From: Fairchild, Susan <Fairchild.Susan@epa.gov>

Sent: Monday, March 11, 2013 9:16 AM

To: 'Higgins, Cortney'

Cc: Barnett, Keith; Okoye, Winifred; VanLare, Paula

Subject: OMB requested documents

Attachments: non-CBI furnace data for OMB request.pptx; Non-CONFIDENTIALChromeSummarySheetKCK.xlsx; Industry conversation notes.docx; INDUSTRY CONVERSATION REGARDING THE 

SOURCE OF HIGH HEX CHROME EMISSIONS.docx; MCCULLOUGH Notes_Redacted.pdf; MW 

Industry conversation notes.pdf; Wiley Online Library_ hi Cr fused cast ref chap 23 ORDER 

35.pdf; 13 August 2012 Telephone conversation with Carlos E. Davisdocx.docx; Hex Chrome in 

Refractories, Volume 34, Numbers 9-10 - SpringerLink.pdf; new high chrome fused cast 

refractory for use in contact with highly corrosive glasses.pdf; Owens Corning, Delmar Plant 

oxyfuel replacement of other furnaces.pdf; Re_ Fw_ WF - PM and Cr Graphs (NEW)-

CWLEE.pdf; Saint-Gobain SEFPRO.pdf; SG Chromium refractories.pdf; Oak Ridge DOE June 

2006 oxyfuel.pdf; DOE refractoriesreportfinal JAN2005.pdf

Cortney,

I am forwarding to you several documents that I think will address your comments to the 

preamble. 

The power point file has is a graph of chrome emissions for each furnace in the source category, grouped by 

both major and area source status. The other charts in this file were provided by industry during the site visits 

and provide general (i.e., not specific) designs of different furnace types. Specific furnace information is 

regarded as CBI, and we could not provide that in addressing some of your questions.

Also you requested 

The DOE study conducted with Oak Ridge National Laboratories

Information regarding our conversations with the plant manager of the highest chrome emitter

Industry discussions regarding the use of chromium refractories

I have also attached a few other documents that are in the docket and that we used in development of this 

proposed rule

Susan Fairchild

Senior Environmental Scientist

(919) 541-5167

USPS Address:

OAQPS/SPPD/MMG

Mail Code D 243-04

Research Triangle Park, NC 27711
13 August 2012 Telephone Conversation with Carlos E. Davis, Certainteed, Kansas City, KS

Susan Fairchild, project lead at EPA, contacted Carlos Davis of Certainteed to get clarification of Mr. 

Davis’ statements in 2011 regarding the chromium content of the K2 furnace refractory versus the 

information submitted by Certainteed in the 114 information response in 2011-12.

Mr Davis reiterated that over 90% of the refractory is chrome bearing, or what EPA may refer to as 

‘chromium compounds’, but that only a percentage of the chromium compounds are actually the 

element chromium. (For instance, Cr2O3 is chromite; the element chromium is 40% of the mineral 

chromite). Thus, when the statement on the 114 response of the chromium-content of the refractory 

was given by the facility, this should be interpreted as the percentage of the element chromium in the 

total chromium bearing constituents of the refractory product. This their interpretation of the 114 

request questions regarding chromium content of the furnace refractory.

Regarding the K2 furnace, if each brick of the furnace is examined, over 90% of the refractories would 

contain chrome. When you look at each piece, however, the chrome content will vary. The varying 

percentage of chrome in the refractory is based upon it’s purpose. Specifically, at the metal / glass linethis area has to be very resistive due to the corrosivity and errosivity of the glass melt in the furnace. 

The metal line is at the top of the ‘cool’ of the glass. The superstructure (entire furnace) will also have 

chromium containing refractories, but the chromium content above the metal/glass line is generally 

lower in chrome content. 

Regarding the K2 furnace, that furnace will not re-start until it is rebuilt. The nature of that 

reconstruction will be identical to the furnace at Chowchilla, which is a different design than the K2 is 

now. The Chowchilla furnace can meet the limits; in fact it emits chromium at about 1% of the proposed 

limit for chromium compounds. 

Regarding the K1 furnace, that furnace meets the proposed limits now.

If a furnace is built using chromium refractories, it will last longer. Therefore it will cost less when

amortized over the life of the furnace; the longer the furnace lasts, the less the total amortized cost. 

High chromium refractories are needed anywhere that contacts glass to achieve this furnace life. This is 

a minimum of a 10-year vs. a 6-year furnace design. 

Regarding the new K2 furnace design, the only way for K2 to meet the proposed limit is to reconstruct. 

The new design will still have high chrome refractories in contact with the glass. Look at the emissions 

of chromium and the Chowchilla furnace: this has an emissions rate at about 1% of the proposed limit. 

The photographs on the cover are an example of an aluminum furnace featured 

on the Harbison Walker web page (top), cathodoluminescence of a fusion-cast 

spinel refractory (center), and cathodoluminescence of a fusion-cast alumina 

refractory (bottom). 

1 

Refractories for Industrial Processing: 

Opportunities for Improved Energy 

Efficiency 

James G. Hemrick,* H. Wayne Hayden,† Peter Angelini,* 

Robert E. Moore,§ and William L. Headrick§

*Oak Ridge National Laboratory

†

Metals Manufacture, Process, and Controls Technology, Inc. 

§

R. E. Moore and Associates 

January 2005 

Prepared for the DOE-EERE Industrial Technologies Program 

Table of Contents 

List of Figures and Tables ............................................................................................................... v

Executive Summary ...................................................................................................................... vii

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

1. Introduction............................................................................................................................. 1

2. Current Refractory Status ....................................................................................................... 1

2.1 Usage in Industry......................................................................................................... 4

2.2 Current R&D Status..................................................................................................... 4

3. Estimated Metrics for Overall Energy Savings Opportunities ............................................... 5

4. Process Technology and Performance Barriers ...................................................................... 7

4.1 Kilns and Calciners...................................................................................................... 7

4.2 Reverberatory Furnaces ............................................................................................... 8

4.3 Boilers and Reactor Systems ..................................................................................... 10

4.4 Gasifiers..................................................................................................................... 13

5. Overall R&D Barriers and Pathways for Refractories.......................................................... 15

6. Specific R&D Areas with Barriers and Pathways ................................................................ 17

6.1 Improved Castable and Monolithic Refractory.......................................................... 17

6.2 Modified Fiber Materials ........................................................................................... 17

6.3 Surface Modification (Coatings, Physical or Thermal Alteration) ............................ 17

6.4 Synthetic Aggregates and Powders............................................................................ 18

6.5 High-Temperature Properties of Refractories............................................................ 18

6.6 In Situ and Graded Refractories................................................................................. 19

6.7 Sensors for Materials Performance............................................................................ 19

7. References............................................................................................................................. 22

Appendix A: Discussion of Refractory Usage by Process Industries........................................... 26

A.1 Aluminum .................................................................................................................. 26

A.2 Cement and Lime....................................................................................................... 30

A.3 Chemical and Petrochemical...................................................................................... 33

A.4 Forest Products .......................................................................................................... 35

A.5 Glass .......................................................................................................................... 37

A.6 Heat-Treating and Annealing..................................................................................... 40

A.7 Iron and Steel............................................................................................................. 41

A.8 Metal Casting............................................................................................................. 44

A.9 References for Appendix A ....................................................................................... 48

Appendix B: Rotary Kiln Analysis Schematic, Conditions, and Results..................................... 50

Appendix C: Reverberatory Furnace Schematics......................................................................... 56

Appendix D: Reverberatory Furnace Analysis Conditions and Results........................................ 58

Appendix E: Boiler Schematic ...................................................................................................... 62

iii 

List of Figures and Tables 

Figures 

ES.1. Temperature regimes for industrial applications of refractories ..................................... vii

1. Long-term energy benefit opportunities estimated for the four analyzed 

furnace types ...................................................................................................................... 6

2. Improving refractory materials for improved energy efficiency ..................................... 16

B.1. Longitudinal section of a typical rotary kiln .. ................................................................. 50

C.1. Reverberatory furnace schematic ..................................................................................... 56

C.2. Typical reverberatory aluminum-melting furnace ........................................................... 57

E.1. Boiler furnace sections .................................................................................................... 62

Tables 

1. Types of furnaces utilized in American industry ............................................................... 2

2. Refractory barriers for furnaces utilized in American industry.......................................... 3

3. Energy analysis by industry with estimated short-term and long-term

energy savings opportunities from improved refractories.................................................. 5

4. Total effect of improved refractories on rotary kilns ......................................................... 8

5. Total effect of improved refractories on reverberatory furnaces...................................... 10

6. Steam system assessment tool single header analysis results for improved 

insulation for improved boiler efficiency ......................................................................... 11

7. Steam system assessment tool single header analysis results for improved 

pipework insulation for improved steam transfer systems ............................................... 11

8. Potential energy savings from improved boiler efficiency due to improved 

refractories and improved transfer pipe-work insulation ................................................. 12

9. Renewable energy generation potential (billion kWh/year) from black liquor 

in the pulp and paper industry in the southeastern United States..................................... 15

10. Recommended R&D pathways for advanced refractories ............................................... 20

A.1. Feedstock for energy use in the cement industry, 1993-2001 .......................................... 32

A.2. Estimated energy use in melting and refining .................................................................. 38

A.3. Conversion of North American glass melters to oxy-fuel firing as 

of 1998 and as of 2004 ..................................................................................................... 39

A.4. Process and energy data for U.S. metals casting industry ................................................ 45

v 

Executive Summary 

Refractory materials with properties suitable for high-temperature applications are used 

throughout U.S. industry as insulation and/or containment vessel linings in high-temperature and 

corrosive environments. Indeed, it would be difficult to identify an industrial process that does 

not use refractory materials in one aspect or another. These materials must not only be capable of 

performing these tasks at elevated temperatures, but may also be called upon to bear mechanical 

loads and transfer heat. Refractories are characterized by two aspects that qualify them as 

strategic materials in support of American industry: (1) their ability to contain heat, chemicals, 

and melts; and (2) the crosscutting nature of their utility. 

Figure ES.1 shows the ranges of usage temperatures for numerous industrial applications of 

refractories. This report and the attached appendices describe the current furnace technologies 

used by U.S. industries with specific highlighted analysis focused on four furnace types—kilns/ 

calciners, reverberatory furnaces, boilers and reactor systems, and gasifiers—which are common 

to multiple industries and are representative of areas where significant energy improvement could 

be made through future refractory improvement. Analyses of these four types of furnaces alone 

suggest potential energy savings of over 400 trillion Btu/year. In addition to identifying areas of 

opportunity, this report describes possible crosscutting research and development (R&D) 

pathways for overcoming refractory-related performance barriers which could lead to improved 

energy efficiency through both the alteration of current materials and practices and the 

development of new materials, furnace designs, or redesigned processes of varying sizes. 

Fig. ES.1. Temperature regimes for industrial applications of refractories. 

vii 

Abbreviations and Acronyms 

BLS black liquor solids 

BOF basic oxygen furnace 

Btu British thermal unit 

DOE U.S. Department of Energy 

EAF electric arc furnace 

EERE Energy Efficiency and Renewable Energy (DOE) 

FCCU fluidized catalyst cracking unit 

ITP Industrial Technologies Program (DOE) 

MECS Manufacturing Energy Consumption Survey 

ORNL Oak Ridge National Laboratory 

R&D research and development 

TBtu trillion Btu 

ix 

Refractories for Industrial Processing 

1. Introduction 

The Industrial Technologies Program (ITP), part of the Department of Energy’s Office of Energy 

Efficiency and Renewable Energy (DOE-EERE), seeks to improve the energy intensity of the 

U.S. industrial sector through a coordinated program of research and development (R&D), 

validation, and dissemination of energy efficiency technologies and operating practices. To 

achieve its mission, ITP partners with industry, equipment manufacturers, and other stakeholders 

to develop technologies that enable industry to use less energy per unit of output. Refractories are 

a class of materials of critical importance to manufacturing industries with high-temperature unit 

processes. This study describes industrial refractory applications and identifies refractory 

performance barriers to energy efficiency for processing. The report provides recommendations 

for R&D pathways leading to improved refractories for energy-efficient manufacturing and 

processing. 

2. Current Refractory Status 

There are many different types of furnaces used in industries in the United States, each with 

different refractory needs and challenges. A discussion of the refractory issues for several energyintensive processing industries can be found in Appendix A. Fifteen classifications of furnaces 

with several common issues can be identified. Typical industrial applications of these fifteen 

furnace types are shown in Table 1. Refractory challenges and problems related to these furnace 

types, as reported by industry personnel and discussed in detail in Appendix A, are shown in 

Table 2. 

Of these fifteen furnace types, four are highlighted in this document: kilns/calciners, 

reverberatory furnaces, boilers and reaction systems, and gasifiers. These were selected because 

they are common to multiple industries and representative of areas where significant energy 

savings [on the order of trillions of Btu (TBtu)/year] could be realized through future refractory 

improvement. Although emphasis was placed on blast furnaces, basic oxygen furnaces (BOFs), 

and other furnace types found in the steel industry, we have chosen not to focus on these furnaces 

because of their design practice of intentionally leaking energy so as to induce a steep thermal 

gradient through the lining. Additionally, BOF refractories are known to last thousands of heats, 

and little improvement is expected in improving refractories for these applications. 

Analysis and discussion will address refractory improvements leading to higher energy efficiency 

and improved product yield for each of these four furnace types. Discussion will focus on the 

energy-intensive industries of the United States including aluminum, cement and lime, chemicals 

and petrochemicals, forest products, glass, heat treatment and annealing, iron and steel, and metal 

casting. 

1 
2 

Table 1. Types of furnaces utilized in American industry 

Furnace type 

Industrial furnace applications 

Aluminum Cement & lime Chemical/ petro­ chemical Forest products Glass Heat treat Iron & steel M

cas

Blast furnace X Basic oxygen furnace X Reverberatory furnace X 

X 

X X 

Multiple hearth Ladle/tundishe/trough X 

X 

X X 

Electrolytic smelting X Kilns X X X X X X Electric arc furnace (EAF) X Cupola X Heat treat X X X X X Still X Reactor/reformer X Boiler X X X X X Gasifier X X Heat exchanger X X X X X X X X a Key to “Other”: 

1. Lead 6. Mining 2. Copper 7. Oxide melting 3. Molten metal handling 8. Various metal products 4. Magnesium 9. Organic and inorganic products 5. Ceramics 10. Virtually all industries 
Table 2. Refractory barriers for furnaces utilized in American industry 

Furnace type 

Refractory challenges 

Abrasion 

Caking/ 

agglomeration 

Contamination/ 

fouling 

Corrosion/ 

wear Erosion 

Mechanical 

degradation 

Thermal 

insulation 

m

de

Blast furnace X 

X 

X X 

Basic oxygen furnace X X X X X Reverberatory furnace X X X X Multiple hearth X 

X 

X 

Ladle/tundishe/trough X X X X X Electrolytic smelting X X X X Kilns X X X X X X X Electric arc furnace (EAF) X X X Cupola Heat treat Still X 

X 

X 

Reactor/reformer X 

X 

X 

Boiler X X X X X X Gasifier X 

X 

X X X 

Heat exchanger X 

X 

X 

3 
Refractories for Industrial Processing 

2.1 Usage in Industry 

Although refractories are used in most, if not all, of U.S. industries, the iron and steel industry 

consumes 70% of the refractories produced globally, while the cement and lime industry 

consumes 7%, the ceramics industry 6%, the glass industry 3–4%, and the oil industry about 4%.1 

In 1998, a survey was conducted by the U.S. Advanced Ceramics Association and Oak Ridge 

National Laboratory (ORNL) to determine opportunities for advanced ceramics to meet the needs 

of manufacturing industries.2

 These industries included aluminum, chemicals and petrochemicals, 

forest products, glass, metals casting, and steel. Additionally, the supporting industry of heat 

treating and annealing was considered. From this survey, a number of critical areas were 

identified in each of these industries where improvements could lead to substantial economic and 

energy savings. 

Additionally, the general practice of most industries is to choose fuel types and sources based 

primarily on cost and availability. This, in many cases, leads to variations in fuel quality and 

purity which will directly impact the performance of the refractory materials used in the 

respective furnaces. This must be considered when choosing furnace materials and when 

evaluating energy efficiency. 

2.2 Current R&D Status 

Substantial changes that have occurred in refractories technology over the past 25 years have 

resulted in crosscutting impacts on a number of manufacturing industries,2

 including the use of 

fibrous and high-strength porous insulation to reduce energy consumption in heat-treating 

furnaces. In the steel industry, the use of resin-bonded MgO-C-metal linings in basic oxygen 

furnaces significantly increases the number of heats between needed relinings, and SiC-based 

refractories have led to increased furnace life for steel blast furnaces. In the glass industry, the 

advent of oxy-fuel melting has led to renewed interest in alternative refractory materials that will 

allow for processing at higher temperatures and greater corrosion resistance to high alkali and 

water-rich environments. Additionally, many more recent ceramic compositions have received 

limited evaluation but have not reached maturity or production status. All of these events are 

representative of industrial process improvements that can be accomplished through advances in 

refractory materials, yet additional improvements could lead to even greater energy savings 

throughout the various industries. 

The choice of refractory materials for a certain application is often dictated by the trade off 

between cost and performance. In the past economic considerations have led to the selection of 

lower cost/poorer performing refractories over higher performance/higher cost materials. Yet, 

with rising energy costs, environmental regulations, and economies of scale; higher priced 

premium refractory materials are receiving renewed interest and may prove cost effective in 

many applications for which they were previously deemed unsuitable. 

4 
Refractories for Industrial Processing 

3. Estimated Metrics for Overall Energy Savings 

Opportunities 

Energy usage in several of the industries and estimates of potential energy savings through 

improvements in refractories are shown in Table 3. The values shown in this table are based 

primarily on DOE Manufacturing Energy Consumption Survey (MECS) data describing 

manufacturing energy consumption in the year 1998.3 The estimates of energy usage affected by 

refractories are based on evaluation of industry roadmap documents and discussions with industry 

leaders. Energy savings goals are based on the needs of the industries and estimates for the four 

typical furnace types (Table 1) used throughout many of the industries, which are presented in 

detail in Appendices B, C, and D. These analyses are based on heat balance studies where the 

energy supplied to the system through combustion is equal to the sum of (1) latent heat in the 

furnace exhaust gases, (2) heat lost through the furnace walls and roofs, and (3) heat required for 

heating reactants to the process temperature and phase transformations at that temperature. The 

achievement of theoretical efficiency would require that items 2 and 3 be equal to zero. From a 

practical point of view, perfect efficiency is not probable, but improvements in furnace 

refractories could help facilitate improvements of furnace efficiency beyond the levels of present 

technology. The projected saving opportunities in the range of 166 TBtu/Yr to 830 TBtu/Yr are 

equivalent to the yearly residential electrical usage of New York state and the yearly residential 

air conditioning usage of the US, respectively. 

Table 3. Energy analysis by industry with estimated short-term and long-term energy savings 

opportunities from improved refractories 

Energy savings goals for 

Industry NAICS 

codes 

Total 

TBtu/yeara 

Usage affected by 

refractories 

Percent TBtu/Yr 

improved refractories 

(TBtu/Yr) 

Near-term 

2% 

Long-term 

10% 

Aluminum 3313XX1 441 90% 397 7.9 39.7 

Chemical 325XXX 3,704 50% 1,852 37.0 185.2 

Fabricated metalsb 33281 445 90% 401 8.0 40.1 

Forest products 321XXX, 3,248 40% 1,299 26.0 129.9 

322XXX 

Foundries 3315XX 233 90% 210 4.2 21.0 

Glass 3272XX 203 70% 142 2.8 14.2 

Iron & steel 33111 1,672 85% 1,421 28.4 142.1 

Petroleum & coal 324XXX 3,622 50% 1,811 36.2 181.1 

products 

Stone, clay products

c 327XXXd 766 100% 766 15.3 76.6 

All 14,334 — 8,299 165.8 829.9 

a TBtu = trillion Btu/year. Source: U.S. Department of Energy, Energy Information Administration, 1998 MECS 

Consumption, Table N3.2, Fuel Consumption, 1998. 

b Includes heat treatment. c Primarily cement and lime. d Excludes category for glass, 3727XX. 

