glassy phase contains 

K2O, CaO, and TiO2

No change 

GEM RL&CL: large and medium sized fused mullite grains in a 

secondary mullite matrix containing amorphous phase 

XRD: crystalline mullite; 

SEM/EDS: large mullite crystals and smaller (3–5 μm) secondary 

mullite crystals in an amorphous matrix, glassy matrix has a similar 

composition to the mullite crystals with small amounts of Na2O, 

CaO, and TiO2

No change 

31 
Table 10. continued 

Brand name Starting microstructure Microstructure after creep 

HF 17 RL&CL: large fused mullite grains in a highly porous matrix of fine 

fused mullite grains plus very tiny alumina crystals 

XRD: mullite and alumina 

SEM/EDS: secondary mullite crystals present as observed as 

agglomerates devoid of glassy bonding phase 

Microstructure appeared 

denser 

MU75AF RL&CL: large fused mullite grains or aggregates with round 

morphology in a fine “granular” mullite and glassy matrix 

XRD: mullite plus trace alumina 

SEM/EDS: the glassy phase contains K2O and Na2O along with 

trace amounts of TiO2 and CaO 

No change 

SL75AD RL&CL: andalusite grains partially converted to mullite along with 

alumina grains and fine fused grain mullite; secondary mullite 

matrix formed around and between partially mullitized andalusite 

grains, large fused alumina grains, and fine fused mullite grains; 

andalusite grains contain free quartz that is converted to 

cristobalite during firing 

XRD: andalusite, mullite, alumina, and cristobalite 

SEM/EDS: glassy matrix contains Na2O, MgO, CaO, K2O, and TiO2

Andalusite grains are 

converted to mullite 

aggregates and the amount 

of glassy phase has 

increased 

UFALA UCR RL&CL: finely crystalline mullite aggregates formed through the 

calcination of andalucite or kyanite in a fine crystalline mullite 

matrix, which contains a large number of micro-fissures and fine 

pores along with a large amount of glassy phase 

XRD: mullite 

SEM/EDS: skeletonized mullite containing elliptical and spherical 

pores characteristic of mullite derived from andalucite; mullite 

present in the matrix has a prismatic morphology and a variable 

crystal size with most crystals being small and needle-like but 

some crystals having a large rectangular shape due to crystal 

growth, the prismatic mullite crystals contain TiO2 impurities and 

the glassy phase contains K2O, CaO, and TiO2

Reduction in porosity, 

elimination of the microfissures, and appearance of 

cristobalite 

ZED FMC RL&CL: large, medium and fine fused mullite grains; the matrix 

between the large grains contains poorly-sintered fine fused mullite 

particles as well as large amounts of micro-fissures along with 

small amounts of alumina and zirconia 

XRD: mullite plus trace alumina 

SEM/EDS: very small amounts of silica glass formation (with CaO 

imourities) along with secondary 1- to 2-μm mullite crystal 

No change 

ZED FM RL&CL: primarily fused mullite grains similar to the ZED FMC and 

trace amounts of alumina and zirconia particles; good particle to 

particle bonding 

XRD: mullite plus trace alumina 

SEM/EDS: strong mullite and glassy bonding, glass contains CaO 

No change 

32 
Fig. 21. Thermal conductivity of mullite refractories as measured by the “hot-disk” method. 

1.E-11

1.E-10

1.E-09

1.E-08

1.E-07

5.7E-04 5.8E-04 5.9E-04 6.0E-04 6.1E-04 6.2E-04 6.3E-04 6.4E-04

1/T (1/K)

BP Mullite

Frimul F

MU75AF

HF 17

GEM

ZED FM

DURITAL S75

UFALA UCR

ZED FMC

SL75AD

Fig. 22. Creep rate measured at 0.6 MPa as a function of reciprocal temperature. 

33 
Activation energies for the various refractory brands ranged from 50 to 200 kcal/mol. Although no 

direct correlation between the magnitude of the activation energy and the amount of creep exhibited 

by a particular brand could be drawn from the data, it is known that materials possessing a high 

activation energy have greater sensitivity to temperature than those possessing a lower activation 

energy. The values for three of the calculated activation energies correspond to the activation energy 

of vitreous silica, which falls in the range of 120 to 170 kcal/mol [20, 21]. This would indicate that 

the glassy phases in these brands affect the overall creep behavior of the material. 

As indicated in Table 10, all of the refractories except the SL75AD, UFALA UCR, and HF 17 brands 

exhibited no major changes after creep testing. The microstructure of the HF 17 appeared to be 

slightly denser. In the case of the SL75AD brand, the andalusite grains were converted to mullite 

aggregates, and the amount of glassy phase increased in the crept specimens. The crept samples of the 

UFALA UCR refractory exhibited a reduction in porosity, the elimination of the microfissures, and 

the appearance of cristobalite. Additional details are provided in Ref. [11]. 

In subsequent aging tests, eight of the ten refractory specimens exhibited similar behavior, with the 

ZED FMC and ZED FM showing extraneous amounts of change (≈8.5%) in mass and volume due to 

thermal exposure. The other eight brands showed less than 0.2% change in mass and less than 0.6% 

change in volume. The changes in volume resulted in less than a 0.5% change in overall density for 

all samples, including the ZED FM and FMC brands. The two Minteq brands along with the Frimul F 

brand all exhibit activation energies associated with vitreous silica. They also exhibit the largest 

changes in mass, which may indicate that the silica plays some role in the temperature-induced mass 

changes. 

Results from the ASTM lid test showed that only minimal amounts of recession occurred for the 

mullite refractories under the testing conditions set forth by the standard. Also, the mass change data 

were found to be ambiguous; some of the specimens had deposits on their surfaces and actually 

gained weight while others lost a small amount of mass. Therefore, those data are not reported here. 

Figure 23 shows the small amounts of refractory recession occurring on the various refractory 

samples. It can be seen that the nonfused grained refractories (Durital S75, SL75AD, and UFALA 

UCR) show a minimal amount of recession and some discoloring while the fused grained materials 

show practically no recession at all. 

5.3 Fusion-Cast Alumina Refractories 

The density profiles (Fig. 24) for the Corhart and Monofrax bulk blocks based on the x and y cores 

(see Fig. 7 for sectioning details) showed a mean bulk density on the order of 3.4 to 3.5 g/cm3

 (212.3 

to 218.5 lb/ft3

) at their surface. This density is fairly constant for the entire x3 core and for the y core 

up to about 5 in. into the block for both cases. At the y core 5-in. mark the densities of both blocks 

begin to fall off due to porosity encountered in the header zones of the blocks. The x7 cores for both 

blocks also showed similar behavior with a fairly constant density in both cases up to around 3.5 in. 

and then a decline in density to around 3.3 g/cm3

 (206.0 lb/ft3) for the Corhart block and to around 3.1 

g/cm3

 (193.5 lb/ft3) for the Monofrax block. Similar trends were seen in the porosity. If the columnar 

regions of the blocks are considered to exist in the outer 1 in. of the samples with the remainder of the 

block exhibiting bulk structure, then average columnar and bulk densities and porosities can be 

determined for each brand based on the previously described measurements. This determination 

34 
DSF FRIMUL F DSF FRIMUL F

SL75AD

GEM

UFALA UCR UFALA UCR

DSF HF-17 DSF HF-17

MU75AF Zed Mullite F Zed Mullite F Zed Mullite FC Zed Mullite FC

BP Mullite BP Mullite Durital S75 Durital S75

ASTM Lid Test ASTM Lid Test

12.5g Na 12.5g Na2CO3 for 24h at 1400°C for 24h at 1400°C

2 in.

Fig. 23. Refractory recession when exposed to sodium carbonate at 

1400o

C for 24 h (according to ASTM C98). 

would result in average columnar density values of 3.45 ± 0.05 g/cm3

 and 3.38 ± 0.02 g/cm3 and 

average bulk density values of 3.45 ± 0.08 g/cm3

 and 3.37 ± 0.14 g/cm3 for the Corhart and Monofrax 

brands, respectively. Average porosity values were found to be 1.38 ± 0.15% and 1.44 ± 0.19% for 

the Corhart columnar and bulk regions, respectively, and 2.64 ± 0.55% and 1.78 ± 0.92% for the 

Monofrax columnar and bulk regions, respectively. These values are close to those published by the 

manufacturers, although the density and porosity values obtained experimentally are a little lower for 

both brands (Corhart bulk density = 3.54 g/cm3

 and porosity = 2.0%, Monofrax bulk density = 3.40 

g/cm3

 and porosity = 1.9%). 

Density measurements were made on ten bulk Monofrax creep samples, ten bulk Corhart creep 

samples, two longitudinal columnar Monofrax creep samples, and two transverse columnar Monofrax 

creep samples. Porosity measurements were only made on five of each brand of bulk specimens and 

on representative samples for each type of columnar specimens. For the Monofrax bulk material, an 

average apparent density of 3.44 ± 0.02 g/cm3 (214.76 ± 1.25 lb/ft3

), and bulk density of 3.44 ± 

0.01 g/cm3

 (214.60 ± 0.89 lb/ft3) was found with an average porosity value of 1.82 ± 0.31%. An 

average apparent density of 3.48 ± 0.01 g/cm3

 (217.26 ± 0.69 lb/ft3) and average bulk density of 

3.48 ± 0.01 g/cm3

 (217.19 ± 0.75 lb/ft3) with an average porosity of 1.47 ± 0.59% was found for the 

Corhart bulk material. These values are in line with those supplied by the manufacturers, but the 

experimental density value is slightly higher and the experimental porosity value is slightly lower for 

the Monofrax material while both the experimental density and porosity levels are slightly lower for 

the Corhart material (Monofrax bulk density = 3.40 g/cm3

 and porosity = 1.9%, Corhart bulk density 

= 3.54 g/cm3

 and porosity = 2.0%). The longitudinal columnar Monofrax material showed an average 

apparent density of 3.42 ± 0.01 g/cm3

 (213.51 ± 0.69 lb/ft3), an average bulk density of 3.42 ± 

0.01 g/cm3

 (213.35 ± 0.50 lb/ft3), and an average porosity of 1.06 ± 0.01%. The transverse columnar 

Monofrax material showed an average apparent density of 3.41 ± 0.01 g/cm3

 (212.89 ± 0.50 lb/ft3), an 

average bulk density of 3.42 ± 0.01 g/cm3

 (213.29 ± 0.33 lb/ft3), and an average porosity of 1.09 ± 

0.21%. 

