Ic 9446 INFORMATION CIRCULAR/1997 

Proceedings: New Technology for Ground Control 

in Retreat Mining 

U.S. DEPARTMENT OF HEALTH AND HUMAN SERVICES 

Publ~c Health Service 

Centers for Disease Control and Prevention 

National Institute for Occupational Saiety and Health CENERS FOR DISEASE CONTROL 

AND PREVENTlON 
Cover photo: Pillar extraction using mobile roof supports. 
Information Circular 9446 

Proceedings: New Technology for Ground Control 

in Retreat Mining 

Christopher Mark, Ph.D., and Robert J. Tuchman 

U.S. DEPARTMENT OF HEALTH AND HUMAN SERVICES 

Public Health Service 

Centers for Disease Control and Prevention 

National Institute for Occupational Safety and Health 

Pittsburgh, PA and Spokane, WA Research Centers 

March 1997 
International Standard Serial Number 

ISSN 1066-5552 
CONTENTS 

Page 

Abstract ...................................................................................... 

A Statistical Overview of Retreat Mining of Coal Pillars in the United States, by C. Mark, Ph.D., F. E. McCall, 

andD.M.Pappas .............................................................................. 

PILLAR DESIGN 

Analysis of Retreat Mining Pillar Stability (ARMPS), by C. Mark, Ph.D., and F. E. Chase ...................... 

Preventing Massive Pillar Collapses in Coal Mines, by C. Mark, Ph.D., F. E. Chase, and R. K. Zipf, Jr., Ph.D. ...... 

Pillar Design and Coal Strength, by C. Mark, Ph.D., and T. M. Barton ..................................... 

A New Laminated Overburden Model for Coal Mine Design, by K. A. Heasley .............................. 

MOBILE ROOF SUPPORTS 

Retreat Mining With Mobile Roof Supports, by F. E. Chase, A. McComas, C. Mark, Ph.D., and C. D. Goble ....... 

Monitoring Mobile Roof Supports, by K. E. Hay, S. P. Signer, M. E. King, and J. K. Owens .................... 

Full-scale Performance Evaluation of Mobile Roof Supports, by T. M. Barczak and D. F. Gearhart .............. 
UNIT OF MEASURE ABBREVIATIONS USED IN THIS REPORT 

cm centimeter 

ft foot 

ft/min foot per minute 

ft2 square foot 

ft3 cubic foot 

GPa gigapascal 

ha hectare 

in inch 

in2 square inch 

Ibf pound (force) 

m meter 

dmin meter per minute 

m2 square meter 

m3 cubic meter 

min minute 

mm millimeter 

MPa megapascal 

psi pound (force) per square inch 

I kg kilogram st short ton 

kipdin kips per inch 

kN kilonewton 

sth short ton per hour 

t ton (metric) 

I kN/cm kilonewton per centimeter % percent 

I kPa kilopascal 0 degree 

I lb pound 

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PROCEEDINGS: NEW TECHNOLOGY FOR GROUND CONTROL 

IN RETREAT MINING 

Compiled by Christopher Mark, Ph.D.,' and Robert J. Tuchman2 

ABSTRACT 

This proceedings volume contains papers presented at technology transfer seminars sponsored by the 

National Institute for Occupational Safety and Health (NIOSH) on New Technology for Ground Control in 

Retreat Mining. The seminars were conducted at five locations: Uniontown, PA (March 26, 1997), Norton, 

VA (April 8, 1997), Pikeville, KY (April 10, 1997), Charleston, WV (April 17, 1997), and Evansville, IN 

(April 22, 1997). 

The papers presented here describe several new, highly practical technologies developed by the NIOSH 

Pittsburgh and Spokane Research Centers3 to improve safety during pillar retreat operations. Two central 

issues are addressed: pillar design and mobile roof supports (MRS's). 

Proper pillar sizing is essential for safe pillar extraction. The Analysis of Retreat Mining Pillar Stability 

(ARMPS) program and its large data base of actual mining case histories are presented. LAMODEL, a second 

computer program, can be used for analysis of multiple-seam and other complex mining situations. Other 

papers address pillar design to avoid massive pillar collapses and the proper role of coal strength testing. 

MRS's have greatly improved safety where they are used for pillar line support. We studied the application 

of MRS's at 20 mines throughout the Eastern United States. Conclusions regarding the most effective section 

layouts, cut sequences, and support placements are reported. Field and laboratory tests of MRS's are also 

described. 

'Mining engineer. 

2~echnical writer-editor. 

Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pinsburgh, PA. 

pi he research described in these papers originated under the former U.S. Bureau of Mines prior to transferring to the National Institute for 

Occupational Safety and Health in 1996. 
A STATISTICAL OVERVIEW OF RETREAT MINING OF COAL 

PILLARS IN THE UNITED STATES 

By Christopher Mark, Ph.D.,' Frank E. McCall,' and Deno M. Pappas2 

ABSTRACT 

The demographics and safety record of the pillar retreat segment of the U.S. coal industry was analyzed 

using statistics collected by the Mine Safety and Health Administration. Pillar recovery is practiced primarily 

by mines in Appalachia and the Midwest. Using 1993 data, the accident rates and productivity of a large 

sample of pillar retreat mines were found to be similar to other room-and-pillar mines in the same geographic 

areas. Pillar recovery apparently accounts for about 10% of all U.S. underground production, but has been 

associated with about 25% of the roof and rib fatalities during 1989-96. However, of the 28 fatalities that were 

analyzed, only 4 occurred for which no citations were issued for violations of mining law. Nearly one-half of 

the fatal incidents occurred during the mining of the last lift or pushout. All four no-citation incidents occurred 

during the removal of the last lift during a "Christmas tree" extraction sequence. 

'Mining engineer. 

'Civil engineer. 

Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 
INTRODUCTION 

Pillar recovery has always been an integral part of U.S. 

underground coal mining. It can be a less capital-intensive, 

more flexible alternative to longwall mining for small. 

irregular reserves [Blaiklock 19921. It'is often employed in 

deeper, high-value seams where recovery rates would be 

unacceptably low if only development room-and-pillar mining 

was conducted. 

The process of pillar recovery removes the main support to 

the overburden and allows the ground to cave. As a result, the 

pillar line is an extremely complex and high-stress rock 

mechanics environment. Historically, retreat mining has 

accounted for a large number of fatal roof fall accidents. 

During 1978-86, 67 roof fall fatalities were attributed to 

retreat mining, representing 29% of the total. Of the pillaring 

fatalities, 49% occurred during the mining of the final stump 

[Montague 19881. Nevertheless, there has apparently never 

been a detailed study of the demographics and safety record of 

pillar retreat mining. This study attempts to fill the gap. 

The overview presented here is based almost entirely on 

information collected by the Mine Safety and Health 

Administration (MSHA). Three primary sources were used: 

MSHA Accident and Employment Data Base: This data 

base contains information on the employment and production 

of all U.S. operating coal mines. It also contains information 

on all accidents reported to MSHA. 

Data Base of Retreat Mines: In 1993, MSHA formed the 

Mine Ventilation Bleeder and Gob Training Committee. Part 

of the committee's work was to survey the nine bituminous 

coal MSHA health and safety districts about the practices of 

their mines. The survey identified 186 nonlongwall mines that 

were maintaining an active gob and that produced more than 

4,500 t (5,000 st) in 1993 [Urosek et al. 19951. These mines 

were approximately evenly divided between those that 

practiced full-pillar recovery and those that were limited to 

partial pillar extraction. An additional 181 mines had 

ventilated, inactive gobs. Some had permanently ceased 

retreat mining, others were developing for pillar recovery 

operations when they were surveyed. Therefore, the 186 

active gob mines represent a large sample of the total retreat 

mine population. The identification numbers of these mines 

were the key to making comparisons using the MSHA 

Accident and Employment Data Base. 

Fatal Accident Reports: Since 1988, a total of 25 accidents resulting in 33 fatalities have occurred during pillar 

recovery operations. MSHA prepared detailed Reports of 

Investigation on all but the most recent of these fatal incidents, 

and the reports were subjected to in-depth analyses. 

DEMOGRAPHICS AND ACCIDENT RATES 

Table 1 compares basic statistics for 1993 for three Table 1 and figures 1-2 show that the sample of retreat 

segments of the U.S. underground coal industry: (1) longwall mines employed 9,129 miners and produced 56 million t (61.7 

mines, (2) all room-and-pillar mines, and (3) the sample of million st) in 1993, representing 18% of the total underground 

186 room-and-pillar retreat mines. production. The 56 million t (61.7 million st) includes both 

Table 1 .- Demographics and accident statistics for U.S. underground coal mines by mine type' 

Average 

No. of No. of mine Tons' Productivity, Total Rooflrib Total Rooflrib 

Mine type em- thousand st/h accident accident days lost days lost mines size (employees ~lovees) st rate rate rate rate 

Room-and-pillar ' . . . 1,014 33,073 33 214,299 3.45 15.92 1.44 45 1 4 1 

Retreat . . . . . . . . . . 186 9.129 49 61,701 3.27 15.58 1.14 432 29 

Longwall . . . . . . . . . . 69 15,419 223 133,132 4.16 13.39 1 .OO 410 29 

Entire industry . . . . . 1,083 48,491 45 347,430 3.69 15.06 1.29 437 37 

'Excludes anthracite mines and mines producing less than 5,000 st. 

'Accident rates are calculated as the total number of injury accidents (severity 1-6) per 200,000 hours worked. 

3Days lost rates are calculated as the total number of days lost due to injury per 200,000 hours worked. 

'Room-and-pillar mines include all nonlongwall mines. 

'Retreat mines are the 186 nonlongwall mines with active gobs identified by Urosek et al. [1995]. 

Source: MSHA Accident and Employment Data Base for 1993. 
30000 

I- ; 25000 

r * 

9 g 20000 

U1 

n 

Z 0' 15000 

II: 

ri 

g 10000 3 

5000 

0 

ALL ROOM LONGWALL 

8 PILLAR MINES 

MINES 

MlNE TYPE 

8 RETREAT MlNE 

SAMPLE ~ 

Figure 1 .-Employment at U.S. underground coal mines in 1993, by mine type. 

ALL ROOM 8 PILLAR 

MINES 

LONGWALL MINES 

MlNE TYPE 

SAMPLE 

Figure 2.-Production at U.S. underground coal mines in 1993, by mine type. 
development and retreat tonnage. A reasonable estimate is 

that pillar recovery operations account for about one-third of 

coal production from retreat mines [Reese et al. 19781. 

Including some contribution from the mines with inactive 

ventilated gobs, it appears that pillar recovery may have 

accounted for about 10% of the 315 million t (347 million st) 

mined underground in 1993. 

An average of 49 miners were employed at each pillar 

retreat mine, slightly more than at the typical room-and-pillar 

mine, but much less than at a longwall mine. The accident 

statistics in figure 3 show that, overall, the injury record of 

retreat mines was similar to that of other mining methods. 

Surprisingly, roof and rib accident rates in figure 4 were 21% 

lower at retreat mines than at other room-and-pillar mines. 

One possible explanation is that roof bolting, which is a 

significant source of roof fall injuries, is seldom employed 

during retreat mining. The rates for days lost from all accidents closely paralleled the overall accident rates. 

Some regional trends are shown in table 2 and figures 5-6. 

It appears that retreat mining was widely practiced throughout 

the Appalachian and midwestern U.S. coal mining areas. The 

only MSHA districts with few active pillar recovery 

operations were District 3 (northern West Virginia), District 

9 (primarily Colorado, Wyoming, and Utah), and District 10 

(western Kentucky). 

The largest number of retreat mines were found in the 

southern Appalachian coalfields (MSHA District 4 in southern 

West Virginia; District 5 in Virginia; District 6 in eastern 

Kentucky; and District 7 in eastern Kentucky and Alabama). 

These four MSHA districts accounted for 156 mines, or 85% 

of the sample. The retreat mines in this region were typically 

smaller than those in other districts, averaging 40 employees 

each, compared with 83 in the average mine outside the 

region. 

Accident rates vary from MSHA district to district, as 

shown in figure 7. Within each district, they tend to be similar 

between the retreat mine sample and all room-and-pillar 

mines. Roof and rib accident rates were lower in 1993 at the 

retreat mines in six of the eight districts. 

Table 3 and figure 8 show that retreat mines tended to be 

larger than the average room-and-pillar mines. Only 15% of 

all small mines were conducting active pillar recovery 

operations, whereas about 40% of all medium and large mines 

were recovering pil~ars.~ Accident rates did not show any 

significant trends with regard to mine size. 

- - 

3~&ill mines are those employing fewer than 50 workers; medium mines 

employ 50 to 150 workers; large mines employ more than 150 workers. 

ALL ROOM 

& PILLAR 

MINES 

RETREAT 

MlNE 

SAMPLE 

LONGWALL 

MINES 

MlNE TYPE 

Figure 3.-Accident rate at U.S. underground coal mines in 1993, by mine type, 
ALL ROOM 8 PILLAR MINES RETREAT MINE SAMPLE LONGWALL MINES 

MINE TYPE 

Figure 4.-Rooflrib accident rates at U.S. underground coal mines in 1993, by mine type. 

Table 2.-Demographics and accident statistics for U.S. underground coal mines by MSHA district' 

Average 

MSHA No. of No. of mine Tons, Total Rooflrib Total . Rooflrib 

District No. mines employees size thousand st Productivity' accident accident days lost days lost st. 

(emplovees) rate ' rate ' rate rate 

RETREAT 

10 . . . . . . . 0 0 0 0 0 0 0 0 0 

ROOM-AND-PILLAR 

2 ........ 52 2,122 4 1 11,134 2.66 26.34 1.15 785 31 

3 ........ 60 1,720 29 11,093 3.38 1 1.50 0.85 306 33 

4 ........ 256 7,490 29 49,797 3.80 16.31 1.64 564 56 

5 . . . . . . . . 164 4,090 25 21,465 2.82 12.97 1.24 456 37 

6 ........ 228 5,447 24 34,356 3.65 16.42 1.82 400 50 

7 . . . . . . . . 192 6,069 32 37,087 3.08 14.34 1.11 309 20 

8 ........ 23 3,443 150 25,435 3.75 17.27 1.41 466 37 

9 ........ 20 836 42 8,061 4.37 10.30 1.08 31 7 36 

10 ....... 19 1,856 98 15,685 4.07 16.88 2.29 41 3 65 

'Excludes anthracite mines and mines producing less than 5,000 st. 

'Accident rates are calculated as the total number of injury accidents (severity 1-6) per 200,000 hours worked. 

'Days lost rates are calculated as the total number of days lost due to injury per 200,000 hours worked. 

'%etreat mines are the 186 nonlongwall mines with active gobs identified by Urosek et al. [1995]. 

5Room-and-pillar mines include all nonlongwall mines. 

Source: MSHA Accident and Employment Data Base for 1993. 
2 3 4 5 6 7 8 9 10 

MSHA DISTRICT 

Figure 5.-Employment at sample retreat mines, by MSHA district. 

2 3 4 5 6 7 8 9 10 

MSHA DISTRICT 

Figure 6.-Number of sample retreat mines, by MSHA district. 
SMALL MEDIUM LARGE 

ROOM & PILLAR MINE SIZE I 8EFXf3FMlNE 

Figure 7.-Employment at retreat and all room-and-pillar mines, by mine size. 

2 3 4 5 6 7 8 9 

MSHA DISTRICT 

Figure 8.-Rooflrib accident rates at retreat and all room-and pillar mines, by MSHA district. 
Table 3.-Demographics and accident statistics for U.S. underground coal mines by mine size' 

Average 

No. of No. of Tons, Total Rooflrib Total Rooflri b 

Mine size mine thousand mines employees size (em- Productivity' accident accident days lost days lost 

st still 

plovees) rate rate rate rate 

RETREAT 

Small . . . . . . . . . 133 2,892 22 20,376 3.50 13.18 1.20 305 27 

Medium . . . . . . . 41 3,481 85 23,785 3.25 17.30 1.04 466 26 

Large ' . . . . . . . . . 12 2,756 230 17,539 3.06 15.83 1.22 519 35 

ROOM-AND-PILLAR 

Small . . . . . . . . . 873 17,253 20 103,912 3.50 14.76 1.44 353 37 

Medium . . . . . . . 113 9,166 81 67,299 3.57 17.07 1.45 54 1 46 

Large ' . . . . . . . . . 28 6,653 238 43,087 3.17 16.84 i .43 539 4 1 

'Excludes anthracite mines and mines producing less than 5,000 st. 

'Accident rates are calculated as the total number of injury accidents (severity 1-6) per 200,000 hours worked. 

3Days lost rates are calculated as the total number of days lost due to injury per 200,000 hours worked. 

4Retreat mines are the 186 nonlongwall mines with active gobs identified by Urosek el al. [1995]. 

5Smali mines are those employing fewer than 50 workers. 

6Medium mines are those employing 50 to 150 workers 

'~arge mines are those employing more than 150 workers. 

aRoom-and-pillar mines include all nonlongwall mines. 

ANALYSIS OF FATAL INCIDENTS 

A total of 25 fatal incidents, resulting in 33 deaths, have 

been attributed to retreat mining during 1989-96. These 

fatalities represent 25% of the I 11 roof and rib fatalities that 

occurred during this period (figure 9). Four of the retreat 

mining fatal incidents (comprising five fatalities) occurred 

during room development with no apparent influence of a gob. 

A report by MSHA has not been completed on the most recent 

incident, a double fatality in Kentucky. The appendix to this 

paper summarizes the information collected on the 20 fatal 

incidents available for analysis. 

Figure 10 shows that, in four incidents, no citations were 

issued for violations of mining law or the mine's roof control 

plan. The remaining 16 fatal incidents were divided into 2 

categories, or classes. Class 1 includes eight incidents where 

gross violation of mining law (and often common sense) was 

deemed to be the chief factor. Class 2 incidents were those 

where a violation contributed to the fatality, but where other 

factors appeared to have played an important role as well. 

Class 3 incidents were those for which no citations were 

issued. 

Figure 11 shows that the States of Kentucky, Tennessee, 

and West Virginia have accounted for 92% of all pillaring 

fatalities. Every incident in Kentucky and Tennessee involved 

a violation. All four of the no-citation incidents occurred in 

West Virginia. 

Geologic factors were cited in eight instances as 

contributing to the fatal incidents. Roof slips and slickensides 

were the most common features. The no-citation incidents 

involved a first fall, a geologic feature, and a multiple-seam 

interaction (figure 12). High vertical stress was a factor in 

three of the class 2 fatal incidents. The types of violations 

cited in the other incidents are shown in figure 13. Mining 

sequence violations were most frequently cited in the class 2 

fatal incidents. 

The mining techniques employed to extract pillars are 

shown in figure 14. All five fatalities during slabbing 

operations occurred on conventional mining sections. Partial 

pillaring or "Christmas tree" methods were used in 82% of the 

incidents where continuous miners were employed. 

In 45% of the fatal incidents, the pushout or last lift was 

being extracted at the time of the fall (figure 15). All four of 

the no-citation fatal incidents had two significant factors 

in common. All employed the Christmas tree extraction 

sequence, and in every case the continuous miner was 

extracting the last lift. 
ROOF AND RIB FATALITIES BY YEAR 

1989 1990 1991 1992 1993 1994 1995 1996 

YEAR 

Figure 9.-Rooflrib fatalities, 1989-96. 

NUMBER OF FATALI'TIES AND FATAL INCIDENTS 

CLASS 2 

FATALITIES FATAL INCIDENTS 

FATALITIES AND FATAL INCIDENTS 

Figure 10.-Fatal rooflrib incidents and fatalities associated with retreat mining, 1989-96. 
FATALITIES BY STATE 

PA V A TN WV KY 

STATE 

Figure 11 .-Retreat mining fatal incidents, by State. 

FACTORS INVOLVED IN FATAL INCIDENTS 

1 CLASS 3 1 

Ci CLASS 2 

II CLASS 1 

FACTORS 

Figure 12.-Contributing factors cited in the retreat mining fatal rooflrib incidents. 
NUMBER OF VIOLATIONS BY TYPE 

TYPE OF VIOLATION 

Figure 13.-Types of violations cited in the retreat mining fatal rooflrib incidents. 

FATALITIES BY MINING TECHNIQUE 

MINING TECHNIQUE 

CLASS 2 

Figure 14.-Pillar extraction techniques employed in the retreat mining fatal rooflrib incidents. 
FATALITIES BY CUT 

PUSHOUTILAST OTHER 

Ll FT 

CUT 

Figure 15.-Lift being extracted when fatal roof fall occurred. 

CONCLUSIONS 

Pillar recovery is employed in many U.S. coal regions. It operations can be largely attributed to violations of existing 

is practiced primarily at many medium and some small mines. mining law. Nearly 50% of fatal incidents have occurred 

The overall accident rates for retreat mines appear to be during the recovery of the final lift (or pushout). Other 

similar to those of other room-and-pillar mines. The number potential problem areas include high stressldeep cover, first 

of fatalities that have occurred during pillar recovery falls, geologic factors, mining sequence, and multiple-seam 

operations seems disproportionately high relative to coal interactions. 

production. Many fatalities that have occurred during retreat 

REFERENCES 

Blaiklock J [19921. ~. A Kentuckian vote for continuous mining. Mining Reese RC, Dash BB, Hamilton PA 119781. Coal recovew from 

Magazine June:365-366. underground bituminous coal mines in the united States, by mining method. 

Monwe 119881. Pillar recovery-the lost art? In: Proceedings of the Washington, DC: U.S. Department of the Interior, Bureau of Mines, IC 8785. 

Urosek JE, Zuchelli DR. Beiter DA [1995]. Gob ventilation and bleeder 

American Mining Congress. Chicago, IL: pp. 531-548. systems in U.S. coal mines. Soc Min Eng preprint 95-78. 
APPENDIX.-INFORMATION ON FATAL PILLAR RECOVERY INCIDENTS 

Table A-1.-InformaUon on fatal plllar recovery lncldents (from MSHA Reports of InvesUgaUon) 

Kentucky: 

Haddix ........ Upper Hignite . . 

Virginia: 

Jawbone ....... 1990: 

Kentucky: 

Creech ........ 

..... High Splint 

.... Pond Creek 

Pennsylvania: 

Pittsburgh ..... Tennessee: 

Jellico ......... 

West Virginia: 

Middle Kittanning 

1991: 

Kentucky: 

Hazard No. 4 ... 

and seam 

Slabbing ...... Conventional .... 54 

Christmas tree Cab CM ........ 60 

Seam Entry Com- Mining 

Extraction Equipment height, width, Cut Fatality occupation pliance sequence 

technique in geology R class violation 

Mod. -weak 

Mod. -weak 

Other 

Pushout 

Cut. mach. oper. 

Roof bolters (2), 

supplyman. 

Yes 

No 

Partial ........ Cab CM ........ 84 Moderate Other Shuttle car oper. 

..... ........ Other Remote CM 36 

..... Split-and-wing Remote CM 60 

..... ........ Partial Remote CM 50 

Moderate 

Weak 

Weak 

Pushout 

Pushout 

Other 

Foreman 

Foreman 

CM helper 

Yes 

No 

Yes 

........ Split-and-wing Cab CM 120 Strong Pushout Roof bolter Yes 

..... Other ........ Remote CM 40 Moderate Other CM operator Yes 

Split-and-wing Cab CM ........ 60 Weak Other Roof bolter 

..... Partial ........ Remote CM 54 Mod. -weak Other CM operator, 

CM helper. 

Timber setter 

Yes 

.... Pond Creek 

West Virginia: 

Dorothy ....... 

Hernshaw ...... 1992: 

Tennessee: 

Sewanee ...... 

West Virginia: 

Upper Dorothy . . 1993: 

Kentucky: 

Hazard No. 4 ... 

..... ........ Partial Remote CM 56 Weak Other Yes 

..... Partial ........ Remote CM 78 

..... Christmas tree Remote CM 52 

Weak 

Weak 

Other 

Pushout 

Foreman 

Foreman 

No 

Yes 

Other Roof bolter, 

utility man. 

... Slabbing ...... Conventional 36 Weak Yes 

Christmas tree Cab CM ........ 74 Strong Pushout Foreman 

........ ........ Partial Cab CM 44 Mod. -weak Other Roof bolter, 

scoop operator. 

Tennessee: 

...... Sewanee II Slabbing ...... Conventional .... 36 Moderate Other Foreman, 

roof bolter. 

Yes 

West Virginia: 

Pocahontas No. 3 

1995: 

West Virginia: 

Coalburg ...... 1996: 

West Virginia: 

Weak Pushout CM helper ........ Christmas tree Cab CM 60 

MRS operator Christmas tree Remote CM ..... 72 Strong Last cut 

..... Christmas tree Remote CM 84 Moderate 20 Pushout CM operator 3 No 
Table A-1.-lnformatlon on fatal pillar recovery Incidents (from MSHA Reports of Investlgatlon)-ConUnued 

Victim under Excessive Excessive Saved Kil

Year, State, unsupported span Geological factor stress First victim location fall by wi

canopy no ca

Kentucky: 

Haddix . . . . . . . . Upper Hignite . . Virginia: 

Jawbone . . . . . . . 1990: 

Kentucky: 

Creech . . . . . . . . High Splint . . . . . 

Pond Creek . . . . Pennsylvania: 

Pittsburgh . . . . . 

Tennessee: 

Jellico . . . . . . . . . West Virginia: 

Middle Kittanning 

1991: 

Kentucky: 

Hazard No. 4 . . . 

Pond Creek . . . . West Virginia: 

Dorothy . . . . . . 

Hernshaw . . . . . . 

1992: 

Tennessee: 

Sewanee . . . . . . 

West Virginia: 

Upper Dorothy . . 