Additionally, one must consider that declines in certain properties may have a direct influence on 

other related properties. This is evident for example in a furnace wall where lack of corrosion 

5 
Refractories for Industrial Processing 

resistance may result in a reduction in thermal efficiency as decreases in wall thickness due to 

corrosion will lead to decreased insulating ability of the furnace wall. The furnace efficiency will 

now not only be impacted, but thinning of the refractory wall may also lead to shut-down of the 

furnace for maintenance causing loss of production and the need for additional energy to reheat 

the furnace. Energy considerations from all these areas are considered in Table 3. 

Figure 1 shows the potential energy benefit opportunities represented by the four furnace types 

analyzed. 

Fig. 1. Long-term energy benefit opportunities estimated for the four analyzed 

furnace types. 

6 
Refractories for Industrial Processing 

4. Process Technology and Performance Barriers 

4.1 Kilns and Calciners 

Kilns and calciners, or the products produced by them, are used in many industries (aluminum, 

cement, chemical, forest products, glass, iron and steel, and mining) and therefore have 

crosscutting interests. Analyses were performed to evaluate cases for both alumina and lime 

calcinations, based on use of rotary-type kilns for various flame temperatures, exhaust 

temperatures, and wall/roof losses (Appendix B). For all the heat supplied by combustion, the 

gross available heat is calculated from the heat content of the flame less the heat content of the 

exhaust gases, and a part of the heat supports wall loses while the remainder is available for the 

work of the kiln/calcinations of feed to product. 

4.1.1 Industrial Applications of Process Technology 

Aluminum — alumina, carbon, lime Glass - lime 

Cement and lime — lime, cement Iron and steel - lime 

Chemical and petrochemical — lime Mining — raw materials 

Forest products — lime 

4.1.2 Current Technology 

Two types of kilns typically found in the industries listed were selected for review. In the case of 

alumina calcination, we considered a kiln measuring 3.5 m in diameter by 80 m long with an 

output of approximately 250,000 tons/year; for lime calcinations, we chose a kiln measuring 5 m 

in diameter by 50 m long with an output of approximately 365,000 tons/year. In the lime 

calcination process, product is often lost through dust loss as loose powder is blown out of the 

kiln along with the exhaust gases. Therefore, in the lime analysis, both the case of no dust and the 

more realistic condition of up to 20% of the calcined product being lost as dust were considered. 

Calculated heat of reactions (at 77°F) of 634 Btu/lb and 1364 Btu/lb were used for the alumina 

and lime calcination processes, respectively, leading to a calculated total heat requirement of 

1728 Btu/lb for the alumina calcination process and 2196 Btu/lb for the lime calcination process. 

4.1.3 Energy Impact 

Energy benefits resulting from refractories in kilns are related to the following operational 

improvements: 

1. An increase in the adiabatic flame temperature of 200°F (as with modified oxy-fuel firing at 

about 20% enrichment) can lead to substantial fuel savings due to thermal efficiency. 

2. A decrease in the exhaust temperature of 200°F (resulting in greater heat retained for the 

actual process) also leads to a substantial fuel savings. 

3. A decrease in wall/roof losses of 200,000 Btu/ton of product can result in substantial energy 

savings (considering half the supplied energy goes up the stack, with the remainder available 

to support work and wall losses). 

4. Process yield improvements will result in substantial energy benefits. 

For the energy analyses, it was assumed that a little over half of the total calcined alumina 

produced in the United States (4.4 million metric tons in 2001) was produced using a rotary kiln. 

Similarly, it was assumed that of the 20 million short tons of lime produced in the United States 

7 
Refractories for Industrial Processing 

in 2002, about 80% of the production was performed using rotary kiln calciners. It was also 

assumed that 80% of the 103 million metric tons of cement produced in the United States in 2002 

was produced using rotary kilns operating in the temperature range of 2600 to 3000°F. Results of 

the analysis are shown in Table 4. 

Table 4. Total effect of improved refractories on rotary kilns 

Energy-intensive 

process industries Product Fuel decrease 

(Btu/ton) 

Wall loss 

decrease 

(Btu/ton) 

Tons affected 

per year 

Energy benefits 

per year 

(TBtu) 

Aluminum Al2O3 413,000 200,000 2,500,000 1.5 

Cement and lime Cement 800,000 200,000 82,400,000 82.4 

Cement and lime, 

chemical and 

petrochemical, 

forest products, 

glass, iron and 

steel 

Lime: 

 No dust 

 20% dust 

413,000 

885,000 

200,000 

250,000 

16,000,000 

16,000,000 

9.8 

18.3 

All 112a 

a

 Without carbon and sulfur contributions. 

Source: Portland Cement Association, Economic Industry Overview, 2003.

4.1.4 Barriers Related to Refractory Materials for Kilns and Calciners 

The central conclusion from the analyses for concerning kilns and calciners is that as flame 

temperatures increase and exhaust temperatures approach process temperatures, greater thermal 

efficiency is realized. Hence, improvements in reducing the thermal conductivity of the 

refractories, and improving the chemical and mechanical properties of the refractories, will have a 

significant payoff in overall fuel consumption. Addressing the following barriers associated with 

the performance of refractory materials used in kilns and calciners would improve energy 

efficiencies. 

1. Unavailability of refractories with sufficient high-temperature strength, erosion/corrosion 

resistance, creep resistance, resistance to large temperature gradients, and stability at higher 

temperatures for use with elevated furnace flame temperatures. 

2. Lack of sufficient high-temperature durability and corrosion/erosion resistance in current 

refractories for improved heat transfer through increased levels of stirring and agitation of the 

material being processed. 

3. Unavailability of refractories with lower thermal conductivity and greater high-temperature 

stability needed to achieve substantial reductions in wall and roof heat losses. 

4. Lack of knowledge concerning operations needed to increase refractory life, reduce product 

sticking, and increase levels of product recovery for the reduction of unscheduled outages, 

improvement of product yield, and the increase of furnace campaign lives. 

4.2 Reverberatory Furnaces 

Analyses of reverberatory furnaces were performed to evaluate cases of aluminum, copper, and 

glass-melting furnaces for various assumed values of flame temperatures, exhaust temperatures, 

and wall/roof losses. For all the heat supplied by combustion, the gross available heat is the heat 

8 
Refractories for Industrial Processing 

content of the flame less the heat content of the exhaust gases. Additionally, part of the gross 

available heat supports wall losses and the remainder supports melting, the real job of the furnace. 

Reverberatory furnace schematics are provided in Appendix C. 

4.2.1 Industrial Applications of Process Technology 

Aluminum — metal melting Metal casting — metal melting 

Glass — glass melting Copper — metal melting 

Iron and steel — metal melting 

4.2.2 Current Technology 

Analyses of reverberatory furnaces were based on a nominal furnace having a capacity for 

melting 20,000 lb of material per hour. This is a typical size used in the aluminum industry and is 

used for the other industrial analyses for comparison purposes. For the fixed melting capacity of 

20,000 lb/h, the energy needed for melting is the same in all aluminum melting cases, 504 Btu/lb. 

In the analysis, wall and roof losses from 0 (really impossible) to 10 million Btu/h, or 0 to 

500 Btu/lb of material, have been assumed. This range is expected to span the full spectrum of 

industrial practice. The gross available heat to do these jobs is the difference in flame and exhaust 

heat contents. It has been assumed that natural gas fuel with an adiabatic flame temperature of 

~3600°F has been used as a base, with increasing flame temperatures that may be achieved 

through regenerators/recuperators or oxygen enrichment. 

Exhaust temperatures ranging from 1400 to 2000°F have been assumed. Increasing exhaust 

temperatures at fixed flame temperatures results in total heat consumptions ranging from 

777 Btu/lb (61% thermal efficiency) to 2261 Btu/lb (22.3% thermal efficiency). Furnaces in the 

aluminum industry generally fall in this range, as summarized below: 

Furnace type Thermal efficiency 

Normal furnace (poor to good practices) 2500–1350 Btu/lb 

State-of-the-art new furnace 1350–1200 Btu/lb 

Furnace builders’ lower limit for reverb. 1200–1000 Btu/lb 

New/radical technology 1000–504 Btu/lb 

Similar efficiencies are expected in the other two industries studied. 

4.2.3 Energy Impact 

Energy benefits that could result from improved refractories in reverberatory furnaces are related 

to the following operational improvements: 

1. An increase in the adiabatic flame temperature of 200°F (e.g., oxy-fuel firing with 20% 

oxygen) leads to substantial fuel savings due to thermal efficiency. 

2. A decrease in the exhaust temperature of 200°F (resulting in greater heat retained for the 

actual process) also leads to a substantial fuel savings. 

3. A decrease in the wall/roof losses of 2.5 million Btu/h also results in substantial energy 

savings (considering half the supplied energy goes up the stack, with the remainder available 

to support melting and wall losses). 

For the calculations, it was assumed that annually in the United States, 9.3 billion lb of aluminum 

(half of all the ingot and mill products in the aluminum industry in 1998) are melted in 

reverberatory furnaces, 0.36 billion lb of refined copper from scrap is processed in reverberatory 

9 
Refractories for Industrial Processing 

furnaces, and an average of the 40 billion lb of glass is produced using reverberatory furnaces 

(based on 1999 numbers). In addition, 20% oxygen enrichment was assumed for the metal 

melting cases, while 100% oxygen enrichment was assumed for the glass melting case. The 

details of the calculations are presented in Appendix D. Results of the analysis are shown in 

Table 5. 

Table 5. Total effect of improved refractories on reverberatory furnaces 

Energy-intensive 

process industries Product 

Fuel 

decrease 

(Btu/lb) 

Wall loss 

decrease 

(Btu/lb) 

Pounds 

affected per 

year 

(× 109

) 

Energy benefits 

per year 

(TBtu) 

Aluminum, copper

a

Glass 

All 

 Metal melting 

Glass melting 

452 

476 

200 

100 

10 

40 

2 

(20% enrichment) 

76 

(100% enrichment) 

78 a Without iron and steel and metal casting contributions 

Sources: American Metal Market, Metal Statistics 1998, 90th ed., 1998; DOE Office of Industrial Technologies, 

Energy and Environmental Profile of the U.S. Glass Industry, 2002. 

4.2.4 Barriers Related to Refractory Materials for Reverberatory Furnaces 

The central conclusion from these analyses is that as flame temperatures increase and exhaust 

temperatures approach process temperatures, greater thermal efficiency is realized. As shown in 

Appendices B, D, and E, the impact of reducing wall losses by 2.5 million Btu/h leads to 

corresponding savings of 4.5 million Btu/h for aluminum (232 Btu/lb), 7.5 million Btu/h for 

copper (333 Btu/lb), and 11.9 Btu/h for glass (595 Btu/lb). Hence, improvements in reducing the 

thermal conductivity of the refractories, and improving the chemical and mechanical properties of 

the refractories, will have a significant payoff in overall fuel consumption. Addressing the 

following barriers associated with the performance of refractory materials used in reverberatory 

furnaces would improve energy efficiencies. 

1. Unavailability of refractories with sufficient high-temperature strength, erosion/corrosion 

resistance, creep resistance, resistance to large temperature gradients, and stability at higher 

temperatures for use with elevated furnace flame temperatures. 

2. Lack of sufficient high-temperature durability and corrosion/erosion resistance in current 

refractories and materials for improved heat transfer through increased levels of stirring and 

agitation of the material being processed. 

3. Unavailability of refractories with lower thermal conductivity and greater high-temperature 

stability. 

4. Lack of knowledge concerning the interactions between refractories and process 

environment/operations which leads to reduced refractory life, product sticking unscheduled 

outages, low product yield, and limited furnace campaign lives. 

4.3 Boilers and Reactor Systems 

Boilers and reactor systems are found in most industrial settings and therefore are of key interest 

as a crosscutting unit process. For the following analysis, boilers for the production of steam are 

considered for the aluminum, cement and lime, chemical and petrochemical, forest products, and 

10 
Refractories for Industrial Processing 

heat treatment industries. (See Appendix E for a schematic illustration of boiler furnace sections.) 

Additional areas for energy efficiency improvements also exist in chemical reactors utilizing 

catalysts and other reactions. 

4.3.1 Industrial Applications of Process Technology 

Aluminum — steam production Forest products — steam production 

Cement and lime — steam production Heat treatment — steam production 

Chemical and petrochemical — steam production 

4.3.2 Current Technology 

For the purpose of the performance analyses of boilers, a standard unit using a single header 

boiler system has been assumed. 

4.3.3 Energy Impact 

Improved refractory practice in boilers will result in energy benefits either from improving the 

boiler efficiency through improved vessel insulation and/or by reduced heat losses during steam 

transfer through improved transfer pipework insulation. 

The DOE Best Practices Steam System Assessment Tool was used to analyze the effects of 

improving boiler efficiency from 85 to 87% through use of the following: 

• Improved insulation for improved boiler efficiency 

• Improved pipework insulation for improved steam transfer systems 

The effects of increasing boiler efficiency by 2% and reducing heat losses during steam transfer 

by 10% are shown in Tables 6 and 7. 

Table 6. Steam system assessment tool single header analysis results for improved insulation 

for improved boiler efficiency 

85% efficient 87% efficient 

Variable boiler case boiler case Improvement 

Fuel cost, $1000/year 4,000 3,908 92 (2.30%) 

Fuel consumption, std ft3/h 71,421 69,799 1,622 (2.27%) 

CO2 emissions, 1000 lb/year 66,403 64,876 1,527 (2.30%) 

NOx emissions, 1000 lb/year 131 128 3 (2.29%) 

Table 7. Steam system assessment tool single header analysis results for improved pipework 

insulation for improved steam transfer systems 

Current 10% 

Variable operation improvement Improvement 

Fuel cost, $1000/year 4,000 3,924 76 (1.90%) 

Fuel consumption, std ft3/h 71,421 70,065 1,356 (1.90%) 

CO2 emissions, 1000 lb/year 66,403 65,142 1,261 (1.90%) 

NOx emissions, 1000 lb/year 131 129 2 (1.53%) 

11 
Refractories for Industrial Processing 

As shown in Table 6, a 2% improvement in boiler efficiency leads to a 2.3% improvement in fuel 

cost and consumption, while the 10% improvement in the efficiency of the transfer pipework 

insulation leads to a 1.9% improvement in fuel cost and consumption. As shown in the tables, 

there are also decreases in emissions associated with improvements in boiler efficiency and 

pipework insulation. If these fuel savings are then applied to the chemical, petroleum refining, 

and pulp and paper industries, estimates can be made concerning the possible energy savings 

resulting from implementation of these improvements in each industry. Table 8 shows the results 

of such an analysis. 

Table 8. Potential energy savings from improved boiler efficiency due to improved refractories 

and improved transfer pipe-work insulation 

Capture Projected 

IOF industry segment TBtu % basis fuel savings 

(%) (TBtu) 

Chemical industry steam opportunities 

Present usage of fuel for generating steam 1,540 

Opportunity for improving boiler efficiency 2.3 

from 85 to 87%

 Fuel saved 35 

Opportunity for reduced line heat losses by 10% 1.9 

 Fuel saved 30 

Chemical industry fuel reduction 65 40 26 

Pulp and paper steam opportunities 

Present usage of fuel for generating steam 2,596 

Opportunity for improving boiler efficiency 2.3

from 85 to 87%

 Fuel saved 60 

Opportunity for reduced line heat losses by 10% 1.9

 Fuel saved 49 

Pulp and paper industry fuel reduction 109 40 44 

Petrochemical steam opportunities 

Present usage of fuel for generating steam 1,675 

Opportunity for improving boiler efficiency 2.3 

from 85 to 87%

 Fuel saved 39 

Opportunity for reduced line heat losses by 10% 1.9 

 Fuel saved 31 

Petrochemical industry fuel reduction 70 40 28 

Total steam opportunity 244 40 98 

According to the DOE-EERE Best Practices report concerning steam usage in the chemical 

manufacturing, pulp and paper, and petroleum industries,4

 1540, 2596, and 1675 TBtu of fuel, 

respectively, is used to generate steam per year by each industry (see Table 8). As can be seen, on 

a 40% capture basis (i.e., 40% of the existing plants will use the improved refractory and 

insulation systems), the levels of projected fuel savings are 26, 44, and 28 TBtu, respectively, for 

the chemical, pulp and paper, and petroleum refining industries. This results in a total projected 

fuel savings of 98 TBtu across the three industries. 

12 
Refractories for Industrial Processing 

4.3.4 Barriers Related to Refractory and Insulating Materials for Boilers and Reactors 

The central conclusion from these analyses is that decreased fuel consumption resulting from 

improved boiler efficiency can be achieved through better insulation practices. As shown, the 

impact of increasing boiler efficiency by 2% and reducing heat losses during steam transfer by 

10% will result in a total fuel savings of 98 TBtu for the industries studied. Addressing the 

following barriers associated with the performance of refractory and insulating materials used in 

boilers and reactors would reduce energy requirements: 

1. Unavailability of insulating refractories with sufficiently low thermal conductivities to lead to 

increases in energy efficiency levels for boiler operations. 

2. Unavailability of improved insulating materials for transfer pipework that could help reduce 

heat losses during steam transfer operations. 

4.4 Gasifiers 

Gasifiers are currently being used in the chemical and petrochemical industry for coal and 

biomass gasification and in the forest products industry specifically for black liquor gasification. 

This technology has the potential for biomass solid fuels to be converted into fuel gas “syngas” 

consisting largely of hydrogen (H2) and carbon monoxide (CO). This gas can then be burned 

cleanly and efficiently in a gas turbine to generate electricity and/or used to synthesize clean 

transportation fuels or chemicals. 

Gasification is currently used in the chemical industry as a means to produce clean energy, highvalue energy products, and various petrochemicals from coal (including the possibility of 

producing hydrogen fuel, high-grade transportation fuels, and industrial-grade methanol). 

Experimental gasifiers are being applied in a few chemical and pulp and paper plants for coal or 

black liquor conversion, while gasification is also being used as a means of using biomass as an 

energy source rather than a waste product. In the future, the forest products industry may use 

gasification to replace Tomlinson boilers. 

A limitation on the use of these experimental gasifiers is the lack of reliable refractories for 

containment. If the refractory requirement could be met, coal, black liquor, and biomass 

gasification could lead to greater reductions in purchased energy and even change these industries 

from net energy importers to net energy exporters.5

 Gasification holds the promise of generating 

175% more electrical power than do current technologies, but the gasification technology 

currently in use requires frequent refractory relines of the vessels, hence reducing the economic 

and energy benefits.5

 Additionally, although chemical or pulp and paper manufacturing facilities 

are not currently used to produce transportation fuels from biomass, they may become a valuable 

asset for establishing an infrastructure for such technology. Future gasification-based 

“biorefineries” might produce a variety of renewable fuels, electricity, and chemicals in 

conjunction with pulp and paper products. 

4.4.1 Industrial Applications of Process Technology 

Chemical and petrochemical — coal and biomass gasification 

Forest products — black liquor gasification 

4.4.2 Current Technology 

A coal gasifier generally consists of a metallic shell protected by a lining made up of two to six 

concentric layers of refractory protecting the interior of the vessel. Coal gasifers are expected to 

13 
Refractories for Industrial Processing 

operate at high temperatures and moderate pressures, and with gas and particle flows containing 

H2, CO, H2O, C, H2S, HCl, hydrocarbons, metallic impurities, and salts. These linings have 

average lifetimes on the order of 7000 run hours and generally fail because of excessive 

mechanical wear (erosion), chemical penetration of slag into the structure of the innermost brick 

(corrosion), or some combination of these two mechanisms. Erosive wear occurs as the feedstock 

itself causes physical decay of refractory hot-face surfaces when injected at high velocities. 