35 
Corhart AB Regular Cast Block X And Y Core

(1in. Segment) Densities

3.0

3.1

3.2

3.3

3.4

3.5

3.6

02468

Mean Depth (in.)

Mean Bulk Density (g/cm3

)

10

Y Core

X 3in Core

X 7in Core Beta Zone

PN 1503 Monofrax M Regular Cast Block X And Y Core 

(1in. Segment) Densities

3.0

3.1

3.2

3.3

3.4

3.5

3.6

02468

Mean Depth (in.)

Mean Bulk Density (g/cm3

)

10

Y Core

X 3in Core

X 7in Core

Beta Zone

Fig. 24. x3-, x7-, and y-core density profiles for the (a) Corhart refractory and 

(b) Monofrax refractory. 

36 
The phase profiles of both refractories were measured by X-ray fluorescence (XRF) and XRD. For 

the x-direction profiles, the Corhart block showed large amounts of β-alumina (much greater than the 

mean pour stream chemistry of the melt) present in the outer 1 to 2 in. of the block and a lower 

(below the mean pour stream of the melt), more consistent, β-alumina content through the remainder 

of the block. In the case of the y core, there was an extremely large amount of β-alumina at the 

bottom block surface and a fairly consistent level of β-alumina in line with what was seen in the 

interior x-cores from ≈ 0.5 to 6.5 in. Beyond the 6.5-in. mark, the header of the block was 

encountered, and the β-alumina content increased to the original high level. The Monofrax block 

showed levels of β-alumina similar to the mean pour stream chemistry of the melt for all three cores 

up to ≈ 1.5 in. Beyond that point, the three cores all showed a lower β-alumina content, similar to 

what was seen in the Corhart block. This low level of β-alumina was sustained for the x3 core but 

showed an increase at ≈ 4.5 in. for the x7 core and at ≈ 6.5 in. for the y core. It is expected that the 

increase in β-alumina content was due to the cores entering the β-alumina-rich header region of the 

block, as seen in the Corhart block. 

Patterns for bulk and columnar regions of Corhart M and Monofrax M could be matched to known 

peaks for corundum (α-alumina) (PDF# 10-0173) and sodium aluminum oxide (β-alumina) (PDF# 

31-1263). Initially, optical microscopy was used to characterize the various components present in the 

microstructure of the fusion-cast alumina blocks. It was found that both the Corhart and Monofrax ascast blocks displayed a similar variable microstructure. As shown in Fig. 25, the bulk region of the 

blocks was composed of blocky α-alumina grains embedded in a continuous matrix of β-alumina 

while the columnar exterior regions were composed of longer, more slender α-alumina grains 

embedded in the continuous 

β-alumina matrix. 

Grain size analysis was performed on characteristic regions (bulk and columnar) of the two regularcast block brands. In-house LabViewTM software that relates the pixels contained in the area of each 

grain to a calculated average grain diameter in microns was used. For the Corhart columnar regions, 

133 grains were analyzed, yielding an average grain size of 82 μm (3.2 mils) ± 43 μm (1.7 mils). The 

Monofrax columnar region had a similar average grain size of 79 μm (3.1 mils) ± 46 μm (1.8 mils) 

based on the analysis of 164 grains. The bulk regions of the Corhart and Monofrax as-cast blocks 

exhibited a larger average grain size and showed more variation between the two brands. Analysis 

was carried out on 225 grains in various Corhart bulk regions, giving an average grain size of 172 μm 

Fig. 25. Optical microscopy of characteristic grain structures for (a) bulk structure, and (b) columnar 

structure.

37 
(6.8 mils) ± 20 μm (0.8 mils). For the Monofrax bulk regions, 179 grains were evaluated, yielding an 

average grain size of 231μm (9.1 mils) ± 47μm (1.9 mils). 

It was originally hypothesized that the c-axes of the hexagonal α-alumina crystals in the columnar 

microstructural zone would all be aligned along the direction of rapid cooling, as indicated by the 

long dimension of the columnar crystals. This is in fact not the case, (i.e., Fig. 26). Through analysis 

of optical micrographs it was found that incomplete α-alumina grains in the microstructure actually 

have their c-axis aligned perpendicular to the cooling direction. The skeletal columnar grains were 

found to be polycrystalline (not single crystals) and to have their c-axes randomly oriented as 

additionally seen in the random habit of the α-alumina phase. The shape of these grains was found to 

therefore be due more to the cooling behavior than the crystallographic direction. 

Other optical images were taken showing the nephelitic (calcium-aluminosilicate) grain boundary 

phase present between the α- and β-alumina phases and in isolated areas of the β-alumina phase. 

These areas appear as grayish, angular features between grains and isolated in the β-alumina phase, as 

shown in Fig. 27. 

Elastic modulus measurements were generated by cyclic compression. Supplemental values were 

generated for comparison by using a nanoindentation technique. At room temperature, an average 

value of 72 ± 4 GPa (10.4 ± 0.6 H 106 psi) was found for the Monofrax bulk material, while the 

Corhart bulk material showed an average value of 109 ± 7 GPa (15.8 ± 0.9 H 106 psi). The 

longitudinal columnar Monofrax samples gave an average value of 112 ± 12 GPa (16.2 ± 1.7 H

106 psi), while the transverse columnar yielded a value of 113 GPa (16.4 H 106 psi) based solely on 

specimen AG18. High-temperature testing was carried out on a sample from each of the above 

categories. Results from the high-temperature testing are shown in Fig. 28. As expected, in general 

the modulus decreased with increasing temperature. Unexpectedly, the modulus of the Monofrax bulk 

at 900°C (1652°F) was higher than that measured at room temperature. This may suggest that the 

room-temperature modulus of the Monofrax bulk material should be higher (on the order of 140 GPa, 

20.3 H 106 psi). 

Nanoindentation measurements were made in each of the three phases (α-alumina, β-alumina, and 

nephelitic grain boundary phase) for both the Corhart and Monofrax bulk materials at room 

temperature. Three analyses were run in α-alumina grains (one Corhart and two Monofrax) and three 

in β-alumina grains (one Corhart and two Monofrax). Six analyses were run in nephelitic grain 

boundary regions (four Corhart and two Monofrax). All results are shown in Table 11. The modulus 

values obtained through nanoindentation of the alumina phases (214 – 442 GPa, 31.0 H 106 – 64.1 H

106 psi) were higher than those previously found for the bulk material by static and measurements 

(62 – 117 GPa, 4.6 H 106 – 17.0 H 106 psi). This may partly be due to the fact that in the 

nanoindentation measurements α - and β-alumina grains are individually sampled where in the other 

measurements, a composite of the bulk material was sampled. Also, it is documented that values 

found through nanoindentation are higher than those found through other methods simply due to the 

technique itself [22]. Regardless, the values are fairly consistent between the different brands of 

material for each sample type and provide a means of comparison between the different components 

of the bulk structure. The α-alumina grains were found to have a much higher modulus than the βalumina grains. The Corhart grain boundary samples also possessed a high modulus on the order of 

those values measured for the β-alumina grains (≈ 260 GPa, 37.7 H 106 psi), but the Monofrax 

boundary specimens had a modulus value more in line with the values measured previously for bulk 

materials with the static techniques (≈ 95 GPa, 13.8 H 106 psi). 

38 
Fig. 26. Optical microscopy of columnar region (micron bar = 200 μm).

Fig. 27. Optical microscopy of nephelitic grain boundary phases.

39 
Young’s Modulus Values for Fusion-Cast Alumina Samples

Fig. 28. Static Young’s modulus values for fusion-cast alumina sample types.

Initial thermal expansion measurements 

were made from room temperature up to 

900°C on high-purity (99.4%) alumina 

(theoretical coefficient of thermal 

expansion and temperature range = 8.2 H

10!6

/°C) using the dual rod dilatometry 

technique to verify its calibration. Three 

measurements were made, yielding an 

average value of 8.7 H 10!6

/°C. This was 

deemed sufficiently close to the 

theoretical value for experimental testing 

to continue. 

Table 11. Fusion-cast alumina nanoindentation results 

Young’s modulus 

Grain (GPa) 

Corhart α-alumina 407 ± 15 

Monofrax α-alumina 442 ± 13 

Monofrax α-alumina 424 ± 14 

Corhart β-alumina 237 ± 5 

Monofrax β-alumina 214 ± 6 

Monofrax β-alumina 231 ± 5 

Corhart grain boundary 296 ± 13 

Testing on Corhart and Monofrax M bulk 

material was initially carried out on 

specimens at 1450°C (2642°F) under the 

following schedule. Cylindrical samples 

were heated to temperature in a 4-h period 

and then held at temperature for 100 h. 