1993: 

Kentucky: 

Hazard No. 4 . . . 

Tennessee: 

Sewanee . . . . . . 

West Virginia: 

Pocahontas No. 3 

1995: 

West Virginia: 

Coalburg . . . . . . 1996: 

West Virginia: 

Beckley . . . . . . . . /I 

No Yes No No No Intersection No Yes 

No No Slickenslided slip No No No No 

lntersection 

No No No Depth No Other No Yes 

Yes Yes No 

Yes No Slickenslided slip 

Yes Intersection No No 

No Intersection No No 

No Other Yes No 

No Other No No 

No Other No No 

No Other No No 

No No Hill seams; shale thickened No No Other No No 

No No No No No Intersection No No 

No No Roof fractures (cutters) Depth No Intersection No No 

No No Slickenslides No No Other No No 

Yes Slickenslided slip 

No No 

No No Other; intersection No No 

Multiple seam No Intersection Yes No 

No No No Depth No Intersection No No 

No Yes Slickenslided slip No No Intersection No No 

No No Slickenslided horseback No No Other Yes No 

No No No No Yes Other No No 

No No No No No Intersection No No 
Table A-1.-lnformatlon on fatal plllar recovery lncldents (from MSHA Reports of Investlgatlon)-J2ontlnued 

Year, State, 

and seam 

1989: 

Kentucky:. 

Haddix . . . . . . . . Upper Hignite . . Virginia: 

Jawbone . . . . . . . 1990: 

Kentucky: 

Creech . . . . . . . . High Splint . . . . . PondCreek . . . . . Pennsylvania: 

Pittsburgh . . . . . Tennessee: 

Jellico . . . . . . . . . West Virginia: 

Middle Kittanning 

1991: 

Kentucky: 

Hazard No. 4 . . . Pond Creek . . . . 

West Virginia: 

Dorothy . . . . . . . Hernshaw . . . . . . 1992: 

Tennessee: 

Sewanee . . . . . . West Virginia: 

Upper Dorothy . . 1992: 

Kentucky: 

Hazard No. 4 . . . 

Tennessee: 

Sewanee . . . . . 

West Virginia: 

Pocahontas No. : 

1995: 

West Virginia: 

Coalburg ..... 

1996: 

West Virginia: 

Beckley . . . . 

Remarks 

The cutting machine did not have a protective canopy. Slabbing created a span that was about 30 ft at the accident site. 

fictims in dangerous location during mining of pushout. No signs of weight on the section. Miner operator was 12 ft inby roof bolts durin

3ib sloughing caused fatal injuries to the shuttle car operator. The shuttle car did not have a car or canopy, as required when the mining 

Visible cracks in the mine roof in the intersection where the accident occurred prior to the day of the accident. Exceeded aft cut by 18 ft 

The victim was in a hazardous position while the continuous miner was being operated. The first pillar fall occurred 45 min before the fall 

4ctual mining far exceeded the approval plan. The roof fell 5 to 6 ft to a plane of weakness above the roof line, which was above the anch

A sufficient amount of coal was not left in place to adequately support the roof or was not removed to allow the roof to fall in a controlled 

Five employees were positioned 10 to DO ft inby permanent roof support. Last cut extended 61 ft; only 20 ft was permitted by roof contro

The canopy over the controls was for another model of roof bolter. NO additional support was installed inby the edge of the slip. 

The hill seams were not adequately supported. Not enough coal was left inby the cut in the NO. 2 pillar; the cut in the No. 3 pillar was sta

The cut taken out of the right side of pillar No. 4 measured 23 ft wide and 39 ft deep. The approved roof control plan states that no cut wi

and 20 ft deep. 

Accident area was developed 6 years earlier. A fatality and roof falls occurred previously with this system of mining. MSHA deemed the r

Approved cut depths were exceeded. Several lifts had been extracted without any supports installed. 

Slabbing created wide unsupported spans. A slickensided slip was not properly supported. Employees routinely traveled inby permanen

Four coal seams had been mined underneath this mine. There were no violations. The shale below the sandstone concealed the cracks 

High pressure was evident in the fall location. This resulted from stresses on the active pillar line and an adjacent pillar area. The size of 

line was not sufficient to prevent pressure from overriding the pillar line. Due to equipment design and pillar orientation, full and partial 

the same pillar line. 

Instead of making splits and lifts, slabbing was done. Slabbing created a span of 30 by 50 ft in the intersection where the accident occur

were not set as required. Loader operator was inby permanent supports when loading lifts. 

There were no violations on the section. A large slickensided horseback fell, causing the accident. The miner operator was saved by a ca

The pressure gauges on the MRS units are small and located in positions requiring operators to stand close to the MRS units to observe 

were being extracted 150 ft from the coal outcrop. The hydraulic gauges took a sudden rise of several hundred pounds immediately be

Some of the pillars were partially mined; others were fully mined in the same row due to poor roof conditions. There was minor floor hea

the accident, but not elsewhere on the section. 
ANALYSIS OF RETREAT MINING PILLAR STABILITY (ARMPS) 

By Christopher Mark, Ph.D.,' and Frank E. Chase2 

ABSTRACT 

The prevention of pillar squeezes, massive pillar collapses, and bumps is critical to safe pillar recovery 

operations. To help prevent these underground safety problems, the Pittsburgh Research Center has developed 

the Analysis of Retreat Mining Pillar Stability (ARMPS) computer program. AMPS calculates stability factors (SF) based on estimates of the loads applied to, and the load-bearing capacities of, pillars during retreat 

mining. The program can model the significant features of most retreat mining layouts, including angled 

crosscuts, varied spacings between entries, barrier pillars between the active section and old (side) gobs, and 

slab cuts in the barriers on retreat. It also features a pillar strength formula that considers the greater strength 

of rectangular pillars. The program may be used to evaluate bleeder designs, as well as active workings. 

A data base of 140 pillar retreat case histories has been collected across the United States to verify the 

program. It was found that satisfactory conditions were very rare when the ARMPS SF was less than 0.75. 

Conversely, very few unsatiifactory designs were found where the ARMPS SF was greater than 1.5. 

Preliminary analyses also indicate that pillar failures are more likely beneath sandstone roof and that the 

ARMPS SF may be less meaningful when the depth of cover exceeds 230 m (750 ft). 

- 

'Mining engineer. 

'Geologist. 

Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 
INTRODUCTION 

The use of remote-control continuous miners, extended 

cuts, and mobile roof supports has increased the productivity 

of room-and-pillar retreat mining (also referred to as "pillaring," "pillar recovery," "robbing," and "second mining"). In 

the southern Appalachian coalfields, many mines are choosing 

room-and-pillar retreat mining because of its lower capital 

cost and greater flexibility [Blaiklock 19921. Unfortunately, 

between 1989 and 1996, 25% of all roof and rib fatalities occurred on pillar recovery sections. 

Roof fall accidents are not the only problem associated 

with retreat mining. Millions of tons of coal are sterilized 

annually because of pillar squeezes, floor heave, pillar line 

roof falls, and pillar bumps. Traditional pillar design 

methods are of little help due to the complex mining 

geometries and abutment pressures that are present during 

pillar extraction. The Pittsburgh Research Center has 

developed the Analysis of Retreat Mining Pillar Stability 

(ARMPS) computer program to aid in the design of pillar 

recovery operations. This paper describes the program and 

presents the findings thus far. 

THE ARMPS METHOD 

The goal of ARMPS is to help ensure that the pillars developed for future extraction (production pillars) are of adequate size for all anticipated loading conditions. The key is to 

be able to estimate the magnitudes of the various loads that the 

pillars might experience throughout the mining process. The 

formulas used in ARMPS are based on those originally developed for the Analysis of Longwall Pillar Stability (ALPS) 

method, which is widely used for longwall pillar design [Mark 

1990, 19921. ALPS was initially derived from underground 

measurements of longwall abutment stresses and was later 

validated by the back-analysis of more than 100 case histories. 

In ARMPS, the formulas have been extensively modified for 

the variety of mining geometries typically found in pillar recovery operations. 

USER INPUT 

The first step in using the ARMPS program is to enter the 

dimensions of the pillars in the worlung section, as illustrated 

in figure 1. The program can accommodate angled crosscuts, 

varied spacings between the entries, and barrier pillars between 

the active section and old (side) gob areas. Slabbing of barriers 

, Boundary of, 

Entry centers (variable) 

Fiqure 1.-Section layout parameters used in ARMPS. 
on retreat can also be included. Other parameters that must be 

defined include depth of cover, mining height, entry width, and 

crosscut spacing. Finally, the user chooses one of four possible 

loading conditions (figure 2). The simplest, loading Condition 1, is development loading only. Loading condition 2 occurs when the active, or "front," panel is being fully retreated 

and there are no adjacent mined-out areas. The total applied 

load is the sum of the development loads and the fiont abutment 

load. Loading condition 3 occurs where the active mining zone 

(AMZ) is adjacent to an old (side) gob and the pillars are 

subjected to development, side abutment, and front abutment 

loads. Where the pillar line is surrounded by gob on three sides 

(sometimes referred to as "bottlenecking"), loading condition 4 

is used. In every case, the extent of each gob is defined by the 

user. 

ARMPS STABILITY FACTOR FOR THE 

ACTIVE MINING ZONE 

The basic output from the ARMPS program is the stability 

factor (SF), defined as 

ARMPS SF = LBCLT, (1) 

where LBC = the estimated total load-bearing capacity of 

the pillars within the AMZ, 

and LT = the estimated total load applied to pillars 

within the AhIZ. 

Figure 3 illustrates the development and front abutment loads 

applied to the AMZ. 

Loading condition 1 Loading condition 2 

Loading condition 3 Loading condition 4 

Figure 2.-The four loading conditions that can be evaluated with ARMPS. 
GL 4 AMZ 4 

KEY 

AMZ Active mining zone H Depth of cover 

B Abutment angle LA Abutment load 

EFW Extraction front width LD Development load 

GL Mined out area 

Figure 3.-Schematic showing the active mining zone, the development load, and the front abutment load. 

The AMZ includes all of the pillars on the extraction front 

(or "pillar line") and extends outby the pillar line a distance of 

five times the square root of the depth of cover (5m). This 

distance was selected because measurements of abutment stress 

distributions [Mark 19901 show that 90% of the front abutment 

load falls within its boundaries (figure 4). 

AMPS calculates the SF for the entire AMZ, rather than 

stability factors for individual pillars, because experience has 

shown that the pillars within the AMZ typically behave as a 

system. If an individual pillar is overloaded, it will normally 

transfer its excess load to adjacent pillars. If those pillars are 

adequately sized, the process ends there. A pillar squeeze occurs only when the adjacent pillars are also undersized. They 

then fail in turn, resulting in a "domino" of load transfer and 

pillar failure. The ARMPS SF is therefore a measure of the 

overall stability of the pillar system. 

PILLAR LOAD-BEARING CAPACITY 

The load-bearing capacity of the AMZ is calculated by summing the load-bearing capacities of all of the pillars within its 

boundaries. The strength of an individual pillar (SP) is determined using a new pillar strength formula (the Mark-Bieniawski formula) that considers the effect of pillar length: 

where S, = in situ coal strength, assumed = 6.2 MPa 

(900 psi), 

w = pillar width, 

h = pillar height, 

and L = pillar length. 

The new pillar strength formula was needed because the pillars used in retreat mining are often much longer than they are 

wide. The strength of rectangular pillars can be significantly 

greater than square pillars due to the greater confinement generated within them. The Mark-Bieniawski formula was derived 

from analyses of the pillar stress distributions implied by empirical pillar strength formulas. A complete discussion of the 

Mark-Bieniawski formula is included in appendix A of this 

paper. The in situ coal strength is assumed to be 6.2 MPa 

(900 psi) in ARMPS; however, this value can be modified by 

the user. 

The load-bearing capacity of the pillars is determined by 

multiplying their strength by their load-bearing area. When 

angled crosscuts are employed, the algorithm still calculates 

accurately each pillar's least dimension, length, and loadbearing area (A,): 
Pillar Pillar Pillar Pillar 

3 4 

9.3m 

Abutment 

stress KEY 

Figure 4.-Distribution of abutment stress, showing that 90% 

of the abutment falls within the distance of (5q from the gob edge. 

1 

where XC = center-to-center crosscut spacing, 

H Overburden depth 

Front abutment load applied 

to the active mining zone 

ECTR = center-to-center entry spacing, 

We = entry width, 

and I$ = angle between the crosscut and the entry. 

The load-bearing capacity of the pillar system is then obtained by summing the capacities of the individual pillars within 

the AMZ. ARMPS calculates the strength and load-bearing capacity of barrier pillars in the same manner as the panel pillars, 

except that their length is limited to the breadth of the AMZ. 

PILLAR LOADINGS 

The loadings applied to the AMZ include development loads, 

abutment loads, and loads transferred from barrier pillars. Table 1 shows the sources of loads and the loading conditions in 

which they occur. 

Table 1 .-Loads applied to the active mining zone In ARMPS 

-- - 

~oadin~ condition 

Source of load .I97 A 

Front abutment .................... XXX 

Side gob abutments ................ X X 

Transfer from barriers between 

active mining zone and side gobs ... X X 

Transfer from remnant barriers 

between front aob and side aobs .... X X 

Development loads are due to the weight of the overburden 

directly above the pillars before any retreat mining takes place. 

The tributary area theory is used in ARMPS to estimate development loads. 

Abutment loads occur as a result of retreat mining and gob 

formation. They are determined by the depth. of cover, the extent of the gobs, the width of the extraction front, and the abutment angles. These parameters are illustrated in two dirnensions in figure 5. The abutment angle determines how much 

load is carried by gob. Measurements of longwall abutment 

stresses indicated that an abutment angle of 21" is appropriate 

for normal caving conditions [Mark 19921. The ARMPS program initializes the abutment angles for all gobs to 21"; 

however, this can be changed by the user. For example, if it is 

known that no caving has occurred, then the abutment angle 

may be set to 90" to simulate zero load transfer to the gob 

[Chase and Mark 19931. 

The abutment stresses are assumed to be distributed following the inverse-square function shown in figure 4. Abutment 

loads are also applied to barrier pillars; however, if a barrier is 

too small to carry its share, then some or all of the excess is 

transferred to the AMZ. 

The front abutment load applied to the AMZ is calculated as 

follows. The volume of the overburden above the mined-out 

active gob is the depth of cover multiplied by the gob area. The 

portion of this volume whose weight is carried by the gob is 

determined by the tangent of the abutment angle, as shown in 

figure 5. This portion is subtracted, and the remainder is shared 

between the AMZ and the unmined coal on the other three sides 

of the gob. It is assumed that barrier pillars (or substantial production pillars) are present on the other three sides of the gob. 

Load applied to the barriers here may be transferred back to the 

AMZ if the barriers are removed later in the mining process. 

The magnitude of the front abutment load applied to the 

AMZ is determined by the extent of the extraction zone and the 

depth of cover. The front abutment is considered fully developed if the gob area is large relative to the depth of cover 

(figure 6A). If only a few rows of pillars have been extracted 

(figure 6B), much of the load will be carried by the back barrier. 

If the full extraction zone is rather narrow (figure 6C), much of 

the load will be carried by the side barriers. 

The side abutment loads are shared by the AMZ and, if it is 

present, the barrier pillar between the AMZ and the side gob. 

The inverse-square stress distribution (figure 4) again is used to 

apportion the load between the barrier and the AMZ. Next, if 

it is determined that the barriers are overloaded, some additional 

side abutment load is transferred to the AMZ. 

To determine whether a barrier pillar can carry the load applied to it, ARMPS estimates the barrier's SF by dividing its 

load-bearing capacity by its load. The total load applied to a 

barrier pillar is the sum of the development load, the front abutment load due to any slabbing, and the side abutment load 

applied to the barrier. If the SF is greater than 1.5, the barrier 

is assumed to be stable. When the barrier's SF is between 1.5 
A 

LDEV 

B 

T 

- 

H tan B DEPAMZ 

I 3 1 LDEV 1 u 

GEXT/2 

- 

DEPAMZ 

KEY 

H Overburden GEXT Gob extent 

LS Abutment load DEPAMZ Breadth of active 

LDEV Development load mining zone 

B Abutment angle 

Figure 5.-Schematic showing the abutment load in two dimensions. A, supercritical gob; B, 

subcritical gob. 

A H tan B < GEXT/2 

and 

H tan B < WT/2 

B GEXT/2 < H tan B 

and 

GEXT/2 < WT/2 

KEY 

H Overburden depth 

B Abutment angle 

WT Width of active 

mining zone 

GEXT Gob extent 

C - WT/2 < H tan B 

and 

WT/ 2 < GEXT/2 

Figure 6.-Illustration of the effect of panel geometry on the front abutment loading in ARMPS. 

A, gob area is supercritical in both width and extent; 6, gob area is subcritical in extent; C, gob area 

is subcritical in width. 
and 0.5, a portion of its abutment load is transferred to the of it is returned to the AMZ. The decision to transfer the load 

AMZ. If the SF is less than 0.5, all of the additional side abut- and how much is based on the remnant barrier's SF. Slabbing 

ment load (but not the development or front abutment load) is of the remnant will also return some abutment load to the AMZ. 

transferred to the AMZ. Further details on the formulas and calculations used in 

The final sources of load on the AMZ are the remnant barrier ARMPS loadings can be found ic the "Help" text that accompillars inby the pillar line (between the front and side gobs). If panies version 4.0 of the program. 

the remnant barriers are too small to carry their load, some part 

VERIFICATION OF THE ARMPS METHOD 

The ARMPS method is being verified through back-ar,a!ysis 

of pillar recovery case histories. To date, 14U case histories 

have been obtained from 10 States (see appendix B of this paper). They cover an extensive range of geologic conditions, 

roof rock cavability characteristics, extraction methods. depths 

of cover, and pillar geometries. Ground condit~ons in each case 

history have been categorized as either satisfactory or unsatisfactory. Pillar failures responsible for unsatisfactory conditions 

were found to include- 

* Pillar squeezes, accompanied by significant entry closure 

and loss of reserves; 

Sudden collapses of groups of pillars, usually accompanied by airblasts; and/or 

Coal pillar bumps (violent failures of one or more 

pillars). 

As figure 7 shows, pillar failures occurred in 93% of the 

cases where the ARMPS SF was less than 0.75. Where the 

ARMPS SF was greater than I .5, 94% of the designs were 

satisfactory. SF values ranging from 0.75 to 1 .SO form a "gray" 

area where both successful and unsuccessful cases are found. 

Curren! research has begun to evaluate other factors that may 

contribult: LU satisfactory conditions when thc ARMPS SF falls 

between 0.75 and 1.5. These includeCoal strength: An extensive data base of laboratory tests of 

the strength of coal was compiled by Mark and Barton [I 9971. 

When compared with the ARMPS data base, no correlation was 

found between coal strength and pillar strength. 

Deprlz of cover: Figure 8 shows that there is a marked reduction in SF as depth of cover increases. When the depth exceeds 

305 m (1,000 ft), the ARMPS SF was below 1.0 for 70% of the 

satisfactory designs. Highly unsatisfactory conditions have also 

been encountered under deep cover, which recently led LO two 

fatalities. Pillar design for retreat mining under deep cover remains an important research issue. 

Sear11 height: A plot of seam height against ARMPS SF 

shows no correlation (figure 9). 

Roof geology: A detailed study of pillar performance was 

conducted at a mining complex in southern West Virginia. 

More than 50 case histories were collected. Analysis showed 

that satisfactory conditions were more likely to be encountered 

under shale roof than massive sandstone roof (figures 10-1 I). 

This implies that better caving occurs with shale, resulting in 

lower pillar loads. 

ARMPS STABILITY FACTOR 

Figure 7.-ARMPS data base. 
0 

0 500 1000 1500 2000 

Depth of Cover (H), ft 

Figure 8.-Relationship between ARMPS SF and depth of cover within the case history data base. 

0 5 I0 15 20 

Seam Height (h), ft 

Figure 9.-Relationship between ARMPS SF and seam height within the case history data base. 
ARMPS Stability Factor 

Figure 10.-Shale roof case histories from mining complex in southern West Virginia. 

ARMPS Stability Factor 

Figure 11.-Sandstone roof case histories from mining complex in southern West Virginia. 
GUIDELINES FOR USING ARMPS 

ARMPS appears to provide good first approximations of the 

pillar sizes required to prevent pillar failure during retreat 

mining. In an operating mine, past experience can be incorporated directly into ARMPS. ARMPS stability factors can be 

back-calculated for both successful and unsuccessful areas. 

Once a minimum ARMPS SF has been shown to provide adequate ground conditions, that minimum should be maintained 

in subsequent areas as changes occur in the depth of cover, coal 

thickness, or pillar layout. In this manner, ARMPS can be 

calibrated using site-specific experience. 

ARMPS is also well suited for initial feasibility studies 

where no previous experience is available. Operators may begin with an SF near 1.5, then adjust as they observe pillar 

performance. ARMPS may also help in optimizing panel designs by identifying pillars that might be needlessly oversized. 

ARMPS may be used to analyze a wide variety of mining 

geometries. For example, most bleeder designs can be analyzed 

by selecting loading condition 3, then setting the extent of the 

active gob to zero. The "Help" text included with version 4.0 

of the program contains many tips on selecting the proper input 

parameters when using ARMPS. 

In some cases, more detail may be desired ihan can be provided by ARMPS. Some complex situations, such as multipleseam interactions, are beyond the capabilities of ARMPS. In 

these instances, the newly developed LAMODEL [Heasley 

19971 may be the appropriate tool to use. 

CONCLUSIONS 

The ARMPS program has already proven to be a useful aid 

in planning pillar recovery operations. It is easy to use, and a 

large number of analyses can be run in a relatively short period. 

The program is sufficiently flexible to be applicable to a wide 

variety of mining geometries. If the user desires, it also provides a full range of intermediate calculations in addition to the 

SF. Many mines throughout the United States and abroad 

already use ARMPS, and the Mine Safety and Health Administration has also made extensive use of the program. 

Current efforts are aimed at improving the interpretation of 

the ARMPS SF. Although pillar failures seem unlikely when 

the ARMPS SF is greater than 1.5, there are apparently many 

cases where SF values as low as 0.75 have been successful. 

Factors such as roof quality, floor strength, and mining method 

may determine whether a pillar design succeeds. These factors 

are now being included in the retreat mining case history data 

base and will be integrated into future design guidelines. 

To obtain a single copy of the ARMPS computer program, 

version 4.0 for Windows, send three double-sided, high-density 

diskettes to: Christopher Mark, Ph.D., NIOSH, Pittsburgh Research Center, Cochrans Mills Rd., P.O. Box 18070, Pittsburgh, 

PA 15236-0070. 

REFERENCES 

Blaiklnck I [!PB2j. .4 Kentuckian vote for continuous mining. Mining 

Magazine June:365-366. 

Chase FE, Mark C [1993]. Ground control design for pillar extraction. Soc 

Min Eng AIME preprint 93-282. 

Heasley KA [1997]. A new laminated overburden model for coal mine 

design. In: Mark C, Tuchman RJ, comp. Proceedings: New Technology for 

Ground Control in Retreat Mining. Pittsburgh. PA: U.S. Department of Health 

and Human Services, Public Health Service, Centers for Disease Control, 

National Institute for Occupational Safety and Health, 1C 9446. 

Mark C [1990]. Pillar design methods for longwall mining. Pittsburgh. PA: 

U.S. Department of the Interior. Bureau of Mines, IC 9247. 

Mark C [1992]. Analysis of longwal! pilla: stability (ALPS): m update. In: 

Iannacchione AT, Mark C, Repsher RC, Tuchman RJ, Jones CC, comp. 

Proceedings of the Workshop on Coal Pillar Mechanics and Design. Pittsburgh, 

PA: U.S. Department of the Interior, Bureau of Mines, IC 9315, pp. 238-249. 

Mark C, Barton TM [1997]. Pillar design and coal strength. In: Mark C, 

Tuchman RJ, comp. Proceedings: New Technology for Ground Conmol in 

Retreat Mining. Pittsburgh, PA: U.S. Department of Health and Human 

Services, Public Health Service, Centers for Disease Control, National Institute 

for Occupational Safety and Health, IC 9446. 
APPENDIX A.-DERIVATION OF THE MARK-BIENIAWSKI PILLAR STRENGTH FORMULA 

Early versions of the ARMPS program, following the ALPS 2. The stress gradient within the yield zone of a given pillar 

program, used the Bieniawslu formula to estimate pillar strength does not change with time or load (i.e., the yielded coal is 

[Bieniawski 19921: perfectly plastic). 

3. The stress distribution is symmetric with respect to the 

Sp = S, [0.64 + (0.36 wh)], (A-1) center of the pillar. 

where S, = pillar strength, 

S, = in situ coal strength, 

w = pillar width (or least plan dimension), 

and h = pillar height. 

The Bieniawski formula was originally developed in the 

1960's from in situ testing of large-scale coal specimens. The 

specimen strengths were determined as the ultimate loadbearing capacity divided by the area. Bieniawski recognized 

that the formula underestimated the strength of rectangular pillars; however, because all of the specimens were square, there 

was no obvious way of estimating a "pillar length" effect. 

It has been recognized that a major disadvantage of empirical formulas, like that of Bieniawski, is that they treat the pillar as a single structural element. In reality, the stress within 

even a relatively small pillar is highly nonuniform. Tests conducted by Wagner [I9741 demonstrated this quite dramatically 

(figure A-1). 