Corrosive attack occurs when the coal slag residues penetrate the refractory hot face surface, thus 

weakening the refractory by forming solid-solution phases. 

The nominal scale of an average pulp and paper mill is 6 million lb/d of black liquor solids 

(BLS)—the equivalent of 1.5 million Btu/h or 438 MW fuel—corresponding to 1900 machine dry 

short tons per day of paper production (1725 metric tons).5

 Pulp mills processing more than 

6 million lb/day BLS account for about one-third of all U.S. capacity today, and this fraction will 

grow over time as mill consolidations continue. A standard sized gasifier for this size mill is 

considered. This unit was assumed to have a BLS concentration of 80% and a steam pressure of 

1870 psia. 

4.4.3 Energy Impact 

For the following analysis, we considered a combined-cycle gasifier typical of those used for coal 

or black liquor gasification with the requirements above. Combined-cycle black liquor 

gasification has the potential of producing more than twice the current output of electricity per 

ton.5

 In many cases, power in excess of a chemical or pulp mill’s needs could be produced and 

exported. Further, if gasified coal or black liquor were converted to a high-quality synthetic gas 

similar to syngas produced from fossil residual oil, the energy and economic benefits would be 

even greater. Conventional processes could then be used to convert the gas to liquid fuels or 

hydrogen. 

According to the economic and energy analysis performed by Larson et al.,5

 full implementation 

of the black liquor gasification initiative could achieve the following energy benefits by 2020: 

1. Production of up to 8 GW of electricity from sustainable, renewable raw materials by 2020 

through black liquor gasification alone. 

2. Production of 16 GW or more of electric power for the application of black liquor 

gasification in conjunction with wood residual gasification by 2020. 

3. Alternatively, displacement of more than 282 million barrels of oil per year if these 

technologies were applied to syngas production to be used for liquid fuels. 

The total of these biomass energy sources consumed at pulp mills in 2002 was estimated at 

1.6 quads (one quad is 1015 Btu).6

 The potential renewable energy generation potential from black 

liquor in the pulp and paper industry in the southeastern United States is shown in Table 9. 

Refractories and insulation materials play key roles in many of the unit operations involved in 

both of these industries. It is estimated that processes affected by refractories account for about 

40% of the industries’ direct energy consumption. Based on subsequent analyses of typical 

furnaces and unit processes, improvement of refractory systems for these two industries could 

lead to energy savings opportunities up to 129 TBtu/year. 

14 
Refractories for Industrial Processing 

Table 9. Renewable energy generation potential (billion kWh/year) from black 

liquor in the pulp and paper industry in the southeastern United States 

Technology 

Potential by 2020 

(billion kWh/year) 

Equivalent energy 

(TBtu/year) 

Conventional coal combustion 726 

(33–35% efficient) (energy needed) 

Combined cycle gasification 660 

(45–50% efficient) (energy needed) 

 Total improvement 66 

Chemical 440 

Forest products 

Tomlinson recovery boiler 4.4 15 

High-efficiency Tomlinson 16.7 57 

recovery boiler 

Combined cycle gasification 36.2 124 

 Total improvement 67 

4.4.4 Barriers Related to Refractory Materials for Gasifiers 

Addressing the following barriers associated with the performance of refractory materials used in 

gasifiers would reduce energy requirements: 

1. Unavailability of advanced refractory materials with improved corrosion/erosion resistance, 

improved reliability, and reduced maintenance costs. 

2. Lack of refractories with resistance to high alkali atmospheres and high levels of stress, 

which lead to gasifier containment issues. 

3. Lack of alternate refractory designs which would provide refractory surfaces (using coatings, 

chemical alteration of refractory surfaces, physical alteration of refractory surfaces, etc.) that 

are more impervious to wear and corrosive attack. 

4. Unavailability of refractory materials with greater thermal compatibility with metallic 

materials used in gasifiers. 

5. Overall R&D Barriers and Pathways for Refractories 

Evaluation of refractory usage, barriers, and needs of these several energy-intensive process 

industries shows that several common themes arise in the pursuit of improving refractory 

materials for greater energy efficiency. These themes relate to improving energy efficiency by 

reducing energy loss and increasing productivity through the development of higher-strength 

refractories capable of operating at higher temperatures, the development of refractories with 

lower thermal conductivity to reduce wall losses at higher service temperatures, the tailoring of 

thermal expansion capabilities of current refractories, and the development of refractories with 

greater resistance to chemical and mechanical wear (corrosion, erosion, abrasion, etc.). This 

concept is shown in Fig. 2. 

15 
Refractories for Industrial Processing 

Fig. 2. Improving refractory materials for improved energy efficiency. 

Through analysis of current refractory use, current barriers, possible improvements in process 

energy efficiency, and R&D refractory needs for several process industries, many issues common 

to multiple industries become apparent. The research focus areas and technical R&D topics below 

have been identified as issues common to many applications of refractories. 

1. Understand the degradation mechanisms of refractories in different environments 

Determine the mechanisms responsible for the chemical and mechanical degradation of 

refractories in the various environments in which they operate. 

2. Achieve improvements in thermal control 

Determine the thermomechanical properties of refractory materials and develop materials that 

will allow furnaces to operate with longer campaigns and with higher levels of insulation to 

reduce heat losses. This may also involve the implementation of new furnace designs, the 

formulation of newly redesigned processes, or the re-engineering of existing industrial 

practices. 

3. Evaluate thermomechanical properties at elevated temperature 

Determine the key properties of refractories at elevated temperatures, including hot modulus 

of rupture, hot crushing strength, creep behavior, refractoriness under load, physical and 

thermal spall resistance, dimensional changes (elastic modulus, thermal expansion, and 

growth from chemical alteration), and thermal conductivity. 

4. Form an expert advisory committee 

Assemble a committee composed of current and retired industry experts to serve in an 

advisory role for industry-focused refractory R&D projects. Additionally, inventory and 

preserve existing industrial equipment and facility resources through systematic surveys of 

refractory experts (especially retirees) so that industrial furnaces and existing laboratories can 

be utilized, when possible, to augment new experimental system requirements. 

16 
Refractories for Industrial Processing 

6. Specific R&D Areas with Barriers and Pathways 

The following text summarizes pathways for overcoming barriers for the six specific R&D areas 

that can yield the greatest returns in refractories. 

6.1 Improved Castable and Monolithic Refractory 

Barrier 

Inadequate resistance to corrosive and mechanical wear, which limits the use of castable and 

monolithic refractories. 

Pathways 

• Fabricate fusion-cast refractories (such as alumina zirconia refractories) that do not include 

silica or other glass formers. These refractories should possess greater corrosion resistance 

and thermoelastic properties. The refractories would be of use to industries such as glass, 

pulp and paper, and many others. 

• Fabricate components for continuous measurement of parameters such as temperature, bath 

chemistry, etc., that will have long-term stability in high-temperature industrial solvents, such 

as cryolite for the electrolytic production of aluminum. Candidate materials such as HfB2, 

HfC, ZrB2, ZrC, and TaC should be considered. 

• Develop oxide/nitride grain boundary additives leading to closure of open porosity in silica, 

mullite, and magnesium aluminates as used by the glass and other industries. 

6.2 Modified Fiber Materials 

Barrier 

Lack of fiber insulation materials with lower thermal conductivity, greater chemical 

resistance, and higher temperature stability limits. 

Pathways 

• Produce frothed latex/fiber mixes for gunning to reduce airborne fibers during installation. 

Enhance adhesion of materials to steel and other metals for application in heat treatment 

furnaces. Develop new inorganic binder systems. 

• Fabricate fibrous materials with greater strength and lower friability for use in heat treatment 

and chemical and petrochemical industries. 

6.3 Surface Modification (Coatings, Physical or Thermal Alteration) 

Barrier 

Lack of techniques and materials for applying surface coatings or inducing physical alteration 

of refractory surfaces through either mechanical or thermal means. 

17 
Refractories for Industrial Processing 

Pathways 

• Modify surfaces or near-surfaces of sintered alumina and mullite refractories by applying 

corrosion/erosion-resistant coatings such as spinel phases (or additional corrosion-resistant 

materials) for use by industries such as glass, pulp, and paper. 

• Determine the effects of “preexisting” refractory heating conditions that, when applied 

concurrently with high-intensity infrared heating, modifies the surface of bonded refractories 

and results in the residual compression of a modified surface at service temperatures. One 

advantage of this could be an increase in the “strain-free” temperature of refractories used in 

process industries. 

• Determine the applicability of radiation-heating attenuation/mitigation, and if applicable, 

simulate refractory structures that would maximize its effect to optimize radiant heat transfer 

(e.g., improve oxy-fuel fired systems for the aluminum, glass, metal casting, and iron and 

steel industries). 

• Investigate the effects of radiative heating and transmittance in refractories on thermal 

conduction for minimizing thermal losses from various types of furnaces used throughout the 

selected industrial areas. 

6.4 Synthetic Aggregates and Powders 

Barrier 

Lack of higher-performance aggregates that would increase the life and durability of 

refractory products used across the various industries and lack of higher-quality matrix 

materials that would increase the durability of refractories. 

Pathways 

• Measure and model strength as a function of aggregate size distribution. Tailor microstructure 

so that flexure and compression strengths are increased. 

• Explore candidate aggregate coatings as “additives” that control or binder grain boundary 

effects (e.g., diffusion, heat transfer, corrosion-resistance, etc.) for improved refractory 

processing. 

6.5 High-Temperature Properties of Refractories 

Barrier 

Lack of information on the high-temperature thermomechanical and corrosion properties of 

current and newly developed refractory materials. 

Pathways 

• Develop realistic test methods and environmental tests systems in order to evaluate elevated 

temperature properties. Evaluate corrosion behavior of refractories for various effects of 

water/CO/CO2/alkali content on the corrosion and phase stability of refractories used in 

gasifiers and in the glass, aluminum, steel, and metal casting industries. Develop lab-scale 

systems to evaluate the degradation of refractories up to 2000°C. 

• Systematically characterize and model reheat-generated permanent expansion effects in 

fusion-cast and bonded refractories, a phenomenon known to occur in many refractory 

systems used throughout the various process industries, but not currently understood. 

18 
Refractories for Industrial Processing 

• Perform postmortem analysis on failed refractory components from several industries and 

identify service-limiting flaw types. Use component samples for measurement of strength 

distribution and compare to strength distribution of virgin materials. Perform probabilistic 

strength-size-scaling analysis using both distributions as input and predict failure time of 

components. 

• Measure the corrosion resistance of bonded refractories as a function of refractory sintering 

temperature and interpret the performance in reference to service temperature. 

• Investigate the effect of specimen size in refractories by measuring strength (modulus of 

rupture) as a function of specimen size in bonded or fusion-cast refractories. Determine if 

variations in strength can be represented by a Weibull distribution and if strength is sizescaled. 

6.6 In Situ and Graded Refractories 

In situ refractories are defined as brick or unshaped products that react internally or with 

furnace atmospheres and/or slag components to enhance their performance. Natural in situ 

formation of Fe-Mg-O spinels in steel-making furnaces permits improved life and durability 

of these furnaces. 

Barrier 

Lack of R&D to improve existing in situ refractory systems and for development of new 

in situ systems for additional industrial applications 

Pathways 

• Model crack initiation and growth as a function of material, depth, and service properties 

using microstructure-based finite element analysis (e.g., μ-FEA, OOF). Simulate and 

optimize hypothetical microstructures that hinder crack growth. 

• Develop high-intensity infrared heating techniques to fabricate optimized graded refractory 

structures, and measure thermomechanical stability and corrosion resistance. 

• Model, simulate, and design refractory microstructures that promote maximum creep and 

erosion resistances. 

• Determine the effect of crystallography of base and in situ phases on the development of 

thermal stresses and the stability of in situ layers. 

• Determine and optimize “substrate” surface geometries (e.g., dimples, corrugated, saw-tooth) 

that will minimize shear stresses in thermo-environmental barrier refractory coatings that are 

subjected to a thermal gradient. 

6.7 Sensors for Materials Performance 

Barrier 

Lack of sensors to monitor refractory-related issues including wall thickness regression, gas 

phases, and other environmental conditions. 

Pathways 

• Develop sensors and methods for the measurement of spectral emissivity for application in 

glass tanks and metal melting furnaces. 

• Develop sensors and methods for the monitoring of wall regression due to corrosion, erosion, 

and wear in glass tanks and metal melting furnaces. 

19 
Refractories for Industrial Processing 

• Develop continuous real-time measurement techniques for parameters such as temperature, 

pressure, melt analysis, flue gas analysis, and environmental gas phases. 

The tasks and the industries which will benefit from the projects described above are shown in 

Table 10. As is evident from the table, these suggested R&D pathways are crosscutting. Although 

there is a current trend for many of these industries moving overseas, it is believed that the R&D 

pathways proposed will directly benefit U.S. industry. Even as processes such as aluminum 

smelting and glass production move off shore, there will still be a need for remelting and other 

production domestically. 

Table 10. Recommended R&D pathways for advanced refractories 

Technical R&D 

potential pathways Aluminum 

Cement 

& lime 

Chem. / 

petrochemical 

Industry segment 

Forest 

products Glass Heat treatment 

Iron 

& 

steel 

Metal 

casting 

Improved castable 

and monolithic 

refractories 

X X X X X X X X 

Modified fiber 

materials 

X X X 

Surface modifications 

(coatings, 

alterations) 

X X X X X X 

Synthetic aggregates 

and powders 

X X X X X X X 

Elevated temperature 

refractory 

properties 

X X X X X X X 

In situ and gradient 

refractories 

X X X X X X X 

Sensors for materials 

performance 

X X X X X X X X 

20 
21 
Refractories for Industrial Processing 

7. References 

1. J. Mosser and G. Karhut, “Refractories at the Turn of the Millennium,” Paper presented at 

UNITECR ’99, Berlin, Germany, pp. xxv–xxx, Sept. 1999. 

2. D. W. Freitag and D. W. Richerson, Opportunities for Advanced Ceramics to Meet the Needs 

of the Industries of the Future, DOE/ORO 2076, U.S. Department of Energy, Office of 

Industrial Technologies, December 1998. 

3. U.S. Department of Energy, Energy Information Administration, 1998 MECS Consumption, 

Table N3.2, Fuel Consumption, 1998. 

4. U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Steam 

System Opportunity Assessment for the Pulp and Paper, Chemical Manufacturing, and 

Petroleum Refining Industries: Main Report, available online at http://www.oit.doe.gov/ 

bestpractices/technical_publications.shtml under “Market Assessments.” 

5. E. D. Larson, S. Consonni, and R. E. Katofsky, A Cost-Benefit Assessment of Biomass 

Gasification Power Generation in the Pulp and Paper Industry, Princeton University, 

October 2003. Available online at http://www.princeton.edu/~energy/publications/pdf/2003/ 

BLGCC_FINAL_REPORT_8_OCT_2003.pdf. 

6. DOE/EIA Renewable Energy Annual 2001, DOE/EIA-0603, 2001 

22 
23 
APPENDICES 

24 
25 
Appendix A — Chemical and Petrochemical 

Appendix A: 

Discussion of Refractory Usage by Process Industries 

A.1 Aluminum 

Current Industry Operations and Refractory Use 

Aluminum production is energy-intensive, consuming about 0.4 quads of electrical energy per 

year in the United States.1 Additionally, the industry faces many of the same competitive 

pressures as the steel industry, including global competition, energy consumption issues, and 

pollution abatement. Therefore, a technology road map has been created to address these issues 

by targeting areas such as energy consumption, productivity, emissions, recycling, quality, 

alternate technologies, and by-products.2 

Many systems are used in the production of aluminum and its alloys, all of which depend on the 

proper selection and management to ensure the cost effectiveness of the metallurgical and 

refractory practices. These systems include steam boilers for Bayer plants, calciners, and carbon 

baking furnaces used for the production of the anodes which are consumed in the Hall-Heroult 

reduction process. In Hall-Heroult electrolytic reduction cells, the key refractories are those used 

to support the cathode hearth and to minimize the penetration of bath constituents into the carbon 

cathode and underlying refractories. Melting furnaces for alloy production by the primary and 

secondary producers include large fossil-fuel-fired reverberatory furnaces as well as a variety of 

induction furnace designs dominated by the coreless and the channel induction types. 

Temperatures and corrodant chemistries, as well as furnace atmospheres, determine the lining 

used for each of these processes. The refractories used for each will be briefly reviewed. 

Primary Aluminum Production 

The production of primary aluminum metal is based on three major steps. The first involves the 

mining of bauxite ore. There are no high-grade deposits in the United States; therefore, 

refractories play little role in this process. 

The second process involves the conversion of bauxite to metallurgical grade alumina. The Bayer 

process is based on digestion of bauxite in caustic solution heated with superheated steam to form 

sodium aluminate. Impurities are removed from the pregnant caustic by chemical reactions with 

dissolved species (primarily SiO2) and filtration of insoluble impurities (primarily iron oxide). 

The saturated solution is then cooled in precipitator tanks in which aluminum trihydrate is 

precipitated and then recovered by filtration. The aluminum trihydrate is converted to alumina by 

calcination. Two methods of calcination are employed in the industry: use of fluid bed calciners 

and use of rotary kilns. 

The third process involves conversion of alumina to aluminum by electrolytic reduction using the 

Hall-Heroult process. In the reduction process alumina is dissolved in a molten bath (cryolite, 

Na3AlF6, with additions of AlF3, CaF2, and sometimes LiF and MgF2). The cell for this process 

utilizes a containment system consisting of a steel shell lined with a castable refractory, a porous 

insulating refractory, a less porous semi-insulating refractory, and finally firebrick. A carbon 

cathode acts as the bottom of the cell for molten aluminum containment. The side-wall 

refractories are generally phosphate-bonded alumina brick or alumina castable materials. Because 

the sides of cell have less insulation, more heat is lost through the walls than the floor. The lower 

26 
Appendix A — Chemical and Petrochemical 

wall temperature is actually beneficial, causing electrolytes in contact with the wall to solidify 

and form a “frozen ledge” which protects the walls of the cell from penetration by molten 

cryolite. This practice is successful, but involves a trade-off between energy efficiency and 

electrolytic cell stability and lifetimes in terms of energy consumption. 

Reduction cells currently employ various refractories just below the carbon cathode which during 

service degrade in various ways, permitting bath constituents to move into the cathode carbon and 

refractories. The systems in place today include fireclay bricks, castable refractories, semiinsulating refractories, and insulating bricks. Systems for primary aluminum production involve 

several different designs depending on the level of amperage and the numbers and array of anode 

blocks. The current intensity for operating reduction cells throughout the world ranges from 

60,000 to 500,000 A. Yet, they all still involve a carbon cathode hearth and consumed carbon 

anodes in which the bauxite ore charge and the systems of alkali fluoride fluxes are placed and 

held while the reduction process is carried out continuously. Typically the reduction process is 

carried out at a temperature in the range of 960 to 970°C. The refractories are key components in 

reduction cells. The atmosphere inside the cavity of a reduction cell is very reactive, containing a 

mixture of CO2, CO, HF, SO2, and particulate fluoride salts. 

In all cases it seems that the formation of a feldspathic composition at the interface between the 

cathode and the barrier is highly desirable but that other bath species can still penetrate the barrier 

and lead to highly shortened cell life. Further, downward movement of alkalis and bath 

constituents through the hearth toward the subhearth insulation eventually results in the failure of 

the furnace, an event which is still quite unpredictable. Because of this, furnace lives vary from 

hundreds to thousands of days. 