Following the soak period, the samples 

were cooled to room temperature through natural cooling. Three tests were run on each commercial 

brand with the Corhart material, yielding an average value of 7.8 H 10!6

/°C, and on the Monofrax 

bulk, yielding an average value of 8.4 H 10!6

/°C. The values published by the manufacturers for 

Corhart M and Monofrax M are 11.1 x 10-6/°C and 7.9 H 10!6

/°C respectively. Theoretically, αalumina is anisotropic due to its hexagonal crystal structure and possesses thermal expansion values 

of 8.3 H 10!6

/°C perpendicular to its c-axis and 9.0 H 10!6/°C parallel to its c-axis. $-alumina has an 

isotropic thermal expansion value of 6.4 H 10!6

/C. 

Corhart grain boundary 269 ± 11 

Corhart grain boundary 249 ± 11 

Corhart grain boundary 240 ± 10 

Monofrax grain boundary 97 ± 5 

Monofrax grain boundary 92 ± 6 

40 
Monitoring of the overall linear sample dimensions during the thermal expansion testing resulted in 

evidence of unrecoverable expansion of the specimens during exposure to temperature. Exposure for 

100 h at 1450°C (2642°F) resulted in a 0.194% change in length for Corhart M specimens and a 

0.131% change in length for Monofrax M specimens. Supplemental testing was performed at 1550°C 

(2822°F) and 1650°C (3002°F) to evaluate the unrecoverable expansion at those temperatures. Again, 

cylindrical specimens were heated to temperature in a 4-h-period and then held at temperature for 

over 100 h before being allowed to return to room temperature through natural cooling. Testing 

revealed that exposure for 100 h at 1550°C (2822°F) resulted in a 0.2151% change in length for 

Corhart M specimens and a 0.357% change in length for Monofrax M specimens. Exposure for 100 h 

at 1650°C (3002°F) resulted in a 0.268% change in length for Corhart M specimens and a 0.396% 

change in length for the Monofrax M specimens. The results for the bulk Corhart and bulk Monofrax 

refractories are summarized in Figs. 29(a) and 29(b), respectively. 

The transient technique was used to make thermal conductivity measurements on both cylindrical 

creep specimens and on sectioned pieces of the original as-cast blocks. Two bulk creep cylinders and 

one sectioned piece with both cut surfaces and original surfaces were sampled for each brand. Results 

from original surfaces are thought to give measurements representative of the columnar structure with 

a parallel orientation to the direction of measurement (longitudinal columnar). Results from the center 

of the cut surfaces are representative of the bulk structure, and measurements on the edges of the cut 

surfaces yield values indicative of the columnar structure with a perpendicular orientation to the 

direction of measurement (transverse columnar). For the creep cylinders, measurements were taken 

on each sample by the single-sided method and between both samples by the standard method. 

Sets of three single-sided measurements on the two bulk creep cylinders (B1 and B2) yielded average 

thermal conductivity values of 5.91 ± 0.10 and 5.97 ± 0.04 W/mK, respectively. The three standard 

measurements taken between the two cylinders gave an average value of 6.00 ± 0.11 W/mK. These 

values agree well with the sectioned block measurements of 6.04 to 6.21 W/mK. Measurements on 

the longitudinal and transverse columnar microstructures yielded average values of 7.10 and 

5.82 W/mK, respectively. The differences in thermal conductivity values between the two columnar 

microstructures and the bulk structure are thought to be due to the varying contributions and 

orientations of the "- and $-alumina grains in the three different cases. 

The piece sectioned from the Corhart block was analyzed along two vertical lines on the cut surface. 

One was taken along the edge of the surface and the other 3 in. in from the edge. Measurements taken 

along the edge of the cut surface (line along surface) correspond to the columnar microstructure with 

a perpendicular orientation to the direction of measurement (transverse columnar). Measurements 

taken along the 3-in. line are representative of the bulk microstructure except for material in about the 

first inch. Measurements along that region contain transverse columnar structure due to interaction 

with surface OS1 (Fig. 30), and the measurements made above 7.0 in. are in the header microstructure. The results indicate that the transverse columnar microstructure had an average thermal 

conductivity ranging anywhere from 5.04 to 7.46 W/mK. The lower values occurred at the top of the 

block, in the vicinity of the header region and were due to the increased porosity in this region and 

therefore are not considered representative of the actual thermal conductivity of the transverse 

columnar structure. Thus the average thermal conductivity values are in the range of 6.50 to 

7.50 W/mK. If it is hypothesized that the "-alumina columnar structure controls the conductivity in 

these regions to varying degrees, then these higher values would make sense because the "-alumina 

grains will possess a higher thermal conductivity than the $-alumina matrix, thus raising the overall 

thermal conductivity of the material. The same phenomenon is seen in the sectioned Monofrax block 

for the longitudinal columnar microstructure. For cut block surfaces, this is considered to be the case 

because the harder "-alumina grains protrude slightly from the block surface (as seen by the intensity 

differences in optical microscopy). The slight protrusion would then cause the "-alumina grains to 

41 
Fig. 29. Specimen elongation as a function of time for (a) the bulk Corhart and 

(b) the bulk Monofrax refractories. The curves represent mathematical fits to the 

experimental data as described in J. G. Hemrick, “Creep Behavior and Physical 

Characterization of Fusion-Cast Alumina Refractories,” Ph.D. dissertation, University 

of Missouri-Rolla, 2001.

42 
Fig. 30. Monofrax thermal conductivity block section denoting test surfaces. Parentheses denote 

hidden surface. Dotted lines designate positions of line scans (os = original surface; cs = cut surface).

dominate the thermal conductivity measurements made on these surfaces and would cause the 

“composite reading” to be higher than expected. 

In the case of the 3-in. line, the bulk microstructure of the Corhart material had an average thermal 

conductivity on the order of 6.25 to 7.10 W/mK with a mean value of 6.78 W/mK. The higher values 

in the bulk region may be due to direct measurements of the "-alumina grains as opposed to 

measurements for areas containing both "- and $-alumina grains. Therefore, it is thought that the 

actual bulk thermal conductivity of the bulk Corhart material is on the order of 6.61 W/mK because 

the continuous $-alumina phase should dominate the conductivity behavior. Due to the high void 

content, the thermal conductivity in the header region fell from approximately 5.00 W/mK down to 

approximately 0.25 W/mK. 

The laser flash method yielded thermal diffusivity/conductivity values from room temperature up to 

1400°C for both the Corhart and Monofrax bulk as-cast materials. Two measurements were taken at 

each temperature value on separate samples with 15 temperatures being measured. Specific heat 

measurements were also made from room temperature up to 1400°C using a calorimeter method. 

Plots of the change in thermal conductivity with temperature are shown for the Corhart and Monofrax 

materials in Fig. 31. The thermal conductivity of the Corhart material at room temperature was higher 

than that measured by the transient method (9.0 W/mK as opposed to 6.6 W/mK). This value then 

decreased to just above 4.0 W/mK at 600°C (1112°F) and then slowly increased to about 7.0 W/mK 

at 1400°C (2552°F). The thermal conductivity of the Monofrax material was much more consistent. 

The room-temperature value, measured to be on the order of 6.0 W/mK was similar to that 

determined from the transient method. The value for the Monofrax sample dropped slightly to around 

4.0 W/mK, as it did for the Corhart material, but at a lower temperature, only 100°C (212°F). The 

thermal conductivity was relatively constant until 900°C (1652°F), at which point the value began to 

increase to around 6.0 W/mK at 1400°C (2552°F). 

The compressive creep curves for the Corhart bulk specimens at 1450°C exhibited a period of rapid 

strain accumulation followed by a leveling off of the accumulated strain. The total accumulated strain 

and the associated length of time over which the rapid strain accumulation occurs was expected to be 

related to the applied stress state. This was not always found to be the case. For example, the creep 

response at 0.8 MPa was significantly lower than the response at 0.6 MPa. In the case of the curves 

obtained at 1550°C, the strain accumulation also exhibited a leveling off with time. However, for a 

given time, the amount accumulated strain increased with increasing stress, predicted by the creep 

equation. The creep curves generated at 1650°C were different from those obtained at the lower 

43 
Fig. 31. Average laser flash thermal conductivity for as-cast bulk refractories: 

(a) Corhart and (b) Monofrax.

temperatures. Both the 0.6-MPa and 0.8-MPa curves exhibited similar behavior, initially showing 

large amounts of accumulated strain followed by a leveling off at 100 and 150 h, respectively. In both 

cases a larger amount of total accumulated strain was seen under these stresses than at either 1450°C 

or 1550°C. The 1.0-MPa curve showed an increase in accumulated strain followed by a region of 

constant negative slope. This was the first test in which traditional creep behavior was exhibited. 

Similar trends were observed for the bulk Monofrax specimens. 

In the case of the Monofrax refractory, creep curves were obtained for longitudinal and transverse 

columnar samples at both the low (1450°C, 6 MPa) and high (1650°C, 1.0 MPa) extreme matrix 

conditions, with an additional transverse columnar sample being tested at 1450°C and 1.096 MPa. For 

44 
the curves at 1450°C, both the transverse and longitudinal columnar samples showed similar behavior 

at 0.6 MPa with a small amount of accumulated strain being followed by a leveling off of 

accumulation after ≈ 150 h. The total amount of accumulated strain was about half as much as that for 

both the Corhart and Monofrax bulk samples under similar conditions. The plot of the transverse 

columnar specimen at 1.0 MPa showed an increase in accumulated strain over that seen at 0.6 MPa 

and a sharp leveling off of the strain accumulation at an earlier point (≈ 50 h). The total accumulated 

strain in this sample (≈ 0.05%) was much lower than that seen in both Corhart and Monofrax bulk 

samples under similar conditions (0.09 and 0.13%, respectively). At 1650°C both 1.0-MPa curves 

showed similar behavior to that seen in the Corhart and Monofrax samples under similar conditions, 

although the rate of accumulated strain was lower in the columnar cases, resulting in lower total 

accumulated strains. This behavior included an increase in accumulated strain followed by a region of 

accumulated strain characterized by a constant slope and traditional creep behavior. 