Modern mechanics-based approaches to coal pillars begin 

with stress distribution. Perhaps the best known is the approach 

proposed by Wilson [1973, 19831. Wilson derived an expression for the vertical stress gradient within the yield zone, which 

he then integrated over the area of the pillar (figure A-2) to 

determine the ultimate pillar resistance (R). The "pillar 

strength" is simply the ultimate pillar resistance divided by the 

pillar area. Numerical models also provide stress distribution 

profiles, although not normally in the form of an equation. 

Mechanics-based approaches can be used to evaluate any pillar 

shape, because the stresses within the pillar are determined by 

laws that are independent of overall pillar geometry. 

Although empirical formulas do not explicitly consider the 

effect of internal pillar mechanics, it is apparent that they imply 

a nonuniform stress distribution because of the shape effect. 

Once the implied stress gradient has been derived, the length effect can be readily determined. The derivation has been published previously [Mark et al. 1988; Mark and Iannacchione 

19921 and is summarized below. 

First, three assumptions are implicit in Wilson's and other 

analytical formulations: 

1. The stress within the yield zone of a given pillar is a continuous function of the distance from the nearest rib. 

Figure A-1.-Pillar stress profiles measured in small coal 

pillars (after Wagner [I 9741). 

Figure A-2.-Determination of pillar load-bearing capacity as 

the integral of the pillar stress distribution. 
The next step in the derivation is to calculate the ultimate 

resistance of a square pillar. Using the Bieniawski formula: 

W 

dR = 4 oV dx dw. i 

(A-2) Equating A-3 and A-4 aid simplifying, we have 

Then, the increase in pillar resistance dR due to an increase 

in cross-sectional area dA = 2w dw (figure A-3A) may be ('4-5) 

calculated by taking the derivative of equation A-2 with respect 

to w: The function that satisfies equation A-5 is 

Equation A-6 is the stress gradient in the yield zone preIn the next step, the assumption that the vertical pillar stress dicted by the Bieniawski formula. Stress gradients have also 

is a continuous function of the rib distance (x) is applied. It been derived for several other common empirical pillar strength 

may be seen (figure A-3B) that formulas [Mark and Iannacchione 19921. 

dw 

Figure A-3.-Determination of pillar stress gradients from a pillar strength formula. A, calculation of dR directly from the formula; B, calculation of dR in terms of the vertical stress gradient. 
To determine the load-bearing capacity of any pillar shape, Dividing by the pillar area (Lw) yields the strength of a strip 

it is now only necessary to integrate equation A-6 over the load- pillar (SJ: 

bearing area of the pillar. For example, the load-bearing capacity of an extremely long strip pillar (R,) is 

RB = 2L S, 0.64 + i! (A-7) Equation A-9 implies that a strip pillar's strength can approach 

150% that of a square pillar, but that the strength difference is 

reduced as the wh ratio is reduced. 

Solving: The ultimate load carried by a rectangular pillar is equivalent 

RE = (Lw) S, (0.64 + 0.14 x) . (A-8) to the load carried by a square pillar of width w plus a section 

\ h ) of a strip pillar of length (L - w), as shown In figure A-4; Cornbining equations A-6 and A-9, the ultimate load carried by n 

rectangular pillar (R,) is 

Simplifying: 

KEY 

L Pillar length 

6L-I 

RECTANGULAR PILLAR 

W Pillar width 

EQU/VALENT 

Tn 

IV 

PLUS 

SQUARE PILLAR SECTION OF 

STRIP PILLAR 

(A- 1 0) 

(A- 1 1) 

Figure A-4.-Pillar stress distributions for square, strip, and rectangular pillars. 
Dividing by the load-bearing area 

formula is obtained: r I \ 

(wL), the Mark-Bieniawski Table A-1 .-Pillar strength from the Mark-Bieniawski 

formula, assuming the strength of a square pillar 

-1 (original Bieniawski formula) as unity (xJ - o.~~ - . (A-12) "'I Lh Pillar uw 1 2 4 

Pillar wlh 

10 20 

Equation A- 12 indicates that the increase in strength in a rec- 1.5 . . . . . . . . 1.06 1.09 1.12 1.14 1.16 

tangular pillar depends on both (wlh) and (w/L). Table A-1 2.0 . . . . . . . . .23 

4.0 ........ 1.14 1.23 1.32 1.41 1.45 

compares the pillar strengths determined by the Mark- , , , , , . , 

Bieniawski formula with those obtained from the Bieniawski 

formula. 

REFERENCES 

Bieniawski ZT [1992]. A method revisited: coal pillar strength formula 

based on field investigations. In: Iannacchione AT, Mark C, Repsher RC, 

Tuchman RJ, Jones CC, comp. Proceedings of the Workshop on Coal Pillar Mechanics and Design. Pittsburgh, PA: U.S. Department of the Interior, Bureau of 

Mines, IC 931 5, pp. 158-165. 

Mark C, Iannacchione AT [1992]. Coal pillar mechanics: theoretical models 

and field measurements compared. In: Iannacchione AT, Mark C, Repsher RC, 

Tuchman RJ, Jones CC, comp. Proceedings of the Workshop on Coal Pillar Mechanics and Design. Pittsburgh, PA: U.S. Department of the Interior. Bureau of 

Mines, IC 9315, pp. 78-93. 

Mark C, Listak JM, Bieniawski ZT [1988]. Yielding coal pillars-field 

measurements and analysis of design methods. In: Proceedings of the 29th U.S. 

Symposium on Rock Mechanics. Minneapolis, MN: University of Minnesota, 

pp. 261-270. 

Wagner H [1974]. Determination of the complete load-deformation characteristics of coal pillars. In: Proceedings of the 3rd International Society for 

Rock Mechanics Congress. Denver, CO: pp. 1076- 108 1. 

Wilson AH [1973]. An hypothesis concerning pillar stability. Min Eng 

(London) 131(14 1):409-417. 

Wilson AH [1983]. The stability of underground workings in the soft rocks 

of the coal measures. Intl J Rock Mech Min Sci 1:91-187. 
APPENDIX B.-ARMPS CASE HISTORY DATA BASE 

Table B-1.-Unsatisfactory pillar retreat case histories 

State and coal seam 

Seam thick- Loading 

SF ness. rn ~ft) Depth. m (ft) 

Alabama: 

Blue Creek .................. 

Blue Creek .................. 

Colorado: 

Cameo ..................... 

D .......................... 

D .......................... 

Kentucky: 

Harlan ..................... 

Harlan ..................... 

Harlan ..................... 

Harlan ..................... 

Hazard No . 4 ................ Hazard No . 4 ................ 

Hazard No . 4 ................ 

Lower Elkhorn (No . 2 Gas) .... 

Lower Elkhorn (No . 2 Gas) ..... 

Ohio: 

Lower Freeport .............. 

Mahoning ................... Mahoning ................... Pennsylvania: 

Lower Kittanning ............. 

Lower Kittanning ............. 

Lower Kittanning ............. 

Pittsburgh ................... 

Pittsburgh ................... Pittsburgh ................... 

Pittsburgh ................... 

Sewickley ................... 

Tennessee: 

BeachGrove ................ 

BeachGrove ................ Utah: 

Blind Canton ................ 

Gilson ...................... 

Gilson ...................... 

Lower O'Connor .............. Virginia: 

Blair ....................... 

Glarnorgan .................. 

Jawbone .................... Jawbone .................... 

Pocahontas No . 3 ............ 

Pocahontas No . 3 ............ 

Pocahontas No . 4 ............ 

West Virginia: 

Beckley .................... 

Coalburg ................... 

Coalburg ................... 

Coalburg ................... 

Coalburg ................... 

Coalburg ................... 

See explanatory notes at end of table 
Table B-1 .-- Unsatisfactory pillar retreat case histories.4 ontinued 

Seam thick- State and coal seam Loading 

SF ness. m (R) Depth. m (R) condition 

West Virginia..< ontinued 

Dorothy .................... 

Dorothy .................... Dorothy (Wlnifrede) ... : ....... 

Dorothy (Winifrede) ........... 

Dorothy (Wlnifrede) ........... 

Dorothy (Winifrede) ........... 

Dorothy (Wlnifrede) ........... No . 2 Gas ................... 

Stockton .................... 

Stockton .................... 

Stockton .................... 

Stockton .................... Stockton .................... 

( ') ......................... 1 ........................ 0.82 1.4 (4.5) 11 5 (375) 1 

NAp Not applicable . 

'Not provided by original reference . 
Table B-2.-Satisfactory pillar retreat case histories 

State and coal seam ARMPS SF Seam thick- Depth. m (ft) Loading 

ness. m (ft) condition 

Alabama: 

Blue Creek ................... 

Colorado: 

Cameo ...................... 

Cameo ...................... 

Cameo ...................... 

D ........................... 

D .......................... !!linois: 

Herrin No . 6 .................. 

Kentucky: 

Harlan ...................... 

Hazard No . 4 ................. 

Kellioka ..................... 

Kellioka ..................... 

Kellioka ..................... 

Kellioka ..................... 

Lower Elkhorn (No . 6 Gas) ...... 

PondCreek .................. 

PondCreek .................. 

PondCreek .................. 

PondCreek .................. 

PondCreek .................. 

Ohio: 

Lower Freeport ............... 

Lower Freeport ............... 

Mahoning .................... 

Pennsylvania: 

L~wer~Freeport ............... 

Lower Kittanning .............. 

Lower Kittanning .............. 

Lower Kittanning .............. 

Lower Kittanning .............. 

Lower Kittanning .............. Pittsburgh .................... 

Pittsburgh .................... 

Sewickley .................... 

Sewickley .................... ilpper Freeport ............... 

Tennessee: 

BeachGrove ................. 

Utah: 

Gilson ....................... 

Virginia: 

Blair ........................ Glamorgan ................... 

Jawbone ..................... 

Jawbone ..................... 

Jawbone ..................... 

Mossy-Haggy ................. 

Pocahontas No . 3 .......... : .. 

Pocahontas No . 3 ............. Pocahontas No . 3 ............. 

Pocahontas No . 4 ............. 

Pocahontas No . 4 ............. 

Pocahontas No . 4 ............. RedAsh ..................... 

Red Ash ..................... Tiller ........................ 

See explanatory notes at end of table . 
Table B-2.-Satisfactory pillar retreat case histories-Continued 

State and coal seam ARMPS SF Seam thick- Depth, m (ft) Loading 

ness, m fft) condition 

West Virginia: 

Beckley ..................... 0.90 1.8 (6.0) 350 (1,150) 4 

Beckley ..................... 1.17 2.7 (9.0) 260 (850) 4 

Coalburg .................... 1.14 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1.30 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1.41 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1 .SO 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1.59 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1.76 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 1.91 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 2.17 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 2.37 2.7 (9.0) 120 (400) NAP 

Coalburg .................... 2.41 2.7 (9.0) 120 (400) NAP 

Dorothy (Winifrede) ............ 2.10 3.4 (1 1 .O) 70 (225) 2 

Dorothy (Winifrede) ............ 1.32 3.0 (1 0.0) 85 (285) 2 

Dorothy (Winifrede) ............ 1.49 3.0 (1 0.0) 100 (325) 2 

Dorothy (Winifrede) ............ 1.72 3.0 (10.0) 70 (225) 2 

Fire Creek ................... 1.24 1.4 (4.5) 260 (850) 2 

Lower Wlnifrede ............... 1.73 2.0 (6.5) 185 (600) 2 

Peerless ..................... 1.56 1.4 (4.75) 2 1 5 (700) 2 

Sewell ...................... 2.55 1.2 (4.0) 105 (350) 2 

Stockton ..................... 1.56 3.0 (10.0) 65 (220) 2 

Stockton ..................... 1.99 3.0 (10.0) 75 (245) 2 

NAp Not applicable. 
PREVENTING MASSIVE PILLAR COLLAPSES 

IN COAL MINES 

By Christopher Mark, Ph.D,' Frank E. Chase,' and R. Karl Zipf, Jr., Ph.D.3 

ABSTRACT 

A massive pillar collapse occurs when undersized pillars fail and rapidly shed their load to adjacent pillars, 

which in turn fail. The consequences of these chain-reaction failures can be catastrophic. One effect of a 

massive pillar collapse can be a powerful, destructive, and potentially hazardous airblast. Thirteen recent 

massive pillar collapses have been documented in West Virginia, Ohio, Utah, and Colorado. Data collected 

at the failure sites indicate that all of the massive collapses occurred where the pillar width-to-height (wlh) ratio 

was 3.0 or less and where the Analysis of Retreat Mining Pillar Stability Factor was less than 1.5. The unique 

structural characteristics of these pillar systems apparently result in sudden, massive pillar failures, rather than 

the more common slow "squeezes." The field data, combined with theoretical analysis, provide the basis for 

two partial-extraction design approaches to control massive pillar collapses. These are the containment 

approach and the prevention approach; practical examples are provided of each. 

'Mining engineer, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 

'Geologist, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 

'Lecturer, Department of Mining and Metallurgical Engineering, University of Queensland, Brisbane, Queensland, Australia. 
INTRODUCTION 

Massive pillar collapses in room-and-pillar mines have also 

been labeled "cascading pillar failures," "domino-type 

failures," or "pillar runs." In this type of failure, when one 

pillar collapses, the load that it carried transfers rapidly to its 

neighbors, causing them to fail, and so forth. This failure 

mechanism can lead to the rapid collapse of very large mine 

areas. In mild cases, only a few tens of pillars might fail; 

however, in extreme cases, hundreds, even thousands, of 

pillars can collapse. 

Massive piliar collapses can have catastrophic effects on a 

mine. Sometimes these effects pose a greater safety risk than 

the underlying ground control problem. Usually, the collapse 

induces a devastating airblast due to the displacement of air 

from the collapsed area. An airblast can totally disrupt the 

ventilation system at a mine by destroying ventilation 

stoppings, seals, and fan housings. Flying debris can seriously 

injure or kill mining personnel. The collapse might also 

fracture a large volume of rock in the pillars and immediate 

roof and floor. In coal and other gassy mines, this fragmentation can lead to the sudden release of large quantities of 

methane gas into the mine atmosphere, creating an explosion 

hazard. Finally, a massive pillar collapse can release 

significant seismic energy that may be experienced on the 

surface as a small earthquake. 

Fortunately, not all pillar failures are sudden, massive 

collapses. Most are slow "squeezes" that develop over days to 

weeks, and because of their slow progress, do not pose as 

great a danger to mining personnel. A central goal of the 

research described in this paper was to identify the physical 

characteristics that distinguish sudden collapses from other 

pillar failures. 

CASE HISTORIES 

The most infamous massive pillar collapse in history 

occurred in 1960 at Coalbrook North Colliery in South Africa. 

Thousands of 12- by 12- by 4.2-m (40- by 40- by 14-ft) pillars 

collapsed over a 305-ha (750-acre) area in 5 min, killing 437 

miners [Bryan et al. 19661. Numerous other, smaller collapses 

have been reported in South Africa since then [Madden 199 11. 

In Australia, the New South Wales Joint Coal Board reported 

eight massive pillar collapses between 1990 and 1993 

[University of New South Wales School of Mines 19941. 

Massive collapses have also occurred in metal and 

nonmetal mines. Zipf and Mark [I9961 documented six 

examples from lead-zinc, copper, silica, and salt mines. The 

largest occurred at a Wyoming trona mine in 1995, where 

I60 ha (400 acres) of 4- by 29- by 6-m (13- by 95- by 19-ft) 

fenders collapsed, resulting in a Richter magnitude 5.3 

earthquake and one fatality underground [Ferriter et al. 19961. 

The ventilation system at the mine was heavily damaged, and 

an estimated 1 million m3 (30 million ft3) of methane was 

liberated on the day of the collapse. Methane release levels 

did not return to normal until 3 months later [Ferriter et al. 

19961. 

In 1992, the former U.S. Bureau of Mines (USBM) was 

asked to investigate a massive pillar collapse and resultant 

destructive airblast that had occurred in a coal mine in Mingo 

County, WV. Subsequent investigations found 12 other 

examples, which were documented by field investigations 

[Chase et al. 19941. Geotechnical evaluations examined the 

competency of the immediate roof, as well as that of the main 

roof and its susceptibility to caving. The Analysis of Retreat 

Mining Pillar Stability (ARMPS) program [Mark and Chase 

19971 was used to determine the pillar stability factors (SF). 

Four examples that illustrate different mining methods and 

effects are described in detail below. 

PILLAR SPLllTlNG (MINE A) 

Mine A is located in Mingo County, WV, and is extracting 

the 2.9-m (9.5-ft) thick Coalburg Coalbed. A 28-m (90-ft) 

thick massive sandstone unit with a compressive strength of 

83 MPa (12,000 psi) formed the roof above the collapsed area. 

The Coal Mine Roof Rating (CMRR) of the immediate roof 

was calculated to be 74. Below the noncleated coalbed is 

10.5 m (34 ft) of competent sandy shale and sandstone units. 

.411 roadways were 6 rn (20 ft) wide. 

In 1991, the panel shown in figure 1 was developed. All 

roadways were driven on 18-m (60-ft) centers and were under 

85 m (275 ft) of cover. After the panel was completed, partial 

pillar recovery was begun. A 6-m (20-ft) wide split was 

mined through the middle of each pillar, and two 3- by 12-m 

(10- by 40-ft) fenders with an ARMPS SF of 0.75 remained. 

Because of the competency of the roof and the support 

provided by the regularly spaced uniform fenders, no caving 

occurred while the panel was being retreat mined. Three 

weeks after the panel had been abandoned, an area measuring 

approximately 140 by 155 m (450 by 500 ft) containing 107 

fenders collapsed. Miners on a nearby section were knocked 

to the floor by the resultant airblast. One miner was bounced 

off of a steel rail and required 26 stitches to his head. 

Fortunately, no miners were near the collapse. However, if 

the failure had occurred 15 min later, two miners would have 
Figure 1.-Failed split-pillar workings in Mine A. 

been rock dusting ribs immediately outby the area that 

collapsed. The airblast blew out 26 cinder block stoppings 

and the fan house weak wall, which closed the mine for days. 

As was the case in many of the other collapses that were 

studied, a number of fenders near the edge of the collapse did 

not fail. There are two possible explanations for this: (1) The 

collapse might terminate as soon as the competent roof units 

were able to bridge the span, or (2) the collapse might terminate where the fenders were shielded from the full load by the 

adjacent abutment. In the second case, the 12- by 12-m (40- 

by 40-ft) pillars with an SF of 2.33 may have provided a hinge 

line, which allowed the roof to cantilever over the first several 

rows of fenders. 

An earlier collapse had occurred at Mine A in partially 

pillared workings under very similar conditions. Damage was 

limited to blown out stoppings, and no one was injured. 

Complete documentation of this case was unavailable. 

After the second collapse, the practice of pillar splitting 

was reexamined at the mine. Several sets of mobile roof 

supports were purchased, and retreat mining continued with 

full pillar extraction. Most recently, some pillar splitting has 

been conducted, with rows of unsplit pillars left as barriers to 

isolate retreated areas. 

PILLAR SPLITTINGIABUTMENT LOAD 

OVERRIDE (MINE C) 

Mine C is located in Logan County, WV, and is extracting 

the 3-m (10-ft) thick Dorothy Coalbed. The immediate and 

main roof throughout the mine is composed of a fine-grained, 
semilaminated sandstone with a CMRR of 64; the floor was 

composed of an extremely firm sandstone. Coalbed cleating 

was nonexistent. All roadways in the mine were 6 m (20 ft) 

wide and were driven on 18-m (60-ft) centers in the relevant 

area. 

In 1992, the operator was splitting pillars in the panel 

shown in figure 2. After the 6-m (20-ft) wide split, two 3- by 

12-m (10- by 40-ft) fenders with an SF of 0.94- 1.15 remained. 

When the operator began to mine the pillar row outby the last 

row split (figure 2), a massive collapse of the fenders in the 

gobbed-out area initiated. The roof bolter operator on the 

section indicated that he and his coworkers were knocked to 

the floor by the resulting airblast, and 103 stoppings were 

destroyed. The pillars where the collapse terminated had an 

SF of 1.97. Overburden in the collapsed area ranged from 53 

to 66 m (175 to 215 ft). 

A subsequent pillar collapse occurred at Mine C, apparently triggered by time deterioration and front abutment 

pressures generated by full pillar extracti'on. Roadways in the 

collapsed area were driven on 15-m (50-ft) centers, and 91 

pillars with an SF of I .08 failed. Pillars with an SF of 1.69 

halted the collapse. These roadways were driven on 18-m 

(60-ft) centers. No stoppings were damaged, and the overburden in the area was 99 m (325 ft). 

Mine C was visited in February 1994 to observe diagonal 

pillar splitting, which is not a common practice. Roadways 

were driven on 15-m (50-ft) centers, and the pillar splits were 

5 m (16 ft) wide. The extraction percentage was 86%. The 

triangular remnant stumps were observed to routinely crush 

out after finishing the pillar row, and the roof caved immediately inby the breakers. The breakers and wedges 

showed no weight. Where the first pillar collapse occurred in 

Mine C using the traditional 6-m (20-ft) wide split through a 

12- by 12-m (40- by 40-ft) pillar, 78% of the coal was 

extracted. This 8% increase in resource recovery, coupled 

with a less stable triangular stump with a smaller perimeter, 

probably explains why the roof caves more readily than in 

traditional pillar splitting. 

SMALL-CENTER MINING (MINE D) 

Mine D is located in Mingo County, WV, and is extracting 

the 3.4-m (1 1-ft) thick Dorothy Coalbed. The roof consisted 

of 76 cm (2.5 ft) of laminated fossiliferous shale and 7 cm (3 

in) of rider coal, and 25 m (80 ft) of cross-bedded sandstone 

was observed in the highwall. The roof had a CMRR of 8 1. 

Below the noncleated coalbed was 1.5 m (5 ft) of sandy shale 

and 28 m (91 ft) of sandstone. All roadways in the mine were 

6 m (20 ft) wide. 

In 1992, ninety-four 6- by 6-m (20- by 20-ft) pillars with 

an SF of 1.15 and thirty-two 9- by 9-m (30- by 30-ft) pillars 

with an SF of 1.45 failed. As shown in figure 3, the pillar 

failures occurred in a panel driven off the mains. The 

resultant airblast blew out 37 stoppings. The only other 

stopping in the mine had a hole in it. Some of these stoppings 

were as far away as 244 m (800 ft) from the perimeter of the 

collapse. In one stopping, it was determined that some of its 

14-kg (30-lb) cinder blocks had been hurled 152 m (500 ft). 

Fortunately, the occurrence was on an idle shift, and no one 

was in the mine. The collapse was halted by pillars in the 

main entries, which were 12- by 12-m (40- by 40-ft) and had 

an SF of 3.33. Cover over the collapsed area was 69 m (225 

ft). 

Scale 

w 

Figure 2.-Location of split-pillar collapse at Mine C. 
LEGEND /I/ 

Figure 3.-Failed small-center development workings at Mine D. 

FLOOR RECOVERY (MINE G) 

Mine G is located in Utah and was extracting the 8-m (25- 

ft) thick Lower O'Connor Seam [Ropchan 19911. There were 

previous workings in the Upper O'Connor above Mine G, 

separated by 18-23 m (60-80 ft) of overburden. The total 

overburden above the collapsed area was about 170 m (550 

ft). 

Room-and-pillar workings were advanced 2.4 m (8 ft) high 

on 18-m (60-ft) centers. The panel was developed nine entries 

wide and 535 m (1,740 ft) long. The pillars were not extracted on retreat, but an additional 3 m (I0 ft) was removed 

from the floor, leaving 5.4-m (1 8-ft) high remnants. Mining 

the floor coal decreased the w/h ratio of the pillars from 5 to 

2.2 and reduced their strength by about 45%. 

The collapse occurred when the section was within two 

crosscuts of being completely retreated. The forcc of the 

airblast hurled three miners for distances of 12-30 m (40- 100 

ft), causing one severe head laceration. A 2-ton shop car was 

blown through a stopping. There was extensive damage to 

ventilation structures; concrete blocks from stoppings were 

scattered up to 30 m (100 ft). The main mine fan was stalled, 

and airflow in the mine was temporarily reversed. There was 

some speculation that a north-south trending fault that 

bordered the panel may have contributed to the collapse. 

SUMMARY OF CASE HISTORIES 

Table 1 summarizes the mining dimensions of 13 examples 

of massive pillar collapses in U.S. coal mines. All occurred 

during the 1980's and 19901s, and all happened suddenly or 

without significant warning. Most resulted in airblasts and 

damage to the ventilation system. 

Analysis of the data reveals some important similarities. 

First. the ARMPS SF was less than 1.5 in every case and less 

than 1.2 in 8 1 % of the cases. This implies that the pillars were 

not sized to carry the full overburden load. Pillar failures are 

not unusual; however, most are slow and nonviolent. What 

apparently distinguishes the sudden collapses from the slow 

squeezes is the pillar's w/h ratio. Every massive pillar collapse 

involved slender pillars with a w/h ratio of less than 3. 

Another common characteristic of the collapses is that the 

overburden was judged to be relatively strong in every case. 