Taberaux3

 has reviewed the failure mechanisms of aluminum reduction cell systems. Fireclay 

brick in one reported case lasted over 3000 days, showing that the material itself can meet the 

challenge, all other factors being optimal in their selection and control. Castables have 

disadvantages in cast-in-place systems, and currently the preference is for precast monolithic 

systems, which have had good success. Two types of dry barrier refractories are currently being 

used. Their advantages rest on ease and rapidity of installation and lack of joints. They include 

aluminosilicates on the one hand and anorthite feldspar on other, chosen because it undergoes 

beneficial in situ reactions and is, in fact, patented. 

Carbon anode baking furnaces have historically been constructed of aluminosilicate bricks. 

These bricks have sufficient refractoriness that efforts to replace them with alumina-reinforced 

versions have not been cost-effective. The challenge to the brick comes from its coking and from 

the alkalis present in the electrode feedstock. Most recently, fireclays with very low alkali content 

have been used for the production of bricks increasingly designed with surface locking systems 

that reduce the bowing of walls, which is the principal reason for the replacement of the 

refractories. Because such large numbers of bricks are used in carbon baking furnaces, attempts 

have been made to recycle them, but past efforts have not resulted in the formulation of a costeffective protocol. 

Schubert4

 recently addressed the basic issues of movement of the flue walls, and points to the 

findings in postmortems that the insulation may be inadequate for the support role that it plays. 

The many challenges to the flue walls include (1) alkali infiltration and attack from bath 

carryover, via recycled anode butts in the recipe; (2) fluoride salt attack from bath carryover, via 

recycled anode butts in the recipe; (3) minor “squatting” from elevated temperatures and load; 

(4) mechanical damage from impact of the anode and/or the crane during loading and unloading; 

(5) thermal shock, especially during the cooling cycle; (6) construction details that aggravate 

27 
Appendix A — Chemical and Petrochemical 

normal wear and tear; (7) various problems due to improper expansion allowances; and (8) coke 

sticking to refractory, aggravated by alkali content. 

Semifabricated and Secondary Aluminum Production 

Secondary aluminum production involves recycling of aluminum metal scrap. This is carried out 

in a variety of furnace designs. When large volumes of alloy are required, reverberatory furnaces 

are traditionally used. These furnaces have seen a constantly evolving refractory practice. 

Monolithics replaced bricks in the distant past, and currently monolithics in the form of precast 

“big blocks” are receiving renewed interest. A recent paper by Canon5

 reports that the real 

advantage of precast furnace components rests in the time savings rather than in the extension of 

service life. 

As in the case of the subhearth refractories used in pot furnaces, it is essential to develop a freeze 

zone in the walls of all the furnaces. This is done by selection of refractories and appropriate 

thicknesses. It is also essential to restrict the penetration of the aluminum alloy into the refractory, 

which is achieved by incorporation of additives, some of which are proprietary. As it is 

impossible to keep oxygen from reacting with the aluminum metal, a so-called bellyband 

develops in all types of melters. Because of the fluctuating level of the metal in some furnaces, 

the bellyband can be quite wide and eventually extend into the furnace to such an extent that it 

must be removed. In the case of the largest reverberatory types of furnaces the removal is 

sometimes carried out using large robotic devices. In many instances the furnace has to be shut 

down to effect adequate removal. In the hearth region of the melter, linings of silicon nitride– 

bonded silicon carbide are used for molten aluminum contact areas. 

After melting, the molten aluminum flows through refractory-lined troughs between the various 

process steps. These troughs are typically lined with oxide ceramics such as fused silica or mullite 

with 10–15% porosity and special additives to resist wetting. These materials are installed in 

board or castable form and are used due to their low thermal expansion and good thermal shock 

resistance. Still, erosion and abrasion issues are of concern. SiC-based materials have been 

considered for this application but are more expensive than oxide refractories. 

Bellyband buildup is the overriding issue in melting furnaces. Quesnel et al. 6 have reviewed the 

problem for aluminum holding furnaces and melters. There is a lack of understanding concerning 

the mechanisms of bellyband formation. The two types of tests typically used in the study of 

candidate refractories are a bellyband corrosion test (which is a simple cup test) and an 

immersion test (which really looks at the penetrability of the refractory by the metal). Long use of 

these tests has produced inconclusive answers about the mechanism of attachment, and it is 

apparent that new approaches are needed. 

Barriers 

Refractories play key roles in most of the unit operations involved in the production of primary 

and secondary aluminum. It is estimated that about 90% of the direct energy use by the industry is 

in processes affected by refractories. Recommended improvements for refractories include 

applications for refining, reduction, carbon anode baking, and melting operations. Based on 

subsequent analyses of typical furnaces and unit processes, improvement of refractory systems 

for the aluminum industry could lead to energy savings opportunities up to 29.2 TBtu/year. The 

specific barriers are as follows: 

28 
Appendix A — Chemical and Petrochemical 

• Aluminum Refining 

Unavailability of improved refractories for superheated steam generation, as well as 

improved refractories for calcining operations, particularly rotary kilns for alumina and lime, 

which could lead to improved energy efficiency for refining operations. 

• Aluminum Reduction 

Lack of understanding concerning the mechanisms and degradation scenarios leading to the 

failure of reduction cell refractories and, ultimately, the reduction cell. Further, there is a lack 

of improved linings and refractories with improved resistance to bath penetration that would 

permit extended lives for reduction cells. 

• Carbon Anode Baking Furnaces 

Unavailability of stronger insulating shapes for components of carbon baking furnaces. 

• Aluminum Melting Furnaces 

Lack of understanding regarding the mechanisms of bellyband buildup formation. This is 

complicated by the inability of current tests to provide conclusive answers regarding the 

mechanism of attachment and to provide sufficient performance projections. 

R&D Pathways 

• Refining 

o Develop improved refractory and insulation systems for high-temperature steam 

generation. 

o Improve refractory materials for calcining operations, particularly for rotary kilns. 

• Reduction 

o Develop a fuller understanding of cell refractory failure through chemical and 

physical modeling of the reactions and transport mechanisms. Develop knowledge 

and control of all aspects of the chemical and physical properties of the refractories, 

and optimized refractory design approaches for reduction cells carrying increasing 

current intensities. 

o Develop a sidewall refractory design that would allow direct contact with the molten 

cryolite bath, permitting elimination of the “frozen ledge.” 

o Develop refractories with improved toughness, impact resistance, and abrasion 

resistance for the hearth region of reduction cells. 

• Carbon Baking 

o Develop understanding of the manner in which insulation components are being 

affected chemically and mechanically by the loads from the overlying charge and 

wall refractories. Perform studies of the mechanical properties of such used 

refractories to provide answers to these and related questions tied to assuring 

thermomechanical stability. 

o Develop improved refractories and furnace designs for reducing energy consumption 

in carbon melting furnaces though modeling of design, process control, quality 

control, and selection of hot-face refractories, not only in the flue walls but in other 

components of the furnaces. 

29 
Appendix A — Chemical and Petrochemical 

• Melting Furnaces 

o Develop more accurate simulations of the melting process. 

o Develop a materials approach that involves the development of graded monolithic 

linings, cast in place or precast, which have much denser surface regions and might 

also contain additives to dissuade penetrants. Determine the key factors which affect 

penetration and attachment such as the permeability of the surface region to the metal 

or altered metal. 

o Develop improved refractories for aluminum melting and casting operations, 

including advanced refractories for reverberatory melting furnaces and prefired 

refractory riser tubes for low-pressure aluminum die casting applications that are 

economical, strong, thermal-shock-resistant, and nonwetting to molten aluminum, 

and that have low gas permeability. 

o Develop efficient, long-life degassing components for aluminum alloy production. 

A.2 Cement and Lime 

Current Industry Operations and Refractory Use 

The United States is the third largest cement-producing country in the world (after China and 

India). The production of cement and its intermediate product, cement clinker, has risen steadily 

over the last decade, with consumption increasing 40% while production has only increased 27%. 

This has led to increased imports in the short term, mainly from Canada and Mexico, but could 

lead to further expansion of the domestic industry in the long term. Currently, there are 115 

cement manufacturing plants in the United States, with about 75% of the total plants owned by 

only 10 large companies: Lafarge North America, Inc.; Holcim (US), Inc.; CEMEX, SA de CV; 

Lehigh Cement Company; Ash Grove Cement Company; Essroc Cement Corporation; Lone Star 

Industries, Inc.; RC Cement Company; Texas Industries, Inc. (TXI); and California Portland 

Cement Company. 

The 1997 Economic Census report (sampling 279 establishments) shows that refractories (clay or 

non-clay) account for 2.5% of the total cost for the cement industry, or approximately $61.5 

million of the total $2.5 billion spent. 

A typical modern cement plant may include the following refractory lined units: (1) a preheater 

equipped with several cyclones as a heat exchanger, (2) a precalciner with a burner for the 

decarbonization process, (3) a kiln, and (4) a kiln cooler. The preheater, cyclones, and heat 

exchangers heat the cement raw meal or slurry before it is passed to the rotary kiln. This aids fuel 

economy by decreasing the amount of heat that the kiln must supply. The precalciner uses 

preheated combustion air from the clinker cooler and/or kiln exit gases with separate burners to 

calcinate up to 95% of the raw material. A precalciner kiln system consisting of a rotary kiln with 

an external furnace may also be used for heating cement raw meal to calcination temperatures. 

This system also often has a multistage cyclonic preheater attached to it for further heating 

efficiency. 

The kiln is where cement raw mix is dried, calcined, and burned into clinker at temperatures of 

2600–3000°F. Kilns can be rotary (rotating at one to three revolutions per minute) or shaft type 

and use coal, oil, gas, or other fuel for heating. A rotary kiln will generally be divided into zones 

30 
Appendix A — Chemical and Petrochemical 

for drying, preheating, calcining, burning, and cooling. Kilns having a preheater and/or 

precalciner do not usually contain the first three zones. The molten batch emerges from the lower 

end of the kiln in the form of very hot, marble-sized chunks known as “clinker.” Once out of the 

kiln, the clinker goes through a cooling process in the kiln cooler. When cool, the clinker is ready 

to pass through a series of grinding and milling processes that result in the gray powder we know 

as cement. 

The kiln is the heart of the plant and is generally lined with fired bricks. Monolithic refractories 

are used only in tapered parts of the inlet chamber, the manhole of the calciner, and the charging 

inlet of the kiln. These monolithics are required to posses high resistance to reaction with cement 

raw materials, high resistance to penetration/infiltration of alkalis, as well as high resistance to 

mechanical stress. Generally, dense low-cement, high alumina castable with service temperatures 

of 1700–1800°C (3092–3272°F) is used here. 

A typical long kiln will include a drying zone, a preheating zone, a calcining zone, a burning 

zone, a transition zone, and a cooling zone to create a temperature profile in the kiln. Generally, 

fireclay bricks or high-alumina bricks are used in the drying and calcining zones. Lightweight 

insulating bricks are used to reduce the heat loss from the kiln. In these zones, refractories are 

subjected to abrasion and structural spalling caused by reactions with alkali from the cement raw 

materials. The temperature is relatively low in these zones, but because of the use of various 

wastes as fuel and constant temperature fluctuations, the refractory wear can become severe. Fuel 

types vary among petroleum coke, tires, solid wastes, etc., resulting in increased concentrations of 

alkalis (Na, K), sulfur (S), and chlorine (Cl) as well as other environmentally detrimental metals 

(Pb, Zn). As a results, the refractories are subjected to harsher alkaline sulfate [(K3Na(SO4)2, 

K2SO4] and alkaline chloride (KCl, NaCl) condensation at cooler zones. Sulfur additionally 

combines with the cement component (CaO) to form CaSO4 cotectic liquid. These liquid and 

vapor phases infiltrate into bricks, resulting in refractory densification, increased thermal 

conductivity, and thermal expansion which ultimately reduces the structural flexibility of kiln 

linings. Additionally, the increase in thermal conductivity often damages the steel shell. 

Monolithic refractories are also used in sections of cement kilns, typically at the outlet where 

wear rates are very high. At these zones, dense silicon carbide–based castables are often used. 

The cyclones in the preheater section of cement plants are also lined with dense castables, due to 

the complexity of their shapes and the wear resistance required. The shell of the exhaust gas duct 

is made from both dense and insulating castable material. In the cooler section of the plant, 

monolithic refractories are used to line the ceiling. The upper and middle cyclones of the calciner 

are generally lined with fireclay castable or equivalent gunning materials possessing a service 

temperature of 1400–1500°C (2552–2732°F). One common problem with this arrangement arises 

when cement raw materials reach the inlet chamber and riser because they adhere (stick) to the 

refractory lining and thus cause clogging. This clog is either mechanically removed or removed 

with high-pressure water. 

In the United States, as in most industrialized countries, two types of manufacturing processes 

have become prevalent in the cement industry; these are known as the “wet process” and the “dry 

process.” Although in many respects these processes are the same, or very similar, in the older 

wet process, ground raw materials are mixed with water to form a thick liquid slurry, while in the 

dry process crushed limestone is used and raw materials are mixed together without the addition 

of water, leading to higher energy efficiency, since drying is not needed. Currently there is a 

general trend in the United States to convert wet kiln technology to dry kiln technology; as a 

result, in 2000 there were 32 wet and 77 dry kilns. 

31 
Appendix A — Chemical and Petrochemical 

The major fuels used in U.S. cement plants are coal and coke, or coal and petroleum, and are used 

for the clinkering process. After 1980, however, other fuels—including tires, solid and liquid 

hazardous wastes, and nonhazardous wastes such as oils and solvents—began to be gradually 

used. The use of these alternative fuel types has significantly increased in recent years, as shown 

in Table A.1. Overall, this suggests that there have been no major changes in energy-efficient fuel 

use in cement kilns since 1990, although more modern plants have been built. The survey, by the 

Portland Cement Association (PCA), suggests slightly lower average energy uses of 4.73 GJ per 

million metric ton of clinker and 4.91 GJ per million metric ton of Portland cement for 2000. 

Further, the 1997 Economic Census report (sampling 279 establishments) showed that energy 

costs (fuel costs and purchased electricity costs) account for 40% of the total cost for the cement 

industry, with approximately $485 million spent on fuel and a similar amount spent on 

electricity.

7 

Table A.1. Feedstock for energy use in the cement industry, 1993-2001 (thousand metric tons) 

1993 1994 1995 1996 1997 1998 1999 2000 2001 % 

change 

Coal 10,034 10,484 8,241 8,764 9.035 9,066 9,206 10,095 10,240 2.05 

Coke 

from coal 

NA NA 455 458 351 432 343 442 420 –7.69 

Petroleum 

coke 

NA NA 1,475 1,295 1,288 1,197 1,622 1,351 1,370 –7.11 

Fuel oilsa 46 49 42 64 86 73 82 124 93 102.2 

Natural 

gasb 

668 650 1,069 710 672 720 653 338 397 –40.5 

Tires 70 120 158 191 277 269 685 374 300 328.6 

Solid 

wastes 

90 74 68 72 68 74 816 1,016 320 255.6 

Liquid 

wastesa 

744 600 885 910 835 1,268 906 929 829 11.4 

a

 Millions of liters. b

 Millions of cubic liters. 

Barriers 

Refractories play key roles in calcining operations, the core production method for the cement 

and lime industries. It is estimated that almost all of the energy consumption is affected by 

refractories. Recommended improvements for refractories include applications for all of the 

thermal zones of typical rotary calciners. Based on subsequent analyses of typical furnaces and 

unit processes, improvements to refractory systems for the cement and lime industries could lead 

to opportunities for energy savings of up to 42.3 TBtu/year. The specific barriers are shown 

below. 

• Unavailability of chrome-free refractories for use in the transition zones of lime kilns, where 

refractories are subjected to large temperature fluctuations by the repeated detachment and 

adhesion of the cement coating. 

• Unavailability of refractories with adequate corrosion resistance and adhesion (coating) 

ability for use in the burning zone of lime kilns, where temperatures reach to above 1450°C 

(2642°F). 

32 
Appendix A — Chemical and Petrochemical 

• Unavailability of refractory monolithics with high abrasion resistance as well as high 

resistance to mechanical and thermal spalling for the inlet chamber and riser of kilns to 

replace less adherent silicon carbide castables and precast shapes. 

• Unavailability of monolithic refractory materials with resistance to alkali reaction for use as 

clinkers in the cooling zone of the kiln to replace low-cement, high-alumina castables used in 

the area. 

R&D Pathways 

• Develop monolithics with better abrasion resistance and low or no adherence properties. 

• Develop methods to modify the surfaces of clinker zone brick, so as to encourage the 

formation of protective coatings by physically altering the refractory surface or through 

chemical alteration of the surface. Develop and improve magnesia-spinel brick and 

demonstrate the effect of zirconia additions. 

• Identify non-chrome-containing refractories which possess increased corrosion resistance to 

the environment present in cement manufacturing. 

• Develop nondestructive techniques for monitoring wear of refractories. 

• Develop and improve silicon carbide castables to produce products with improved thermal 

properties and even less adherence to cement raw materials. 

A.3 Chemical and Petrochemical 

Current Industry Operations and Refractory Use 

The chemical industries have processing equipment that spans a broad range of types, with 

associated processing temperatures ranging from as low as 370–425°C (700–800°F) to as high as 

1870°C (3400°F). The refractory-lined units in a typical full-spectrum refinery can include 

fluidized catalyst cracking units (FCCUs), catalytic reformers, fired heaters, boilers, hydrogen 

reformers, ethylene furnaces, incinerators, and carbon furnaces. Some of these units are designed 

to have substantial insulating capability in their linings, while others are low-temperature-process 

units, like FCCUs that contain only a single layer of refractory designed to withstand the 

extremely severe abrasion caused by the catalyst. 

The estimated cost of refractory lining installation and maintenance work for the chemical and 

petroleum industries is relatively low (as low as 1% of the total operating cost) compared to other 

equipment construction and maintenance costs; yet, refractory lining degradation and failure are 

leading factors for frequency and duration of shutdowns. Due to lower operating temperature, as 

compared to other industries, refractory is not normally affected by high temperatures, yet 

refractory degradation may be caused by erosion, corrosion, and mechanical stresses caused by 

thermal cycling, depending on location of the refractory within the unit. Another problem that can 

affect refractory reliability is the manner in which the refractory is installed (i.e., pneumatically 

gunned, shotcreted, vibration cast, pump cast, and/or hand packed). Additionally, with operating 

time, minor imperfections can become locations of exacerbated lining degradation. Lining design 

(i.e., anchoring, lining thicknesses, and design of construction/expansion joints) can also add to 

lining degradation. 

33 
Appendix A — Chemical and Petrochemical 

Also, the total tonnage of refractories in an average-sized full-spectrum refinery can be as high as 

1000 tons of refractory for a single lining that is intended to last for at least a decade: a typical 

refinery requires approximately 2500 tons of refractory. Some of the more critical units are put on 

maintenance, with annual turnaround, and the replacement and repair of refractories for a single 

unit can amount to 250 tons, or a fourth of the lining. 

Virtually all of the refractories used by chemical plants are monolithics. These include castables, 

cast vibration mixes, gunning mixes, low- and no-cement castables, plastics, and a variety of 

insulating refractories. Because many of the temperatures of refinery processes are too low to 

promote a ceramic bond to replace the hydraulic bond of calcium aluminate–bonded castables, 

the refractory practice has been based largely on phosphate-bonded plastics, which develop the 

needed properties at or below the process temperatures. Refractory practice varies widely from 

one company to another. Some companies have incorporated castables, especially vibratables, 

into their practice, and today companies are evaluating shotcretable refractories in locations 

where a ceramic bond can be developed at the process temperature. 

However, the place where advances in refractories may have the greatest impact on efficiency in 

the chemical industry is in the boilers used to produce steam for power generation. Improvements 

in the thermal efficiency of boilers themselves, through improved insulation, and in the insulation 

used on transfer pipework could lead to reduction in heat loss and lower fuel consumption. 