Based on the above observations, it is clear that nontraditional creep behavior was exhibited by 

fusion-cast alumina under most test conditions (those below 1650EC/1.0 MPa). This behavior was 

thought to be due to the unrecoverable expansion phenomenon shown in Fig. 29. Based on those 

premises, plots were constructed incorporating both the expansion effects and creep effects seen by 

each material under the various test conditions and the composite effect resulting from the 

combination of these two concurrent mechanisms. Such plots are shown below for both Corhart 

(Fig. 32) and Monofrax (Fig. 33) bulk specimens. These plots were constructed by subtracting a 

similar exponential fit of the measured expansion curve as determined by dilatometry at each test 

temperature from the modeled creep curves at each stress of the representative same temperature. 

Fig. 32. Combined plots of expansion and creep effects in Corhart bulk samples.

45 
Fig. 33. Combined plots of expansion and creep effects in Monofrax bulk samples. 

The new curves predict a greater creep resistance in the Monofrax bulk material at 1450°C (better by 

≈ 0.20%), but a greater creep resistance in the Corhart bulk material at 1550°C (better by ≈ 0.15%) 

and at 1650°C (better by ≈ 0.10%). However, this may not be an accurate portrayal of the creep 

behavior exhibited by these materials. It must be remembered that the behavior exhibited by these 

materials in service would still be represented by the measured creep curves. Therefore, by 

performing the analysis in this fashion, it may be possible to produce creep curves that possess a 

steady-state, non-zero creep regime from which a model can be conceived by using a traditional 

method such as the NBA equation. Doing so will not capture the actual behavior exhibited by this 

material, however, because any model must incorporate the effect of the expansion along with the 

creep. A model is needed that can handle the high degree of nonlinearity seen in the measured creep 

curves generated for fusion-cast alumina while also being sensitive to the transient nature of the data 

because no equilibrium state is in effect. 

Several samples were analyzed after being crept, and results were compared with those previously 

obtained from virgin samples. Analysis included thermal conductivity, optical microscopy, and 

SEM/EDS chemical analysis. The thermal conductivity values for the Monofrax post-creep cores 

were similar to those obtained earlier from the virgin cores, which had thermal conductivities in the 

range of 5.9 W/mK. The value obtained from the Corhart post-creep core was significantly lower than 

the average value of 6.61 W/mK found earlier for the Corhart bulk material. 

Initial optical microscopy and SEM/EDS analyses were performed on Corhart and Monofrax bulk 

samples exposed to 1550°C and 1.0 MPa. This work was conducted to determine whether any 

46 
changes in the microstructure or chemistry could be detected at that intermediate condition. 

SEM/EDS analysis was then used to determine whether analysis needed to be performed on samples 

at the lower temperatures and stresses. No discernible deformation and only slight chemical changes 

were seen in the samples. 

Subsequent optical microscopy and SEM/EDS analyses were performed on Corhart and Monofrax 

bulk samples exposed to the extreme conditions of 1650°C and 1.0 MPa. The structure of the Corhart 

sample was found to resemble the virgin structure, while the Monofrax crept structure displayed a 

smaller 

α-alumina grain size and a tighter α-alumina grain structure with more equiaxed angular grains than 

previously observed in the virgin structure. More severe chemical changes were also seen between the 

virgin and crept samples. 

5.4 Fusion-Cast Spinel Refractories 

Engineering creep data for selected refractory materials was generated to a temperature of 1650°C in 

ambient air. Tests for the fusion cast material at 1650o

C and 3.0 MPa. One test was run on bulk 

material and one was run on columnar (with the structure running parallel to the loading direction). 

There is some initial movement due to the application of the load [the samples were heated to 

temperature under low load (30–50 lb) and were allowed to dwell for 24 h before the main load 

(768.909lb) was applied], then very little creep over the time of testing. The test frame was refitted 

with a larger air cylinder and load cell to enable testing at higher loads in an attempt to get this 

material to creep. Also examined was the effect of using smaller samples to obtain higher loads. 

The behavior exhibited by the transverse columnar structure was not as expected (Fig. 34). It was 

expected to be similar to the behavior of the longitudinal columnar sample with perhaps even less 

accumulated strain. Instead, the behavior is more similar to that of the bulk structure with about the 

same amount of total accumulated strain but a slower initial rate of strain accumulation. Curves from 

two separate samples both show the same initial strain accumulation, showing that the results are 

repeatable. Rather than analyzing the final 100 h of the test, all the data available past 100 h of testing 

were analyzed (Table 12). This analysis provided a better representation of the apparent “steadystate” creep rates exhibited by these materials. There was a plot of this analysis included as well. The 

conclusion was that the fusion-cast material does not creep (even at 8.5 MPa). There was more initial 

strain due to application of the load (~ 3% more, which corresponds with the increase from 3 to 8.5 

MPa), but there was not a change in steady-state creep rate with the increased load. The creep rate for 

the longitudinal columnar structure was lower, but the transverse columnar structure showed a creep 

rate similar to that of the bulk structure. The bonded material showed significant creep and was far 

inferior to the fusion-cast material, as shown in Figs. 35, 36, and 37. 

Thermal diffusivity measurements were made by the laser flash method. Density and thermal 

diffusivity are required as input into the calculation of thermal conductivity. The density of the fused 

cast spinel was measured as 3.47 g/cc with a porosity of 2.3%. The density of the columnar structure 

and bulk structure were 3.47 g/cc and 3.5 g/cc with porosities of 2.1% and 2.5%, respectively. The 

density of the bonded spinel was measured as 3.04 g/cc with a porosity of 14.2%. The calculated 

thermal conductivity results are shown in Fig. 38. 

47 
Table 12. Calculated strain rates 

Material Temperature (°C) Stress (MPa) Strain rate 

1.27 × 10-6 Fused spinel longitudinal Columnar 1650 3.0 /h 

4.27 × 10-6 Fused spinel transverse columnar 1650 3.0 /h 

5.66 × 10-6 Fused spinel bulk 1650 3.0 /h 

5.00 × 10-6 Fused spinel bulk 1650 8.5 /h 

48.0 × 10-6 Bonded spinel bulk 1650 0.07 /h 

8.89 × 10-6 Bonded spinel bulk 1550 0.06 /h 

23.8 × 10-6 Bonded spinel bulk 1550 0.3 /h 

Fusion Cast Spinel (1650oC)

-3

-2.5

-2

-1.5

-1

-0.5

0

0 50 100 150 200 250 300

Time (hours)

Accumulated Strain (%)

Bulk (3.0 MPa)

Longitudinal Columnar (3.0 MPa)

Transverse Columnar (3.0 MPa)

Bulk (8.5 MPa)

Fig. 34. Fused spinel accumulated strain curves. 

Bonded Spinel (1650oC, 0.07 MPa)

-2.5

-2

-1.5

-1

-0.5

0

0 5 10 15 20 25

Time (hours)

Accumulated Strain (%)

Application of 3.0 MPa Load

Failur

Fig. 35. Bonded spinel accumulated strain curve at 1650°C. 

48 
Bonded Spinel (1550o

C, 0.06 MPa)

-0.1

-0.09

-0.08

-0.07

-0.06

-0.05

-0.04

-0.03

-0.02

-0.01

0

0 5 10 15 20 25

Time (hours)

Accumulated Strain (%)

Fig. 36. Bonded spinel accumulated strain curve at 1550°C. 

Bonded Spinel (1550oC, 0.3 MPa)

-1.2

-1

-0.8

-0.6

-0.4

-0.2

0

0 50 100 150 200 250 300 350

Time (hours)

Accumulated Strain (%)

Discontinuity due to gage problem

Fig. 37. Bonded spinel accumulated strain curve at 1550°C. 

49 
0

2

4

6

8

10

12

14

16

0 200 400 600 800 1000 1200 1400

Temperature (C)

Thermal Conductivity (W/mK)

Bonded Fused Columnar Fused Bulk

0

0.005

0.01

0.015

0.02

0.025

0.03

0.035

0.04

0.045

0.05

0 200 400 600 800 1000 1200 1400

Temperature (C)

Thermal Diffusivity (cm2/sec)

Bonded Fused Columnar Fused Bulk

Fig. 38. Thermal diffusivity of fused and bonded spinel. 

The creep specimens were subjected to detailed microstructural and phase analysis. One objective of 

this task was to better understand how creep and corrosion behavior relates to microstructure and 

microchemistry. X-ray diffraction data showed little differences in the patterns between the “as 

supplied” and “exposed” samples. The spinel peaks were slightly shifted due to heat treatment and 

stress relief. There were no additional phases found after exposure. 

The sintered spinel brick was made from high-purity large fused spinel grains. The matrix contained 

16% to 18% porosity but was well sintered. No impurity phase (forsterite, monticellite, merwinite or 

melilite) was identified/recognized at the magnification scale of RL/CL microscope used. This does 

50 
not, however, exclude the presence of small amounts of glassy or ternary phases, which would coat 

the smaller spinel grains in the matrix. In general, the fused spinel block consisted of spinel and 

magnesia. The presence of free magnesia or magnesia inclusions suggests that the liquid was rich in 

MgO. The composition of spinel was also expected to be rich in MgO. Due to the rapid cooling or 

crystallization columns of spinel skeletons were formed and the magnesia was segregated as tiny 

inclusions within these skeletal spinel columns. Such rapid crystallization appeared as linear or 

locally oriented features in the fused block. The orientation of these skeletal spinel columns was, in 

general, also random. In a single section both transversal or longitudinal microstructures were 

observed. For example, Fig. 39 shows a diagonal (NW to SE) crystallization feature. Although spinel 

(green CL) crystals still exhibit random orientation, overall there is a diagonal (in these micrographs) 

feature developed during crystallization of spinel liquid. These micrographs illustrate the 

microstructure of rapidly crystallized spinel columns sectioned perpendicularly along the axis of the 

column. 