Finally, the collapsed areas wcre all at least 1.6 ha (4 acres), 

and the minimum dimension of a collapsed panel suffering 

major damage was 110 rn (350 ft). 
Table 1 .-Massive pillar collapses in coal mines 

Dept Pillar size, m (tt) ARMPS SF w/h ratio Collapsed area, Case history State m (ft ha (acres) Collapse size, m (tt) 

A . . . . . . . . . . W 84 (275) 3 by 12 (10 by 40) 0.86 1.05 2.3 (5.7) 150 by 150 (500 by 500) 2

01 ......... W 73 (240) 3 by 12 (10 by 40) 0.96 1 .OO - - 3

3 by 18 (10 by 60) 1.10 1.00 

02 . . . . . . . . . W 75 (245) 3 by 12 (10 by 40) 0.94 1.00 1.7 (4.1) 100 by 150 (350 by 500) 4

03 ......... W 85 (280) 9 by 9 (30 by 30) 1.46 3.00 2.8 (6.8) 180 by 180 (600 by 600) 7

6 by 12 (20by 40) 1.47 2.00 

C1 . . . . . . . . . W 60(195) 3by12(10by40) 1.19 1.00 2.1 (5.2) 140 by 150 (450 by 500) 1

C2 . . . . . . . . . W 99 (325) 9 by 9 (30 by 30) 1.15 3.00 1.9 (4.8) 100 by 180 (350 by 600) M

D.......... W 69 (225) 6 by 6 (20 by 20) 1.15 1.82 1.7 (4.3) 100 by 160 (350 by.540) 3

9 by 9 (30 by 30) 1.42 2.73 

E 1 . . . . . . . . . W 91(300) 3by12(10by40) 0.79 1.42 7.4 (18.2) 240 by 290 (800 by 950) M

E2 . . . . . . . . . W gl(300) 3by12(10by40) 0.71 1.11 6.7 (16.6) 220 by 275 (720 by 900) M

F . . . . . . . . . . OH 76(250) 2by12(7by39) 0.66 2.12 2.0 (4.9) 90 by 21 5 (300 by 700) M

G.......... UT 168 (550) 12 by 12 (40 by 40) 0.95 2.29 7.9 (19.4) 150 by 490 (480 by 1,620) M

O.......... W - - 1.03 2.50 1.8 (4.5) 120by150(400by500) - 

R . . . . . . . . . . CO 120 (400) 4 by 24 (12 by 80) 0.57 1.71 2.8 (6.8) 180 by 150 (600 by 500) M

NOTE.- Dash indicates no data available. 
MECHANICS OF MASSIVE PILLAR COLLAPSES 

A conceptual model of a massive pillar collapse can be 

described as follows. Undersized, regularly spaced remnant 

pillars help the stiff and competent roof to bridge a relatively 

wide span. A pressure arch is created, with much of the 

overburden load being transferred by the stiff roof to the 

barrier pillars surrounding the extraction area. Within the 

pressure arch, the pillars are shielded from the full weight of 

the overburden. Eventually, any one of a number of mechanisms may cause the pressure arch to break down: 

- The extraction area becomes so large that it exceeds the 

bridging capacity of the roof. 

Mining approaches a fault or other discontinuity. 

The roof weakens over time. 

The remnant pillars weaken over time. 

Once the pressure arch breaks down and additional 

overburden load is shifted to the pillars, their structural 

characteristics are such that a sudden, massive collapse can 

occur. Slender pillars have little residual strength and shed 

load rapidly as they fail. When one fails, the weight it 

transfers can overload adjacent pillars, and a rapid "domino" 

failure of adjacent pillars can ensue. Pillars that are more 

squat retain most of their load even after failure. Such pillars 

will squeeze slowly, rather than collapse. 

Laboratory tests have shown that the residual strength of 

coal specimens depends on their wlh ratio [Das 19861. 

Specimens with a wh ratio of less than 3 typically have little 

residual strength, which means that they shed almost their 

entire load when they fail (figure 4). As the specimens 

STRAIN, lo-' 

become more squat, their residual strength increases. Once 

the w/h ratio reaches 8-10, the specimens become "strainhardening." which means that they never shed load, and 

sudden collapse is impossible. 

Figure 5 summarizes available postfailure modulus data for 

large in situ coal specimens and full-scale coal pillars. The 

dashed line indicates a conservative envelope for these limited 

in situ data. In general, the laboratory postfailure moduli 

exceed the large-scale test values. 

The importance of the pnstfailure stiffness is further 

explained by the theory of local mine stifiess, first proposed 

by Salamon [I9701 and discussed by Zipf [1992, 19961. The 

theory states that if the pillar's postfailure modulus (K,) is less 

than the stiffness of the mine roof (the local mine stiffness, or 

K,), the failure is stable and gradual (figure 6B). If K, exceeds K,, on the other hand, the failure is sudden and violent 

(figure 6A). The local mine stiffness depends on the modulus 

of the immediate roof; floor and pillar materials; and the 

layout of pillars, mine openings, and barrier pillars. The postfailure stiffness, K,, depends on the w/h ratio of the coal pillar, 

as shown in figure 5. Using a boundary-element method program similar to the USBM's MULSIMNL program, it is 

possible to simulate both massive pillar collapses and stable, 

progressive pillar failures [Zipf 19961. The behavior of computer simulations changes depending on whether the model 

satisfies or violates the local mine stiffness stability criterion. 

WIDTHIHEIGHT RATIO 

Figure 5.-Postfailure modulus of coal pillars, in situ coal 

Figure 4.-Complete stress-strain curves for Indian coal specimens, and laboratory samples. Darkened circles represent 

specimens, showing increasing residual strength with increasing laboratory tests, remaining symbols represent in situ tests [Chase 

wlh ratio (after Das [1986]). et al. 19941. 
A Unstable failure 

"soft" load system 

B Stable fail~~re 

"stiff" load system 

CONVERGENCE 

Figure 6.-Illustration of the local mine stiffness concept. A, local mine stiffness (K,) is less than postfailure stiffness of the pillar (K,), 

resulting in unstable failure. B, local mine stiffness (KJ exceeds the pillar's postfailure stiffness (K,), resulting in slow and stable failure. 

DESIGN APPROACHES TO CONTROL MASSIVE PILLAR COLLAPSE 

In coal mining, small-center mining and partial pillaring 

are methods to achieve high extraction without full pillar 

recovery. Both leave significant remnant pillars in the minedout areas. For example, mining on 15-m (50-ft) centers using 

6-m (20-ft) entries leaves about 35% of the coal in 9- by 9-m 

(30- by 30-ft) pillars. Splitting pillars developed on 18- by 

18-m (60- by 60-ft) centers leaves about 22% of the coal. 

Both techniques can be adapted to avoid massive pillar collapses following the strategies of prevention or containment. 

In the prevention approach, the panel pillars are designed 

so that collapse is highly unlikely. This can be accomplished 

by increasing either the SF of the pillars or their w/h ratie. In 

the containment approach, high extraction is practiced within 

individual compartments that are separated by barriers. The 

small pillars may collapse within a compartment; however, 

because the compartment size is limited, the consequences are 

not significant. The barriers may be true barrier pillars, or 

they may be rows of development pillars that are not split on 

retreat. The containment approach has been likened to the use 

of compartments on a submarine. 

Full extraction can be another strategy to avoid massive 

pillar collapses. Mining all of the coal removes the support to 

the main roof, thereby limiting the potential width of the 

pressure arch. Although some "first falls" behind longwalls 

and other full-extraction systems have been destructive, they 

generally involve areas smaller than massive pillar collapses. 

SMALL-CENTER MINING: A PREVENTION 

APPROACH 

Square pillars are generally used in small-center mining. 

Table 1 indicates that three collapses involved 9-m (30-ft) 

square pillars, and one involved 12-m (40-ft) square pillars. 

Square pillars may be designed to be collapse-resistant in two 

ways. The first is to increase their w/h ratio. Because no 

collapses have been documented in which the w/h ratio was 

greater than 3.0, a design w/h ratio of 4.0 is suggested to 

provide an adequate margin of safety. 

Pillar collapses may also be avoided by maintaining a 

sufficiently high SF. The ARMPS case history data base 

[Mark and Chase 19971 suggests that normally an ARMPS SF 

of 1.5 is sufficient to limit the probability of pillar failure. 

Where slender pillars are being employed and their failure 

may result in a massive collapse rather than a slow squeeze, it 
18 ft Wide Entries 

ARMPS SF = 2.0 ...-.-*-.-I. 

-. - 

KEY 

-Seam Ht. = 6 ft 

-Seam Ht. = 8 ft 

- 0- Seam Ht. = loft 

20 ft Wide Entries 

ARMPS SF = 2.0 

0 100 200 300 400 500 

DEPTH OF COVER, ft 

Figure 7.-Suggested minimum square pillar size to avoid massive pillar collapse. A, 5.5-m (18-ft) entries; 8, 6-m (20-ft) 

entries. 
might be prudent to increase the SF to 2.0. The SF can be 

increased by increasing the pillar width, decreasing the 

extraction ratio, or both. These two design criteria have been 

combined to develop guidelines for small-center mining. 

Figure 7 was developed assuming square pillars with an SF of 

2.0 or a w/h ratio of 4.0. 

When using 6-m (20-ft) wide entries, the minimum 

suggested pillar sizes are increased by about 6%. Also note 

that these design criteria are only for controlling massive pillar 

collapses. At greater depths, pillar sizes may need to be 

increased beyond a w/h ratio of 4 to maintain an adequate SF. 

The faiiure of piliars with a wfn ratio greater than 4 should be 

a slow squeeze rather than a sudden collapse. 

PILLAR SPI,ITTING: A CONTAINMENT 

APPROACH 

Fenders left from pillar-splitting operations have failed at 

even shallow depths. For example, 3- by 12-m (10- by 40-ft) 

fenders in a 3-m (10-ft) seam have an SF of 1.5 at only 55 m 

(180 ft) of cover. The potential for a destructive massive 

collapse can be reduced by limiting the size of the gob area. 

To separate the gob areas, rows of unsplit development pillars 

can be left as barriers. This strategy is based on two 

assumptions: 

By limiting the span above the mined-out area, a 

bridging failure of the strong overburden is less likely. 

By minimizing the size of the potential collapsed area, 

any airblast resulting from a collapse would be less powerful. 

Table 1 shows that no major collapses have been 

documented in which the gob area was less than 1.5 ha (4 

acres). In the five cases where the gob area was between 1.5 

and 1.9 ha (4 and 5 acres), about 60% of the incidents resulted 

in major damage. Additionally, no damaging incidents 

occurred when the minimum dimension of the mined-out area 

was less than 100 m (350 ft). Using these data, acceptable 

dimensions of a pillar-splitting operation might be a maximum 

area of 1.2 ha (3.2 acres), with a minimum dimension of less 

than 90 m (300 ft). For example: 

Assuming 18- by 18-m (60- by 60-ft) centers in a nineentry system with four rows split, the mined-out area would 

have a minimum dimension of 72 m (240 ft) and an area of 

about 1.1 ha (3 acres), as shown in figures 8A and 8B. 

Assuming the same pillar size in a six-entry system with 

five rows split, the minimum dimension would be 90 m (300 

ft) and the area would be about 1 ha (2.5 acres), as shown in 

figures 8C and 80. 

The next question is: how many unsplit rows should be left 

between these mined-out areas? The goal is to leave enough 

of a "barrier" so that the failure of one gob area does not 

initiate failure in adjacent areas. ARMPS was used to evaluate 

the loading on unsplit pillars between two mined-out areas. 

The program was modified so that two "front" gobs could be 

applied to the unsplit pillars. The analyses were run with 

abutment angles of 90°, which assumes that none of the load 

is carried by the gob, but instead is transferred to the barriers. 

In the first set of analyses, two rows of full-sized pillars 

were used as the barrier. An ARMPS SF of 1.5 was deemed 

necessary to prevent the collapse of one gob area triggering 

the collapse of an adjacent area. Three rows of pillars were 

used in the second set of analyses; the SF was reduced to 1.0 

because of the greater stiffness of the barrier. Pillars on 18- 

by 18-m (60- by 60-ft) centers were used in all cases. 

Other parameters that were varied included the number of 

rows that were split (three, four, and five), the entry width (5.5 

and 6 m (1 8 and 20 ft)), the seam height (2, 2.5,and 3 m (6, 

8, and 10 ft)), and the number of entries in the section (five, 

seven, and nine). The results are presented in figure 9, which 

shows the suggested maximum depth of cover for each 

combination of parameters. In general, considering 5.5-m 

(18-ft) entries in a 2.5-m (8-ft) seam, it appears that two rows 

of unsplit pillars are an adequate barrier at depths less than 

about 300 ft and that three rows are acceptable to about 170 m 

(550 ft) of cover. 

Barriers must also be left between extracted panels. These 

can be unsplit development pillars or solid coal. If unsplit 

development pillars are used, the analysis in figure 9 should 

apply. For solid coal barriers, figure 10 shows the suggested 

widths, using the same loading assumptions. For a 2.5-m (8- 

ft) seam, a 17-m (55-ft) solid barrier appears to be appropriate 

at 75 m (250 ft) of cover, and 23 m (75 ft) might be needed at 

120 m (400 ft). 
Figure 8.-Possible pillar-splitting plan for airblast control. A, nine-entry system, two rows of unsplit pillars for 

barrier. 13, nine-entry system, three rows of unsplit pillars for barrier. C, six-entry system, two rows of unsplit pillars 

for barrier. D, six-entry system, three rows of unsplit pillars for barrier. 
A Entry Width = 18; 2 unsplit rows as Barrier 

. 1 I- - 1 

400 450 500 550 600 650 700 750 800 

ALLOWABLE DEPTH OF COVER, ft 

KEY 

-+ h = 6'; 3 Rows Split + h = 6'; 4 Rows Split + h = 6'; 5 Rows Split 

++ h = 8'; 3 Rows Split -rc h = 8'; 4 Rows Split - h = 8'; 5 Rows Split 

-h = 10'; 3 Rows Split -h = 10'; 4 Rows Split -h = 10'; 5 Rows Split 

Figure 9.-A, suggested maximum depth for two rows of unsplit pillars as barrier between gob areas, 5.5-171 (18-ft) entry, 18- 

by 18-m (60- by 60-ft) pillars. B, suggested maximum depth for three rows of unsplit pillars as barrier between gob areas, same 

entry and pillar sizes. 
9 Required Barrier Pillar Width vs Depth of Cover 

9 

KEY . -6' Seam 

. -8' Seam 

-c 10' Seam 

. 

DEPTH OF COVER, fi 

Figure 10.-Suggested solid coal barrier width between two areas where pillars have been split. 

CONCLUSIONS 

The potential for massive pillar collapses should always be 

considered when designing room-and-.pillar mining operations. A collapse can occur when one pillar fails suddenly, 

overstresses its neighbors, causing them to fail, and so forth, 

in very rapid succession. Very large mining areas can collapse 

via this mechanisrri within seconds with little or no warning. 

The collapse itself can pose serious dar:ger to nearby mining 

personnel. Additionally, the collapse can induce a violent 

airblast that disrupts or destroys the ventilation system. 

Further critical danger to miners exists if the mine atmosphere 

becomes explosive or contaminated as a result of the pillar 

collapse. 

Research has found that massive collapses in coal mines 

have the following common characteristics: 

Slender pillars (w/h ratio less than 3.0). 

Low SF (less than 1.5). 

Competent roof strata. 

Collapsed area greater than 1.6 ha (4 acres). 

Minimum dimension of the collapsed areas greater than 

110 m (350 ft). 

Two alternative strategies may be successful in preventing 

massive pillar collapses. For small-center mining, prevention 

may be applied by increasing either the w/h ratio or the SF. 

Containment is appropriate for pillar splitting and requires 

leaving barriers or rows of unsplit pillars to limit the area of 

potenria! collapses. A final strategy is to go to full pillar 

extraction. By removing the support provided by the remnant 

fenders left during traditional pillar splitting, the bridging 

capacity of the roof should be substantially reduced. 

Finally, it is important to note that the massive pillar collapses discussed in this paper are not to be confused with coal 

bumps or rock bursts. Although the outcomes may appear 

similar, the underlying mechanics are entirely different. 

Bumps are sudden, violent failures that occur near coal mine 

entries and expel large amounts of coal and rock into the 

excavation [Maleki 199.51. They occur at great depth, affect 

pillars (and longwall panels) with large w/h ratios, and are 

often associated with mining-induced seismicity. The design 

recommendations discussed here for massive pillar collapses 

do not apply to coal bump control. 
REFERENCES 

Bryan A, Bryan JG, Fouche J [1966]. Some problems of strata control and 

support in pillar workings. Min Eng (London) 123:238-254. 

Chase FE, Zipf RK, Mark C [1994]. The massive collapse of coal pillars - 

case histories from the United States. In: Proceedings of the 13th Conference on 

Ground Control in Mining. Morgantown, WV: West Virginia University, 

pp. 69-80. 

Das MN [1986]. Influence of widthheight ratio on postfailure behavior of 

coal. Intl J Min Geol Eng 4:79-87. 

Ferriter RL, Zipf RK, Ropchan DM, Davidson J [1996]. Report of technical 

investigation underground nonmetal mine collapse accident, Solvay Mine. U.S. 

Department of Labor, Mine Safety and Health Administration. 

Madden BJ [1991]. A re-assessment of coal pillar design. J South Afr lnst 

Min Metall 91(1):27-37. 

Maleki H [1995]. An analysis of violent failure in U.S. coal mines--case 

studies. In: Maleki H, Wopat PF, Repsher RC, Tuchman RJ, eds. Proceedings: 

Mechanics and Mitigation of Violent Failure in Coal and Hard-Rock Mines. 

Pittsburgh, PA: U.S. Department of the Interior, Bureau of Mines, SP 01-95, 

pp. 5-25. 

Mark C, Chase FE [1997]. Analysis of retreat mining pillar stability 

(ARMPS). In: Mark C, Tuchman RJ, comp. Proceedings: New Technology for 

Ground Control in Retreat Mining. Pittsburgh, PA: U.S. Department of Health 

and Human Services, Public Health Service, Centers for Disease Control. 

National Institute for Occupational Safety and Health, IC 9446. 

Ropchan D [1991]. Ground support evaluation-Belina No. 2 Mine ID No. 

42-01280. Denver, CO: U.S. Department of Labor, Mine Safety and Health 

Administration. 

Salamon MDG [1970]. Stability, instability, and design of pillar workings. 

Intl J Rock Mech Min Sci 7:613-631. 

University of New South Wales School of Mines [1994]. Pillar 

performance. Strata Control for Coal Mine Design Apr:5. 

Zipf RK Jr. [1992]. Analysis of stable and unstable pillar failure using a 

local mine stiffness method. In: lannacchione AT, Mark C, Repsher RC, 

Tuchman RJ, Jones CC, comp. Proceedings of the Workshop on Coal Pillar 

Mechanics and Design. Pittsburgh, PA: U.S. Department of the Interior, Bureau 

of Mines, 1C 9315, pp. 128-143. 

Zipf RK [1996]. Analysis and design methods to control cascading pillar 

failure in room-and-pillar mines. In: Bieniawski ZT, ed. Milestones in Rock 

Engineering, Balkema, pp. 225-264. 

Zipf RK, Mark C [1996]. Design methods to control violent pillar failures 

in room-and-pillar mines. In: Proceedings of the 15th Conference on Ground 

Control in Mining. Golden, CO: Colorado School of Mines, pp. 571-585. 
PILLAR DESIGN AND COAL STRENGTH 

By Christopher Mark, Ph.D.,' and Timothy M. Barton' 

ABSTRACT 

A comprehensive data base was created that includes more than 4,000 individual uniaxial compressive 

strength test results from more than 60 coal seams. These data were compared with 100 case studies of in-mine 

pillar performance from the Analysis of Retreat Mining Pillar Stability (ARMPS) data base. 

Statistical analysis found no correlation between the ARMPS stability factor of failed pillars and coal 

specimen strength. Pillar design was much more reliable when a uniform coal strength of 6.2 MPa (900 psi) 

was used in all case histories. The conclusion is that laboratory testing should not be used to determine coal 

strength for ARMPS. 

Other analyses provided evidence of why laboratory strength does not correlate with pillar strength. The 

data showed clearly that the "size effect" observed in laboratory testing is related to coal structure. The widely 

used Gaddy formula, which predicts a significant strength reduction as the specimen size is increased, was 

found to apply only to "blocky" coals. For friable coals, the size effect was much less pronounced, or even 

nonexistent. Laboratory tests do not account for large-scale discontinuities, such as roof and floor interfaces, 

which apparently have more effect on pillar strength than small-scale structure. 

'~ining engineer, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA 
BACKGROUND 

The uniaxial compressive strength of coal was one of the 

first issues addressed by early rock mechanics researchers. 

Bunting [ 19 1 11 observed that "to mine without adequate pillar 

support will result, sooner or later, in a squeeze; the inherent 

effects of which are crushing of the pillars, the caving of the 

roof, and the heaving of the bottom." By testing anthracite 

specimens of various sizes and shapes in the laboratory, Bunting and his collaborators hoped to aid mine operators in 

"establishing the width of chambers and pillars." They soon 

found that "the crushing strength of small cubes is greater than 

that for large cubes; and, with a constant base area, the crushing strength becomes less as the height increases" [Daniels and 

Moore 19071. Bunting apparently concluded that these two issues, the "size effect" and the "shape effect," prevented the direct use of laboratory strength results in design. His design 

equation was the first U.S. empirical coal pillar strength 

formula: 

where S, = pillar strength, 

S, = coal strength parameter, 

w = pillar width, 

and h = pillar height. 

Bunting used the laboratory results to determine the shape of 

the formula (figure 1). The coal strength parameter was determined from analysis of in situ pillar failure ("actual 

squeezes" in figure 1). For anthracite pillars, it was set at 7 

MPa (1,000 psi). 

The basic approach employed by Bunting and his colleagues remained the state of the art for much of the 20th 

century. For example, Zern presented the following equation 

in the 1928 edition of the "Coal Miner's Pocketbook": 

Zern's suggested value of the coal strength parameter is 4.8-7 

MPa (700- 1,000 psi). 

More than 20 years later, Gaddy [I9561 attempted to provide the link between laboratory specimens and field strength. 

He attacked the size effect by testing coal cubes of various 

sizes from five seams. Gaddy concluded that the strength 

decrease with increasing specimen size could be expressed as 

where k = Gaddy constant 

= estimated strength of a 2.5-cm (I-in) cube, 

S = coal specimen strength, 

and d = specimen dimension (in). 

His work led to the widely used Holland-Gaddy pillar strength 

formuia [Hoilanci and Gaciciy 14561: 

The Holland-Gaddy formula appears to have been the first in 

the United States to employ a seam-specific strength parameter determined from laboratory testing. 

In situ testing of full-scale pillars in South Africa during 

the 1960's resulted in the concept of a "critical" specimen size 

beyond which the strength is constant [Bieniawski 19681. The 

Bieniawski pillar strength formula below employed this 

concept: 

where S, = in situ coal strength. 

Following Hustrulid 119761, Bieniawski recommended that 

the in situ strength be determined from laboratory tests and 

that the Gaddy formula be used to reduce the strength to that 

of a 1-m (36-in) critical-sized specimen [Bieniawski 19841. 

Others proposed versions of the Holland-Gaddy and ObertDuvall (Bauschinger) formulas that employed the in situ 

strength parameter [Bieniawski 19841. It may be noted that 

the in situ coal strength in equation 5 is functionally equivalent to the "coal strength parameter" in equations 1 and 2. 

Despite the fact that textbooks have considered laboratory 

testing an integral part of pillar design for nearly 30 years, it 

has remained controversial. One reason is that coal remains 

notoriously difficult to test. Coal contains many types of discontinuities, including microfractures, cleats, bedding planes, 

partings, shears, and small faults. Three sources of unreliability have been identified: 

1. Material variability within a particular seam: Unrug 

et al. [I9851 tested multiple layers of the Warfield and the 

Coalburg Seams and found that the strongest layers were six 

times stronger than the weakest in each seam. Newman and 

Hoelle [I9931 reported similar results from the Harlan Seam. 
3 0 Mine "A" 

\ Mine "B" 

\ 

\ A Mine "c" 

\ A Mine I'D" 

\ x Actual squeezes 1 

HElG HT - h 

RAT'o LEAST LATERAL DIMENSION 

- - 

W 

Figure 1.-Data used to develop the first U.S. pillar strength formula (after Bunting, [1911]). 

2. Variation in sampling, specimen preparation, and testing techniques: Townsend et al. [I9771 found that small cylindrical specimens were typically 30% weaker than cubical 

specimens of the same cross-sectional area. Khair 119681 documented large effects due to platen friction. 

3. Variation in size and shape effects between seams: 

Panek [I9941 and Mrugala and Belesky [1989], among others, 

have speculated that Gaddy's size effect exponent of -0.5 may 

be the mmimum and not universally applicable. The shape effect has been the subject of numerous studies. 

Some have held that these difficulties and the resulting high 

variability in results are enough to largely invalidate laboratory testing. Another school of researchers in the Republic 

of South Africa, Australia, and the United States have argued 

that, although the strength of laboratory-sized specimens varies widely, the in situ coal strength may fall within a narrow 

range [Salamon 1991; Galvin 1995; Mark 19901. In each 

case, their conclusions were based on analysis of in-mine 

pillar failures. Salamon and MUNO [I9671 originally analyzed 

27 pillar collapses and 92 intact cases. Their formula, perhaps 

the most widely used in the world, explained the data very 

well without reference to individual seam strengths. In 1991, 

Madden reanalyzed an updated version of their data set. 

Although he found some differences in strength between 

seams, he concluded again that the average strength could 

represent all seams. Galvin [I 9951 conducted a probabilistic 

analysis of 30 collapsed and stable bord-and-pillar workings 

from Queensland and New South Wales, Australia. He concluded that "pillar strength in the field is only marginally 

dependent on the seam strength once the wlh exceeds 2." In 

the United States, Mark and Chase [I9971 presented data from 

140 case histories, which were analyzed using the Analysis of 

Retreat Mining Pillar Stability (ARMPS). ARMPS estimates 

pillar strength using a slightly modified version of the 

Bieniawski formula; the analyses assumed a uniform in situ 

coal strength. Mark and Chase [I9971 found that pillar failures occurred in 83% of cases when the ARMPS stability factor (SF) was less than 0.75, but only 8% of cases when it 

greater than 1.5 (figure 2). 