In the petrochemical and synthesized gas industries in particular, new designs in the catalyst 

support media and the installation of refractory insulation have increased operation life and 

efficiencies of reactors of various types. New refractory designs for support media have allowed 

increased flow rate and longer productive life for expensive catalyst materials. More attention 

being paid to the use of sprayable refractory ceramic fiber lining materials as well as new 

microporous silica backup materials. More attention is also being paid to the use of precast parts 

for component of the various unit processes, and wider use of precast technology is anticipated. 

Because it is so important to maintain shell temperatures of vessels and reactors above a threshold 

temperature to avoid condensation onto the shell, designers are trying to refine the models for 

predicting temperature gradients and cold face temperatures. Lack of data on the thermal 

conductivity of many commercial monolithic refractories is holding back progress in this area. 

Barriers 

Refractories and insulation materials play key roles in most of the varied unit operations involved 

in the production of chemical and petrochemical products; it is estimated that about 50% of the 

industries’ direct energy consumption is in processes affected by refractories. Based on 

subsequent analyses of typical furnaces and unit processes, improvement of refractory systems 

for the chemical and petrochemical industries could lead to energy savings opportunities of up to 

266.4 TBtu/year. The specific barriers are shown below. 

• Lack of understanding of high-temperature properties such as creep rupture, short-time 

strength, thermal conductivity, thermal expansion, ductility, and stability for the monolithic 

refractories used in this industry. 

• Lack of understanding of the corrosion properties of refractories used in the chemical and 

petrochemical industries, which subject materials to conditions ranging from hot-oxidizing 

environments to environments that contain hot halogen gases or gases with entrained 

particles. 

34 
Appendix A — Chemical and Petrochemical 

• Lack of understanding regarding turnaround times and maintenance costs, which have 

become increasingly more important cost elements. Lack of understanding regarding the 

reduction of energy consumption and air pollution. 

• Lack of understanding regarding high-temperature and corrosive environments which limit 

refractory life and increase maintenance costs in numerous chemical and petrochemical 

reactors. 

R&D Pathways 

• Develop improved refractories for minimizing mechanical damage to refractory-lined 

refining vessels through reduction of the amount of mechanical stresses on the lining. 

• Perform work to optimize the physical properties of refractories (i.e., thermal insulation, 

erosion resistance, corrosion resistance, thermal shock resistance, strength, porosity, and 

permeability) under known required operating conditions for specific unit types. Develop 

realistic and reproducible product development and quality control testing that can be used to 

relate step improvements in needed physical properties of the current refractory lining 

materials in service. 

• Develop improved methods of analysis for high-temperature equipment through techniques 

for online refractory inspection. 

• Develop longer-life refractories that are field-repairable, less brittle, and more shockresistant. 

• Acquire thermophysical, kinetic and mechanical data on refractories. 

• Develop ceramic-metal-matrix composites with improved stability and explore surface 

modification and/or coatings for use in corrosive environments. 

• Develop materials to withstand high temperatures (1000–3000°F) while retaining superior 

strength, ductility, corrosion and wear resistance to resist damage due to corrosion and high 

temperatures present in the chemical process industry. 

• Develop materials with enhanced resistance to organic acid environments and improved 

materials for chlorine-based processes, to improve plant operations and maintenance 

requirements and reduce costs. 

• Develop refractories and refractory coatings for high-temperature furnaces to reduce energy 

and maintenance costs. Develop high-temperature nonstick surfaces that do not degrade or 

become volatile at high temperatures to help reduce needed maintenance of chemical process 

equipment. 

A.4 Forest Products 

Current Industry Operations and Refractory Use 

Although the pulp and paper industry is not a substantial user of refractories, the forest products 

industry is the third largest consumer of energy in the United States, at approximately 3.3 quads 

per year.8

 To its credit, the industry generates 63% of its own energy using recycled woody waste 

products and other renewable sources for fuel (bark, wood, and pulping liquor). Since 1972, the 

industry has reduced its use of fossil fuels and purchased energy by about 2% while increasing 

total production by nearly 64%. In the future, black liquor and biomass gasification will lead to 

35 
Appendix A — Forest Products 

greater reductions in purchased energy and could even change the pulp and paper industry from a 

net energy importer to a net energy exporter. Gasification holds the promise of generating 175% 

more power than do current Tomlinson furnaces, but the gasification technology currently in 

place requires yearly refractory relines of the vessels.9 

Lime recovery kilns are used in most forest products industries as an integral component of the 

Kraft pulping process, with lime recovery kilns operating at a level of 7 million Btu/ton of lime 

recovered. Yet, these kilns have the disadvantages of cyclic production rates, poor product 

quality, dust cycles, rings, material buildups, and blockages. Most of these deficiencies can be 

related to premature brickwork failure and buildups, with loss in production of up to 35 days per 

year being attributed to ringing. 

Barriers 

It is estimated that about 40% of direct energy consumption by the forest products industry is in 

processes affected by refractories. Based on subsequent analyses of typical furnaces and unit 

processes, improvement of refractory systems could lead to energy savings opportunities of up to 

129 TBtu/year. Specific barriers are shown below. 

• Lack of understanding regarding problems associated with sticking of reactants and products 

to the refractory linings of lime kilns, limiting their operating lives. 

• Unavailability of materials with corrosion resistance to the harsh, high-alkali black smelt 

environment (consisting of steam, sulfur, hydrogen, and CO) and abrasion/erosion resistance, 

spall resistance, and thermal compatibility with metallic materials used in the gasifiers. 

• Unavailability of improved materials for containment which do not react with the high-alkali 

atmosphere present in the gasifier. Such reactions result in slabbing, which limits the lining 

life. 

• Lack of a database on the properties of the castable materials used in this industry, 

information which is not provided by the refractory producers. 

• Unavailability of improved refractories for Tomlinson recovery boilers to replace fireclay, 

alumina brick, or similar monolithic compositions used in the floor. Improved refractories 

could decrease the amount of furnace downtime needed for repairs and would decrease 

energy usage due to shutdown and restart cycles. 

• Lack of refractory systems with sufficient alkali corrosion resistance for storage container 

linings used for hot pulping chemicals. 

R&D Pathways 

• Develop improved refractories for lime kilns to reduce downtime due to ringing and buildups, 

thus leading to increases in energy efficiency due to better thermal management and increases 

in lime kiln lining lifetimes. 

• Develop advanced refractories to provide suitable materials for use in black liquor and 

biomass gasifiers. Advanced refractories could lead to substantial improvements in energy 

efficiency for the forest products industry, which currently requires yearly refractory relines 

of the vessels. 

• Develop materials with greater resistance to the corrosion in molten salt environments, 

greater abrasion/erosion resistance, and greater spall resistance. In addition, develop alternate 

refractory schemes to provide refractory surfaces which are more impervious to wear and 

36 
Appendix A — Forest Products 

corrosive attack through coatings, chemical alteration of refractory surfaces, physical 

alteration of refractory surfaces, etc. 

• Develop refractory materials which are more thermally compatible with metallic materials 

used in gasifiers. 

A.5 Glass 

Current Industry Operations and Refractory Use 

The glass industry produces a wide variety of products using different process technologies. 

Perceived differences across the segments of the glass industry remain a barrier to broad-based 

industry collaboration. These differences have led individual segments to select different melting 

technologies, prefer different refractories, and operate furnaces that differ in size and scale. 

Capital and operating costs for the individual segments reflect these product and process 

differences. Further, differences in glass types and industry segments and the needs of large 

versus small furnaces limit the ability to define common interests. There are also concerns with 

project definition and management for cooperative efforts: costs, timing, and sharing of technical 

and financial risks and the intellectual property generated. 

Melting of the batch is accomplished using different types and sizes of furnaces, with refining 

occurring in the melting chamber. Furnaces may be discontinuous (<5 ton/day) or continuous 

(large operations, estimated to last up to 10–12 years today). Continuous furnaces may be directfired (units fired with natural gas and producing 20–150 ton/day), recuperative furnaces (directfired furnaces fitted with recuperators to recover heat from exhaust gases), regenerative furnaces 

[the most commonly used furnaces, with capacities of 100–1000 ton/day, and with the furnace 

heat collected in a regenerator that is used to preheat combustion air as high as 1260°C (2300°F) 

and achieve higher energy efficiency] or all-electric melters. 

Commonly, alumina-zirconia-silica refractories are used as sidewalls, with zirconia content 

ranging from 33 to 41%. Chrome-containing refractories are also used in the fiberglass industry 

because of their additional corrosion and erosion resistance, but their use is decreasing because of 

environmental concerns about chrome. Refractories containing chrome are not used in the flat, 

container, or pressed/blown glass industries because chrome impurities can discolor the glass. 

Silica is commonly used for the crown of the furnace, or sometimes mullite if oxy-fuel firing is 

employed. Due to chemical incompatibilities, a layer of alumina-zirconia-silica is often used as a 

buffer between the crown and sidewall.1 

Significant progress has been made in reducing energy intensity in areas of the glass industry 

over the last ten years. This increase in efficiency has been accomplished mostly through 

improved process control systems, the development and use of advanced refractory materials, and 

technologies such as oxy-fuel firing and electric boost, which increase production capacity. 

Advanced technology has reduced the fuel consumed per ton of glass melted by 25% since the 

early 1980s, and energy use has declined in the fiberglass sector by 30% since 1978.10 

Still, energy efficiency depends on the size of the operation and the production rate. For example, 

a large (120 ton/day) all-electric furnace uses electricity at 66% overall efficiency, while a small 

all-electric furnace (4 ton/day) has about 34% overall efficiency.

11 Electric glass-melting furnaces 

are more thermally efficient than oxy-fuel because there is less mass (melt basin only) and less 

surface area for heat losses. Average oxy-fuel melters consume 3–5 MMBtu from fossil fuel per 

37 
Appendix A — Glass 

ton of glass melted. This is in comparison with the average electric consumption of all-electric 

direct resistance glass melters of about 2.7 MMBtu/ton. 

Yet, modern glass melting still requires a high-temperature device that consumes a significant 

quantity of energy, with melting energy representing about 15% of the manufacturing costs. 

Although there are some proven energy-reduction technologies for melting that are not currently 

being implemented and some concepts which have been identified to further reduce energy 

consumption (improvements in corrosion-resistant refractories, insulation, combustion, process 

control), the economic incentive for adopting these proven technologies and developing these 

new concepts may not currently be sufficient to justify the cost and the effort. This may be 

because the industry views minimizing the energy cost per ton of glass produced as more 

important than reducing the energy content as measured in thermal units and because the value of 

the energy saved at the current cost of energy, with an assumed low rate of cost escalation, is not 

sufficient to justify many energy-saving technologies. In particular, technologies that require any 

significant capital investment are usually discarded by the glass industry or looked upon 

unfavorably, with many glass manufacturers expecting a one- to two-year payback for capital 

investments.12 Specific energy usage by sector is shown in Table A.2. 

Table A.2. Estimated energy use in melting and refining 

Av. specific Annual 

Glass sector energy use energy use 

(106

 Btu/ton) (10 TBtu) 

Flat 8.6 42.9 

Container 5.5 53.1 

Pressed/blown 7.3 18.1 

Fiber 8.4 25.6 

Source: DOE, Energy and Environmental Profile of the US Glass Industry, 

DOE/GO -102002-1590, Office of Industrial Technologies, April 2002. 

Oxy-fuel firing is also still a viable technology and could improve efficiency in the glass industry. 

The trend towards using oxy-fuel firing is steadily increasing because an oxy-fuel furnace can 

produce the same amount of glass as with gas and/or air, but at a lower fuel input.13 Additionally, 

when regenerative furnaces are converted to oxy-fuel firing, the regenerator refractory structure is 

not needed, eliminating the exhaust volume by as much as 75%.

14 As of 1998, only 20% of the 

glass furnaces in the United States had been converted from air-fuel to oxy-fuel technology, with 

the greatest conversion rate found in the fibers sector.15 To date, it is estimated that the number of 

conversions has increased only to around 25% of the total industry.15 A breakdown of the number 

of conversions to oxy-fuel technology completed by 1998 in each sector of the glass industry is 

shown in Table A.3, along with estimates of the number of conversions made to date. Therefore, 

there is still a great opportunity for further improvements in the energy efficiency of the glass 

industry through further utilization of oxy-fuel fired furnaces. Further improvements in refractory 

technology are needed to fully utilize the benefits of oxy-fuel technology. 

Boosting with oxygen firing will result in increased production but will also increase the melting 

cost per ton of glass. This trade-off is accepted by the industry because the additional capacity 

gained is at a much lower capital cost than that associated with building new furnaces. Sampling 

of the industry indicates that to date, 125–150 of 500–550 total glass furnaces have been 

38 
Appendix A — Glass 

Table A.3. Conversion of North American glass melters to oxy-fuel firing as 

of 1998 and as of 2004 

Glass sector No. of oxy-fuel 

conversions 

Total no. of 

glass tanks 

% converted 

Containers 23 210 10 

Fibers 41 110 37 

 Textiles 

 Insulation 

(30) 

(11) 

(70) 

(40) 

(42) 

(27) 

Flat 2 45 5 

Specialty 54 235 23 

 All sectors 120 600 20 

Containers 27 180 15 

Fibers 46 100 46 

 Textiles 

 Insulation 

(35) 

(11) 

(62) 

(38) 

(56) 

(29) 

Flat 3 45 7 

Specialty 60 225 27 

 All sectors 136 550 25 

As of 1998 

As of 2004 (estimated) 

converted to oxy-fuel firing. This is only 25% of all the furnaces in the industry, leaving a large 

number of furnaces that could be converted to the more efficient technology, leading to pollution 

abatement and higher energy efficiency, if improved alternate refractories with sufficient 

corrosion resistance and thermal efficiency could be found. Of the various segments making up 

the glass industry, the fiber area has made the most use of oxy-fuel firing, with nearly 50% of 

these furnaces having been converted; in second place is the specialty area, with over 25% of the 

furnaces converted. Great opportunities exist in the flat and container glass segments, which have 

conversion rates on the order of 5 to 15%. 

Barriers 

Refractories and insulation materials play key roles in many of the unit operations involved in the 

glass production; it is estimated that about 70% of the direct energy consumption is in processes 

affected by refractories. Based on subsequent analyses of typical furnaces and unit processes, 

improvement of refractory systems for the glass industry could lead to energy savings 

opportunities of up to 13.9 TBtu/year. Specific barriers are shown below. 

• Unavailability of durable and inexpensive materials for use for refractories in constant 

contact with molten glass, combustion gases, and hot glass products in glass furnaces which 

operate continuously at temperatures exceeding 1400°C in the molten bath and 1600°C in the 

combustion atmosphere for operating lives of 8–10 years. 

• Unavailability of parts and installation equipment to replace parts and materials that fail 

during production campaigns, without major furnace shut-downs and loss production. 

• Limitations on energy efficiency and degrading glass quality, particularly in oxy-fuel 

environments, due to limitations of refractory performance. 

• Lack of knowledge that would improve the performance and quality of existing refractories 

and lead to increased production life and lower installation costs. 

39 
Appendix A — Glass 

• Lack of modeling of corrosion mechanisms that would help determine solutions for reducing 

corrosion in refractories and for providing adequate on-line sensors that would optimize 

operation control, leading to longer-life, corrosion-resistant refractories, thereby reducing 

capital costs for glass melting and forming applications. 

• Lack of knowledge in the form of a materials property database, identification of new melting 

techniques and their associated refractories, and better methods for the evaluation and repair 

of refractories in service. 

R&D Pathways 

• Develop integrated furnace models for the prediction of methods for improving thermal 

efficiency, for developing new glassmaking technologies, and for projecting performance of 

alternate refractory materials. 

• Develop refractories (particularly for furnace crowns) having sufficient strength and 

corrosion resistance at higher operating temperatures when using oxygen enrichment. 

• Develop monolithic and/or coated refractories that will be more resistant to penetration of 

corrosive liquid and gaseous species. 

A.6 Heat-Treating and Annealing 

Current Industry Operations and Refractory Use 

Heat-treating is a $20 billion industry in the United States and represents $75 billion or more in 

value-added manufacturing globally.

16 It is also crosscutting, enabling technology that affects 

most of the industries found in the United States, including aluminum, glass, iron and steel, and 

metal casting. The goal of the Heat Treating Industry Vision 2020 is to reduce energy usage in the 

heat-treating industry by 80%. To achieve this goal, cooperation is needed between all the related 

industrial groups, including the heat treating community, metal producers, foundry and forging 

groups, fabricators, government, and academia. 

One of the most important research needs for this industry is the development of integrated 

process models. Heat treating today is viewed as an experience-based art. However, in the future 

it must function in a scientific, predictive environment, where processes are truly understood and 

capable of being simulated. For example, it has been demonstrated that the solutionizing time for 

a cast aluminum alloy could be reduced from 12 to 2 h with no loss of quality or properties. This 

magnitude of heat treatment cycle time reduction could result in increased productivity and/or 

reduced energy consumption of greater than 50%. Validated databases for multicomponent alloys 

and predictive models will enable comparable results to be achieved for a wide range of alloys 

and applications.17 

Barriers 

Refractories and insulation materials play key roles in various furnace systems used in the heattreating and annealing industries; approximately 90% of direct energy consumption is in 

processes affected by refractories. Based on subsequent analyses of typical furnaces and unit 

processes, improvement of refractory systems for the heat treatment and annealing industries 

could lead to energy savings opportunities up to 45 TBtu/year. Specific barriers are shown below. 

40 
Appendix A — Heat Treating and Annealing 

• Unavailability of high-performance refractories for the improvement of thermal efficiency in 

furnace systems that would lead to reduced cycle times, lower required energy input, and 

lower processing costs. 

• Unavailability of gradient refractory systems with lower thermal conductivity and hightemperature strength and corrosion/erosion resistance for reducing heat losses and improving 

chemical compatibility. 

• Unavailability of refractory materials with longer lifetimes due to improved corrosion 

resistance, extended temperature resistance, and mechanical thermal stability, while also 

possessing reduced costs. 

R&D Pathways 

• Develop improved insulation materials to achieve higher operating temperatures, reduce heat 

losses, and eliminate reheating while also reducing process time and production costs. 

• Develop higher-efficiency insulation materials to increase the thermal efficiency of current 

processes and simultaneously lower heat-treatment times, leading to significant reduction of 

energy consumption levels. This could then lead to the development and application of 

accelerated heating technologies, which will require alternative insulation materials with 

greater thermal shock resistance and high-temperature stability. 

A.7 Iron and Steel 

Current Industry Operations and Refractory Use 

In the steel industry, competitive pressures such as worldwide overcapacity, the high cost of 

energy, environmental and safety regulations, competing materials, customer demand for high 

quality, and the high cost of capital have led to a need for sweeping changes across the industry. 

Therefore, four critical areas for improvement or targets of opportunity were identified: 

(1) process efficiency through improved throughput, quality, and energy efficiency, (2) increased 

recycling, (3) new processes to avoid or reduce air and water pollution, and (4) product 

development.18 

As part of the approach to address these areas, refractories currently used throughout the steel 

making process were analyzed. Oxide-based ceramics of aluminum, calcium, zirconium, 

chromium, magnesium, and carbon are primarily used in steel contact in an effort to reduce 

contamination and interaction with the highly alkaline slags. Nonoxide ceramics consisting of 

combinations of silicon, titanium, tungsten, and aluminum are used for hot handling applications 

where thermal shock resistance, creep resistance, wear resistance, and resistance to acidic species 

is desired. These materials may be in the brick or castable form. In many areas of steel furnaces 

refractory life is extended by splashing slag against the refractory walls. This slag solidifies on 

the wall, forming a protective layer. This process, though highly successful, is very inefficient 

thermally. 

Today, steelmaking from pig iron commences in the transfer ladle ahead of the basic oxygen 

furnace (BOF) and continues after the BOF to the ladle stations and on to the tundish, which 

distributes the steel into the continuous casting or stationary molds. Refractories for all these unit 

processes are selected on the basis of their longevity and the cleanliness of the steel that passes 

through them. The cost per pound of refractories for all the various components, with few 

41 
Appendix A — Iron and Steel 

exceptions, has increased markedly with the awareness that advanced refractories can meet the 

goals of greatly extended service and contribute to the quality of the steels. 