Fig. 39. RL (left) and CL (right) images of fused spinel. 

51 

6. Accomplishments 

6.1 Research-Based Accomplishments 

This study led to significant improvements in the resolution of the techniques commonly used to 

measure thermal expansion and creep of refractories. For example, the “Standard Test Method for 

Load Testing Refractory Brick at High Temperatures” involves loading a refractory brick for a 

prescribed duration and temperature and then determining the material’s resistance to deformation by 

comparing its before- and after-test axial dimensions. A shortcoming of this test is the introduction of 

error in the analysis if the refractory brick’s ends “bed-down.” The procedures and equipment 

recommended in the “Standard Test Method of Measuring the Thermal Expansion and Creep of 

Refractories Under Load” are similar to those used in the present study. A shortcoming of this 

procedure involves the recommended measurement of dimensional change in the creep specimen. The 

use of LVDTs to measure the contraction of the specimen ends is advocated but can be problematic 

for reasons described previously. The extensometry system described in Fig. 2 of ASTM C832 

closely matches the extensometry system used in the present study; however, in the present study a 

capacitance gage is used while in the standard the use of an LVDT for the measurement is shown. 

The load that activates an LVDT can be enough to affect the intimacy of the extensometer contact 

with the specimen and thus can yield a misleading creep measurement if there is insufficient 

horizontal load applied to the specimen. Because of its differing operating principle, the capacitance 

gage on the extensometry systems used in the present study requires no load for activation and does 

not introduce uncertainties associated with the horizontal load and the use of an LVDT. 

This study also identified new mechanisms responsible for deformation of fusion-cast alumina and 

silica refractories in the absence of an applied stress. For both refractories these mechanisms led to 

expansion of the test specimens in the absence of an applied stress. 

6.2 Technology Transfer 

The results of this project have established hardware and a test methodology that can be used by the 

refractories community to evaluate the creep and aging behaviors of most refractory materials. Of 

particular merit is the development of specialized extensometry, which offers extremely high 

resolution required for the measurement of small deformations. The project staff made presentations 

at national meetings and published results in industry technical journals. 

6.3 Publications and Patents 

[1] A. A. Wereszczak, T. P. Kirkland, G. A. Pecoraro, and R. A. New, “Compressive Creep 

Behavior of Fusion-Cast Alumina Refractories,” pp. 401–406 in Ceramic Transactions, 

Vol. 82, Advances in Fusion and Processing of Glass II, ed. A. G. Clare and L. E. Jones, 

American Ceramic Society, Westerville, Ohio (1998). 

[2] A. A. Wereszczak, T. P. Kirkland, and W. F. Curtis, “Compressive Creep Resistance of MgO 

Refractories at Temperatures ≥ 1400°C (2550°F),”pp. 407–412 in Ceramic Transactions, 

Vol. 82, Advances in Fusion and Processing of Glass II, ed. A. G. Clare and L. E. Jones, 

American Ceramic Society, Westerville, Ohio (1998). 

[3] A. A. Wereszczak, T. P. Kirkland, and W. F. Curtis, “Creep of CaO/SiO2 - Fluxed MgO 

Refractories,” J. Mater. Sci (1998) in press. 

53 
[4] A. A. Wereszczak, M. Karakus, K. C. Liu, B. A. Pint, T. P. Kirkland, and R. E. Moore, 

Compressive Creep Performance and High Temperature Dimensional Stability of Conventional 

Silica Refractories, ORNL/TM-13757, Oak Ridge National Laboratory (1999). 

[5] J. G. Hemrick, A. A. Wereszczak, M. Karakus, K. C. Liu, H. Wang, B. A. Pint, T. P. Kirkland, 

and R. E. Moore, Compressive Creep and Thermophysical Performance of mullite Refractories, 

ORNL/TM-2002/84, Oak Ridge National Laboratory (2002). 

[6] J. G. Hemrick and A. A. Wereszczak, “Compressive Creep of Fusion-Cast Alumina as a 

Function of Microstructure,” Glass Industry (November 1999). 

[7] J. G. Hemrick, “Creep Behavior And Physical Characterization Of Fusion-Cast Alumina 

Refractories,” a Dissertation Presented to the Faculty of the Graduate School of the University 

of Missouri-Rolla in Partial Fulfillment of the Requirements for the Degree of Doctor of 

Philosophy in Ceramic Engineering (2001). 

6.4 Commercialization 

The data generated in this project were provided to participating refractory manufacturers and are 

available from those manufacturers to any users of these types of refractories. It has been confirmed 

with industrial partners involved in this project that creep data and other thermomechanical property 

information generated as part of this project was used in analysis of glass furnaces currently in service 

and for the planning and design of refractory systems for future furnace constructions. The data has 

not only been used by glass companies in their analysis and planning of furnace systems, but by 

refractory companies and industrial consultants as well. The information generated was said to be 

“invaluable” by one industrial participant and more data of similar scope is felt to be of general 

interest to the industry as a whole. 

54 
7. Summary and Conclusions 

7.1 Silica 

The creep and corrosion performances of the six refractories were examined. The bulk characteristics 

of the refractories were then characterized before and after their creep testing in an attempt to 

correlate their changes to the measured creep responses. Bulk density, phase content, microstructure, 

and secondary phase composition were examined, and their changes as a function of test temperature 

were analyzed. Lastly, the effect of temperature on the dimensional stability of these refractories was 

examined. Specimens were aged at the same temperatures that the creep tests were performed. The 

procedures used and the results generated are discussed. 

• The amount of compressive creep of Gen-Sil, Stella GGS, Vega, Vega H, and Harbison-Walker’s 

new developmental brand (flux factors less than 0.5% or “Type A” for all five of these brands) 

conventional silica was negligible at temperatures between 1550 and 1650°C (2820–3000°F) and 

at compressive stresses from 0.2 to 0.6 MPa (29–57 psi). Only one specimen per condition was 

tested, so the authors were unable to statistically conclude that any one of the five brands had 

superior creep resistance to the others; however, if differences do indeed exist, then they are 

believed to be insignificant. 

• The compressive creep of SI96AU (flux factor greater than 0.5% or “Type B”) conventional 

silica was negligible at 1550°C. However, SI96AU deformed (contracted) by more than 20% at 

1600°C at stresses between 0.2 and 0.6 MPa. This behavior was inferior to that of the other five 

tested brands. SI96AU was not creep tested at 1650°C; however, its creep deformation was 

expected to still be inferior to the other five silica brands. 

• The compressive creep rates of all six brands were not able to be represented as a function of 

temperature and compressive stress by the conventionally used NBA creep equation (i.e., the 

Arrhenius power-law creep model). Concurrently active mechanisms other than creep, resulted in 

larger or oppositely anticipated dimensional changes than those produced by creep; this effect 

limited the ability to identify or interpret the lesser-active creep mechanism in these silica 

refractories. Models other than NBA (or equivalent “creep” models that represent creep rate as a 

function of temperature and stress) should be used to represent or predict the long-term 

dimensional stability of conventional silica superstructures that are subjected to compressive 

stresses less than 0.6 MPa and temperatures between 1550 and 1650°C. 

• The change in dimensions of the compressively crept specimens indicated that their size (both 

diameter and length) had actually increased as a consequence of the creep test conditions. The 

relatively large amount of expansion of the specimen diameters could not have been a sole 

consequence of “compressive creep barreling” because some specimens did not contract at all, 

and those that did not contract enough to cause a Poisson effect. All six brands showed this 

expansion effect. The increases in diameter and length of the creep specimens were between 0.35 

and 1.0%. Only a fraction of the expansion was detected during the compressive creep testing. 

This indicates that most of the permanent expansion exhibited by the silica specimens occurred 

during their heating from ambient to the creep-testing temperature or during the cooling from test 

temperature to ambient. 

• The secondary phase constituents remained in all brands when they were tested at 1550°C 

(cumulative time at temperature approximately 250 h). A fraction of these phases visibly evolved 

from the specimens at 1600°C (slight glass bubbling on the specimen and fixturing). This 

phenomenon was quite severe at 1650°C. The density changes in the crept specimens (a net effect 

of the volume expansion and loss in secondary phase mass) ranged from a decrease of 1.6 to 

55 
3.9% with a subtle trend of greater density decreases at higher test temperatures. The density 

decreases among the six brands were statistically equivalent within the data scatter. 

• The cause of the volume expansion was not identified; however, because of the absence of quartz 

it can be concluded that the expansion was not due to quartz (a relatively high-density phase) 

converting to tridymite or cristobalite (lower-density phases). The authors speculate that the 

volume expansion is due to the high-temperature process of the conversion of tridymite to 

cristobalite and the consequential growth of the residual pores and the original cristobalite 

aggregates. 

• Unstressed or “aged” silica refractory specimens from all six materials also exhibited a loss of 

mass, dimensional increases, and density decreases between 1550 and 1650°C. 