These researchers have all determined that the value of the 

in situ coal strength falls between 5.4-7.4 MPa (780-1,070 

psi). The range is remarkably small, considering that it was 

determined from three data sets that span the globe. On the 

other hand, at least one South African seam has been shown 

by back-calculation to be significantly weaker than the average [Van der Merwe 19931. In India, researchers concluded 

from back-analysis of 43 pillar case histories that coal strength 

should be considered in design [Sheorey et al. 19871. 
0 0.25 0.5 0.75 1.0 1.25 1.5 1.75 2.0 2.25 2.5 2.75 

ARMPS STABILITY FACTOR 

Figure 2.-ARMPS data base. 

Interest in the uniaxial compressive strength of coal has also waned over the past 15 years because researchers have devoted their energy to analytic pillar strength formulas and numerical models. These theories are developed from the principles of mechanics rather than curve-fitting to test data. The 

shift in emphasis has been related to the recent focus on pillar 

design for longwall mining. Longwalls employ pillars that are 

much more "squat" than those traditionally used in room-andpillar operations. Few compressive strength tests have ever 

been conducted where the specimen width-to-height (wlh) ratio exceeded 4; however, longwall pillars often employ w/h 

ratios of 10, 20, or greater. 

Obviously, the very concept of pillar failure takes on a different meaning for squat pillars. The wide range of conflicting theories about the mechanics of squat pillars and the substantial difficulties with obtaining field data to confirm or 

disprove any of them have been described elsewhere [Mark 

and Iannacchione 19921. On the other hand, Mark et al. 

[I9941 have shown that longwall tailgate performance can be 

accurately predicted without reference to seam-specific coal 

strength. There is clearly overwhelming,evidence, theoretical 

and empirical, that the uniaxial compressive strength is irrelevant to the strength of a squat pillar. 

Longwall mines, however, account for only 45% of the 

coal mined underground in the United States. Much of the remainder comes from small room-and-pillar mines usually 

operating at relatively shallow cover. These mines use many 

"slender" pillars, and traditional pillar failures still occur. The 

ARMPS data base contains 60 instances of pillar squeezes, 

bumps, or collapses that have taken place in recent years. 

About one-half of these occurred at depths of less than 150 m 

(500 ft) and involved pillars whose w/h ratio was less than 5. 

The failures occurred in a variety of seams. Because some 

seams appear blocky and strong, and others seem weak and 

extremely friable, it is reasonable to expect that these obvious 

structural differences might affect pillar strength. As figure 2 

shows, successful and unsuccessful designs occur in approximately equal proportions in the ARMPS SF range of 0.75 

to 1.5. Might seam-specific laboratory coal strength data explain some of this variability? That was the question this 

research was initiated to answer. 

DISCUSSION OF RESEARCH 

Despite the large volume of coal strength testing reported more than 30 references. All of the data have been entered inin the literature, it had never been compiled into a single data to a spreadsheet and are readily accessible for a wide variety 

base. The Pittsburgh Research Center, therefore, undertook of statistical studies. 

the task. The Coal Strength Data Base now contains the re- Two types of data are included. For about 2,300 tests, insults from more than 4,000 individual uniaxial compressive formation was provided on single specimens. These data were 

strength tests covering more than 60 seams and obtained from entered individually, then grouped by reference, seam, 
specimen geometry, and specimen size. Each group, or suite, 

of tests was placed on a separate page within the data base. A 

"summary line" containing the mean compressive strength and 

standard deviation for the suite was also generated. The summary lines were collected and placed in the summary table. 

The summary table also includes lines representing about 

1,700 tests that were reported in summary form in the original 

reference. The summary table contains information on about 

380 suites of tests. The structure of the Data Base of Uniaxial 

Coal Strength (DUCS) is illustrated in figure 3. 

A single copy of the DUCS may be obtained by sending 

three formatted, double-sided, high-density diskettes to: 

Timothy M. Barton, NIOSH, Pittsburgh Research Center, Cochrans Mill Rd., P.O. Box 18070, Pittsburgh, PA 15236-0070. 

Please specify whether you prefer .XIS, .wk3, or commaseparated values format. 

A table of average U.S. coalbed strengths was derived from 

the summary statistics (table 1). To minimize size and shape 

effects, this table uses only specimens whose w/h ratio is approximately 1.0 and whose smallest dimension is approximately 5-8 cm (2-3 in). The average coalbed strength is calculated as the weighted mean of all of the summary lines for 

a particular seam that meet these geometric criteria. In addition to strength data, the Coal Strength Data Base also includes a variety of coal quality information for each seam 

tested. The most relevant is perhaps the Hardgrove Grindability Index (HGI), which is a measure of the relative 

grindability of coal. Larger HGI values imply easier grindability and greater friability. The HGI is almost universally required by utilities that purchase coal, so the information is 

readily accessible. Representative values of the rank, carbon 

content, volatile content, ash content, and heating value are 

also included. Because the coal quality data were collected independently of the coal strength data and 'rom different 

sources, they are approximations for comparative purposes only. 

During the past 6 years, coal samples measuring about 0.003 

rn2 (0.1 f?) have also been collected from 45 seams. These were 

classified using the following simple system: 

Composition: 

Bright (>90% bright coal) 

Semibright (60%-90% bright coal) 

Intermediate (40%-60% bright coal) 

Semidull (60%-90% dull coal) 

Dull (>90% dull coal) 

structure: 

Blocky (major cleat spacing > 8 cm (3 in)) 

Serniblocky (major cleat spacing 3-8 cm (1 -3 in)) 

Friable (cleat spacing < 3 crn (1 in)) 

Shearing: Yes or no. 

Figure 3.-Structure of the Coal Strength Data Base. 

UNGROUPED-SINGLE SPECIMEN 

TEST DATA 

(SPECRAT 1 .XLS) 

GROUPED-SINGLE SPECIMEN 

TEST DATA 

(BY SUITE) 

(SPECDAT2.XLSl 

GROUPED- SUMMARY 

TEST DATA 

(GRPDATA. XLS) 

4? Suite 

SUMMARY STATISTICS 

FOR ALL DATA 

(SUMSTAT. XLS) 

- 

I 

SEAM- BY- SEAM 

AVERAGE STRENGTH 

(SEAMAVG.XLS) 
Table 1 .--Unconfined compressive strength of U.S. coal seams (5- to Scm (2- to 3-in) specimens) 

... 

Alma .......... 27.7 (4,024 ) 55 30 ((~ower Kkanning ... 14.6 (2[1 17j 90 96 

Seam average Typical No. of Coalbed strength, 

MPa (psi) HGI tests 

Allen .......... 10.8 (1.570) 100 11 

B ............. 

Bakerstown ..... Beckley ........ Blind Canyon .... 

Blue Creek ...... ChiRon ......... 

Clintwwd ....... Coalburg ....... D ............. 

Darby ......... Douglas ........ 

E ............. 

Eagle .......... 

Elkhorn No. 4 .... 

Geneva ........ 

Harlan ......... 

Hazard No. 4 .... 

Hemshaw ...... 

Henin No. 6 ..... 

Island Creek .... 

Jawbone ....... Kellioka ........ 

Kentucky No. 9 ... 

Kentucky No. 11 . . Kentuckv No. 12 . . 

- 

Seam average 

Coalbed strength, Typical No. of HGI tests MPa (psi) 

KentuckvNo.13 ... 26.8 (3.890) 60 37 

The ARMPS data base contains the best available information Pillar squeezes, with significant entry closure and loss of 

on the in situ strength of U.S. coal pillars. ARMPS SF have been reserves. 

back-calculated for 140 case histories (figure 21, covering an ex- Sudden collapses of groups of pillars, usually accompanied 

tensive range of geologic conditions, extraction methods, depths by airblasts. 

of cover, and pillar geometries [Mark and Chase 19971. Ch-ound Coal pillar bumps (violent failures of one or more pillars). 

conditions in each case history have been categorized as either 

satisfactory or unsatisfactory. Unsatisfactory conditions included61 

12 

30 

54 

10 

6 

40 

124 

10 

22 

7 

4 

10 

24 

3 

88 

67 

10 

102 

8 

3 

49 

46 

52 

5 

RESULTS 

Marker .......... 44.9 (6,509) 

Mary Lee ........ 7.8 (1,135) 

No. 2 Gas ....... 12.4 (1,801) 

Pittsburgh ........ 160 (4,330) 

Pocahontas No. 3 . . 10.5 (1,528) 

Pocahontas No. 4 . . 19.9 (2,892) 

Pocahontas No. 5 . . 14.7 (2,127) 

Pond Creek ...... 32.0 (4,635) 

Powellton ........ 13.8 (2,008) 

Redstone ........ 20.2 (2,932) 

Sewell .......... 1 6.5 (2,386) 

Sewickley ........ 27.6 (4,000) 

Stockton ......... 47.2 (6,844) 

Sunnyside ....... 26.6 (3,856) 

Tiller ........... 15.3 (2,215) 

Upper Banner. .... 9.6 (1,391) 

Upper D ......... 46.5 (6,746) 

Upper Freeport .... 10.3 (1.493) 

Upper Hiawatha ... 37.6 (5,446) 

Upper Kittanning ... 10.5 (1,519) 

Warlield ......... 22.7 (3,295) 

Waynesburg ...... 30.9 (4,474) 

Welch .......... 13.1 (1,902) 

W~nfrede ........ 43.8 (6,345) 

York ............ 18.9 (2,735) 

ARMPS CASE HISTORY DATA BASE 

Coalbed specimen strength data were available for approximately 100 case histories in the ARMPS data base. The case 

histories are about evenly split between successes and failures. 

In figure 4, the ARMPS SF are plotted against coal strength. 

All AMPS SF were calculated assuming the in situ strength 

was 6.2 MPa (900 psi). If pillar strength was related to specimen strength, low-strength seams would be expected to fail at 

greater SF than high-strength seams. Instead, no meaningful 

correlation between SF and coal strength is apparent in the 

data. The best discrimination is achieved at an ARMPS SF of 

1.55, with a misclassification rate of 20%. Only one failure is 

included among the misclassifications; which is highly significant from a practical standpoint. 

In a second analysis, the ARMPS SF were recalculated 

using individual seam strengths instead of the uniform in situ 

strength. The seam strengths were divided by 4, as suggested 

by the Gaddy formula for a 6.5-cm (2.5-in) specimen, resulting in a mean SF that is about the same as in the first 

analysis. 

The results are shown in figure 5. Now there is a strong 

correlation between specimen strength and SF, with "stronger" 

coals requiring higher SF to avoid failure. The best misclassification rate, at an SF of about 1.7, is 37%. Also, the 

misclassifications now include 10 failures. In other words, 

when seam-specific strengths are used, the SF becomes almost 

meaningless. 

A third analysis applied seam-specific size-effect exponents 

to the coal strength data, as defined later in the "Size Effect" 

section of this paper. The correlation between seam strength 

and SF was still apparent, as in figure 5. Although the misclassification rate improved to 33%, it was still 50% greater 

than in the uniform seam strength analysis. 
COAL SPECIMEN STRENGTH, psi 

0 1,000 2,000 3,000 4,000 5,000 6,000 7,000 

I I I I I I 

0 

0 0 

- 0 0 - 

9 0 

oO 0 0 - 0 

000 *. 0 Oo 

Q 

8 

O 0 OO 0 0 

- m 8 . 0 

m m.0 8- 

. H. 0 P 

0 0 L% 8 

- . P u0 8 o o m8 ¤ m ¤ . r 

8 o I Key 3! 

I 

- I .m o Success - 

Fai iure 

I I I I I I 

7' 14 2 1 28 35 42 49 

COAL SPECIMEN STRENGTH, MPa 

Figure 4.-ARMPS SF compared with specimen strength data. 

0 7 14 2 1 2 8 3 5 4 2 49 

COAL SPECIMEN STRENGTH, MPa 

Figure 5.-Recalculated ARMPS SF compared using seam-specific coal strength data. 

COAL SPECIMEN STRENGTH, psi 

0 1,000 2,000 3,000 4,000 5,000 6,000 7,000 4.0 

3.5 

LL 

cn 3.0 

V) 

L 

5 2.5 

a 

P 

w 2.0 + 

a 

J 

3 1.5 

0 

J 

a y 1.0 

a 

0.5 

- I I I I I 10 U 

- Key 0 - 

o Success 

Failure - 00 

C 

0 a3 8 - . 

08 

O 0 O 0 

OOm- 

. - 0 n = - 

0 8 0 . 0 8 0. - . 

- - 0 H I 0 . - ZP . - 

.8 O O 8 ou 

B O. .. 

.b: 0 - 0 .. - 

P .o 8 I 

I I I I I I 
The data indicate that there is no meaningful correlation 

between the specimen strength and the in situ strength of U.S. 

coal seams. Knowledge of the specimen strength does not 

improve the accuracy of the design formula's prediction; it reduces it. A uniform coal strength provides a much more reliable prediction of pillar performance. Based on these results, 

laboratory test results are explicitly not recommended for use 

in ARMPS. 

exponents is the different structure of the coalbeds. In a 

blocky coalbed, like the Pittsburgh (figure 6), a small sample 

will be largely free of cleats and fractures. As the specimen 

size increases, the density of cleating increases until it finally 

approaches in situ. In contrast, the fracture density of even a 

small sample of a friable seam like the Pocahontas No. 3 is 

nearly as great as in situ (figure 7). The following relationship 

between size effect and HGI was found (r2 = 0.75): 

SIZE EFFECT a = 0.0063 HGI - 0.75. (7) 

The Coal Strength Data Base contains information from 10 

seams where a wide range of specimen sizes have been tested. 

Five of these were the seams originally tested by Gaddy. 

To determine the size effect, only specimens with a wh ratio of approximately 1:l were used. Figures 6 and 7 show 

how power curves were fit to the data of the form: 

The implications of seam-specific size effects are quite important. It appears that the Gaddy equation underestimates the 

in situ strength of most seams, sometimes by a factor of 3 or 

more. Extremely costly and inefficient mining plans have 

certainly been the result. 

COAL STRUCTURE 

where a = size effect exponent. 

The results are summarized in table 2. Gaddy's a of -0.5 

was found to apply to four seams: the Blind Canyon, Elkhorn, 

Pittsburgh, and Taggart-Marker. At the other extreme, the two 

Pocahontas Seams displayed negligible size effect. The other 

four seams had intermediate size effects. The r2 values indicate that the size effect typically explains about 50% of the 

variability in the test results, which is higher than expected 

considering all of the potential sources of variation in these 

data. The explanation for the substantial range in size effect 

Several analyses explored the relationship between coal 

structure and specimen strength. Figure 8 shows U.S. coalbed 

strengths plotted against HGI. It shows that specimens from 

all seams with HGI > 70 have strengths less than 20 MPa 

(3,000 psi). These seams include all of the medium- and lowvolatile seams in the data base. Likewise, 85% of seams with 

HGI < 50 have a strength exceeding 20 MPa (3,000 psi). For 

the large number of seams between these extremes, the HGI 

is a poor predictor of strength. Many of these intermediate 

HGI seams are high-volatile A in rank. The r2 for the power 

curve fit to the entire data set is 0.33. 

SPECIMEN SIZE, in 

0 

5 10 15 20 25 

SPECIMEN SIZE, cm 

Figure 6.-Size effect in the blocky Pittsburgh Seam. 
SPECIMEN SIZE, in 

0 2 4 6 8 10 12 

Figure 8.-Specimen strength and HGI of U.S. coalbeds. 

17.5 

15.0- 

B 

I 

-12.5 

I 

Ig 10.0 

W 

K 

ti 7.5 

E 

5.0 0 

W 

3 

2.5 

0 

49 (*I I I 1 I I I I ** 

42 - - 

- 

- 

@ - 

** A 

6 

- 

& 

V) ** 

7 -* ** $ - 

0 I I I I I I I I 

I I I I I 

- 

- ** w 

- ** 4 - 

4 

- - 

- 

- 

- 

I I I I 1 

7,000 

6,000 

5,000 

4,000 

3,000 

2,000 

1,000 

0 

0 5 10 15 20 25 30 

SPECIMEN SIZE, cm 

Figure 7.-Size effect in the friable Pocahontas No. 3 Seam. 

40 50 60 70 80 90 100 110 120 130 

HARDGROVE INDEX 
Table 2.-Size effect exponents for 10 U.S. coal seams 

Maximum 

Coalbed No. of No. of HGI Size effect specimen specimens references ( a) k P 

size, cm (in) 

Blind Canyon ....... 126 2 29.5 (1 1.6) 46 -0.54 7,045 0.90 

Clintwood .......... 88 1 18 (7) 63 -0.31 3,686 0.93 

Elkhorn ........... 69 1 16.0 (6.3) 42 -0.55 7,302 0.52 

Harlan ............. 129 2 18 (7) 44 -0.29 6,491 0.31 

Herrin No. 6 ........ 150 5 34.5 (13.6) 56 -0.38 4,293 0.33 

Taggarl-Marker ..... 60 1 18 (7) 47 -0.45 9,837 0.99 

Pittsburgh .......... 272 7 22.9 (9.0) 55 -0.52 5,601 0.48 

Pocahontas No. 3 ... 140 5 29.5 (1 1.6) 110 -0.03 1,670 0.01 

Pocahontas No. 4 ... 74 1 18 (7) 100 -0.13 3,238 0.58 

Umer Banner. ...... 78 1 20.8 (8.2) 84 -0.29 1.730 0.34 

The second analysis compared the structure of the hand Compressive strength and sample structure data were 

samples obtained from the mines with the HGI. In this case, available for 26 seams. The specimen strength of all eight 

every seam rated "blocky" had an HGI less than 60. The HGI blocky seams exceeded 23 MPa (3,500 psi), but so did that of 

of the "semiblocky" seams was less than 80. "Friable" seams four friable and one semiblocky seam. Another 13 friable and 

were found throughout the range of HGI. semiblocky seams were intermixed below 23 MPa (3,500 psi). 

CONCLUSIONS 

The results of this study cast doubt on many textbook assumptions about the value of coal strength testing. The data 

clearly show that specimen strength and the "size effect" are 

highly seam-specific and related to coal structure. The widely 

used Gaddy formula, which applies a uniform strength reduction for all seams as specimen size is increased, only applies 

to "blocky" coals with cleats spaced more than 8 cm (3 in) 

apart. For friable coals, the size effect was much less pronounced, or even nonexistent. 

Case histories of failed pillars are the best available data on 

in situ coal strength. This study found no correlation between 

the ARMPS SF of failed pillars and coal specimen strength in 

the ARMPS data base. In current ARMPS practice, pillars are 

designed assuming an in situ strength of 6.2 MPa (900 psi) for 

all seams. When the specimen strength was used instead, the 

reliability of the ARMPS design method decreased substantially. Australian and South African studies have also found 

that pillar strength in the field is largely independent of 

specimen strength. 

It should be noted that the coal strength tests were only 

matched with the seams in the case histories, not with the individual mines. It is also possible that some of the case histories involved roof or floor failure rather than pillar failures. 

Using a different pillar strength formula might also have 

changed the results somewhat. However, the data base is so 

large and the trends so strong that it is highly unlikely that the 

study is unrepresentative. 

The most likely explanation for the results of the study is 

that specimen and in situ strengths are determined by different 

parameters. Laboratory tests, particularly those of blocky 

coals, require a significant amount of fracturing of intact coal. 

Pillars contain so many cleats and other discontinuities that 

their failure can occur almost entirely along preexisting fractures. The laboratory tests measure a parameter-the intact 

coal strength-that is apparently irrelevant to the in situ 

strength. 

The study did not prove that the in situ strength of all U.S. 

coals is uniform. It only showed that a uniform strength is a 

better approximation than one based on laboratory testing. 

There is still significant variability in the ARMPS SF range of 

0.75 to 1.5. Recent model studies have indicated that features 

such as roof and floor interfaces, bedding planes, partings, or 

weak coal layers have the greatest effects on in situ strength 

[Iannacchione 1990; Su and Hasenfus 19961. A rock mass 

classification, such as the one proposed by Kalamaras and 

Bieniawski [1993], may prove to be an effective way of evaluating these effects in the field. In the meantime, laboratory 

uniaxial coal strength test results should not be used for pillar 

design with ARMPS. 
REFERENCES 

Bieniawski ZT [1968]. The effects of specimen size on the compressive 

strength of coal. lntl J Rock Mech Min Sci 5:325-335. 

Bieniawski ZT [I984]. Rock mechanics design in mining and tunneling. 

Balkema, Rotterdam, pp. 55-92. 

Bunting D [I91 11. Chamber pillars in deep anthracite mines. Trans AIME 

42:236-245. 

Daniels J, Moore LD [1907]. The ultimate crushing strength of coal. Eng 

Min J, Aug 10:263-268. 

Gaddy FL [1956]. A study of the ultimate strength of coal as related to the 

absolute size of cubical specimens tested. VA Polytechnic Bull 112, Aug: 1-27. 

Galvin JM [1995]. Roadway and pilIar mechanics workshop: stage 2 - 

design principles and practice (Course notesj. Sydney, New South Wales, 

Australia: University of New South Wales. 

Holland CT, Gaddy FL [1956]. Some aspects of permanent support of 

overburden on coal beds. In: Proceedings of the West ~irginia Coal Mining 

Institute, pp. 43-65. 

Husttulid WA [1976]. A review of coal pillar strength formulas. Rock 

Mechanics 8:115-145. 

Iannacchione AT [1990]. The effects of roof and floor interface slip on 

coal pillar behavior. In: Hustrulid WA, Johnson GA, eds. Proceedings of the 

31st U.S. Symposium on Rock Mechanics. Golden, CO: Colorado School of 

Mines, pp. 153-160. 

Kalamaras GS, Bieniawski ZT [1993]. A rock mass strength concept for 

coal seams. In: Proceedings of the 12th Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, pp. 274-283. 

Khair AW [1968]. The effects of coefficient of friction on strength of 

model coal pillar [Thesis]. Morgantown, WV: West Virginia University, Department of Mining Engineering. 

Madden BJ [1991]. A re-assessment of coal-pillar design. J South Afr lnst 

Min Metall 91:27-37. 

Mark C [1990]. Pillar design methods for longwall mining. Pittsburgh, 

PA: U.S. Department of the Interior, Bureau of Mines, IC 9247. 

Mark C, Iannacchione AT [1992]. Coal pillar mechanics: theoretical 

models and field measurements compared. In: Iannacchione AT, Mark C, 

Repsher RC, Tuchman RJ, Jones CC, comp. Proceedings of the Workshop on 

Coal PiIlar Mechanics and Design. Pittsburgh, PA: U.S. Department of the 

Interior, Bureau of Mines, IC 9315, pp. 78-93. 

Mark C, Chase FE, Molinda GM [1994]. Design of longwall gate entry 

systems using roof classification. In: Mark C, i'uchman RJ, Repsher RC, 

Simon CL, comp. New Technology for Longwall Ground Control; Proceedings: U.S. Bureau of Mines Technology Transfer Seminar. Pittsburgh, PA: 

U.S. Department of the Interior, Bureau of Mines, SP 01-94, pp. 5-18, 

Mark C, Chase FE [1997]. Analysis of retreat mining pillar stability 

(ARMPS). In: Mark C, Tuchman RJ, comp. Proceedings: New Technology 

for Ground Control in Retreat Mining. Pittsburgh, PA: U.S. Department of 

Health and Human Services, Public Health Service, Centers for Disease 

Control, National Institute for Occupational Safety and Health, IC 9446. 

Mrugala MJ, Belesky RM [1989]. Pillar sizing. In: Proceedings of the 

30th U.S. Symposium on Rock Mechanics. Morgantown, WV: West Virginia 

University, pp. 395-402. 

Newman DA, fioelie JL i1993j. Tne impact of variabiiity in coai suengrh 

on mine planing and design - a case history. In: 12th Conference on Ground 

Control in Mining. Morgantown, WV: West Virginia University, pp. 237- 

243. 

Panek LA [1994]. Scaling mine pillars' size and shape with the psi function. Soc Min Eng AIME preprint 94-52, pp. 1277-1280. 

Salmon MDG [1991]. Behavior and design of coal pillars. Australian 

Coal J 32:ll-22. 

Salmon MDG, MUNO AH [1967]. A study of the strength of coal pillars. 

J South Afr lnst Min Metall, Sept:55-67. 

Sheorey PR, Das MN, Barat D, Prasad RK, Singh B [1987]. Coal pillar 

strength estimation from failed and stable cases. lntl J Rock Mech Min Sci 

Geomech Abstr 24(6):347-355. 

Su DWH, Hasenfus GJ [1996]. Practical coal pillar design considerations 

based on numerical modeling. In: 15th Conference on Ground Control in 

Mining. Golden, CO: Colorado School of Mines. 

Townsend JM, Jennings WC. Haycocks C, Ned GM, Johnson, LP [1977]. 

A relationship between the ultimate compressive strength of cubes and cylinders for coal specimens. In: Symposium for Rock Mechanics, Keystone, 

CO, pp. 4A6-1 to 4A6-6. 

Unrug KF, Nandy S, Thompson E[1985]. Evaluation of the coal strength 

for pillar calculations. SME AIME preprint 85-314. 

Van der Merwe JN [1993]. Revised strength factor for coal in the Vad 

basin. J South Afr Inst Min Metall 93(3), pp: 71-77. 