In steel ladles, maintenance costs are high and heat efficiency is poor. Currently, high-alumina 

and magnesia-based castable refractories are preferred for ladle linings at and below the slag line 

due to the high process temperatures, erosion from the stirred bath, and the high corrosiveness of 

the slag. Life is short for these refractories; they typically last for only 30 to 50 heats. Highertemperature-capable oxide/carbon mixtures have been evaluated for use to provide a balance 

between purity, thermal shock resistance, and erosion resistance, along with other alternative 

refractories with longer life, lower cost, and safer disposal. Yet, no material of choice has been 

identified. 

Monolithics (refractory concretes and gunning mixes) comprise at least 60% of all refractories 

used in iron and steelmaking in the United States today (over 70% in Japan), and there is a strong 

trend to increased use of precast versions of monolithics. Installation practice has undergone 

some major advances which translate to much more rapid original construction, replacement of 

working linings, and repair. The adaptation of shotcreteing methods from construction concrete 

technology has resulted in significant reductions in downtime and in some cases longer service 

life as well. In the ladle lining practice, it has been possible to successively repair the original 

lining such that complete relining is no longer necessary. 

Improved materials for runners (used to flow the molten steel from a BOF or an electric arc 

furnace to the refining ladles) are needed. These components are generally constructed from 

fireclay brick, alumina castable, or chromia castable. They experience corrosion, wear, and 

thermal stresses that limit their lives to one to two weeks. Similar improvements are needed for 

water-cooled pipes and stirrers enclosed in refractory. They also experience corrosion, wear, and 

thermal stresses that limit their life and require weekly repair. 

Energy consumption in the various unit processes found in the iron and steel industry has 

changed substantially due to advances in refractory practice, but it is difficult to identify the role 

of refractories in energy savings because many changes are being made in designs and 

combustion practices as well as in metallurgical and slag practices. Some unit processes like blast 

furnaces and BOF melters are designed to leak energy to induce a steep thermal gradient through 

the lining. In an extreme example of this process design concept, the walls and roofs of electric 

arc furnaces are water-cooled to cause the steepest possible thermal gradient. Consequently, it is 

in the other unit processes where refractory practice contributes to energy savings in several ways 

(e.g., more efficient insulation, extension of the time between maintenance shutdowns, reduction 

in the amount of refractory scrap and reduction in metal scrap through improved cleanliness of 

the final steel products). 

As mentioned, blast furnaces and BOF melters are designed to leak energy, and even in the case 

of most ladles, insulation is minimized in the walls and bottoms in an effort to induce a thermal 

gradient in the interest of reducing the refractory wear. Ladle covers, on the other hand, are 

absolutely necessary to conserve the temperature of the contained metals and are made from 

increasingly advanced materials. There are many opportunities for direct energy savings in all of 

the unit processes and components in iron and steel plants that are not designed to incorporate 

steep thermal gradients. There are also opportunities to extend refractory life through the use of 

lining thickness monitors and other sensors and instruments which improve the combustion 

efficiency and reduce stack losses as well as release of undesirable products of the processing of 

the iron and steel. 

42 
Appendix A — Iron and Steel 

Significant improvements are needed for meeting operating challenges for monolithic 

refractories. Monolithic linings possess many inherent problems. Ideally, steelmakers would like 

to have monolithic linings for all of the ladles and even for the basic oxygen furnace. The socalled basic lining has received considerable research attention, but no reports of success in a 

BOF have been reported to date. Problems exist with the hydratability of the basic components of 

the products, which are based on MgO, magnesite, doloma (calcined dolomite, CaCO3·MgCO3), 

and spinel. Further, monolithics for BOF linings need to incorporate substantial amounts of 

carbon in the form of graphite, and research is ongoing to resolve the problem that derives from 

the interference of the high volumes of graphite with the binder systems of the monolithics. 

Barriers 

Refractories and insulation materials play key roles in the smelting, refining, and melting 

operations used in the iron and steel industries; it is estimated that 85% of the industries’ direct 

energy consumption is in processes affected by refractories. Based on subsequent analyses of 

typical furnaces and unit processes, improvement of refractory systems for the iron and steel 

industries could lead to energy savings opportunities up to 148.5 TBtu/year. Specific barriers are 

shown below. 

• Lack of adequate corrosion resistance for refractories in contact with molten metal, which 

leads to undesirable equipment lifetimes. This is especially important for refractories used in 

parts of furnaces, ladles, and tundishes. 

• Unavailability of improved materials for runners (used to flow the molten steel from a BOF 

or an electric arc furnace to the refining ladles), water-cooled pipes, and stirrers enclosed in 

refractory. 

• Lack of utilization of advanced monolithics in ladles due to problems with their explosive 

tendencies and lack of data on some of the key input properties and parameters for most of 

the monolithic products that meet the requirement for steel ladle practice. 

• Unavailability of improved impact pads at the bottom of ladles, which experience high wear, 

thermal shock, and abrasion. 

• Unavailability of thin-walled alumina refractory ladle shrouds that are used to reduce oxygen 

pick-up during pouring with lifetimes of greater than ten heats in the extreme conditions of 

this application. 

R&D Pathways 

• Develop methods for suppressing the hydration of magnesites. 

• Develop refractory linings with improved life and reduced maintenance requirements, to 

reduce the yearly or even monthly maintenance shutdowns generally necessary to repair 

refractory problems. Reducing the frequency of or eliminating these shutdowns would 

provide significant cost reduction, not to mention associated energy savings. 

• Develop advanced materials (such as refractories with in situ coatings or externally applied 

coatings) for resisting corrosion/erosion damage. 

• Develop new materials along with improved materials databases for members of the industry 

to adapt to new materials in their operations. 

• Develop improved refractories for use in various parts of furnaces, heat processing 

equipment, ladles, and tundishes. 

43 
Appendix A — Iron and Steel 

o Develop improvements in runners, water-cooled pipes and stirrers, which all 

experience corrosion, wear, and thermal issues that can limit their life and require 

weekly repair. 

o Develop improvements in materials and installation methods for ladle bottoms and 

impact pads by using ceramic coatings or alternate localized materials. 

o Develop long-lived, economically feasible, prefired refractory crucible linings for 

induction melting of stainless steel in air-melt operations. 

o Develop improved ceramic/refractory materials for net shape steel casting, 

eliminating rolling mill operations. 

A.8 Metal Casting 

Current Industry Operations and Refractory Use 

Energy consumption for the metals casting industry is on the order of 0.2 quads per year.20 This 

resulted in an estimated energy cost for the foundries industry in the United States of $1.2 billion 

in 19977

 and $1.5 billion in 1998,20 or about 10% of the total material cost put into production 

and 4–5% of the value of total shipment.21 Therefore, the industry wishes to optimize melting, 

holding, and heat-treating furnaces to conserve energy (i.e., to reduce energy usage by 20%) and 

minimize pollutant emissions. Further, the industry wishes to increase the life of refractories and 

other ceramic components that are exposed to molten metals and/or furnace environments and 

develop in-line sensors that are stable in the molten metal environment.10 

The unit processes utilized in the metal casting industry are lined with a wide range of refractory 

compositions (including silica, alumino-silicate, high alumina, zircon, magnesia, spinel, chrome, 

and magnesia-carbon) and forms (monolithics, precast shapes, and bricks). Most metals casting 

industry melting and holding furnaces are lined with ceramic refractories which are selected to 

minimize reaction with the specific metal being processed. Major refractory lined units include 

reverberatory furnaces, crucible (pot) furnaces, channel induction furnaces, coreless induction 

furnaces, electric arc furnaces, and pouring ladles. These furnaces are lined with wide ranges of 

refractory compositions, including silica, alumino-silicate, high alumina, zircon, magnesia, spinel, 

chrome, and magnesia-carbon. 

Substantial problems do still exist with molten material containment and handling. For example, 

cupola linings for ferrous alloy melting usually require weekly repairs because of corrosion and 

mechanical damage. An alternative to refractory linings is “skull melting” (in titanium melting, 

for example) where in a mechanism similar to the slag splashing in the steel industry or the 

formation of a “frozen ledge” in the aluminum industry, a protective layer of metal is solidified 

on the containment vessel surface, creating a corrosion-resistant layer but lowering the thermal 

efficiency of the system. 

Iron foundries are the major energy consumer for this sector, representing about 12% of the 

industry, followed by steel foundries at about 10%. Within each process, approximately 55% of 

the total energy cost is used for the melting of the metal, with other processes (core making, mold 

making, heat treatment, and post-casting processes) accounting for the other 45%. Although the 

total energy cost (fuel + electricity) accounts for only about 10% of the total material cost, it can 

44 
Appendix A – Metal Casting 

be as high as 20% or more for many foundries depending on efficiency and the materials 

processed. Specific energy usage by the metals casting industry is given in Table A.4. 

Table A.4. Process and energy data for U.S. metals casting industry 

Process Energy consumption 

Cupola melting Total: 8.6 × 106 Btu/ton 

Coke: 5.8 × 106 

Electricity: 1.1 × 106 

Natural Gas: 1.6 × 106 

Oxygen: 0.1 × 106 

Core making and mold making 7–20% energy, 1–3 × 106 Btu/ton 

Electric melting Induction furnace: 4.3–7.6 × 106 Btu/ton 

Electric arc furnace: 4.3–5.2 × 106 Btu/ton 

Electric-resistance heated reverberatory 

furnace: 5.2–7.9 × 106 Btu/ton 

Gas-fired melting furnace Crucible: 1.8-6.8 × 106 Btu/ton 

Reverb.: 2.5-5.0 × 106

 Btu/ton 

Refining and pouring processes Electricity: 0.4 × 106 Btu/ton 

Natural gas: 0.2 × 106 Btu/ton 

Cleaning and finishing processes 1.0 × 106 Btu/ton 

Investment casting Electricity: 3.0 × 106 Btu/ton 

Natural gas: 4.0 × 106 Btu/ton 

Lost foam casting 2.5 × 106 Btu/ton 

Die casting Electricity: 4.0 × 106 Btu/ton 

Natural gas: 5.0 × 106 Btu/ton 

Other casting processes Permanent: 2.5 × 106 Btu/ton 

Shell mold: 2.5 × 106 Btu/ton 

Plaster mold: 3.0 × 106

 Btu/ton 

Source: Energy and Environmental Profile of the U.S. Metal Casting Industry. 

Although the service life of refractory linings for all foundry practices has been extended greatly 

in the last decade, many of the problems identified above still persist. For example, most channel 

and coreless induction furnaces in iron and steel foundries use spinel-forming dry-vibratable 

castable, a shift from the wet ramming practiced in the past. This extended the service campaign 

of the foundries, but metal penetration and chemical corrosion by slag still persists. Also, there 

are few case studies that show good refractory maintenance management, new refractory 

development, or proper selection methods resulting in great energy savings and reduced melting 

cost. With the ever-increasing demand for quality cast metal, alloy composition, and innovative 

melting technologies, refractories sections have been pushed to more challenging positions. 

Research programs in refractories science and technology in the United States have never been as 

large and as comprehensive as R&D programs for other industry and technology sectors. 

Nonetheless, the metal casting industry is still one of the largest industry sectors in the United 

States, and the refractories industry is essential to it. 

For steel casting, wear of refractories and reaction of the refractories with highly basic (calciacontaining) slag can generate defect-forming inclusions. High-alumina refractories are generally 

used throughout this area of the process, with application of zirconia-based materials in areas of 

high wear and thermal shock and magnesia-based materials in areas of high slag corrosion. 

45 
Appendix A – Metal Casting 

Carbon materials cannot be used for this application, as they react with oxide-based refractories 

to form CO and with aluminum in steel to form alumina. Similarly, silica-based refractories 

cannot be used because of contamination issues. 

There is a serious lack of data on the mechanical, thermal, and chemical properties of monolithic 

refractories; and much of the data that exists is lacking in detail or is obsolete. There is only one 

reference source for such key data as strength, modulus of elasticity, thermal conductivity, heat 

content, permeability, thermal shock and creep behavior; and this source was first available 15 

years ago. Hundreds of new and updated versions of monolithics are on the market today, but 

there is very little design and other technical data available aside from the product data sheets, 

which do not list much beyond bulk properties and some typical values (not to be used in design 

but only as guide for selection). 

Barriers 

Refractories are used throughout the various unit processes found in the metal casting industry; 

approximately 90% of the direct energy consumption is in processes affected by refractories. 

Based on subsequent analysis of typical furnaces and unit processes, improvement of refractory 

systems for the metal casting industry could lead to energy savings opportunities up to 

8.5 TBtu/year. Specific barriers are shown below. 

• Lack of methods and materials to reduce or eliminate refractory failure due to physical 

erosion and chemical corrosion by molten metal and slags in nearly all stages of metal 

melting, refining, and pouring processes where refractories make contact with molten metal. 

• Lack of methods and materials to reduce or eliminate refractory failure related to fusion of 

refractories and chemical corrosion by MnO and FeO in steel foundries where the molten 

steel is generally tapped at high temperatures [up to 2150°C (3900°F)], and the common 

practice is to use low-cost silica, zircon, or alumino-silicate ramming mixes and plastics, 

especially in those foundries that apply acid melting practices in smaller furnaces. 

• Lack of methods and materials to reduce or eliminate clogging or build-up formation in 

channels and throats is problematic for most metal casting operations. 

• Lack of methods and materials to reduce or eliminate refractory wear in electric arc furnaces 

for the melting of steel, where reduced refractory wear and energy savings would be 

especially beneficial. 

• Lack of methods and materials to reduce or eliminate wear and reaction of refractories with 

highly basic (calcia-containing) slags for steel casting. 

• Lack of data on the mechanical, thermal, and chemical properties of monolithic refractories, 

with much of the data that exists lacking in detail or being obsolete. 

R&D Pathways 

• Develop improved refractories that will resist physical erosion and chemical corrosion by 

molten metal and slags. Specifically, develop material systems for steel containment that are 

chemically stable with the molten metal and that provide mechanical integrity against wear 

and abrasion. 

• Develop improved refractories which are resistant to corrosion by MnO and FeO to prevent 

fusion of refractories in steel casting. 

46 
Appendix A – Metal Casting 

• Develop refractories which will resist the adhesion of molten metal that results in clogging or 

build-up formation in channels and throats. 

• Develop improved refractories for use in electric arc furnaces for the melting of steel that will 

resist wear and provide energy savings though improved insulation. 

• Determine the mechanical, thermal, and chemical properties of monolithic refractories used 

in the metal casting industry. 

• Develop novel low-cost and high-performance monolithics (in particular for iron, steel, and 

aluminum foundries), basic gunning mixes, and MgO-graphite castable for rapid repair of 

electric arc furnaces. 

• Develop and model slag chemistry compatible with refractory linings and maximize the slag 

foaming properties, along with efforts to reduce or eliminate repair and patching of electric 

arc furnaces due to refractory wear. 

47 
Appendix A – References 

A.9 References for Appendix A 

1. D. W. Freitag and D. W. Richerson, Opportunities for Advanced Ceramics to Meet the 

Needs of the Industries of the Future, DOE/ORO 2076, U.S. Department of Energy, Office 

of Industrial Technologies, December 1998. 

2. Aluminum Industry Technology Roadmap, The Aluminum Association, February 2003. 

3. A. T. Taberaux, “Recent Development in Refractories for Aluminum Cells,” pp. 271–81 in 

Proceedings of the International Symposium: Advances in Refractories for the 

Metallurgical Industries II, ed. M. Rigaud and C. Allaire, Montreal, Quebec, August 4–9, 

1996. 

4. D. Schubert, “Refractories for Primary Aluminum Production,” pp. 74–80 in Proceedings of 

the Thirty-Seventh Symposium on Refractories, St. Louis, Missouri, March 23, 2001. 

5. J. Canon, “Aluminum Furnace Design: Reduction of Metal into Metal Out Time,” pp. 237– 

53 in 39th Symposium on Refractories, STL Section of the ACerS, April 9–10, 2003. 

6. S. Quesnel, C. Allaire, and S. Afshar, “Corrosion of Refractories at the Bellyband of 

Aluminum Holding and Melting Furnaces,” pp. 321–27 in Proceedings of the International 

Symposium: Advances in Refractories for the Metallurgical Industries II, ed. M. Rigaud 

and C. Allaire, Montreal, Quebec, August 4–9, 1996. 

7. Economic Census 1997 

9. U.S. Environmental Protection Agency, Profile of the Pulp and Paper Industry, 2nd ed., 

EPA/310-R-95-015, 2002. Available online at http://www.epa.gov/compliance/resources/ 

publications/assistance/sectors/notebooks/pulp.html. 

9. E. D. Larson, S. Consonni, and R. E. Katofsky, A Cost-Benefit Assessment of Biomass 

Gasification Power Generation in the Pulp and Paper Industry, Princeton University, 

October 2003. Available online at http://www.princeton.edu/~energy/publications/ 

pdf/2003/BLGCC_FINAL_REPORT_8_OCT_2003.pdf. 

10. Glass: A Clear Vision for a Bright Future, U.S. Department of Energy, Office of Industrial 

Technologies, 1996. Available online at http://www.eere.energy.gov/industry/ 

glass/pdfs/glass_vision.pdf. 

11. All-Electric vs. Oxyfuel Glass Melting, EPRI Technical Report TR-112788, Electric Power 

Research Institute, April 1999. 

12. GMIC Advancing Glass Melting Technologies: A Technical and Economic Assessment. 

13. Energy and Environmental Profile of the US Glass Industry, DOE/GO -102002-1590, U.S. 

Department of Energy, Office of Industrial Technologies, April 2002. 

14. M. J. Damsell, M. L. Joshi, J. R. Latter, and D. B. Wishnick, Glass Technology 37, 114 

(1996). 

15. Personal correspondence with refractory and glass industry representatives. 

16. ASM Heat Treating Society, An R&D Plan for the Heat Treating Community, 2001. 

Available online at http://www.asm-intl.org/pdf/RD_Plan_Heat_treat_Community.pdf. 

17. WPI Integrated Heat Treatment Model for Aluminum Castings, available online at 

http://www.wpi.edu/Academics/Research/MPI/News/twoawards.html. 

48 
Appendix A – References 

18. Steel Technology Roadmap, American Iron and Steel Institute, December 2001. Available 

online at http://www.steel.org/mt/roadmap/roadmap.htm. 

19. Metal Casting Industry Profile, 2004. Available online at http://www.eere.energy.gov/ 

industry/metalcasting/profile.html. 

20. Metal Casting: Annual Report, 2000. Available online at http://www.eere.energy.gov/ 

industry/metalcasting/annualreport.html (2004 report). 

49 
Appendix B 

Appendix B: 

Rotary Kiln Analysis Schematic, Conditions, and Results 

Fig. B.1. Longitudinal section of a typical rotary kiln. Reprinted from Modern Refractory 

Practice, 3rd ed., Harbison Walker Refractories, Pittsburgh, 1950. 