- The mass losses of the six brands at 1550 and 1600°C were equivalent; however, some 

silica brands lost more mass than others at 1650°C: Harbison-Walker’s developmental 

brand lost approximately 0.1% of its mass; SI96AU lost ≈ 0.2%, Vega H lost ≈ 0.3%, 

Gen-Sil lost ≈ 0.5%, Vega lost ≈ 1.0%, and Stella GGS lost ≈ 1.5%. Most of the mass 

loss for all six brands occurred in less than 25 h. 

- The dimensional increases of the silica specimens were a function of temperature. 

Additionally, some brands expanded more than others. The ranges of expansion were 

≈ 1–2.5% at 1550°C, ≈ 2.25–3.75% at 1600°C, and ≈ 3.5–6% at 1650°C. 

Harbison-Walker’s developmental brand expanded the least amount at all three 

temperatures, followed by Gen-Sil. The ascending order of expansion for the other four 

brands varied with temperature. Stella GGS expanded the most at 1550°C (≈ 2.5%), 

SI96AU expanded the most at 1600°C (≈ 3.75%), and Vega expanded the most at 

1650°C (≈ 6%). Dimensional increases were greater than axial expansion of the crept 

silica specimens at the same temperatures, indicating that the axially applied compressive 

load during creep testing resulted in some specimen contraction. 

- The density decreases of the silica specimens were a function of temperature. 

Additionally, some brands became less dense than others. The changes in density are a 

consequence of the above-described changes in mass and volume. The ranges of density 

decreases were 

≈ 1–2.75% at 1550°C, ≈ 2.25–4.0% at 1600°C, and ≈ 3.5–7.0% at 1650°C. The density of 

Harbison-Walker’s developmental brand changed the least of the six brands at all three 

temperatures. The ascending order of density for the other five brands varied with 

temperature. The density of Stella GGS decreased the most at 1550°C (≈ 2.75%), that for 

SI96AU decreased the most at 1600°C (≈ 4%), and that for Vega decreased the most at 

1650°C (≈ 7%). 

• The corrosion resistances of the six silicas were statistically equivalent when they were exposed 

to sodium carbonate at 1400°C (2550°F) for 24 h. The amount of recession increased linearly 

with temperature between 1250 and 1450°C for the Gen-Sil silica. 

7.2 Mullite 

Compressive creep testing of ten commercially available mullite refractories was performed at 1300 

to 1450o

C and at static stresses between 0.2 and 0.6 MPa. The mullite refractories were examined 

because they are used in borosilicate glass furnace crowns and superstructures as well as in sidewall 

applications. Additionally, despite their high cost (≈$500/ft3

), they are cheaper than other refractories 

such as chrome alumina (≈ $3000/ft3

) or fusion-cast alumina (≈$900/ft3), which are used as 

replacements for traditional silica refractories in harsh oxy-fuel environments. Corrosion resistance 

was also evaluated. In addition, measurements were made that tracked their dimensional stability, 

phase content, microstructure, and composition as a function of temperature and time. The techniques 

used for these characterizations and their respective analyses are described. 

56 
An intent of this study was to provide objective and factual results whose interpretations were left to 

the reader. The salient observations and conclusions are as follows. 

• The amount of compressive creep for the BP mullite, Durital S75, and Frimul F brands was found 

to be low at temperatures between 1300 and 1450o

C (2370 to 2640oF) and at compressive stresses 

between 0.2 and 0.6 MPa (29 to 87 psi). Creep rates were on the order of 10!11 s!1

 at the lower 

test temperature and on the order of 10!10 s!1

 at the higher test temperature. Only one specimen 

per condition was tested, so the authors were unable to statistically conclude that any one of the 

three brands had superior creep resistance to the others; however, if differences do indeed exist, 

then they are believed to be insignificant. 

• The amount of compressive creep for the GEM, HF 17, MU75AF, and ZED FM brands was 

found to be slightly higher than that for the three refractories listed above, but still low at 1300 to 

1450o

C and 0.2 to 0.6 MPa . Creep rates were on the order of 10!10 s!1 at the lower test 

temperature and on the order of 10!9

 s!1 at the higher test temperature. The GEM brand specimen 

showed higher amounts of creep than the other three brands in this group and a low stress 

exponent (0.4). This places its behavior closer to that described next for the brands exhibiting 

significant amounts of creep. Again, only one specimen per condition was tested so the authors 

were unable to statistically conclude that any one of the other three brands had superior creep 

resistance to the others; however, if differences do indeed exist, then they are believed to be 

insignificant. 

• The amount of compressive creep for the SL75AD, UFALA UCR, and ZED FMC brands was 

found to be significant at 1375 to 1450o

C and at 0.4 and 0.6 MPa. Creep rates were on the order 

of 10!9

 s!1 at the lower test temperature and on the order of 10!8 s!1 at the higher test temperature. 

All three of these brands exhibited low stress exponents (< 0.5) and changes in chemistry or 

microstructure due to the creep testing conditions. Only one specimen per condition was tested, 

so the authors were unable to statistically conclude that any one of the three brands had superior 

or inferior creep resistance to the others; however, if differences do indeed exist, then they are 

believed to be insignificant. 

• Durital S75 was found to have a stress exponent of 2.4, indicating that the rate-controlling 

mechanism may be grain boundary creep. This is opposed to diffusion-controlled or Coble creep, 

which is expected to be the rate-controlling mechanism in the brands that exhibited a stress 

exponent of approximately unity (Frimul F, HF 17, and MU75AF). 

• It is believed that the low-stress exponents seen for the BP mullite, GEM, SL75AD, UFALA 

UCR, ZED FMC, and ZED FM brands are due to a non-steady-state condition existing during the 

entire extent of testing. This condition is a consequence of contraction or time-hardening effects 

occurring during testing and of the samples never reaching a state of equilibrium. The result is 

measurement of creep due to the applied compressive stress coupled with contraction of the 

sample due to time-hardening effects. Therefore, the assumption of a single deformation 

mechanism being active at all temperatures and stresses is not valid, and multilinear regression 

falsely yields extremely low stress exponent values. 

• The brands that possessed high levels of matrix porosity (HF 17, UFALA UCR, and ZED FMC) 

all showed compaction of the microstructure due to creep testing. This is believed to be the 

reason for the poor creep resistance of the UFALA UCR and ZED FMC brands. In addition, the 

UFALA UCR sample was found to possess a fine grain size, which could have contributed to the 

increased creep rate. 

• SL75AD was found to be composed of andalusite grains containing free quartz, which was 

converted to cristobalite during firing. Further, the andalusite grains are fully converted to mullite 

during creep testing, and the amount of glassy phase increased. These events in combination may 

lead to the brand’s poor creep resistance. 

57 
• Activation energies for the various refractory brands ranged from 50 to 223 kcal/mol. Although 

no direct correlation between the magnitude of the activation energy and the amount of creep 

exhibited by a particular brand could be drawn from the data, it is known that materials 

possessing a high activation energy have greater sensitivity to temperature than those possessing 

a lower activation energy. Also, the values for many of the calculated activation energies 

correspond to the activation energy of vitreous silica. This indicates that the glassy phases in 

these brands affect the overall creep behavior of the material. 

• The distinction of a brand being fused-grained or non-fused-grained mullite did not appear to 

have a bearing on the creep resistance of the individual brands. 

• Room-temperature thermal conductivity values were similar for all ten brands of mullite with the 

average being on the order of 3.1 W/mK [21.5 (Btu*in.)/(h*ft2

*°F)]. Frimul F was found to have 

the highest thermal conductivity at just below 3.5 W/mK [24.3 (Btu*in.)/(hr.*ft2

*°F)]. SL75AD, 

UFALA UCR, and ZED FM were found to have the lowest thermal conductivities at around 2.5 

W/mK [17.4 (Btu*in.)/(hr.*ft2

*°F)]. 

• Eight of the ten refractory brands showed less than 0.2% change in mass and less than 0.6% 

change in volume due thermal aging in the absence of an applied load. The two exceptions, ZED 

FM and ZED FMC, exhibited extraneous amounts of change (≈8.5%) in both mass and volume 

due to thermal exposure. All samples showed a less than 0.5% change in bulk density. 

• The corrosion rates of the ten mullite brands could not be determined through the use of ASTM 

C987, but it was determined visually that only minimal amounts of recession occurred due to 

exposure as defined by the ASTM lid test. 

7.3 Fusion-Cast Alumina 

The creep and corrosion performances of the two brands of M-type fusion-cast alumina refractories 

(Corhart Refractories of Louisville, Kentucky, and Monofrax, Inc., of Falconer, NewYork) were 

examined. All specimens were fabricated from regular fusion-cast alumina blocks of approximate 

dimensions 0.60 H 0.50 H 0.30 m (24 H 20 H 12 in.). Key findings are summarized below.

• As expected, both the density and porosity of the as-cast refractory blocks were highly consistent. 

The phase profiles for the α- to β-alumina ratios also showed similar trends to those seen in the 

density/porosity profiles. XRD showed that both brands consisted of only peaks identifiable as 

corundum (PDF# 10-0173) or sodium aluminum oxide (PDF# 31-1263). Closer examination of 

the XRD patterns showed that there were some differences in ratios of α- to β-alumina between 

the two brands and between the bulk and columnar microstructures. The microstructures of the 

two brands of refractory were also comparable with a columnar region existing in the outer 11⁄2 in. 

of the blocks and then transitioning to a more equiaxed structure of α-alumina grains floating in a 

β-alumina matrix as sampling moves into the center of the block. Average grain sizes were found 

to be on the order of 172 μm and 82 μm for the Corhart bulk and columnar regions, respectively. 

Average grain sizes were found to be on the order of 231 and 79 μm for the Monofrax bulk and 

columnar regions, respectively. 

• With exception of the Monofrax bulk material, all of the elastic modulus values obtained by static 

testing (cyclic compression) at room temperature were similar, ranging from 109 to 113 GPa. 