Zem EN [1928]. Coal miner's pocketbook. McGraw-Hill, NY, 12th ed., 

pp. 641-645. 
A NEW LAMINATED OVERBURDEN MODEL 

FOR COAL MINE DESIGN 

By Keith A. Heasley' 

ABSTRACT 

In the past, numerous boundary-element models of stratified rock masses have been proposed using a 

homogeneous isotropic elastic overburden. In this paper, it is postulated that a laminated overburden model 

might be more accurate for describing the displacements and stresses in these stratified deposits. In order to 

investigate the utility of using a laminated overburden in a boundary-element model, the fundamental 

mathematical basis of the laminated model is presented and graphically compared with the fundamental 

behavior of homogeneous isotropic elastic overburden and with field data. Specifically, the stresses and 

displacements surrounding an idealized longwall panel as determined from the laminated overburden model 

are presented and compared with results from the homogeneous isotropic overburden and with measured 

abutment stress data. Additionally, the remote displacements and surface subsidence as calculated by the 

laminated overburden model are compared with homogeneous isotropic calculations and with measured 

subsidence data. Finally, the new laminated boundary-element program, LAMODEL, is used to model the 

underground stresses and displacements, the topographic stresses, and the interseam interactions at a field site. 

The results of this investigation show that the laminated overburden is more supple, apt to propagate 

displacements and'stress further, and better able to fit observed data than the classic homogeneous isotropic 

overburden. Ultimately, it is suggested that the laminated model has the potential to increase the accuracy of 

displacement and stress calculations for a variety of mining situations. 

'Mining engineer, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA 
INTRODUCTION 

If one wishes to perform a mechanical analysis of the 

geologic structure of a mining operation, there are several 

broad mathematical techniques available. For instance, one 

may choose finite-element, boundary-element, discreteelement, finite-difference techniques, and/or hybrid combinations of these techniques. In general, these mathematical 

techniques have strengths and weaknesses when applied to a 

specific geologic envuonment, mining geometry, and material 

behavior. Naturally, in each practical application the 

mathematical technique best suited to the prevailing conditions 

should be applied. 

To analyze the displacements and stresses associated with 

the extraction of tabular deposits, such as coal, potash, and 

other thin vein-type deposits, the displacement-discontinuity 

version of the boundary-element technique is frequently the 

method of choice. In the displacement-discontinuity approach, 

the mining horizon is treated mathematically as a discontinuity 

in the displacement of the surrounding media. Thus, only the 

planar area of the seam is discretized in order to obtain the 

solution. Often this limited analysis is sufficient, because in 

many applications only the distributions of stress and 

convergence on the seam horizon are of interest. In addition, 

by limiting the detailed analysis to only the seam, the 

displacement-discontinuity method provides considerable 

computational savings compared with other techniques that 

discretize the entire body (such as finite-element, discreteelement, or finite-difference). It is a direct result of this 

computational efficiency that the displacement-discontinuity 

method is able to handle problems involving large areas of 

tabular excavations. 

In the original mathematical formulations [Berry 1960; 

Salamon 19621 and computer implementations [Plewman et al. 

1969; Crouch and Fairhurst 1973; Sinha 19791 of the 

displacement-discontinuity variation of the boundary-element 

method, the media surrounding the seam were assumed to be 

homogeneous, isotropic, or transversely isotropic elastic. This 

basic behavior of the surrounding media provided fairly good 

seam-level displacement and stress results for South African 

gold mines [Salamon 1964; Cook et al. 19661 and for U.S. coal 

mines [Kripakov et al. 1988; Zipf and Heasley 1990; Heasley 

and Zclank~ '9921. However, it was noted early in the 

application of the displacement-discunti~~uity method that thc 

homogeneous isotropic overburden model does not fit 

measured subsidence data [Berry and Sales 19611. As an 

alternative, it was proposed that a laminated model might be 

more suitable for describing the behavior of stratified coalmeasure rocks [Salamon 196 1, 19631. Recently, a laminated 

overburden model was found to give good results for 

predicting surface subsidence [Salamon 1989a; Yang 19921. 

Because the source of surface subsidence is convergence in the 

seam, it seems reasonable that a laminated overburden model 

might also be able to provide more accurate predictions of inseam displacements and stresses. 

If the utility of using a laminated overburden in a 

displacement-discontinuity model is to be determined, the 

fundamental mechanical behavior of the laminated model 

needs to be investigated and compared with both the classic 

homogeneous isotropic model and field data. In this paper, the 

stresses and displacements surrounding an idealized longwall 

panel as determined from the laminated overburden model, the 

homogeneous isotropic overburden model, and field data are 

presented and compared. In addition, the remote displacements and surface subsidence as calculated by the two 

overburden models are compared with measured subsidence 

data. Lastly, the laminated overburden model is applied to a 

site study. 

FUNDAMENTAL BEHAVIOR 

MATHEMATICAL FOUNDATION 

The mathematical basis for the laminated model was originally proposed by Salamon in 1961-62 and more recently updated in 1991. Conceptually, the media in the laminated model 

consist of a horizontal stack of homogeneous isotropic layers 

where the interfaces between the layers are parallel and free of 

shear and cohesive stresses, and the vertical stresses and displacements are continuous across the layers. In the "homogeneous stratification" version of the model, all layers have the 

same elastic modulus (E), Poisson's ratio (v), and thickness (t). 

Thus, the homogeneous stratification formulation does not 

allow (or need) the specification of the properties for each 

individual layer, yet it provides the desired suppleness of the 

basic laminated model (compared with a homogeneous isotropic 

elastic model). In addition, the behavior of the rock mass in the 

laminated model is effectively characterized by two parameters, 

the elastic modulus and the lamination thickness, whereas the 

homogeneous isotropic model only has a single effective parameter, the rock mass modulus (the Poisson's ratio has a minor 

effect in both models). 

The mathematical foundation of the laminated model is the 

theory of thin plates [Salamon 19911. From this theory, the 

relationship between the vertical deflection (w) of the middle 

plane of a horizontal plate and the resultant transverse pressure 

(p) acting on the plate is defined by 
D V4 W(X,Y) = P(x,Y), (1) 

where D is the flexural rigidity of a plate: 

and denotes the biharmonic operator in the xy plane, 

specifically: 

From equation 1, the convergence in the seam (S) can be related 

to the induced stress (a,) in the overburden laminae by the 

following second-order, partial differential equation [Salmon 

1991.1: 

where the laminae-related value, A, is defined as 

and oi is the vertical, or transverse, stress on the laminae at seam 

level induced by mining. 

PANEL CONVERGENCE 

The first step in investigating the fundamental behavior of 

the laminated model was to analyze the convergence across a 

two-dimensional slot. This slot can be viewed as an idealized 

longwall panel with no gob support and rigid ribs. From 

equation 4, the seam convergence across a two-dimensional slot 

for the laminated model (S,) as a function of the distance from 

the panel centerline (X) can be determined as 

Here, L is the half-width of the slot; q is the primitive vertical 

stress at the mining horizon, which for an open panel is equal to 

the induced stress (a,). In solving equation 4, it was assumed 

that the convergence value at the rib side is zero and that the 

convergence distribution is symmetric about the panel centerline. Also, in this result and the result in equation 7, the stressfree ground surface was ignored. 

Jaeger and Cook [I9791 provide a comparable equation for 

the roof-to-floor convergence across a two-dimensional slot 

with homogeneous isotropic elastic overburden (S,): 

The fundamental difference between these two equations is that 

the convergence in the laminated model is proportional to the 

square of the panel span, while the convergence in the 

homogeneous model is linearly proportional to the span. 

In order to plot and compare the convergence computed 

from equations 6 and 7, some "typical" values were assumed for 

the geometric and rock mass parameters: a panel width of 

200 m (656 ft)(L = loo), an overburden depth (H) of 160 m 

(525 ft)(q = 4 MPa (580 psi)), a seam height (M) of 2 m (6.6 ft), 

an elastic modulus of the rock mass (E) of 20 GPa (2.9 million 

psi), a Poisson's ratio (v) of the rock mass of 0.25, and a 

lamination thickness (t) of 15 m (49 ft). Using these values for 

the parameters, the convergence across the slot for both the 

laminated and the homogeneous overburden is plotted in figure 

1. As expected from the nature of the equations, the laminated 

overburden is considerably more flexible. In fact, with the 

given parameters, the laminated overburden exhibits six times 

the convergence of the homogeneous isotropic overburden. 

ABUTMENT STRESS 

The next step in investigating the fundamental behavior of 

the laminated model was to analyze the abutment stress at the 

edge of a two-dimensional slot. If the seam is assumed to be 

linear elastic with a modulus of E, and Poisson's effect is 

ignored, then the induced stress in the seam for the laminated 

model is 

Then, from equation 4, the in-seam convergence in the laminated model is defined by 

d2Sp 2Es 

S, =o, 

dX2 ELM 

which has a solution (for positive X values): 

The associated induced vertical abutment stress (0,) in the 

unmined seam bounding the panel is 

(1 1) 

ELM 

The total abutment stress at the edge of a two-dimensional 

slot in a homogeneous isotropic elastic model (a,) is given by 

[Salmon 19741: 
Distance from Center of Panel (m) 

- 

- 

Figure 1 .-Comparison of longwall convergence computed from the laminated and homogeneous isotropic models. 

It is noteworthy that this equation is independent of material 

properties. During its derivation, it was assumed that the seam 

material was identical to the surrounding media. Thus, to be 

consistent, the elastic modulus of the seam (E,) in equation 11 

is assumed to be equal to the modulus of the overburden (E), 20 

GPa (2.9 million psi). 

In addition, numerous fieid measurements of abutment load 

have been tabulated by Mark [1990], where he found that the 

measured distribution of induced abutment stress (0,) follows 

the equation: 

I I I I 

A I I I I I I I 

J Homogeneous Elastic Overburden 

where L, is the total side abutment load (which in our case 

without any gob load is equal to qL), and D is the maximum 

horizontal extent of the abutment stress from the panel edge, 

which was determined from field measurements [Mark 19901: 

/ 

- / - 

/ 

K / / 

0 

- 0 

0 

0 

0 10 0 0 

0 

- 0 

0 -- 

0 

0 

0 

/ 

0 

m / / - 

// 

I- - -**( 

Laminated (t.15) Overburden I 

The total abutment stress curves, as calculated from the 

laminated and homogeneous isotropic models and from the 

empirical formula (equations 11, 12, and 13), are plotted in 

- 

m 

m 

figure 2. (It should be noted that equations 11 and 13 calculate 

induced stresses and equation 12 calculates the total stress. 

Therefore, in the following plots, the virgin overburden stress 

(q) has been added to the results from equations 11 and 13 to 

provide a valid comparison with the total abutment stress values 

from equation 12.) In figure 2, it can be seen that the 

homogeneous isotropic abutment stress has a relatively sharp, 

infinite peak at the edge and approaches zero asymptotically 

with increasing distance from the panel. In contrast, the 

abutment stress in the laminated overburden is finite at the 

panel edge and approaches virgin overburden stress (q) rapidly. 

Neither of these mathematical models (using the assumed 

parameters) comes very close to matching the empirical 

abutment stress. 

However, if the abutment stress level in the laminated model 

and that obtained from the empirical formula are equated at the 

edge of the seam (X = L), then the lamination thickness (t) that 

ensures this equality can be determined: 

For a typical seam modulus (E,) of 2 GPa (290,000 psi) in the 

laminated model, equation 15 provides a fitted lamination thickness of 157 m (515 ft). The plot of the abutment stress curve 

for the laminated overburden model with a fitted lamination 

thickness of 157 m (515 ft) is shown together with the empirically determined abutment stress in figure 3. The degree of 

agreement between the two curves is very good and serves to 
Figure 2.--Comparison of longwall abutment stress computed from the laminated and homogeneous isotropic 

models and from the empirical formula. 

0 

0 10 20 30 40 50 60 

Distance from Edge of Panel (m) 

- 

- 

- 

50 

h 2 40-1 

s 

- U) 30 

U) 

0 I 

I 

I 

I 

I 

I 

I 

I 

I 

I 

0 10 20 30 40 50 60 

Distance from Edge of Panel (m) 

/ I 

I 

I 

I 

I 

I 

I 

I 

I 

I 

. \ Laminated (t=15) Overburden 

-! \ - 

\ 

'\ 

\ 

i \ -100 o 50 

--I Homogeneous Elastic Overburden - 

\ \ 

Figure 3.-Plot of the laminated abutment stress fitted to the empirical formula. 

- 

- - - - 

-100 0 50 

- - - - 

. rn Empirical Abutment Stress - - 

. -------, 

Laminated (t=157) Overburden I I I I I 

I 

I 

I 

I 

I 

I 

I 

I 

I 

I 

- - - 
highlight the numerical flexibility provided by a variable lamination thickness parameter in the laminated model. (If a compressible seam is used, it may be possible to fit the abutment 

stress distribution in the homogeneous isotropic overburden to 

the empirical curve by varying the seam and overburden 

moduli. However, that calculation is beyond the scope of this 

paper.) 

REMOTE DISPLACEMENTS 

The next step was to analyze the remote displacements in the 

overburden generated by seam convergence. For the !aminated 

overburden model, the kernel, or influence function, which 

relates the seam convergence (S,) to the vertical displacement 

(Wt) of the overburden, was derived by Salmon 11962, 1989b] 

and Yang [1992]: 

Here, the magnitude of the convergence, S,, is assumed to occur 

over a unit element of the seam; the values of X and Y are the 

horizontal and vertical distances between the centroid of the 

converged seam element and the point in the overburden at 

which the displacement is desired. 

Similarly, the kernel for the homogeneous isotropic overburden, which relates the seam convergence (S,) to the vertical 

displacement (W,) of the overburden, was derived by Crouch 

[I9761 (see equation 17). Again, the magnitude of the convergence, s,, is assumed to occur over a unit element of the 

seam, and the coordinates X and Y were defined previously in 

conjunction with equation 16. Note that the expression in equation 17 is again independent of elastic moduli. 

Plots of the overburden displacements generated by these 

models are depicted in figure 4 for Y = 20 m (66 ft) and Y = 

50 m (164 ft). In computing this illustration, a unit convergence 

spread over a seam element of unit length was assumed. Thus, 

the volume of convergence is identical in the two models. 

However, consistent with the greater suppleness of the 

laminated model, the stratified overburden (with a 15-m (49-ft) 

lamination thickness) appears to concentrate the displacement 

more tightly over the panel. This feature is particularly noticeable as the distance from the seam is increased. 

This difference in remote displacement behavior is even 

more obvious in figure 5, which shows the 1-cm (0.39-in) displacement contours for both models generated above a unit 

volume convergence in a seam element. In this figure, the 1-cm 

(0.39-in) contour, or displacement "bulb," for the laminated 

model (with a 15-m (49-ft) lamination thickness) is broader and 

extends almost twice the distance into the overburden as the 

contour from the homogeneous elastic model. However, if the 

lamination thickness in the laminated model is increased to 

28 m (92 ft)(as shown ~n figure 51, then the vertical extent of the 

1-cm (0.39-in) displacement contour is equal between the two 

models, although the laminated displacement bulb is still 

broader. For most practical purposes (lower lamination thicknesses), both figures 4 and 5 indicate that seam displacements 

and stresses for a laminated overburden would propagate further 

and in a tighter pattern than the displacements and stresses from 

the homogeneous overburden. This greater remote response, 

coupled with the tendency for the laminated model to produce 

greater seam convergence, should greatly increase the remote 

displacements and stresses associated with a laminated 

displacement-discontinuity model. 

SURFACE SUBSIDENCE 

The final step in investigating the fundamental behavior of 

the laminated model was to analyze the surface subsidence over 

a longwall panel. In this analysis, the surface subsidence curves 

for the laminated and homogeneous isotropic models were 

calculated by taking the panel convergence from equations 6 

and 7 and numerically integrating the surface subsidence using 

equations 16 and 17. These calculated subsidence curves are 

then compared in figure 6 with the results of the U.S. Bureau of 

Mines subsidence model [Jeran et al. 19861, which essentially 

represents an empirically derived "average" of 11 different 

subsidence curves from the Northern Appalachian Coa! Basin. 

For the subsidence curve of the laminated model in figure 6, 

the lamination thickness was optimized to provide the best fit 

with the empirical curve. This resulted in a lamination 

thickness of 5.3 m (17 ft), and from the figure, it can be seen 

that the laminated model provides a good fit to the empirical 
Homogeneous 

50 m above seam 

Laminated (t=15) 

-60 -40 -20 0 20 40 60 

Horiz. Dist. from Center of Seam Element (m) 

Homogeneous 

20 m above seam /- 

- - 

- - 

I 

Overburden 

Figure 4.-Remote displacement due to a unit volume displacement. 

curve. In earlier work where the laminated model was fit to the empirical curve. However, long before the maximum 

several different individual sets of subsidence data [Yang 19921, surface subsidence from the model matched the maximum 

a factor, o: empirical surface subsidence, the convergence in the seam 

exceeded the seam thickness. The homogeneous elastic surface 

subsidence actually plotted in figure 6 was determined using an 

= pE (18) elastic modulus of 1 GPa (145,000 psi). From this curve, it is 

v 2A, clear that the homogeneous isotropic surface subsidence is 

was found to be fairly constant at an average of 6.9. (Here, H 

is the overburden depth, and A is defined in equation 5.) For the 

fitted subsidence curve from the laminated model in figure 6, 

the value of o is 7.1, which agrees very well with this previous 

work. 

For the subsidence curve from the homogeneous elastic 

overburden, the elastic modulus was lowered in an attempt to fit 

naturally much shallower and broader than the empirical data, 

and with only one effective variable parameter (E), the 

homogeneous isotropic model cannot be accurately fitted to the 

Northern Appalachian data. This result further confirms earlier 

indications that the homogeneous isotropic elastic overburden 

could not be made to fit subsidence data in the United Kingdom 

[Berry and Sales 1961; Salmon 19631. 
Horiz. Dist from Center of Seam Element (m) 

I Laminated (t.15) 

/ 

/ 

/ 

/ 

/ 

/ 

/ 

Homogeneotts 

Elastic Ovetburden 

Figure 5.-1-cm (0.39-in) displacement contours associated with a unit volume displacement. 
-200 -100 0 100 200 

Distance from Center of Panel (m) 

Figure 6.-Plot of the laminated and homogeneous isotropic subsidence fitted to the empirical curve. 

THE LAMINATED MODEL PROGRAM, LAMODEL 

IMPLEMENTATION FEATURES 

The iaminated overburden model, as presenled in the 

previous sections of this paper, has been implemented into a 

modern boundary-element computer program called 

LAMODEL. This implementation has numerous practical 

features, including- 

* Single- and multiple-seam simulations. 

Numerous individual excavation steps. 

Infinite media or surface effects for shallow seams. 

A constant overburden or a variable topography. 

Seam-level convergence and stress calculation, with each 

of the individual stress components (overburden, material, interseam, and surface) separately tabulated. 

User-defined laminae properties (elastic modulus, Poisson's ratio, and thickness). 

Up to 26 different in-seam materials can be specified from 

a selection of material models, which include elastic, elasticplastic, strain-softening, bilinear strain-hardening, and exponential strain-hardening. 

User-defined convergence criteria. 

Grid sizes limited solely by the computation requirements 

(practical limit: 300 by 300). 

Either rigid or symmetric boundary conditions. 

Graphical pre- and postprocessors for simplified input entry and output analysis. 

CASE STUDY 

As part of the initial investigation and validation of this new 

implementation, the underground stresses, displacements, 

topographic stresses, and interseam interactions were modeled 

at a field site. This case study site is a multiple-seam situation 
in eastern Kentucky. The geology in this area is fairly typical topographic relief of over 600 m (2,000 ft). At the case study 

of the Southern Appalachian Coal Basin, with various site, the overburden averages about 240 m (800 ft), but ranges 

sedimentary layers of sandstones, siltstones, shales, and from 90 m (300 ft) at the southeastern corner of the site to over 

numerous coal seams. The topography in the area is very 360 m (1,200 ft) at the northwestern comer (figure 7). (Because 

rugged, with various steep ridges and valleys that have a of the steep topography, it was critical to include the 

, 

Figure 7.-Map of case study mines in eastern Kentucky. 
topographic stress effects in the model to obtain accurate 

results.) 

The overlying mine operates in the Upper Darby Seam, 

which typically averages about 2.0 m (6.5 ft) thick; however, in 

the model area, the extraction thickness had increased to over 

2.7 m (9 ft). The lower mine operates in the Kellioka Seam, 

which averaged about 1.5 m (5.0 ft) thick in the study area. The 

interburden between the mines averages around 14 m (45 ft) 

and consists of interbedded sandstones and shales. The core 

logs nearest to the study site indicate about 3.5 to 5 m (12 to 15 

ft) of thinly laminated shaley/carbonaceous sandstone (stack 

rock) directly over the Kellioka Seam. This is then overlain by 

7.5 to 10.5 m (25 to 35 ft) of interbedded sandstones and shales, 

with shale primarily forming the floor of the Upper Darby 

Seam. 

Both mines are room-and-pillar drift mines and utilize 

continuous miners for coal extraction. In some production 

sections, depending on local mining conditions, the mines 

remove the pillars on retreat for full extraction. In the study 

area, the lower mine had driven a seven-entry-wide set of main 

entries from north to south with pillars on 2 1 - by 24-m (70- by 

80-ft) centers and 6-m (20-ft) wide entries. Subsequently, the 

upper mine drove a seven-entry-wide set of panel development 

entries roughly perpendicular across the lower mains (figure 7). 

Relatively short (one- to two-crosscut) production rooms were 

driven to the north of the upper mine development entries 

during advance. At this point, no appreciable stress interaction 

was observed. Then, as the upper mine was pulling out of the 

section, long (seven- to eight-crosscut) production rooms with 

pillars on 18- by 18-m (60- by 60-ft), and smaller, centers were 

driven on the south side of the development entries (figure 7). 

At the extent of mining shown in figure 7, the upper mine began 

to experience major problems with pillar failure and floor heave 

and was forced to abandon the section. 

Coincident with the failures in the upper mine, the lower 

mine experienced ground control problems in areas directly 

underlying the boundaries of the upper panel. These problems 

were primarily manifested as increased pillar spalling for 

approximately 30 m (1 00 ft) of entry and major roof cracking 

at overmined intersections. Both of these ground control 

problems were mitigated by supplemental bolting and cribbing. 

The new laminated model, LAMODEL, was applied at this 

site to both quantify the magnitude of the stress interaction 

between the seams and to correlate the model results with inmine ground control problems for subsequent predictions of 

mining conditions in future mine planning analysis. In the 

model, the seams were discretized with 3-m (10-ft) elements on 

150-by-150 grids with the extent as shown in figure 7. 

Symmetrical seam boundary conditions were set, and no freesurface effects were included. The interburden was set at 14 m 

(45 ft), and the rock mass was simulated with a modulus of 

20,700 MPa (3 million psi) and 15-m (50-ft) thick laminations. 

A strain-softening material was used for the in-seam coai, and 

the peak strength of the coal was varied until the pillars in the 

upper seam had just reached failure. 

Additionally, because of the high topographic relief at the 

site, the topography was discretized with 15-m (50-ft) elements 

for an area extending 300 m (1,000 ft) beyond the limits of the 

displacement-discontinuity grids. The importance of including 

the topographic stress effects in the model is clearly evident in 

figure 8, which shows the topographic stress at the level of the 

upper mine. It is interesting to note in this figure the amount to 

which the topographic stress is "smoothed" with depth in 

comparison with the original topography shown in figure 7. 

Also, it should be observed in figure 8 that near the boundary of 

the upper mine (the area of interest) the topographic stress 

varies about 1.4 MPa (200 psi), or 30%, across the pillars in the 

lower mine. 

The primary results of this multiple-seam modeling effort are 

shown in figure 9. Figure 9C shows the stress concentrations 

on the lower seam resulting from the pillar failure in the upper 

seam. In this image, stress concentrations up to 4.5 MPa (650 

psi) and with functional widths of between 9 to 37 m (30 to 120 

ft) can be seen. This model response correlates well with the 

underground observations. The calculated seam interaction 

stress results in increases of the average pillar stress in the lower 

mine up to 55% (figures 9A and 9B). By correlating this 55% 

increase in pillar stress with the observed ground control 

problems underground (figure 9), the magnitude of future 

ground control problems at this site can now be more accurately 

determined using LAMODEL. 
Legend (MPa) 

2.0 3.4 4.8 6.2 7.6 9.0 

Figure 8.-Topographic stress on lower seam. 
mmm &;a 

Z'LC S'PC L'CC 0'6 2' '8 Z'LC S'PC L'CC 0'6 
SUMMARY AND CONCLUSIONS 

The investigation of the fundamental behavior of the 

laminated overburden presented in the first part of this paper 

has produced a number of significant results. First, it is clear 

that the laminated overburden (with a low lamination thickness) 

is more flexible or supple than the homogeneous isotropic 

overburden. This increased suppleness is evident in the greater 

convergence across a longwall panel, in the larger extent of the 

remote displacement contours, and in the nature of the surface 

subsidence. Second, due to the larger extent of the remole 

displacement contours, it is clear that a multiple-seam mine 

model using the laminated overburden will show increased 

interseam displacements and stresses compared with a 

homogeneous isotropic overburden. Third, because the 

overburden in the laminated model is effectively described with 

two parameters (as opposed to one parameter in the homogeneous isotropic model) and therefore provides two degrees of 

freedom for fitting observed data, the laminated model was 

capable of closely matching the observed abutment stresses and 

surface subsidence. In fact, the laminated model was easily fit 

to the observed surface subsidence, whereas the subsidence for 

the homogeneous isotropic overburden was fundamentally 

different from the observed subsidence. 