Assumptions for alumina calcination, rotary kiln 

Plant production 500,000 tons Al2O3/year 

 Assume two kilns 

Output per kiln, no downtime 277,778 tons Al2O3/year 

With 90% uptime 250,000 tons Al2O3/year 

Output per day 685 tons Al2O3/d 

Output per hour 29 tons Al2O3/h 

Kiln diameter 3.5 m (11 ft) 

Kiln length 80 m (262 ft) 

Kiln exterior surface 880 m2 (9468 ft2) 

Reaction 2 Al(OH)3 → Al2O3 + 3 H2O 

Heat of reaction at 77°F 634 Btu/lb Al2O3

Assume reaction temperature of 1100°C 

2012°F 

 say 2000°F 

Heating products to 2000°F 

Al2O3 562 Btu/lb Al2O3

 3 H2O 532 Btu/lb Al2O3 

Total heat requirement 1728 Btu/lb Al2O3 

3,456,824 Btu/ton Al2O3 

Heat of reaction 

35.91 cal/mol Reaction 634 

64,638 Btu/#mol Heating 562 Al2O3 

634 Btu/lb Heating 532 H2O 

Total 1,728 Btu/lb Al2O3 

50 
Appendix B 

Results of alumina rotary kiln analysis 

(Temperatures in °F) 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 3,456,824 39.84% 

Wall Loss Energy 400,000 4.61% 

Exhaust Energy 4,821,030 55.56% 

Total Energy 8,677,853 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 3,456,824 42.46% 

Wall Loss Energy 400,000 4.91% 

Exhaust Energy 4,285,360 52.63% 

Total Energy 8,142,183 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 3,456,824 37.87% 

Wall Loss Energy 600,000 6.57% 

Exhaust Energy 5,071,030 55.56% 

Total Energy 9,127,853 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 3,456,824 40.36% 

Wall Loss Energy 600,000 7.01% 

Exhaust Energy 4,507,582 52.63% 

Total Energy 8,564,406 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 3,456,824 44.81% 

Wall Loss Energy 400,000 5.19% 

Exhaust Energy 3,856,824 50.00% 

Total Energy 7,713,647 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 3,456,824 47.17% 

Wall Loss Energy 400,000 5.46% 

Exhaust Energy 3,471,141 47.37% 

Total Energy 7,327,965 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 3,456,824 42.61% 

Wall Loss Energy 600,000 7.39% 

Exhaust Energy 4,056,824 50.00% 

Total Energy 8,113,647 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 3,456,824 44.85% 

Wall Loss Energy 600,000 7.78% 

Exhaust Energy 3,651,141 47.37% 

Total Energy 7,707,965 100.00% 

Savings for alumina 

Btu/ton Nature of saving saved 

Effect of 200°F increase in flame temperature 472,621 

Effect of 200°F decrease in exhaust temperature 912,268 

Effect of reducing wall loss energy from 600,000 to 

400,000 Btu/ton 413,056 

51 
Appendix B 

Assumptions for lime calcination, rotary kiln 

Plant production 730,000 tons CaO/year

 Assume two kilns 

Output per kiln, no downtime 405,556 tons CaO /year 

With 90% uptime 365,000 tons CaO /year 

Output per day 1,000 tons CaO /d 

Output per hour 42 tons CaO /h 

Kiln diameter 5 m (16 ft) 

Kiln length 50 m (164 ft) 

Kiln exterior surface 785 m2 (8,454 ft2) 

Reaction CaCO3 → CaO + CO2

Heat of reaction at 77°F 1364 Btu/lb CaO 

Assume reaction temperature of 1000°C 

1832°F 

 say 1800°F 

Heating products to 1800°F 

CaO 371 Btu/lb CaO 

CO2 391 

Total heat requirement 2126 Btu/lb CaO 

4,252,259 Btu/ton CaO 

Heat of reaction 

57.29 cal/mol Reaction 1364 

103,122 Btu/#mol Heating 417 CaO 

1,839 Btu/lb Heating 415 CO2 

Total 2,196 Btu/lb CaO 

52 
Appendix B 

Results of lime rotary kiln analysis (neglecting dust loss) 

(Temperatures in °F) 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 4,252,259 40.62% 

Wall Loss Energy 400,000 3.82% 

Exhaust Energy 5,815,323 55.56% 

Total Energy 10,467,582 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 4,252,259 43.30% 

Wall Loss Energy 400,000 4.07% 

Exhaust Energy 5,169,176 52.63% 

Total Energy 9,821,435 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 4,252,259 38.95% 

Wall Loss Energy 600,000 5.50% 

Exhaust Energy 6,065,323 55.56% 

Total Energy 10,917,582 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 4,252,259 41.51% 

Wall Loss Energy 600,000 5.86% 

Exhaust Energy 5,391,399 52.63% 

Total Energy 10,243,657 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 4,252,259 45.70% 

Wall Loss Energy 400,000 4.30% 

Exhaust Energy 4,652,259 50.00% 

Total Energy 9,304,517 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 4,252,259 48.11% 

Wall Loss Energy 400,000 4.53% 

Exhaust Energy 4,187,033 47.37% 

Total Energy 8,839,291 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 4,252,259 43.82% 

Wall Loss Energy 600,000 6.18% 

Exhaust Energy 4,852,259 50.00% 

Total Energy 9,704,517 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 4,252,259 46.12% 

Wall Loss Energy 600,000 6.51% 

Exhaust Energy 4,367,033 47.37% 

Total Energy 9,219,291 100.00% 

Savings for lime, neglecting dust loss 

Btu/ton Nature of saving saved 

Effect of 200°F increase in flame temperature 567,631 

Effect of 200°F decrease in exhaust temperature 1,095,660 

Effect of reducing wall loss energy from 600,000 to 

400,000 Btu/ton 413,056 

53 
Appendix B 

Results of lime rotary kiln analysis (assuming 20% dust loss) 

(Temperatures in °F) 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 5,102,710 39.00% 

Wall Loss Energy 400,000 3.06% 

Exhaust Energy 7,581,767 57.94% 

Total Energy 13,084,478 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 5,102,710 43.30% 

Wall Loss Energy 400,000 3.39% 

Exhaust Energy 6,283,012 53.31% 

Total Energy 11,785,722 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 2000 

Reaction Energy 5,102,710 37.39% 

Wall Loss Energy 600,000 4.40% 

Exhaust Energy 7,944,267 58.21% 

Total Energy 13,646,978 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 2000 

Reaction Energy 5,102,710 41.51% 

Wall Loss Energy 600,000 4.88% 

Exhaust Energy 6,589,678 53.61% 

Total Energy 12,292,389 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 5,102,710 43.87% 

Wall Loss Energy 400,000 3.44% 

Exhaust Energy 6,127,936 52.69% 

Total Energy 11,630,647 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 5,102,710 48.11% 

Wall Loss Energy 400,000 3.77% 

Exhaust Energy 5,104,439 48.12% 

Total Energy 10,607,150 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Reaction Energy 5,102,710 37.39% 

Wall Loss Energy 600,000 4.40% 

Exhaust Energy 7,944,267 58.21% 

Total Energy 13,646,978 100.00% 

Parameter Btu/ton % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Reaction Energy 5,102,710 46.12% 

Wall Loss Energy 600,000 5.42% 

Exhaust Energy 5,360,439 48.45% 

Total Energy 11,063,150 100.00% 

Savings for lime, assuming dust loss 

Btu/ton Nature of saving saved 

Effect of 200°F increase in flame temperature 1,565,167 

Effect of 200°F decrease in exhaust temperature 965,410 

Effect of reducing wall loss energy from 600,000 to 

400,000 Btu/ton 885,374 

54 
55 
Appendix C 

Appendix C: 

Reverberatory Furnace Schematics 

Fig. C.1. Reverberatory furnace schematic. 

56 
Appendix C 

Fig. C.2. Typical reverberatory aluminum-melting furnace. Reprinted from Modern 

Refractory Practice, 3rd ed., Harbison Walker Refractories, Pittsburgh, 1950. 

57 
Appendix D 

Appendix D: Reverberatory Furnace Analysis Conditions and 

Results 

Results of aluminum reverberatory furnace analysis 

Temperature (oF) 

Wall/roof Loss 

Case Flame Exhaust (Btu/h) 

Case 1 3600 1600 5,000,000 

Case 2 3800 1600 5,000,000 

Case 3 3600 1600 7,500,000 

Case 4 3800 1600 7,500,000 

Case 5 3600 1800 5,000,000 

Case 6 3800 1800 5,000,000 

Case 7 3800 1600 7,500,000 

Case 8 3800 1800 7,500,000 

Results of aluminum reverberatory furnace analysis 

(Temperatures in °F) 

Case 1 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 1600 

Melting Energy 10,095,271 37.2% 

Wall Loss Energy 5,000,000 18.4% 

Exhaust Energy 12,076,217 44.4% 

Total Energy 27,171,489 100.0% 

Case 3 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 1600 

Melting Energy 10,095,271 31.9% 

Wall Loss Energy 7,500,000 23.7% 

Exhaust Energy 14,076,217 44.4% 

Total Energy 31,671,489 100.0% 

Case 5 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Melting Energy 10,095,271 33.4% 

Wall Loss Energy 5,000,000 16.6% 

Exhaust Energy 15,095,271 50.0% 

Total Energy 30,190,543 100.0% 

Case 7 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 1800 

Melting Energy 10,095,271 28.7% 

Wall Loss Energy 7,500,000 21.3% 

Exhaust Energy 17,595,271 50.0% 

Total Energy 35,190,543 100.0% 

Case 2 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 1600 

Melting Energy 10,095,271 38.7% 

Wall Loss Energy 5,000,000 19.2% 

Exhaust Energy 10,978,379 42.1% 

Total Energy 26,073,651 100.0% 

Case 4 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 1600 

Melting Energy 10,095,271 33.2% 

Wall Loss Energy 7,500,000 24.7% 

Exhaust Energy 12,796,561 42.1% 

Total Energy 30,391,833 100.0% 

Case 6 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Melting Energy 10,095,271 35.2% 

Wall Loss Energy 5,000,000 17.4% 

Exhaust Energy 13,585,744 47.4% 

Total Energy 28,681,016 100.0% 

Case 8 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 1800 

Melting Energy 10,095,271 30.2% 

Wall Loss Energy 7,500,000 22.4% 

Exhaust Energy 15,835,744 47.4% 

Total Energy 33,431,016 100.0% 

58 
Appendix D 

Savings for aluminum reverberatory furnaces 

Furnace change 

Btu/h 

saved 

Btu/lb Al 

saved 

Effect of 200°F increase in flame temperature 1,411,637 70.58 

Effect of 200°F decrease in exhaust temperature 3,486,046 174.30 

Effect of reducing wall loss energy from 7.5 to 5 million 4,642,045 232.10 

Case 1 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2200 

Melting Energy 6,272,301 21.6% 

Wall Loss Energy 5,000,000 17.2% 

Exhaust Energy 17,713,616 61.1% 

Total Energy 28,985,917 100.0% 

Case 3 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2200 

Melting Energy 6,272,301 17.7% 

Wall Loss Energy 7,500,000 21.2% 

Exhaust Energy 21,642,188 61.1% 

Total Energy 35,414,489 100.0% 

Case 5 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2400 

Melting Energy 6,272,301 18.5% 

Wall Loss Energy 5,000,000 14.8% 

Exhaust Energy 22,544,602 66.7% 

Total Energy 33,816,904 100.0% 

Case 7 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2400 

Melting Energy 10,095,271 24.4% 

Wall Loss Energy 7,500,000 18.2% 

Exhaust Energy 23,721,632 57.4% 

Total Energy 41,316,904 100.0% 

Results of copper reverberatory furnace analysis 

(Temperatures in °F) 

Savings for copper reverberatory furnaces 

Case 2 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2200 

Melting Energy 6,272,301 23.4% 

Wall Loss Energy 5,000,000 18.7% 

Exhaust Energy 15,499,414 57.9% 

Total Energy 26,771,715 100.0% 

Case 4 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2200 

Melting Energy 6,272,301 19.2% 

Wall Loss Energy 7,500,000 22.9% 

Exhaust Energy 18,936,914 57.9% 

Total Energy 32,709,215 100.0% 

Case 6 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2400 

Melting Energy 6,272,301 20.5% 

Wall Loss Energy 5,000,000 16.3% 

Exhaust Energy 19,323,945 63.2% 

Total Energy 30,596,246 100.0% 

Case 8 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2400 

Melting Energy 6,272,301 16.8% 

Wall Loss Energy 7,500,000 20.1% 

Exhaust Energy 23,609,659 63.2% 

Total Energy 37,381,960 100.0% 

Furnace change 

Btu/h 

saved 

Btu/lb Al 

saved 

Effect of 200°F increase in flame temperature 3,018,769 150.94 

Effect of 200°F decrease in exhaust temperature 5,791,405 289.57 

Effect of reducing wall loss energy from 7.5 to 5 million 6,662,946 333.15 

59 
Appendix D 

Results of glass reverberatory furnace analysis 

(Temperatures in °F) 

Case 1 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2800 

Melting Energy 13,771,643 16.3% 

Wall Loss Energy 5,000,000 5.9% 

Exhaust Energy 65,700,751 77.8% 

Total Energy 84,472,394 100.0% 

Case 3 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 2800 

Melting Energy 13,771,643 14.4% 

Wall Loss Energy 7,500,000 7.8% 

Exhaust Energy 74,450,751 77.8% 

Total Energy 95,722,394 100.0% 

Case 5 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 3000 

Melting Energy 13,771,643 12.2% 

Wall Loss Energy 5,000,000 4.4% 

Exhaust Energy 93,858,216 83.3% 

Total Energy 112,629,859 100.0% 

Case 7 

Parameter Btu/h % Usage 

Flame Temperature 3600 

Exhaust Temperature 3000 

Melting Energy 13,771,643 10.8% 

Wall Loss Energy 7,500,000 5.9% 

Exhaust Energy 106,358,216 83.3% 

Total Energy 127,629,859 100.0% 

Case 2 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2800 

Melting Energy 13,771,643 19.3% 

Wall Loss Energy 5,000,000 7.0% 

Exhaust Energy 52,560,601 73.7% 

Total Energy 71,332,244 100.0% 

Case 4 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 2800 

Melting Energy 13,771,643 17.0% 

Wall Loss Energy 7,500,000 9.3% 

Exhaust Energy 59,560,601 73.7% 

Total Energy 80,832,244 100.0% 

Case 6 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 3000 

Melting Energy 13,771,643 15.4% 

Wall Loss Energy 5,000,000 5.6% 

Exhaust Energy 70,393,662 78.9% 

Total Energy 89,165,305 100.0% 

Case 8 

Parameter Btu/h % Usage 

Flame Temperature 3800 

Exhaust Temperature 3000 

Melting Energy 13,771,643 13.6% 

Wall Loss Energy 7,500,000 7.4% 

Exhaust Energy 79,768,662 78.9% 

Total Energy 101,040,305 100.0% 

Savings for glass reverberatory furnaces 

Furnace change 

Btu/h 

saved 

Btu/lb glass 

saved 

Effect of 200°F increase in flame temperature 19,521,102 976.06 

Effect of 200°F decrease in exhaust temperature 31,173,901 1,558.70 

Effect of reducing wall loss energy from 7.5 to 5 million 11,906,250 595.31 

60 
61 
Appendix E 

Appendix E: Boiler Schematic 

Fig. E.1. Boiler furnace sections. Reprinted from Modern Refractory Practice, 3rd ed., 

Harbison Walker Refractories, Pittsburgh, 1950. 

62 
63 
Copyrighted page(s) removed. Copyrighted 

material is not available in Regulations.gov since 

it may not be reproduced without consent of the 

copyright holder. 

Removed pages can be viewed at the USEPA 

Docket Center, Public Reading Room. 

Address: 1301 Constitution Ave, NW 

Room 3334 

Washington, DC 20004 

Telephone: 202-566-1744 

Email: docket-customerservice@epa.gov
INDUSTRY CONVERSATION REGARDING THE SOURCE OF HIGH HEX CHROME EMISSIONS

Carlos E. Davis, Certainteed/St. Gobain 14 April 2011

This call was made because I was looking over the test report that BTEC did on KCK furnace...K2 

furnace..at Kansas City. I was wondering if there was a problem with the testing, and if the test was 

done correctly. Mr. Davis confirmed that the test was done correctly; that the math was checked on the 

calculations. I led in that I was interested in where all this chrome was coming from. By the testing, and 

the industry’s test agreement that the tested furnace was representative of the K1 furnace, I calculated 

over 900 pounds a year total chrome from the KCK plant. This was higher than Mr. Davis remembered, 

but confirmed that they had a lot of chrome. He was not aware that 92% of the chrome was hexavalent.

In a glass furnace, you feed the fine sands and by that I mean actual sand with other refined and purified 

raw materials in a mixture into the furnace. It’s a continuous furnace. The furnace temperature is 

about 15600 C. At the surface of the glass, is the metal-line glass contact.

Now the furnace itself extends above the surface of the glass..the metal line glass contact. We have 

constructed the K2 furnace from what is called ‘High Chrome’ or sometimes also called ‘Jade Green’ 

refractories. They are just that color, a deep beautiful green color. They are that color because they are 

made of between 98-96% chrome. We construct our furnace using noncorrosive materials because you 

get a much longer life from your furnace. In older designs, you have to rebuild a glass furnace at least 

every 7-10 years. You can add on to the outside of the furnace (meaning additional refractories) up to a 

point (2 layers), but after that, you have to shut it down and tear it out. 

In the K2 furnace you have glass to a depth of a little more than a meter deep. Then to the top of the 

straight sides of the furnace, it’s about 2 meters, so there’s this additional straight wall up to the 

beginning of what we call the ‘crown’, this kind of domed curved top of the furnace. That’s the ‘furnace 

crown’. The surface of the glass is about 70 square meters (70m2

).

The molten glass is very corrosive. So there is high corrosivity and errosivity. And as the molten glass 

moves through the furnace, it starts to eat away at the refractories, and it doesn’t really eat evenly, it 

eats little holes in the refractory wall inside the furnace. So these are like little caves, little holes. And 

so the surface area of the refractory wall is actually increasing with the life of the furnace. Because the 

molten glass continues to move through these caves or holes, it continues to erode even larger areas 

inside the refractory wall from the inside of the furnace reaching out. So when that outer wall gets hot, 

you just add some more refractory onto the outside. Otherwise you’d have a real bad situation, a 

breakthrough and you can’t have that. So your emissions of chromium are actually increasing, not 

decreasing over the life of the furnace due to this ever-increasing surface area being exposed to the 

molten glass. So the degradation accelerates with the age of the furnace. We didn’t realize that this was 

almost all hexavalent chromium, we’ve only tested total chrome. That’s a big surprise. 

K2 furnace at Certainteed in KCK has a planned shutdown date of 2013. This furnace is on it’s final end 

run. 
Update 1 from Region 7:

Certainteed is planning to shift all products to the K2 line, but that line will only hold up for another 

year.

Update 2 from Region 7:

We found some 2005 testing in the NEI. We don’t know the test method they used, and the test report 

is not in our files, but the results of that testing were 0.0166 TPY of total chrome, which is about 300 

pounds a year, facility-wide. We don’t know what fraction of that total chrome is hexavalent chrome; 

we don’t know which furnace was tested, or if one was tested as representative of the other. So there 

are a lot of unknowns with this test, but this shows that even in 2005, these high levels were being 

emitted.
Notes on Differences in Collection Operations in the Mineral Wool Industry

23 June 2011 Steve Edris, Thermafiber

The VERTICAL COLLECTION design.

The collection chamber at Thermafiber is 10-15 feet long, about 6 feet wide.

At the end of the chamber is a vertical perforated belt moving belt, running at about 300 feet per minute (this is the 

fastest in the bonded lines)

The cupola produces at the same steady rate, pouring a thin stream of molten rock onto the spinners. Binder is sprayed 

onto the fibers through ports on the circumference edge of the spinners. Air flow comes from behind (and under 

cupola) to blow the fibers away from the spinners as they are formed. 

A stronger air flow pulls negative on the screen, sucking the now cooled, binder sprayed fibers against the moving 

vertical screen, forming a mat that is about 1-1.5 inches. This screen moves much faster than any other in the industry 

because it is vertical. Otherwise, you’d have the fiber mat slumping along the bottom curve. 

This design makes a cleaner product that is low in ‘shot’. Shot is ok for some products, but can’t be present in 

horticultural applications (tomatoes need a uniform density product without ‘shot’, as do other vegetables). The CHOH 

is not trapped. 

The HORIZONTAL COLLECTION design.