Contrary to this, the Monofrax bulk material showed a significantly lower modulus at room 

temperature (72 GPa). These values are substantiated by dynamic values obtained by resonant 

ultrasound spectroscopy. Although the room-temperature value is vastly different and the 

magnitude of the modulus values at elevated temperatures for the bulk Monofrax is higher, the 

general trend of decreasing modulus with increasing temperature holds as it does in the other 

microstructures. Furthermore, the modulus value of the bulk Monofrax at 1400°C is approaching 

that of the other sample types, indicating that at even higher temperatures (like those experienced 

58 
by these materials in service), the modulus values for the four microstructures may be virtually 

the same. 

• Thermal expansion values obtained for the two brands of fusion-cast alumina by the dual-rod 

dilatometry technique were found to be different from those published in the manufacturers’ 

literature, but both materials showed similar values of expansion. The Corhart material was found 

to have lower expansion (by ≈ 30%) than published; while the Monofrax material was found to 

have higher expansion (by ≈ 6%) than published. Based on the published values and the 

experimental results, it would seem that both materials have an experimental thermal expansion 

of around8.0 H 10!6

/°C over the 500 to 1200°C temperature range. 

• Room-temperature thermal conductivity measurements produced by the transient plane source 

(“hot disk”) technique show the Monofrax bulk material to possess a lower thermal conductivity 

(on the order of 6.00 to 6.25 W/mK) than the Corhart bulk material (mean value of 6.78 W/mK). 

The columnar structures in the Corhart block show a range of thermal conductivities between 

6.50 and 7.50 W/mK. This is similar to the values obtained for the longitudinal columnar 

measurements made on the Monofrax block (average = 7.10 W/mK). Values on the transverse 

columnar structure of the Monofrax block (average = 5.82 W/mK) were lower than those of the 

two bulk microstructures and lower than those of the Corhart columnar and Monofrax 

longitudinal columnar microstructures. These results indicate that the columnar structure is not as 

well oriented in the Corhart blocks as in the Monofrax blocks. 

• Laser-flash thermal diffusivity/conductivity values show a gradual trend moving from the low- to 

elevated-temperature data. A change in thermal conductivity from decreasing to increasing slope 

was seen, as is common in materials where the mechanism controlling thermal conductivity 

changes with temperature. It is expected that the mechanism controlling the sign of the curves’ 

slope changes from conduction to radiation. Further, this behavior is characteristic of a material 

that contains microcracking. The presence of microcracks initially lowers the thermal 

conductivity of the material, but as the material is exposed to elevated temperatures, the “healing” 

of the microcracks will cause thermal conductivity values to increase. 

• The Monofrax material was found to be slightly more creep-resistant than the Corhart material at 

1450°C and 0.6 and 1.0 MPa. The advantage in creep resistance is only slight (0.03% strain at the 

most extreme) and can be considered negligible in relation to the magnitude of deformation it 

would cause in a furnace crown. Behavior of the 0.8-MPa samples is similar with both brands of 

material, showing an initial period of increased strain accumulation followed by a sustained 

period of decreasing strain accumulation. The final amount of accumulated strain for both brands 

is also similar, causing the overall trend of the 0.8-MPa curves to show considerably less total 

accumulated strain than both the 0.6- and 1.0-MPa curves. 

• Similar behavior occurs at 1550°C, where the Corhart material initially accumulates strain at a 

more rapid rate but the total accumulated strains for both materials are nearly identical, further 

negating any advantage the Monofrax material may hold over the Corhart material. 

• At 1650°C, the behavior of both brands is nearly identical for both 0.6- and 0.8-MPa stresses; all 

of the curves overlap one another, and neither material shows even a slight advantage. All three 

of the 1.0-MPa curves display traditional creep behavior, which is not seen in any of the other 

bulk test conditions. 

• Both columnar orientations show much less accumulated strain than the bulk samples under 

similar stresses. This is due to the elongated needle-like grains, which are oriented perpendicular 

to the direction of the applied compressive load, acting to inhibit creep in the transverse columnar 

samples. In the longitudinal columnar case, it is expected that the unrecoverable expansion effect 

plays a role in the reduced strain accumulation. The columnar samples at 1650°C show less 

accumulated strain than the bulk sample under similar stress, but the behavior between the bulk 

and columnar samples is more similar. Both microstructures show traditional creep behavior 

under 1.0 MPa of stress and much larger amounts of accumulated strain than the bulk samples at 

0.6 and 0.8 MPa. 

59 
• The three bulk 1650°C/1.0-MPa samples all have a similar creep rate; the rate for the Corhart 

material is on the order of 5 H 10!6

 h!1, and that for the Monofrax material is on the order of 6 H

10!6

 h!1. The columnar structures at 1650°C/1.0 MPa have a lower creep rate, on the order of 4.25 

to 4.50 H 10!6

 h!1. Creep rates for the other samples were not obtainable because nontraditional 

creep behavior occurred. Stress exponents and activation energies could not be calculated from 

this data because steady-state creep rates were only obtained at one stress and temperature. 

• The combined expansion/creep plots predict a greater creep resistance in the Monofrax bulk 

material at 1450°C (better by ≈ 0.20%), but a greater creep resistance in the Corhart bulk material 

at 1550°C (better by ≈ 0.15%) and at 1650°C (better by ≈ 0.10%). These predictions may not be 

an accurate portrayal of the creep behavior exhibited by these materials. Behavior exhibited in 

service would still be represented by the measured creep curves. 

• Unrecoverable expansion was seen in all materials, and is thought to be connected to the 

nontraditional creep behavior seen in the compressive creep testing at temperatures up to 1650°C 

and loads up to 0.8 MPa. As the temperature increases, the amount of unrecoverable expansion 

also increases such that when samples are heated and immediately cooled with no hold time at 

temperature, no unrecoverable expansion occurs. The mechanism causing the unrecoverable 

expansion effect occurs during the prolonged exposure of the sample to high temperatures and 

therefore would be of great importance during creep testing and during service of a brick in a 

furnace campaign. The occurrence of this unrecoverable expansion is thought to be caused by 

thermal expansion hysteresis due to microstresses or microcracking. 

• The data generated in this work for fusion-cast alumina refractories are unique in that very little 

data of that kind has been generated. The behavior exhibited by this material is highly transient 

due to the non-steady-state nature of the material (it fails to reach equilibrium). Consequently, the 

generated creep curves could not be represented by the empirical NBA creep equation. Instead, an 

empirical fit utilizing an exponential expansion series with multiple time constants was used to 

represent the behavior exhibited by the fusion-cast alumina refractories. The same exponential fit 

used to model the raw creep curves was also utilized to describe this composite behavior. 

• Several samples were analyzed after being crept, and results were compared with those 

previously obtained from virgin samples. Analysis included thermal conductivity, optical 

microscopy, and SEM/EDS chemical analysis. The thermal conductivity values for the Monofrax 

post-creep cores were similar to those obtained earlier from the virgin cores, which had thermal 

conductivities in the range of 5.9 W/mK. The value obtained from the Corhart post-creep core 

was significantly lower than the average value of 6.61 W/mK found earlier for the Corhart bulk 

material. Initial optical microscopy and SEM/EDS analyses were performed on Corhart and 

Monofrax bulk samples exposed to 1550°C and 1.0 MPa. This work was conducted to determine 

whether any changes in the microstructure or chemistry could be detected at that intermediate 

condition. This analysis was then used to determine whether analysis needed to be performed on 

samples at the lower temperatures and stresses. No discernible deformation and only slight 

chemical changes were seen. Subsequent optical microscopy and SEM/EDS analyses were 

performed on Corhart and Monofrax bulk samples exposed to the extreme conditions of 1650°C 

and 1.0 MPa. The structure of the Corhart sample was found to resemble the virgin structure, but 

the crept Monofrax structure displayed a smaller α-alumina grain size and a tighter α-alumina 

grain structure with more equiaxed angular grains than previously observed in the virgin 

structure. More severe chemical changes were also seen between the virgin and crept samples. 

• Considering that the behavior of fusion-cast alumina in service would be represented by the 

measured creep curves, any model that is used to capture its true behavior must incorporate both 

the effect of the expansion and the effect of creep. This means that a model is needed that can 

handle the high degree of nonlinearity seen in the measured creep curves generated for fusioncast alumina while also being sensitive to the transient nature of the data due to the lack of an 

equilibrium state. 

60 
7.4 Fused Spinel 

The fused spinel materials had negligible creep rates up to 1650°C at 8.5MPa of stress. The bonded 

spinel, surprisingly, crept substantially compared to the fused material. However, these findings must 

be compared with other common crown materials, the total creep of the bonded material is less than 

0.5% over 100 h at 1550°C and 0.3 MPa, which are fairly standard temperature and pressure for 

testing refractory materials The creep of silica materials is approximately 0.4% over 100 h at 1450°C 

and 0.3 MPa. The spinel materials did not react with soda in the exposure tests. There was some 

limited reaction with silica and calcia from the gasifier atmosphere. Thermal conductivity was 

determined by laser flash and was found to decrease as temperature increases from 14 (100°C) to 8 

(700°C) W/mK for the fused spinel and 9.5 (100°C) to 4 (1300°C) for the bonded spinel. A model of 

the glass tank crown was developed, which takes into account nonlinear changes in the refractory 

material due to time, temperature, and corrosion. 

61 

8. Recommendations 

• To truly model the deformation behaviors of fusion-cast alumina and silica refractories as a 

function of temperature, time, and stress, the composite effects of unrecoverable expansion (aging 

behavior) and creep response must be known. Given that these processes may be highly 

nonlinear, the traditional equations, which assume that steady-state deformation rates are 

dominant, are not appropriate. New models must be derived from basic theoretical principles that 

can describe the nonlinear, transient, nonsteady equilibrium behavior observed in this work. 