In the second part of this paper, a new laminated 

displacement-discontinuity program, LAMODEL, was 

presented. This new program, in addition to the laminated 

overburden, also implements a number of innovative features, 

including topographic stress calculations, various in-seam 

material models, and variable boundary conditions at the seam 

level. In order to evaluate the accuracy and utility of the new 

model, it was used in a case study of a multiple-seam mining 

sitilation in s!.eep topography. At this site. the ability of 

LAMODEL to include topographic stress effects, strainsoftening coal, and symmetric boundary conditions greatly 

increased the realism and accuracy of the model. By correlating 

the LAMODEL results with the observed ground control 

problems, mine management will be better able to design and 

plan for future multiple-seam interact~ons. Because of the 

realistic flexibility of the laminated overburden model and the 

utility of the numerous practical features implemented in the 

new program, it appears that LAMODEL can provide realistic 

stress and displacement calculations for a wide range of mining 

situations. 

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Beny DS, Sales TW [1961]. An elastic treatment of ground movement due 

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Cook NGW, Hoek E, Pretorious JPG, Ortlepp WD, Salamon MDG [1966]. 

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Crouch SL, Fairhurst C 119731. The mechanics of coal mine bumps and the 

interaction between coal pillars, mine roof, and floor. U.S. Department of the 

Interior, Bureau of Mines, OFR 53-73. 

Heasley KA, Zelanko JC [I 9921. Pillar design in bump-prone ground using 

numerical models with energy calculations. In: lannacchione AT, Mark C, 

Repsher RC, Tuchman RJ, Jones CC, comp. Proceedings of the Workshop on 

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Interior, Bureau of Mines, IC 9315, pp. 50-60. 

Jaeger JC, Cook NGW [1979]. Fundamentals of rock mechanics. London: 

Chapman and Hall. 

Jemn PW, Adamek V, Trevits MA [1986]. A subsidence prediction model 

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Pittsburgh. PA: pp. 101-1 12. 

Kripakov NP, Beckett LA, Donato DA, Durr JS [198X]. Computer-assisted 

mine design procedures for longwall mining. Pittsburgh. PA: U.S. Department 

of the Interior, Bureau of Mines, R1 9172. 

Mark C [1990]. Pillar design methods for longwall mining. Pittsburgh, PA: 

U.S. Department of the Interior, Bureau of Mines. 1C 9247. 

Plewman RP, Deist FH, Ortlepp WD 119691. The development and 

application of a digital computer method for the solution of strata control 

problems. J South Afr Inst Min Metall 70(9):33-44. 

Salamon MDG [1961]. An introductory mathematical analysis of the 

movements and stresses induced by mining in stratified rocks. Durham, United 

Kingdom: King's College, University of Durham, Department of Mining, 

Research Report 9:Bull 3. 

Salamon MDG [1962]. The influence of strata movement and control on 

mining development and design [Thesis]. Durham, United Kingdom: University 

of Durham, Department of Mining. 

Salamon MDG [1963]. Elastic analysis of displacements and stresses 

induced by the mining of seam or reef deposits, part I. J South Afr Inst Min 

Metal1 63(4): 128-149. 

Saiamon MDG iI964j. Elastic analysis of displacements and stresses 

induced by the mining of seam or reef deposits, part 11. J South Afr Inst Min 

Metall 64(6): 197-2 18. 

Salamon MDG [1974]. Rock mechanics of underground excavations. In: 

Proceedings of the 3rd Congress, International Society for Rock Mechanics. 

Denver. CO: National Academy of Sciences, l(b)951-1099. 

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model. In: Proceedings of the 30th U.S. Symposium on Rock Mechanics. 

Morgantown, WV: West Virginia University, pp. 891-898. 

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model. In: Proceedings of the 30th U.S. Symposium on Rock Mechanics. 

Morgantown, WV: West Virginia University, pp. 503-510. 

Salamon MDG [1991]. Deformation of stratified rock masses: a laminated 

model. J South Afr Inst Min Metall 91(1):9-26. 

Sinha KP [1979]. Displacement discontinuity technique for analyzing 

stresses and displacements due to mining in seam deposits [Thesis]. 

Minneapolis, MN: University of Minnesota. 

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longwall coal mining using a laminated model [Thesis]. Golden, CO: Colorado 

School of Mines. 

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Hustrulid WA. Johnson GA, eds. Proceedings of the 31st U.S. Symposium on 

Rock Mechanics. Golden, CO: Colorado School of Mines, pp. 947-954. 
RETREAT MINING WITH MOBILE ROOF SUPPORTS 

By Frank E. Chase,' Allen McComas,' Christopher Mark, Ph.D.,3 and Chester D. Goble4 

ABSTRACT 

Mobile roof supports (MRS's) are shield-type support units mounted on crawler tracks. MRS's are used 

during retreat mining and eliminate the setting of roadway, turn, and crosscut breaker posts that are required 

during pillar recovery operations. Mobiles are a more effective ground support than timbers, and their usage 

enhances the safety of section personnel and reduces material handling injuries. MRS usage is rapidly 

increasing, and approximately 40 U.S. coal mines have successfully employed this relatively new technology. 

This paper addresses the practical aspects of MRS usage in underground coal mines. 

During this study, nearly one-half of the U.S. mines that have utilized mobiles were visited. This report 

depicts the more common pillar extraction methods that operators have found successful. The "Christmas tree" 

and outside lift methods are described and illustrated. Roof control plans that do not require breaker posts or 

allow pillar extraction with fewer than four mobiles are also examined. In addition, operators' experiences 

with setting pressures, loads, and rates of loading during pillar extraction are addressed. Mining and support 

strategies to more effectively control hillseams, weak roof, and gob overrides are also discussed. 

'Geologist, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 

*president, Mobile Mining Supports, Inc., Peach Creek, WV. 

'Mining engineer, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 

'Vice president, Mobile Mining Supports, Inc., Peach Creek, WV. 
INTRODUCTION 

Mobile roof supports (MRS's) are shield-type support units 

mounted on crawler tracks (figure 1). MRS technology was 

pioneered by the former U.S. Bureau of Mines during the 

1980's [Thompson and Frederick 19861. Commercial units are 

currently manufactured by J. H. Fletcher and Co., Huntington, 

WV, and Voest-Alpine Mining and Tunneling, Pittsburgh, PA. 

Fletcher refers to its units as "Fletcher Mobile Roof Supports" 

(FMRS). Voest-Alpine has designated its units as "Alpine 

Breaker Line Supports" (ABLS). For the purposes of this 

paper, the generic term "MRS," or "mobile," is used to identify 

both manufacturers' units. In 1988, the Donaldson Mine in 

Kanawha County, WV, was the first U.S. operation to use 

mobiles. Since then, approximately 40 U.S. mines in 5 States 

have utilized mobiles. These States include Illinois, Kentucky, 

Pennsylvania, Virginia, and West Virginia. Mobiles have been 

employed in more than 15 different U.S. coalbeds ranging 

from 1.7 to 4 m (5.5 to 13 ft) thick. MRS units are primarily 

used in coalbeds thicker than 2.4 m (8 ft). Currently, there are 

approximately 100 units in use in the United States. Mobiles 

have been used primarily during full and partial pillar recovery 

operations. However, the first longwall face shield recovery 

operation to utilize mobiles instead of walking shields was in 

a southern West Virginia coal mine in 1996. Operators and 

others have also been discussing mobile employment in 

longwall headgates and tailgates. 

MRS's provide improved safety to section personnel 

compared with a conventional timber plan. However, the first 

fatality on an MRS section occurred in 1995 in Mingo County, 

WV. During the fatality investigation, questions arose regarding (1) p~llar extraction methods, (2) setting pressures, 

(3) loads and rates of loading on the machines, and (4) proper 

positioning of mobiles and face personnel. In order to 

determine the current state of the art in MRS usage, researchers 

from the Pittsburgh Research Center visited 20 U.S. mines with 

different geologic and mining conditions. Personnel with 

practical hands-on experience were questioned at each 

operation, and operational advantages and issues were discussed. In addition, all Mine Safety and Health Administration 

(MSHA) approved MRS roof control plans were examined. 

The findings are summarized below. 

Figure 1 .-Full pillar extraction using mobile roof supports. 
SAFETY ADVANTAGES 

MRS's are used in lieu of roadway, turn, and crosscut 

breaker posts during pillar recovery operations. Eliminating the 

setting of these posts enables miners to remain further outby the 

pillar line and reduces their exposure to gob overrides and rib 

spalling. Mobiles are active supports, whereas wooden posts 

are strictly passive. Mobiles also provide better roof coverage. 

Roof coverage depends on the manufacturer; however, the unit 

with the least canopy dimensions provides 3.3 m2 (36 ft2) 

of pressurized roof coverage, compared with less than 0.1 m2 

(1 ft2) for a wood post. The improved stability of mobiles is 

also a major advantage. Wood posts will fail, sometimes with 

little or no warning, at less than 2.5 cm (1 in) of convergence. 

Mobiles will displace approximately 2.5 cm (I in) before 

yielding and have the ability to yield through a few meters of 

displacement without becoming unstable. Mobiles are much 

better suited to handle eccentric load conditions (i.e., horizontal 

and lateral loading), which are common during pillar extraction, 

compared with wood posts, which suffer reduced stability for 

anything but pure axial (vertical) loads [Barczak and Gearhart 

19971. Gob sliding and rib rolls will commonly kick out 

breaker and turn posts. 

MRS units are available in two different support capacities. 

With one exception, all U.S. mines employ mobiles that can 

exert up to approximately 5,338 kN (600 tons) of force against 

the roof. These units each have a load-bearing capacity equivalent to six 20-cm (8-in) diameter hardwood posts [Barczak and 

Gearhart 19971. One deep-cover operation in Virginia is using 

mobiles that have a 7,118-kN (800-ton) support capacity. 

Based on the above, MRS's are superior to wood posts for pillar 

extraction. At every operation visited, personnel expressed the 

opinion that mobile usage enhances pillar line stability and 

safety. 

Mobiles reportedly reduce material handling injuries and free 

personnel for other less strenuous assignments. In a mine 

visited in Mingo County, WV, the operator noticed a significant 

reduction in back injuries with mobile usage. Prior to mobiles, 

scores of posts 3.4 m (1 1 ft) long and weighing 80 kg (175 lb) 

had to be set to recover each pillar. Because of reduced 

roadway clearance, shuttle cars were constantly knocking out 

the posts, which then had to be reset. Three miners were 

required to set each post. One miner had to climb a stepladder 

to drive in the wedges. This miner summarized by saying 

"we're not setting posts, we're planting trees." In another mine 

visited in Mingo County, WV, 105 posts were typically set for 

every pillar mined, compared with only 8 breaker posts when 

mobiles were employed. The mine worked three shifts, and it 

required one miner on each shift to haul in enough timbers to 

keep up with pillar line advancement. This same operator 

reported a reduction in cost of $0.65/t ($0.60/st) over 

conventional timbering and an 18% increase in production. A 

reduction in cost of $2.20/t ($2.00/st) was reported by an 

operator in Boone County, WV. Mobiles also increase the selfesteem of reassigned miners. Miners have traded in their axes 

and bow saws for more modern technology. 

PILLAR EXTRACTION METHODS 

CHRISTMAS TREE METHOD 

The "Christmas tree" method (also called left-right, fishbone, 

or treetopping) is the most commonly used full pillar extraction 

method with MRS's (see figures 2A through D). This method 

is generally employed under deep cover when pillars on 18- or 

24-m (60- or 80-ft) centers are required to maintain necessary 

pillar stability factors. Figures 2A through D depict a common 

sequence in which lifts are extracted during barrier and 

production pillar extraction. As shown in figure 2A, mobile unit 

4 is trammed approximately 2.1 m (7 ft) outby and pressurized 

prior to mining lift 2. Prior to mining lift 3 (figure 2B), unit 3 

is trammed 4.3 m (14 ft) outby and pressurized. This process 

continues until the breakers are set, as indicated in figure 2B. 

After the breakers are set, unit 3 is moved to position F and unit 

4 is trammed to position G. When referring to a particular 

mobile, units 1 through 4 are designated as shown in figure 2B 

by convention. After mining the barrier pillar, the mobiles 

tramming down the entry are referred to as the "No. 1" (left 

side) and "No. 2" (right side) units. Mobiles maneuvered 

through the crosscut are designated as the "No. 3" (pillar line 

side) and "No. 4" (solid pillar side) units. 

The size and shape of the pillar remnants, back wing, and 

pushout stump (figures 2C and D) can vary from pillar to pillar. 

The riskiest process during pillar extraction is pushout removal. 

Some operators routinely try to extract 60% or more of the 

pushout, conditions permitting; others do not attempt to remove 

the pushout. This decision is based on mining conditions, past 

experiences with equipment entrapments, and safety considerations. In general, the more competent the roof, the more 

likely the pushout is removed. Operators who typically remove 

the push will abandon it if the stump shows signs of excessive 
1,w \. 

k-63 

. . . . Barrier ' ' ' ' b pillar 

Figure 2.-Christmas tree extraction method. A, lifts 1-2A. B, lifts 1-8A. C, lifts 1-push; pushout removal with units 1 and 2 in tandem. 

D, lifts 1-push; pushout removal with units 1 and 2 staggered. 

weight. It should be noted that when attacking the pushout, 

most operators situate the mobiles so that the roadway is from 

3.7 to 4.3 m (12 to 14 ft) wide. The clearance is so restrictive that the ripper head has bent mobile canopies. Bit marks 

were also observed on a few canopies. If the push is wider than 

the ripper head, the sump will normally be taken through the 

middle of the push. 

Based on operators' experiences and underground data 

obtained by Hay et al. [1997], the area most prone to roof falls 

during pillar extraction is the intersection. Operators sometimes 

refer to the intersection as the "critical area." During the study 

conducted by Hay et al. [1997], units 1 and 2 were situated in 

the entry just inby the intersection during back wing and 

pushout removal. Significant roof deflection and higher roof 

bolt loads were monitored, compared with an adjacent 

instrumented intersection, where timbers were used to extract a 

pillar. Therefore, Hay et al. [I9971 concluded that units 1 and 

2 should be placed in the intersection as much as possible to 

protect miners and equipment. It is important to note that MRS 

units I and 2 were positioned inby the intersection when the 

1995 fatality occurred in Mingo County, WV (figure 3). The 

fatal accident happened during the mining of the last lift in the 

pillar. To enhance intersection stability, some operators 

position mobiles 1 and 2 in tandem, as shown in figure 2C, prior 
, 

I 

I Victim 

u 

XO. I 1 1- 1 I 

Figure 3.-MSHA drawing of fatal roof fall accident [Vance et 

al. 19951. 

to pushout removal. Other operators stagger units 1 and 2 side 

by side in the intersection as much as possible (see figure 20). 

In figures 2C and D, lift 11 is a 14-m (46-ft) extended cut. In 

certain MSHA districts, the length of an extended cut (measured 

from the rib-side column of bolts) cannot exceed 14.8 m (40 ft). 

A variation of the usual Christmas tree lift sequence is 

illustrated in figure 4, where lift 1 is taken from the back wing 

(also called bottom of the block). Mobile units 3 and 4 are then 

moved to locations 3B and 4B and pressurized. Some operators 

indicated that this positioning enhances intersection stability 

prior to the mining of the left and right wings (lifts 2-5). 

However, some State and Federal roof control specialists 

expressed the concern that the removal of the back wing lift first 

actually reduces intersection stability, which is critical. 

OUTSIDE LIFT METHOD 

The extended-cut outside lift method (figures 5A through C) 

generally has been used under less than 120 m (400 ft) of cover. 

Entry spacings are typically about 15 m (50 ft) with crosscuts 

on 25-37 m (80-120 ft) centers. One complaint from operators 

concerning this method is that the smaller pillars contain less 

coal; therefore, the equipment spends more time moving from 

Figure 4.-Modified Christmas tree method. 

pillar to pillar. As a result, some operators prefer that the pillars 

be as long as possible, with the major constraint being 

ventilation on development. A few operators have oriented the 

pillars in what is sometimes referred to as the "laid-down" 

position. In other words, the long axis of the pillar is 

perpendicular, not parallel, to panel development. This panel 

design increases the number of stoppings and roof bolts 

required on development, but is sometimes chosen because it 

reduces belt move and haulage time. 

Some operators prefer the outside lift method because most 

power cables exit continuous miners in a right rear position. 

Therefore, miner operators tend to position themselves on the 

right side of the continuous miner. During the outside lift pillar 

extraction, the miner operator's positioning gives him or her an 

excellent line of vision. A few operators believe that this 

method is safer than the Christmas tree method because the 

continuous miner operator is less tempted to move further 

forward to observe the mining of the left wing. 

Under weak roof conditions, some operators also prefer the 

outside lift method because the unsupported span (width) of the 

mined-out area is smaller compared with the area opened up 

with the Christmas tree method. In addition, the outside lift 

method provides added protection to the continuous miner 

because a solid coal pillar is nearby. Further support can be 
Scale Ej Scale 

61 Scale 

Figure 5.-Outside lift method. A, lifts 1-6; 6, lifts 1-7A; C, lifts 1-push. 
Figure 6.-Herringbone panel design. 

obtained under weak roof conditions if remnants are left at the 

back end of the lifts. One disadvantage to the outside lift 

method compared with Christmas treeing is that the lift lengths 

are usually longer (deeper). Prolonged exposure while mining 

deeper lifts subjects the continuous miner to greater risk. 

EXTRACTION WITH CONTINUOUS HAULAGE 

The Christmas tree and outside lift methods have been used 

in combination to extract the parallelogram-shaped pillars 

(figure 6) that are developed using continuous haulage. 

Crosscuts are driven on approximately 60" angles to facilitate 

the movement of bridges and carriers. This panel configuration 

has been referred to as the "herringbone" or "turkey foot" 

design. Common entry centers range from 16 to 18 m (52 to 58 

ft), with crosscuts on 25- to 27-m (81- to 90-ft) centers. Barrier 

and production pillars can be extracted as shown in figure 6, or 

pillars can be mined from right to left, then from left to right. 

After the mining of lifts 1 and 1A (figure 7A), unit 4 is trammed 

2.1 m (7 ft) outby and pressurized. Lifts 2 and 2A are then 

removed from pillar 2. These lifts are removed to reduce the 

length of lift 7 shown in figure 7B. Mobile units 3 and 4 are 

positioned and pressurized at locations E and F, respectively, 

during breaker installation (figure 7A). 

If poor ground conditions occur, they normally develop in 

the center or belt entry because the center pillars being mined 

are surrounded by gob on three sides, as shown in figure 7C. 

Figure 7C also displays mobile positioning during pushout 

removal in pillar 7B. A single pushout is usually removed from 

either pillar 7A or 7B. The push is taken from whichever 

fender is least loaded. The other fender (pillar 7A remnant) is 

left intact to function as a breaker during equipment removal. 

If both fenders exhibit severe loading, no pushout is taken. A 

laid-down version of this design has also been tried by one 

operator. A high incidence of roof falls in the belt entry 

prompted this operator to change to three-way belt entry 

intersections (figure 8). Three U.S. operators have employed 

mobiles in conjunction with continuous haulage. In one 

operation, shift production exceeded 5,400 t (6,000 st) during 

barrier pillar slabbing. 

PARTIAL PILLAR RECOVERY 

Mobiles were used during partial pillar recovery operations 

in two mines visited. Managers at one mine chose partial 

pillaring because the shale roof was so weak that they believed 

that full pillar extraction was not feasible. At this operation, 

pillars were initially developed on 30-m (100-ft) centers. On 

retreat, the pillars were "L-slabbed," with 7- to 10-m (20- to 30- 

ft) cuts taken from the entry and crosscut (figure 9). The 

remnant stumps measured approximately 19 by 19 m (62 by 62 

ft). Two continuous miners were used on the section, each 

working with a pair of mobiles. As a row of pillars was 

extracted, one of the miners worked from the crosscut, while the 

other moved from entry to entry. Mine officials were quite 

satisfied with the results obtained from this method. 

In the other mine practicing partiai recovery, a massive 

sandstone roof subjected the mobiles to excessive loading 

during full pillar recovery. At this operation, a 4.9-m (16-ft) 

wide diagonal split was cut through a 9- by 9-m (30- by 30-ft) 

pillar, which left two triangular stumps. Miners call these 

stumps "coal cribs" because they provide enough short-term 

support for the equipment to be moved safely to the next pillar 

before they yielded and crushed out. 
Scale 

0 6 rn 

0 2b f! 

Scale 

0 6 m 

C 0 

- 

20 ft 

Figure 7.-Herringbone design: A, lift sequence for pillars 1 and 2; B, lift sequence for pillars 1-3; C, lift sequence for pillars 7A and 7%. 
re 8.-Three-way inters iection belt entry des iign. 

- c- - --T 

Figure 9.-Partial pillar extraction by slabbing. 
OPERATIONAL ISSUES 

ROOF FALLS 

Roof falls or major rib rolls (side hits) can foul or bury 

mobiles. "Fouling" is a common term used by miners when 

mobiles are covered up and it requires less than 2 hours to 

recover them. Usually a 3.2-cm (1.25-in) high-strength steel 

cable or 2.5-cm (1-in) chain is attached to the mobile, and it is 

pulled out with the continuous miner or another mobile. A few 

operators with weak roof reported that weak "drawrock" 

sometimes fell prematurely and fouled the continuous miner 

while still mining in the lift. To remedy this, two mines that 

were visited employed fenders that separated the lift being 

mined from the gob. At an operation in Pennsylvania, the 

fender was initially fairly substantial, but it was extracted as the 

lift was completed (figure 10). An operation in Kentucky left 

0.6-m (2-ft) fenders between lifts. These fenders were not 

extracted, and they provided enough support to complete the lift 

and remove the continuous miner before the fenders crushed 

out. 

Major roof falls that bury mobiles may necessitate extensive 

cribbing and rebolting to rehabilitate the area, drilling and 

shooting, and a miner retriever (crab). A common mistake is to 

lower the canopy of the buried mobile in an attempt to tram it 

out from under the rock. This practice usually aggravates the 

problem, and mine operators have suggested that one or two 

cribs should be set along each side of the mobile before the 

canopy is lowered. Major roof falls under massive sandstone 

roof rock, especially during first cave conditions, can subject 

the mobiles to impact (shock) loads. This shock loading can 

cause hydraulic cylinders to swell, mushroom, or even bend. 

This damage occurs when the rock burst valves are unable to 

release sufficient hydraulic fluid to prevent the excessive 

buildup of hydraulic pressure. If sufficient fluid is released, the 

mobile becomes inoperative. Shock loading has also sheared 

lemniscate pins. 

Numerous operators mentioned problems associated with 

hillseams (also called mountain cracks and surface breaks). 

Hillseams usually occur under shallow cover in sandstone roof 

rock. Parallel and intersecting hillseams can segment the roof 

into huge isolated blocks. Massive roof falls can occur when 

:he coal pil!ars supporting these blocks are extracted. Several 

machines have yielded, been buried, or been severely damaged 

because of hillseams. Mining strategies under these conditions 

include leaving sufficiently sized pillar remnants to support the 

roof. Also, because cover is normally shallow, operators room 

out on 12- by 12-m (40- by 40-ft) centers near the outcrop and 

do not recover the pillars. One mobile operation in eastern 

Kentucky was mining under a massive sandstone main roof that 

would not break. First attempts at retreat mining were 

terminated due to squeeze conditions. Pillaring plans, which 

factored in hillseam locations, were later developed. As 

reported by Unrug et a]. [1991], controlled systematic caving 

was achieved. 

OPERATOR TRAINING 

All of the coal mine operators visited indicated a need for 

practical, hands-on training. Training is required around the 

clock and can take from 2 to 6 weeks depending on the crew 

and conditions. One of the most common mistakes during 

training is mobile advancement. Many operators reported that 

mobiles best work in pairs, thereby providing protection to one 

Figure 10.-Outside lift pillar extraction using temporary support fenders. 
another. When moving mobiles during the mining of a pillar, 

it has been recommended that the machine be advanced no 

more than one-half the length of a canopy, or approximately 2.1 

m (7 ft) from the canopy tip of the adjacent machine, and then 

pressurized. The second mobile, which is now protected by the 

first, can then be lowered and trammed forward. Mobile 

operators tend to advance one mobile too far and not leapfrog 

them, which has caused mobiles to be buried. Also, the canopy 

should not be lowered more than necessary to clear roof 

obstructions during advancement. 

OPERATOR POSITIONING 

After section personnel become familiar and comfortable 

with mobiles, it is not uncommon to hear miners say that "a 

mobile can pick up the whole mountain." It is important that 

miners do not position themselves in potentially hazardous 

areas because of this overconfidence and false sense of security. 

Leaning against the mobiles or standing in the intersection 

during the later stages of pillar recovery has been discouraged. 

During the final stages of recovery, some operators have made 

it a company policy that mobile and other nonessential section 

personnel be outby the intersection, in the entry of the next 

pillar row to be mined. At one mine visited, the superintendent 

remarked that "all it took was one gob override outby the 

canopies to convince the miners to stay out of the intersection." 

PRESSURE GAUGES 

Two pressure gauges are mounted on each mobile (figure 1). 

These gauges are visually monitored by operating personnel to 

determine loads and rates of loading on the units. Since the first 

fatality on an MRS section, considerable emphasis has been 

placed on the size of gauges. Some m.obile operators have 

mentioned that it is necessary to stand close to the mobiles to 

read standard gauges. Other operators expressed that even the 

10-cm (4-111) diameter gauges are of adequate size and can be 

read easily from 9 m (30 ft) away. Gauges are constantly 

monitored to determine loads and load development rates on the 

mobiles. Common sense modifications, which have made 

gauges easier to read, include the mounting of flood lights to 

illuminate the gauges and/or the adherence of reflective tape on 

the gauges' glass covers to mark a critical load threshold. At 

one operation, the positioning of the reflective tape depends on 

the roof rock type being mined under. The Spokane Research 

Center, MRS manufacturers, and others are currently working 

to develop lighting systems that can be seen from farther 

distances. Green, yellow, and red pulsating lights will indicate 

different total load levels or stages of loading. Additionally, a 

light-emitting diode bar graph will indicate rates of loading on 

individual units. Prototype lighting systems will initially be 

placed on the MRS units during the underground testing phase. 