Everything is pretty much the same right up to the collection process. In the horizontal collection, the horizontal mat is 

much thicker. That belt moves at about 50 feet per minute, and CHOH is trapped in a thicker mat. The belt doesn’t have 

to move fast because you are not working against gravity as you are in the vertical design. However, there are some 

products you can’t make on a horizontal design. First, the shot remains in the bonded product. That’s okay for the 

products they make at that company, because that’s mostly insulation products where some amount of shot does not 

affect the final application. But you couldn’t sell to the horticultural markets with that product.

The DRUM COLLECTION design.

Again, it’s pretty much the same right up to the collection process. But then this is basically a totally different approach. 

Instead of spinners and a moving belt, imagine a piece of equipment that looks like your washing machine, except 

turned. The molten rock is poured from the cupola into this big spinning drum, like the inside of your washer. And the 

little holes in the inside where the water empties out of the washing machine....that’s where the air flows out of the 

spinning drum carrying with it the threads of mineral wool. Unlike the other two designs, the air flow is already 

contained within the drum, and it can be directed as an air stream that lends itself well to controls. Out of the holes 

flow the forming fibers coming out as a loose long flow of fibers that is pulled from the spinning drum. The air stream 

from the collection process can then be directed to either the hottest part of the cupola stack or to the RTO on the 

curing oven. Either way, you achieve really low emissions of the organics in the binder. 
Notes on the Origins and Cause of COS Emissions in Mineral Wool Industry

Chris Bullock, IIG 28 June 2011

IIG is not ‘afraid’ of adding pollution controls. They do NOT support NAIMA’s approach on the SBAR panel. They said 

the industry group will look stupid and ineffectual going forward with the proposed limits. They do not think industry 

should get a pass just because they claim to be small. IIG can meet the limits EPA will be proposing.

They wouldn’t have to install a new RTO. They route their collection emissions to the hottest part of their cupola stack. 

This causes the formaldehyde, MeOH and phenol to incinerate in the stack before being emitted to the atmosphere. 

IIG has done research to show the source of sulfur is the SLAG....COKE is the source of CO. Coke is not so much a 

contributor to formation of COS and SO2. They purchase low-sulfur slag locally so their costs are still low.

Chris Bullock, IIG 28 June 2011

We mine and melt our own rock so we have better control of our inputs. We can melt at higher temperatures. But we 

have a 50% cost increase to us on our raw materials.

Low sulfur slag is not an option (?)

Oxygen injection will increase the COS because it burns hotter (?)

We don’t have a sulfur compounds being emitted from our raw materials, because we don’t put sulfur containing raw 

materials in to start with. In order to have a problem with SO2 or COS, you have to have sulfur in your raw materials. 

We know exactly what we are melting. All the other companies do too, or they ought to. Everybody keeps records of 

their raw material inputs because you have to do that to keep a consistent and reliable end product. 

Where is the leveling of the playing field? We are good environmental stewards. Again, we don’t think industry should 

get a pass on regulation just because this industry is full of companies that say they are small.













Notes on Differences in Collection Operations in the Mineral Wool Industry

23 June 2011 Steve Edris, Thermafiber

The VERTICAL COLLECTION design.

The collection chamber at Thermafiber is 10-15 feet long, about 6 feet wide.

At the end of the chamber is a vertical perforated belt moving belt, running at about 300 feet per minute (this is the 

fastest in the bonded lines)

The cupola produces at the same steady rate, pouring a thin stream of molten rock onto the spinners. Binder is sprayed 

onto the fibers through ports on the circumference edge of the spinners. Air flow comes from behind (and under 

cupola) to blow the fibers away from the spinners as they are formed. 

A stronger air flow pulls negative on the screen, sucking the now cooled, binder sprayed fibers against the moving 

vertical screen, forming a mat that is about 1-1.5 inches. This screen moves much faster than any other in the industry 

because it is vertical. Otherwise, you’d have the fiber mat slumping along the bottom curve. 

This design makes a cleaner product that is low in ‘shot’. Shot is ok for some products, but can’t be present in 

horticultural applications (tomatoes need a uniform density product without ‘shot’, as do other vegetables). The CHOH 

is not trapped. 

The HORIZONTAL COLLECTION design.

Everything is pretty much the same right up to the collection process. In the horizontal collection, the horizontal mat is 

much thicker. That belt moves at about 50 feet per minute, and CHOH is trapped in a thicker mat. The belt doesn’t have 

to move fast because you are not working against gravity as you are in the vertical design. However, there are some 

products you can’t make on a horizontal design. First, the shot remains in the bonded product. That’s okay for the 

products they make at that company, because that’s mostly insulation products where some amount of shot does not 

affect the final application. But you couldn’t sell to the horticultural markets with that product.

The DRUM COLLECTION design.

Again, it’s pretty much the same right up to the collection process. But then this is basically a totally different approach. 

Instead of spinners and a moving belt, imagine a piece of equipment that looks like your washing machine, except 

turned. The molten rock is poured from the cupola into this big spinning drum, like the inside of your washer. And the 

little holes in the inside where the water empties out of the washing machine....that’s where the air flows out of the 

spinning drum carrying with it the threads of mineral wool. Unlike the other two designs, the air flow is already 

contained within the drum, and it can be directed as an air stream that lends itself well to controls. Out of the holes 

flow the forming fibers coming out as a loose long flow of fibers that is pulled from the spinning drum. The air stream 

from the collection process can then be directed to either the hottest part of the cupola stack or to the RTO on the 

curing oven. Either way, you achieve really low emissions of the organics in the binder. 
Notes on the Origins and Cause of COS Emissions in Mineral Wool Industry

Chris Bullock, IIG 28 June 2011

IIG is not ‘afraid’ of adding pollution controls. They do NOT support NAIMA’s approach on the SBAR panel. They said 

the industry group will look stupid and ineffectual going forward with the proposed limits. They do not think industry 

should get a pass just because they claim to be small. IIG can meet the limits EPA will be proposing.

They wouldn’t have to install a new RTO. They route their collection emissions to the hottest part of their cupola stack. 

This causes the formaldehyde, MeOH and phenol to incinerate in the stack before being emitted to the atmosphere. 

IIG has done research to show the source of sulfur is the SLAG....COKE is the source of CO. Coke is not so much a 

contributor to formation of COS and SO2. They purchase low-sulfur slag locally so their costs are still low.

Chris Bullock, IIG 28 June 2011

We mine and melt our own rock so we have better control of our inputs. We can melt at higher temperatures. But we 

have a 50% cost increase to us on our raw materials.

Low sulfur slag is not an option (?)

Oxygen injection will increase the COS because it burns hotter (?)

We don’t have a sulfur compounds being emitted from our raw materials, because we don’t put sulfur containing raw 

materials in to start with. In order to have a problem with SO2 or COS, you have to have sulfur in your raw materials. 

We know exactly what we are melting. All the other companies do too, or they ought to. Everybody keeps records of 

their raw material inputs because you have to do that to keep a consistent and reliable end product. 

Where is the leveling of the playing field? We are good environmental stewards. Again, we don’t think industry should 

get a pass on regulation just because this industry is full of companies that say they are small.
Chromium Emission Rate (lb/ton) by Furnace Type

1

Oxy FuelOxy FuelOxy Fuel

Air GasAir GasAir Gas

Electric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric Melt

ESSESSESSESSESSESSESSESSESS

Oxy FuelOxy FuelOxy FuelOxy FuelOxy FuelOxy FuelOxy Fuel

Electric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric MeltElectric Melt

ESSESSESSESSESSESSESSESSESSESSESSESSESSESSESSESSESSESS

Cr Emission Rate (lb/ton)

Major Sources Area Sources 

Furnaces out of 

graph range 0.00020

0.00015

0.00010

0.00006 

0.00005

 0

Furnace Type

Notes: - Proposed chromium emission limit is 0.00006 (6x10-5) lb/ton

 - Some furnaces were tested in 2010 and 2012. For valid tests, both data points are displayed.
Air Gas Furnace

2

Deliberative Document – Do Not Cite or Quote
Oxyfuel Furnace

3

Deliberative Document – Do Not Cite or Quote
Cold Top Electric Furnace

4

Deliberative Document – Do Not Cite or Quote
Steel Shell Furnace

5

Deliberative Document – Do Not Cite or Quote
Furnace type Glass Contact -- 

Side Wall

Glass Contact -- Floor 

Paving

Breast Wall Crown Back Wall Portneck/ 

Stack

Throat CrVI Emissions 

(lb/hr)

Detection Limit 

Reported as lb/hr 

Emissions? 

(Yes/No)

CrVI Emissions 

(lb/year)

Tot. Chrome 

Emissions 

(lb/hr)

Detection Limit 

Reported as lb/hr 

Emissions? 

(Yes/No)

Tot. Chrome 

Emissions 

(lb/year)

1 Oxy-fuel * 30-60% Chrome - Chrome Chrome Chrome Chrome N/A N/A N/A N/A N/A N/A

2 Oxy-fuel 30-50% 50% 30% 30% 30-50% 50% 50% 0.00023 No 2.0148 0.00026 No 2.2776 KCK ct/SG 5

3 Oxy-fuel 30-50% 50% 30% 30% 30-50% 50% 50% 0.00015 No 1.3140 0.00027 No 2.3652

4 Oxy-fuel 50% Chrome - - - - Chrome N/A N/A N/A N/A N/A N/A

5 Oxy-fuel 50% Chrome Chrome 50% 50% Chrome Chrome 0.047 No 411.7200 0.051 No 446.76 KCK ct/SG 2

6 Oxy-fuel 49.5% 49.5 - 53.7% 49.5% 49.5% 49.5% 49.5 - 92.0% 50 - 92% 0.00014 No 1.2264 0.00056 No 4.9056

7 Oxy-fuel 30-50% Chrome Chrome - Chrome Chrome Chrome N/A N/A N/A N/A N/A N/A KCK ct/SG 1

8 Steel shell 30-60% 30% - - - - - 0.000011 Yes 0.0964 0.00011 No 0.9636

9 Steel shell 95% 30% - - - - 95% 0.0000066 No 0.0578 0.000023 No 0.2015

10 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

11 Steel shell 95% 30% - - - - 95% 0.000012 No 0.1051 0.00003 Yes 0.2628

12 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

13 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

14 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

15 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

16 Steel shell 30-50% 30% - - - - - 0.0000021 Yes 0.0184 0.000045 No 0.3942

17 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

18 Steel shell 30-60% 30% - - - - - 0.000012 Yes 0.1051 0.000027 No 0.2365

19 Steel shell 95% 30% - - - - 95% 0.0000056 Yes 0.0491 0.00009 No 0.7884

20 Steel shell 30-60% 30% - - - - - N/A N/A N/A N/A N/A N/A

21 Cold top electric 50% 30% - - - - 50-95% 0.000014 No 0.1226 0.000018 Yes 0.1577

22 Cold top electric 50-60% 50-60% - - 50-60% - 95% 0.00000078 Yes 0.0068 0.00000021 Yes 0.0018

23 Cold top electric 50% Chrome - - - - Chrome N/A N/A N/A N/A N/A N/A

24 Cold top electric 50-60% 50-60% #VALUE! - 50-60% - 95% N/A N/A N/A N/A N/A N/A

25 Cold top electric 50% 30% - - - - 50-95% 0.0000076 No 0.0666 0.000032 No 0.2803

26 Cold top electric 49.5% 49.5 - 53.7% - - - - 50-92 0.0000079 Yes 0.0692 0.000035 No 0.3066

27 Cold top electric 50% 30% - - - - 50-95% 0.000015 No 0.1314 0.000026 Yes 0.2278

28 Cold top electric 50-60% 50-60% - - - - 95% 0.0000022 Yes 0.0193 0.000073 No 0.6395

29 Cold top electric 50% 30% - - - - 50-95% 0.0000055 No 0.0482 0.000039 No 0.3416

30 Cold top electric 50-60% 50-60% - - 50-60% - 95% 0.0000052 No 0.0456 0.00004 Yes 0.3504

31 Air gas 30-50% 30% - - - - 50-95% 0.0000013 Yes 0.0114 0.0000021 Yes 0.0184

32 Air gas 30-50% 30% - - - - 50-95% 0.0000014 No 0.0123 0.000004 No 0.0350

33 Air gas 30-50% 30% - - - - 50-95% 0.0000033 No 0.0289 0.0000046 No 0.0403

34 Air gas 28-95% 28-60% - - 28-95% 95% 75% N/A N/A N/A N/A N/A N/A

35 Air gas 28-95% 28-60% - - 28-95% 95% 75% N/A N/A N/A N/A N/A N/A

36 Air gas 30-50% 30% - - - - 50-95% 0.0000026 No 0.0228 0.0000037 No 0.0324

37 Specialty - - - - - - - 0.00017 No 1.4892 0.000064 Yes 0.5606

38 Specialty 30.0% 30.0% - - 30.0% - - 0.0002 No 1.7520 0.00045 No 3.9420

*

BTEC Emisssions Test Data

All oxy-fuel furnaces use a dry electrostatic precipitator as a control device.
 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 
DOCUMENT AVAILABILITY

Reports produced after January 1, 1996, are generally available free via the U.S. 

Department of Energy (DOE) Information Bridge. 

Web site http://www.osti.gov/bridge 

Reports produced before January 1, 1996, may be purchased by members of the 

public from the following source. 

National Technical Information Service 

5285 Port Royal Road 

Springfield, VA 22161 

Telephone 703-605-6000 (1-800-553-6847) 

TDD 703-487-4639 

Fax 703-605-6900 

E-mail info@ntis.fedworld.gov 

Web site http://www.ntis.gov/support/ordernowabout.htm 

Reports are available to DOE employees, DOE contractors, Energy Technology Data 

Exchange (ETDE) representatives, and International Nuclear Information System 

(INIS) representatives from the following source. 

Office of Scientific and Technical Information 

P.O. Box 62 

Oak Ridge, TN 37831 

Telephone 865-576-8401 

Fax 865-576-5728 

E-mail reports@adonis.osti.gov 

Web site http://www.osti.gov/contact.html
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), 

classified information, information subject to export control classification, or other information not 

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. 

Copyrighted page(s) removed. Copyrighted 

material is not available in Regulations.gov since 

it may not be reproduced without consent of the 

copyright holder. 

Removed pages can be viewed at the USEPA 

Docket Center, Public Reading Room. 

Address: 1301 Constitution Ave, NW 

Room 3334 

Washington, DC 20004 

Telephone: 202-566-1744 

Email: docket-customerservice@epa.gov
From: Chun-Wai Lee

To: Susan Fairchild

Cc: Reeves, David; Larke Williams; Bill Linak

Subject: Re: Fw: WF - PM and Cr Graphs (NEW)

Date: 03/28/2012 05:01 PM

Attachments: WF PM and Cr Grpahs 3.22.12 v2.xlsx

The first graph shows units which exceed the proposed Cr emission limit (0.00006

lb/ton) are all oxy-fuel furnaces either tested previously (# 16, 17, 38, 39, and 40)

or in this year (#23 and 40) with one exception (#26) for a recoup gas fired

furnace. Emissions from the other oxy-fuel furnaces below the limit are also high

relative to those of other units. 

The high Cr emissions shown in the graph support Dr. Brosnan's conclusions

that flame temperature is the most important variable in furnace operation

influencing the quantity and speciation of chromium emissions from wool fiberglass

furnaces employing chromium refractory linings. Refractory linings of a oxy-fuel

furnace are well known to be much hotter than those of the conventional air fired

furnaces as a result of the exceedingly high temperature produced by oxy-fuel

combustion (5,035 o

F compared to 3,562oF produced from conventional air-fired

furnaces) greatly enhances radiative heat transfer. He also quotes references which

show volatility of Cr2O3 increases rapidly above 2,400 o

F , and the high oxygen

atmosphere (as it is existed in a oxy-fuel furnace) also promotes the volatilization of

Cr2O3 (page 7 in his letter). Furthermore, the high oxygen atmosphere

(thermodynamic effect) coupled with the high temperature (kinetic effect) in a oxyfuel furnace provide a favorable conditions to promotes oxidation of Cr2O3 existed

in the linings to CrO3 (CrVI). All these factors discussed above provide a strong

scientific basis that oxy-fuel furnaces with Cr refractory lining have higher potential

for Cr, toxic CrVI in particular, emissions. 

CW 

************************************

C.W. Lee, Ph.D.

U.S. Environmental Protection Agency (E-305-01)

Office of Research and Development

National Risk Management Research Laboratory

109 T.W. Alexander Drive

Research Triangle Park, NC 27711

E-mail: lee.chun-wai@epa.gov

Voice: (919) 541-7663

Fax: (919) 541-0554

************************************

▼ Susan Fairchild---03/28/2012 11:29:01 AM---Dear CW and Larke- The contractor

has compiled the existing and new chrome test data into a spreadsh

From: Susan Fairchild/RTP/USEPA/US

To: Larke Williams/DC/USEPA/US@EPA, Chun-Wai

Lee/RTP/USEPA/US@EPA

Cc: "Reeves, David" <dwreeves@rti.org>

Date: 03/28/2012 11:29 AM

Subject: Fw: WF - PM and Cr Graphs (NEW)
Dear CW and LarkeThe contractor has compiled the existing and new chrome test data into a

spreadsheet and has graphed the data. This is attached.

The oxyfuel furnaces are in red.

As you can tell, there are several types of furnaces:

cold top electric

Electric Steel Shell (ESS), aka Steel shell electric (we believe these are the same

types of furnaces)

gas fired electric boost

Johns Manville HERM melters

other HERM melters (I believe there may be distinct differences, but don't know

what they are)

Recoup gas fired

air gas

We have not answered all the questions about the furnaces that we need to answer,

but do these data help in addressing the NAIMA comments and Dr. Brosnan's

attached letter supporting those comments?

Thanks,

Susan Fairchild

Environmental Scientist

U.S. Environmental Protection Agency

direct line: 919.541.5167

fax: 919.541-3207

email: fairchild.susan@epa.gov

Overnight mail address:

US EPA

Mail Code D243-04

109 TW Alexander Drive

RTP, NC 27709

----- Forwarded by Susan Fairchild/RTP/USEPA/US on 03/28/2012 11:15 AM -----

From: "Reeves, David" <dwreeves@rti.org>

To: Susan Fairchild/RTP/USEPA/US@EPA, Jeff

Telander/RTP/USEPA/US@EPA

Date: 03/25/2012 10:26 AM

Subject: FW: WF - PM and Cr Graphs (NEW)

Hi Susan and Jeff.  I am forwarding the PM and Cr graphs that Cindy finished on
Friday that include the new 2012 test data for fiberglass furnaces.  Note that there

are 8 or 9 furnaces with 2 data points where we had old test data from the facilities

and then the new testing was done this year.  As you will see, there are no clear

data breaks or obvious subcategory differences for major sources.  Because of the

funding issues, I’ve told Cindy to hold off doing any more analysis – pending your

discussion and decisions.

Dave Reeves

RTI International

dwreeves@rti.org

(919) 316-3739

From: Hancy, Cindy Wang 

Sent: Friday, March 23, 2012 5:02 PM

To: Reeves, David

Subject: WF - PM and Cr Graphs (NEW)

Dave,

Here are new graphs. I based the PM tech review value on K2 since it was the closes

to being ~0.4 lb/ton.

Thanks!

Cindy Wang Hancy

RTI International

P.O. Box 12194

Research Triangle Park, NC 27709

Phone:  (919) 541-6623

Fax:  (919) 316-3420
Copyrighted page(s) removed. Copyrighted 

material is not available in Regulations.gov since 

it may not be reproduced without consent of the 

copyright holder. 

Removed pages can be viewed at the USEPA 

Docket Center, Public Reading Room. 

Address: 1301 Constitution Ave, NW 

Room 3334 

Washington, DC 20004 

Telephone: 202-566-1744 

Email: docket-customerservice@epa.gov