• Creep testing needs to include simultaneous measurement of specimen elongation/contraction 

occurring in the absence of an applied stress. Such testing could be conducted using a system 

consisting of furnace of sufficient size to accommodate multiple mechanical load trains. During 

testing one load train would be used to monitor elongation/contraction in the absence of any 

appreciable stress. The stress range employed with the other load stations must be large enough to 

promote significant and measurable changes in the deformation response. It is recommended that 

the stress range cover at least one order of magnitude (e.g., 5 to 50 MPa). Testing times must also 

be much longer (>1000 h) to ensure that representative deformation data are generated. A second 

approach for distinguishing creep from aging effects is to use the stress relaxation technique [23] 

for measuring creep response. Because this method is able to generate an entire creep-rate vs 

stress curve in a matter of a few hours, one could use it to measure creep behavior at periodic 

intervals during a standard aging test. Thus a single sample could be used to assess both aging 

and creep behavior. 

• To facilitate model development, the microstructural changes responsible for 

elongation/contraction in the absence of an applied stress must be identified. 

• Creep testing and physical characterization of salvaged fusion-cast alumina refractory bricks from 

oxy-fuel furnaces should be carried out. Results from such work could be compared with results 

generated in this work and would be useful in validating any finite-element models generated 

from the data collected in this work. 

• Work needs to be pursued in the area of nonlinear finite element analysis of furnace crowns. This 

kind of analysis would provide a better prediction of furnace life and long-term refractory 

behavior. 

• Technical efforts should be undertaken to measure and model the composite effects of 

unrecoverable expansion (aging behavior) and creep response. Given that these processes may be 

highly nonlinear, the traditional equations, which assume that steady-state deformation rates are 

dominant, are not appropriate. New models must be derived from basic theoretical principles that 

can describe the nonlinear, transient, nonsteady equilibrium behavior observed in this work. 

Creep testing needs to include simultaneous measurement of specimen elongation/contraction 

occurring in the absence of an applied stress. Such testing could be conducted using a system 

consisting of a furnace of sufficient size to accommodate multiple mechanical load trains. During 

testing one load train would be used to monitor elongation/contraction in the absence of any 

appreciable stress. The stress range employed with the other load stations must be large enough to 

promote significant and measurable changes in the deformation response. It is recommended that 

the stress range cover at least one order of magnitude (e.g., 5 to 50 MPa). Testing times must also 

be much longer (> 1000 h) to ensure that representative deformation data are generated. A second 

approach for distinguishing creep from aging effects is to use the stress relaxation technique for 

measuring creep response. Because this method is able to generate an entire creep-rate vs stress 

curve in a matter of a few hours, one could use it to measure creep behavior at periodic intervals 

during a standard aging test. Thus a single sample could be used to assess both aging and creep 

behavior. 

63 
• To facilitate technology transfer and commercialization, software tools should be developed that 

allow users to readily link the key properties to finite element codes used in furnace construction. 

This software would allow one to quickly assess the effects of specific materials upon long-term 

thermal and structural reliability. It should also include a routine that allows the user to optimize 

performance of a given material by changing the material properties; these altered properties 

would serve as new targets for those involved in materials development. 

• Any follow-on effort must also address new materials. 

64 
9. References 

[1] Crown Refractories for Glass Manufacturing with Oxy-Fuel Combustion, Teltech Research 

Services, Minneapolis, Minnesota (1996). 

[2] M. Gridley, “Philosophy, Design, and Performance of Oxy-Fuel Furnaces,” Ceram. Eng. Sci. 

Proc., 18 1–14 (1997). 

[3] R. W. Schroeder, “Development of Oxy-Fuel Technology in the Glass Industry,” The Glass 

Researcher 8 (1), 1–4 (1998). 

[4] A. A. Wereszczak, T. P. Kirkland, G. A. Pecoraro, and R. A. New, “Compressive Creep 

Behavior of Fusion-Cast Alumina Refractories,” pp. 401–406 in Ceramic Transactions, Vol. 82, 

Advances in Fusion and Processing of Glass II, ed. A. G. Clare and L. E. Jones, American 

Ceramic Society, Westerville, Ohio (1998). 

[5] A. A. Wereszczak, T. P. Kirkland, and W. F. Curtis, “Compressive Creep Resistance of MgO 

Refractories at Temperatures ≥ 1400°C (2550°F),”pp. 407–412 in Ceramic Transactions, Vol. 

82, Advances in Fusion and Processing of Glass II, ed. A. G. Clare and L. E. Jones, American 

Ceramic Society, Westerville, Ohio (1998). 

[6] A. A. Wereszczak, T. P. Kirkland, and W. F. Curtis, “Creep of CaO/SiO2 - Fluxed MgO 

Refractories,” J. Mater. Sci (1998) in press. 

[7] Glass: A Clear Vision for a Bright Future, U.S. Dept. of Energy, Office of Industrial 

Technologies, Advanced Industrial Materials Program (January 1996). 

[8] Refractories for Industrial Processing: Opportunities for Improved Energy Efficiency, U.S. 

Dept. of Energy, www.eere.energy.gov/industry/ imf/pdfs/refractoriesreportfinal.pdf. 

[9] “Standard Practice for Vapor Attack on Refractories for Furnace Superstructures,” 

ASTM C987, Vol. 15.01, American Society for Testing and Materials, West Conshohocken, 

Pa. (1998). 

[10] A. A. Wereszczak, M. Karakus, K. C. Liu, B. A. Pint, T. P. Kirkland, and R. E. Moore, 

Compressive Creep Performance and High Temperature Dimensional Stability of Conventional 

Silica Refractories, ORNL/TM-13757, Oak Ridge National Laboratory (1999). 

[11] J. G. Hemrick, A. A. Wereszczak, M. Karakus, K. C. Liu, H. Wang, B. A. Pint, T. P. Kirkland, 

and R. E. Moore, Compressive Creep and Thermophysical Performance of Mullite 

Refractories, ORNL/TM-2002/84, Oak Ridge National Laboratory (2002). 

[12] J. G. Hemrick and A. A. Wereszczak, “Compressive Creep of Fusion-Cast Alumina as a 

Function of Microstructure,” Glass Industry (November 1999). 

[13] J. G. Hemrick, “Creep Behavior And Physical Characterization Of Fusion-Cast Alumina 

Refractories,” a Dissertation Presented to the Faculty of the Graduate School of the University 

of Missouri-Rolla in Partial Fulfillment of the Requirements for the Degree of Doctor of 

Philosophy in Ceramic Engineering (2001). 

[14] “Standard Classification of Silica Refractory Brick,” ASTM C416, Vol. 15.01, American 

Society for Testing and Materials, West Conshohocken, Pa. (1998). 

[15] S. Shin and O. Buyukozturk, Material Property Development for Refractories, US DOE Report 

ORNL/Sub/79-07862/02 (1990). 

[16] F. H. Norton, The Creep of Steel at High Temperature, McGraw Hill, New York (1929). 

65 
[17] M. Karakus and R. E. Moore, “CLM—A New Technique for Refractories,” Ceramic Bulletin

77 55–61 (1998). 

[18] S.E. Gustafsson, “Transient Plane Source Techniques for Thermal Conductivity and Thermal 

Diffusivity Measurements of Solid Materials,” Review of Scientific Instruments. 62 (3), 797–

804 (1991). 

[19] “Standard Test Method for Thermal Diffusivity of Solids by the Flash Method,” ASTM E1461-

92, Vol. 15.01, American Society for Testing and Materials, West Conshohocken, Pa. (2000). 

[20] G. Hetherington, J. H. Jack, and J .C. Kennedy, “The Viscosity of Vitreous Silica,” Physics and 

Chemistry of Glasses 5, 130–136 (1964). 

[21] F. E. Wagstaff, “Crystallization and Melting Kinetics of Cristobalite,” Journal of the American 

Ceramic Society 52, 650–654 (1969). 

[22] M. F. Doerner and W. D. Nix, “A Method for Interpreting Data from Depth-Sensing 

Indentation Instruments,” Journal of the Materials Research Society 1 (4) (1986). 

[23] M. G. Jenkins and M. K. Ferber, “Determination of High-Temperature Creep Parameters in 

Structural Ceramics From Constant-Displacement, Stress-Relaxation Tests,” pp. 94–101 in 

Proceedings of the 1991 SEM Spring Conference on Experimental Mechanics, Milwaukee, 

Wisconsin, June 9–13, 1991, Society of Experimental Mechanics, Bethel, Connecticut (1991). 

66 
67 

Appendix: Publications 

A. A. Wereszczak, J. Heide, and T. P. Kirkland, “High Temperature Deformation of an AZS 

Refractory,” Advances in Fusion and Processing of Glass II, 379–384. 

A. A. Wereszczak, T. P. Kirkland, and W. F. Curtis, “Creep of CaO/SiO2–Containing MgO 

Refractories,” Journal of Materials Science 34 (1999): 215– 227. 

A. Wereszczak, H. Wang, M. Karakus, W. Curtis, V. Aume, and D. VerDow, “Postmortem Analyses 

of Salvaged Conventional Silica Bricks from Glass Production Furnaces,” Glastech. Ber. Glass 

Sci. Technol. 73 (2000), No. 6: 165–174. 

J. G. Hemrick, C. W. Kistler Jr., A. A. Wereszczak, and M. K. Ferber, “Thermal Conductivity 

of Alumina Measured with Three Techniques,” 1–13.