MACHINE RELIABILITY 

During this study, operators indicated that mobiles were very 

reliable. Most operators cited cut power cables due to roof 

spalling or rib rolls as the most frequent cause of downtime. 

When mining in thicker reserves, high ribs are usually more 

hazardous and troublesome than the roof. At one mine visited, 

a 3-m (10-ft) high rib rolled onto the continuous miner's 

deadman switch while the miner was deep in a lift. While 

recovery efforts were underway, the roof deteriorated and a 

major roof fall occurred. Also, cable handling and hanging is 

more cumbersome in high coal, and concern has been expressed 

regarding the mobile operator's positioning when cables are 

detached from the mine roof after the lift has been taken. 

Breakaway cable-holding devices have recently become 

available. Dialog has begun with equipment manufacturers on 

permissible battery-powered mobiles, which has greatly 

interested mine operators. 

RECOVERY AND PRODUCTION CONCERNS 

Mobile usage has enabled a few operators to retreat mine 

reserves that could not be mined previously due to poor ground 

conditions. A mine operator in Boone County, WV, mentioned 

that previous attempts to retreat mine using a timber plan had 

failed because of poor roof conditions. He stated that caving 

would occur so quickly that it endangered the miners. Attempts 

to pillar using mobiles have proven successful in controlling 

premature caving. Another operator noted that past pillaring 

attempts with a conventional timber plan had failed because of 

weak floor conditions. The timbers punched into the floor, and 

the section would go on a squeeze. These same reserves were 

later successfully pillared using mobiles, which exert less 

ground-bearing pressure than wood posts because the crawler 

tracks distribute the load to the mine floor more uniformly. 

Coal recovery rates as high as 85% to 95% have been reported 

when using mobiles. 

Mobile usage also assists operators in meeting production 

goals. Contrary to popular opinion, most of the mines contacted 

reported that shift production is usually higher on advance than 

during retreat when timbers are used. This is especially true 

when supersections are used on development. One major coal 

producer estimated that during retreat mining with timber 

supports, 20% less coal is mined per shift than during panel 

development. Some of the production decrease is attributable 

to downtime while setting timbers, practicing caution, and 

waiting for the roof to cave. Most of the operators indicated 
that when mobiles are employed, shift production can improve reported a slight production decrease when they switched to 

to approximately 90% to 115% of what it is during panel mobiles. This occurred because the pillars are typically smaller 

development. Decreased production during retreating opera- and, therefore, contain less coal, and considerable time is lost 

tions has prompted some operators to run supersections with tramming the units. These operators have continued mobile 

mobiles or continuous haulage in conjunction with mobiles. A usage, however, because of their safety advantages. 

few operators using the extended-cut outside lift method have 

PROCEDURAL VARIATIONS 

Procedures governing MRS usage are addressed in each 

mine's roof control plan. In some MSHA districts, if one 

mobile becomes inoperative, the section is down until the unit 

is again functional. In these districts, the mine operator should 

consider having an approved conventional timbering plan so 

that mining can continue. At one mine visited, a faulty solenoid 

on a mobile idled the pillar line. The section squeezed and all 

the mobiles were entrapped. In one district, pillaring can 

continue with three mobiles; however, mining of the back wing 

and pushout are prohibited. Eight posts or two cribs are set in 

lieu of the No. 4 machine. Mobile No. 3 and the posts or cribs 

are set just outside of the intersection in the crosscut prior to 

mining the left or right wing. The district's position is that this 

scenario is safer than setting 3-m (10-ft) timbers, and operators 

have applauded this decision. In four MSHA districts, mines 

have approved plans to use only two mobiles in conjunction 

with timbers. Figure 11 shows mobile and post positioning 

during back wing removal for an outside lift plan. At least one 

approved combination mobile/timber plan has been approved 

for the Christmas tree pillar extraction method. Mobile units 1 

and 2 are positioned in the same locations, as shown in 

figure 1 I. Two rows of posts on 1.2-m (4-ft) centers are also 

set so that the roadway width into the push is 5 m (16 ft) or less. 

In two districts, plans have been approved that do not require 

entry breaker posts. To prevent miners from wandering into the 

gob, one of these districts requires a recoverable, permissible, 

battery-powered pulsating light to be mounted on a tripod at eye 

level midway down the entry. The other district requires that at 

least eight roof bolts on 1.2-m (4-ft) centers be installed in lieu 

of breakers. These bolts must be at least 0.3 m (1 ft) longer 

than those installed immediately outby and anchored at least 0.3 

m (1 ft) into competent strata. In addition, access to the 

or barrier ribbons. Reflectorized warning streamers, at least 5 

crn (2 in) wide by 0.9 m (3 ft) long, are suspended from the roof 

at or near the restrictive devices. The other districts require the 

setting of eight entry breaker posts. These breakers are usually 

knocked out with the mobiles prior to mining the first lifts. 

Most of the operators expressed dissatisfaction with the 

requirement that entry breaker posts be installed when using 

mobiles. They insisted that breakers served no function, 

especially in competent roof, and that setting breakers subjected 

mine personnel to unnecessary risk. Two operators who 

believed that breakers were helpful were mining under weak 

roof. Their decision was based on previous experiences with 

the gob overriding the breakers, which necessitated the 

abandonment of lifts. 

gobbed-out area must be restricted at the outby end of the Figure 11 .--Combination mobilehimber plan. unrnined pillars by devices, such as chain link fencing, roping, 
SETTING PRESSLIRES 

The active loading capabilities of an MRS can be a 

significant advantage if properly used. Ideally, the support 

should be set against the roof with just enough force to close 

any gaps within the immediate roof structure. Excessive force 

has damaged the bolted horizon and promoted roof beam failure 

outby the canopy tip. The Spokane Research Center conducted 

an extensive field study monitoring MRS and roof bolt loading 

during pillar extraction [Hay et al. 19971. Results indicated that 

adjacent roof bolt loads decreased when the mobiles were 

pressurized, which was expected. However, when the canopies 

were lowered, load was transferred to the bolts, and load 

increases up to 1 1 kN (2,500 lbf) were recorded. This increase 

in bolt load did not cause the bolts to yield in this particular 

mine, but the possibility exists, especially on a pillar line. 

Hydraulic setting pressures on the instrumented mobiles were 

normally 10 MPa (1,450 psi). 

Setting pressures ranged from 6.9 to 30.3 MPa (1,000 to 

4,400 psi) at the mines visited. The support force applied to 

the roof by an MRS is the product of the hydraulic pressure 

multiplied by the leg cylinder area. For example, an FMRS 

with a bore (inside cylinder) diameter of 25.4 cm (10 in) will 

exert a force of 2,803 kN (3 15 tons) with a setting pressure of 

13.8 MPa (2,000 psi). An ABLS with a bore diameter of 21.8 

cm (8.6 in) will require 18.8 MPa (2,730 psi) to provide the 

same setting force. If the same hydraulic setting pressure is 

used, the support force is equivalent regardless of the ABLS 

model or whether the hydraulic cylinders are two- or three-stage 

[Hatch 19961. Regarding the FMRS units, the set force is 

different for two- versus three-stage hydraulic cylinders given 

the same set pressure [Howe 19961. 

Some mines have found that weak shale top can he damaged 

by high setting forces. Hairline cracks can develop, mechanical 

bolts can "pop" and fail, and loose rock can become dislodged. 

At one mine, an MRS operator was injured by spalling shale 

roof while pressurizing the canopy. The operator was standing 

close to the machine, and company policy now mandates that 

MRS operators stand back when the canopy is being set. In 

addition, the setting pressures were reduced from 13.8 MPa 

(2,000 psi) to 10.3 MPa (1,500 psi), and no problems have 

occurred since. It has been recommended that mobile operators 

periodically check the setting pressure. Pressure bleed-off can 

occur if the floor is soft or if there is excessive floor gob. If 

bleed-off occurs, the MRS should be repressurized. 

LOADS AND LOADING RATES 

Operators cited several factors and conditions that have 

imposed increased loads on mobiles. In two operations visited, 

mining had been previously conducted in an upper seam. 

Excessive machine loading occurred while mining under 

remnant barrier pillars. Another reported source of mobile 

loading was underdesigned pillars, which had low stability 

factors. During retreat mining, pillars are subjected to 

development, front abutment, and sometimes side abutment 

loading. If pillars are to be mined safely and efficiently, all 

anticipated loading conditions should be considered during the 

design process [Mark and Chase 1997). MRS loading can also 

be induced by excessive floor heave. 

In general, mobiles operating under weaker shale roof rock 

have experienced lower loads. Pressure increases after the 

canopy was set were in the range of 3.4 to 4.1 MPa (500 to 600 

psi) and, at this point, the roof normally breaks. Conversely, 

numerous mobiles operating under massive sandstone roof, 

which tends to cantilever, have yielded. At one mine visited, 

mobiles have been used under both sandstone and shale roof 

rock. The operator stated that, on average, load pressures are 

1.4-2.1 MPa (200-300 psi) higher under sandstone roof rock, 

other factors being equal. 

During the mining of a pillar, most of the operators indicated 

no significant load development on the No. 1 and No. 2 units 

during the mining of the first few lifts. Hay et al. [I9971 

monitored increased loads on units 1 and 2 as mining 

progressed during a lift and as the pillar size was reduced during 

successive lifts. Operators experienced maximum loading 

during pushout removal, with pressure increases ranging from 

3 to 3 1 MPa (500 to 4,500 psi). The No. 3 unit sustained the 

greatest loads and suffered the most damage because of side 

hits. The No. 4 unit, which is normally situated along an 

unmined pillar, experienced the least loads and damage. 

Periodically, some operators switch the No. 3 and No. 4 units. 

During the mining of the first rows of pillars in a panel, prior to 

establishing a first cave, most operators reported no significant 

mobile load development. However, a few operators noted load 

increases of 2-7 MPa (300-1,000 psi) higher than normal. In 

order to establish a first cave and minimize outby loading, 

pillars at these operations are mined as completely as possible, 
which included pushout removal. Some of these operators 

discontinue pushout removal after the first cave. 

Technical specialists and operators agree that the critical 

factor is not the amount of load, but rather the loading rate. 

Loading rates and magnitudes varied from mine to mine. Most 

operators commented that the decision to remove section 

personnel and equipment based on steady, rapid loading 

conditions is a judgment call based on past experiences. One 

operator reported that it was time to back out the continuous 

miner, shut down, and listen when 1.4-2.1 MPa (200-300 psi) 

pressure increases occurred every 5 min over a 15-min interval. 

Another operator indicated that if the pressure jumps 1.4 MPa 

(200 psi) once, and then again, it is time to lower the canopies 

and tram out because a fall normally always occurs. Other 

operators cited instances during pushout removal when, in a 

matter of seconds, 21-28 MPa (3,000-4,000 psi) pressure 

increases occurred. The continuous miner and personnel were 

removed, and a fall never occurred. The mobiles later inched 

their way out. During the mobile field monitoring investigation 

conducted by Hay et al. [1997], no significant MRS load 

increase was detected when the canopy of the adjacent mobile 

was lowered. In other mines, operators noticed immediate 

increases in the 2-14 MPa (300-2,000 psi) range. Significant 

load transfer to the adjacent machine can cause it to yield. 

Whether or not load is transferred is a function of roof geology 

and the pillars' stiffness. Some operators reported that the 

pressure gauges seldom showed any change during the mining 

cycle, and a few operators even observed pressure decreases. 

No pressure change implies that the force exerted by the roof 

has not exceeded the setting force applied by the MRS. A 

decrease in MRS pressure might occur if the machine settled 

into the floor or roof, or if there was a leak in the hydraulic 

system. 

When mobiles are subjected to consistent load increases, 

operators practice one of two strategies. Some operators 

remove section personnel, the continuous miner, and lastly the 

mobiles prior to the cave. Other operators leave the mobiles in 

place until after the fall to breaker it off. These operators 

indicated that when the mobiles start creaking and taking 

weight, they are performing their function. Even when the yield 

valves open and the hydraulic fluid is spewing out, these 

operators insist that it is still not time to lower a canopy. In fact, 

one of the advantages of an MRS is that it continues to maintain 

a high support load even though it is yielding. Some operators 

reported that lowering the canopy removes the support from the 

roof and allows the roof cave to override the MRS's. 

By replacing roadway, turn, and crosscut breaker posts, MRS 

usage enhances the safety of section personnel by (1) providing 

a more effective ground support, (2) reducing worker exposure 

near the gob edge, and (3) eliminating a major cause of material 

handling injuries. Compared with wood posts, mobiles provide 

better roof coverage and improved stability. To achieve the full 

advantages of mobiles, they must be employed properly. 

Training and careful attention to standard operating procedures 

are essentiai. Throiigh ihe experience at different mines, 

strategies have been developed to cope with the hazards posed 

by hillseams, weak roof, and first falls. 

During the later stages of pillar recovery, the area most prone 

to a roof fall is the intersection. Placement of mobile units 1 

and 2 in the intersection during pushout removal enhances roof 

stability. Positioning of nonessential personnel outby the 

intersection better ensures their safety. 

Mobiles have been employed during full and partial pillar 

recovery operations. Most pillars are recovered using either the 

Christmas tree or outside lift method. In some MSHA districts, 

roof control plans have been approved that do not require 

breaker posts. A few districts have approved plans that allow 

pillar extraction with fewer than four mobiles. 

Setting pressures in the mines visited ranged from 6.9 to 30.3 

MPa (1,000 to 4,400 psi). The most common setting pressure 

fcr mcbiles is 10 MPa (1,500 psi). In some instances, higher 

setting pressures have damaged weak shale roof rock. 

Operators reported that maximum loading occurred during 

pushout removal. MRS loading is typically higher under 

sandstone roof rock than shale. 
REFERENCES 

Barczak TM, Gearhart DF [1997]. Full-scale performance evaluation of 

mobile roof supports. In: Mark C, Tuchman RJ, comp. Proceedings: New 

Technology for Ground Control in Retreat Mining. Pittsburgh, PA: U.S. 

Department of Health and Human Services. Public Health Service, Centers for 

Diwe Control, National Institute for Occupational Safety and Health. IC 9446. 

Hatch JW [19%]. Personal communications between J. W. Hatch, Product 

Manager, Voest-Alpine Mining and Tunneling, Pittsburgh, PA, and F. E. Chase, 

Pittsburgh Research Center, National Institute for Occupational Safety and 

Health, Centers for Disease Control, Public Health Service, U.S. Department of 

Health and Human Services. 

Hay KE. S~gner SP, King ME. Owens JK i1997j. Monitoring mobile roof 

supports. In: Mark C, Tuchman RJ, comp. Proceedings: New Technology for 

Ground Control in Retreat Mining. Pittsburgh, PA: U.S. Department of Health 

and Human Services, Public Health Service, Centers for Disease Control, 

National Institute for Occupational Safety and Health, IC 9446. 

Howe LC [1996]. Personal communications between L. C. Howe, Product 

Manager, J. H. Fletcher and Co., Huntington, WV, and F. E. Chase, Pittsburgh 

Research Center, National Institute for Occupational Safety and Health, 

Centers for Disease Control, Public Health Service, U.S. Department of Health 

and Human Services. 

Mark C, Chase FE [1997]. Analysis of retreat mining pillar stability 

(ARMPS). In: Mark C, Tuchrnan RJ, comp. Proceedings: New Technology for 

Ground Control in Retreat Mining. Pittsburgh, PA: U.S. Department of Health 

and Human Services, Public Health Service. Centers for Disease Control, 

National Institute for Occupational Safety and Health, IC 9446. 

Thompson RT, Frederick JR [1986]. Design and field testing of a mobile 

roof support for retreat mining. In: Proceedings of the 5th International 

Conference on Ground Control in Mining. Morgantown, WV: West Virginia 

University, pp. 73-79. 

Unmg KF, Tussey IJ, Moore R [1991]. Using mobile roof supports for 

pillar extraction at Martin County Coal. Min Eng Ocr:1215-1218. 

Vance C Jr., Cybulski JA. Gray WJ [1995]. Report of investigation 

(underground coal mine): fatal roof-fall accident, Big Branch Mine (ID No. 46- 

05978). Eastern Mingo Coal Company, Naugatuck, Mingo County. West 

Virginia. Mount Hope, WV: U.S. Depamnent of Labor, Mine Safety and Health 

Administration. District 4. 
MONITORING MOBILE ROOF SUPPORTS 

By Kenneth E. Hay,' Stephen P. Signer,2 Michael E. King,3 and John K. Owens4 

ABSTRACT 

Researchers from the Spokane Research Center conducted a field study to assess the safety of remotely 

controlled mobile roof supports (MRS's) in a retreat pillar mining operation. Data were collected to provide 

the Mine Safety and Health Administration with criteria needed to develop guidelines for MRS use and to 

determine if precursors could be identified that would alert miners to imminent roof falls. 

Two test sites at which two different support methods-MRS's and posts-were used were monitored to 

obtain information on entry stability. Pressure transducers and string potentiometers were installed on all four 

MRS's to obtain loading and displacement information. Roof bolt load cells, sonic probes, extensometers, and 

survey targets were installed in the surrounding entries to obtain information on ground behavior. 

Results showed a larger increase in roof bolt loading and roof movement when MRS's were used, especially 

in the intersection area. Roof bolt loads in the entries showed decreases when the MRS's were set and increases 

of up to 1 1.1 kN (2,500 lbf) when the MRS's were unloaded. Unloading of one MRS in a pair did not 

significantly increase load on the other. MRS's 1 and 2 usually had the higher loads; these loads increased as 

the pillars on each side were being mined. MRS 3 normally had lower loads than 1 and 2; however, it also 

experienced some very high loads when in the last position near the pushout. MRS 4 usually had the lowest 

loads, primarily because it was located near the solid pillar that was not being mined. 

'Supervisory civil engineer. 

'Mining engineer. 

'Electronics technician 

'Mechanical engineer. 

Spokane Research Center, National Institute for Occupational Safety and Health, Spokane, WA. 
Mobile roof supports (MRS's) (figure l), also known as 

breaker line supports, were developed by the former U.S. 

Bureau of Mines (USBM) in the 1980's [Thompson and 

Frederick 19861. Currently, 2 mining equipment companies 

manufacture commercial units, and 21 units (4 machines per 

unit) are in operation in U.S. coal mines. MRS's have been in 

use in Australia since 1987 [Follington et al. 19921, where 

they replace posts during full or partial pillar extraction. 

MRS's were developed to reduce the high number of 

injuries and fatalities in retreat pillar mining operations [Chase 

and Mark 19931. Most of these accidents occur during the 

preparation and installation of the required turn and breaker 

posts used in retreat sections. 

The Mine Safety and Health Administration (MSHA) is 

assessing safety associated with the use of MRS's and has 

asked us for assistance in establishing guidelines for their use. 

There is also a need for a method to detect imminent roof 

failure. Previous methods of using loading on posts and other 

indicators are no longer applicable when using MRS's. 

Use of MRS's requires an understanding of how an active 

roof support behaves. To further this understanding, the 

Spokane Research Center conducted a field test in which instruments were installed on four MRS's and in entries 

surrounding two pillars. One pillar was mined using posts; the 

other, using an MRS as the support. These instruments included load cells on point-anchor bolts and fully grouted bolts 

to measure loading and unloading of roof bolt supports, sonic 

probes to measure displacement in the roof to a depth of 6 m 

(20 ft), and extensometers and survey targets to measure roof 

sag. Instruments installed on the MRS's included pressure 

transducers on all hydraulic lines to measure load on each of 

the support cylinders and string potentiometers to measure leg 

closure. 

MINE LAYOUT 

The field test was conducted at a mine in southwestern directly above the Lower Cedar Grove Seam, and the HernWest Virginia, where the Lower Cedar Grove Coalbed is be- shaw Seam lies 91 m (300 ft) above. The depth of cover varing mined. This seam is approximately 2.8 m (10 ft) thick. ies from approximately 91 m (300 ft) near the outcrop to over 

The Upper Cedar Grove Seam lies 24 to 30 m (80 to 100 ft) 305 m (1,000 ft). 

Figure 1 .-Mobile roof supports in operation. 
The pillars for the retreat section were developed on 21- by 

27-m (70- by 90-ft) centers and 6-m (204) wide entries. Normal roof support consisted of 1.2-m (4-ft) mechanically anchored roof bolts on' 1.2-rn (4-ft) centers. Breaker posts were 

still required between the gob and the pillar line. 

The test site was located in the second panel being mined 

with MRS's (figure 2). Each row in the panel had seven or eight 

pillars, but the number varied depending on whether or not the 

banier pillar between the second and first panel had been mined. 

The test pillars were the third ones from the end of the pillar 

rows. 

The immediate roof is a dark gray shale that contains 

numerous plant fossils. Its high clay content results in a soft, 

moisture-sensitive rock with poorly bonded bedding planes. 

The shale grades upward into a moderately hard, strong 

sandstone located 1 to 2 m (4 to 6 ft) above the opening. 

The pillar recovery method used with the MRS's was 

"Christmas tree" extraction (figure 3A). With posts only, the 

pillar recovery method was pocket-and-wing (figure 3B). 

Figure 2.- Retreat panel layout. 
B 

KEY 

Breaker post . 

. nn 

1 Cut sequence 

Post with cap . • 

. 

• 

rn Entries, rooms, and crosscuts are . 

rn supported during development . 

Figure 3.-Mining plan using (A) mobile roof supports and (B) posts. 
MOBILE ROOF SUPPORT INSTRUMENTS 

HYDRAULICPRESSURETRANSDUCERS 

Pressure transducers were installed on the hydraulic lines to 

obtain information on the loading patterns on each of the four 

support cylinders, as well as on total loading of each MRS. 

Displacement devices were also installed to measure leg 

closure. Results showed that the front and rear hydraulic legs 

were not consistently set to the same loads, but that similar 

setting loads occurred in the. left and right cylinders of the front 

and rear legs, which were plumbed together. A dial indicator 

indicated loading on the front pair and the rear pair of leg 

cylinders, so that the operator could set the desired load for each 

pair of cylinders. 

During pillar mining in the test area, the hydraulic pressures 

were normally set at 10,000 kPa (1,450 lbf). Shortly after 

setting each MRS, the load would decrease on each leg. In one 

case, on MRS 2, it decreased by approximately 80% of the load 

(figure 4). This decrease could be attributed to bleed-off of 

hydraulic pressure, soft roof and/or floor conditions, loose 

debris (coal and rock) on the floor, transfer of load because of 

mining, or any combination of these conditions. The data show 

that the decrease in load became less as the MRS was advanced 

in the entry. Thus, it would be expected that the load decrease 

resulted from debris on the floor, because that was the only 

condition that changed. 

I I 

MRS 3 

- - 

- 

0 0 

0 1 2 3 

TIME, hours 

I I 

MRS 4 

- 

KEY - - Leg 1 - --- Leg 2 

---- Leg3 -. - . . -. - .I .I - 

,- A Fy\ocatio: a= 

0 

0 1 2 3 

270 TIME, hours 

Figure 4.-Loading patterns on mobile roof supports. 
The highest loads were attained in the entry with MRS's 1 

and 2. Typically, loads on MRS's 1 and 2 increased as mining 

progressed during a cut and as the pillar size was reduced 

during successive cuts. The data show that all legs on MRS 1 

reached their maximum load at location D in the entry (figure 4). The highest loads were on cylinders 1 and 3, which indicated that loading was greater on the pillar side on the left. 

On MRS 2, the highest loading occurred at location D in the 

entry. The highest loads were on cylinders 2 and 4, which were 

closest to the pillar side on the right. There was no significant 

increase in loads on MRS's 3 and 4 until the first end cut (6) 

was being made (figure 3B). The maximum loads on MRS's 3 

and 4 were almost the same at both the first and second locations. However, on one occasion, while several other pillars 

were being mined, MRS 3 exceeded the yield load. The instruments were not collecting data at this time. On MRS 3, the 

highest loads were on cylinders 1 and 3, which were closest to 

the pillar being mined. MRS 4 had the lowest overall load of 

the four MRS's. 

To determine if unloading of one MRS increased loading on 

the MRS next to it, loads were compared when MRS 1 was first 

moved. The unloading of MRS 1 did not increase the load on 

MRS 2 (figure 5). In evaluating data during other moves of 

MRS pairs, there were no significant increases in load on one 

MRS when another was unloaded. 

I I I I I 

KEY - 

- - - MRSI - - - 

4 

MRS2 - 

- 

_II-I-.-- --LC - 

4 

Unloading Setting- - - 

% - I I I I I 

Data were compiled on the MRS loads for a consecutive 7- 

day period following the first test to identify maximum loads 

over an extended time (figure 6). Overall, loading was similar 

to that found at the first test site. MRS's 1 and 2 had higher 

maximum total loads than MRS's 3 and 4. However, average 

total loads on MRS 3 were higher than those on MRS's 1 and 2. 

MRS 4 continued to take less load than the others. 

Monitoring of the MRS's continued during most of the panel 

mining. One month after the first test, another pillar was observed during extraction and was documented in detail to provide information for an analysis of the MRS data. Loading on 

each of the MRS's was similar to that at the first test site. 

0123456701234567 

TIME, days 

200 

5 150 

n 

g 100 

V) 

50 

0 

Figure 6.-Total loads on mobile roof supports. 

45 

40 , 

35 mg 

30 ? 

25 0- 

20 3 

l5 , 

10 5 

5 

0 

0 4 8 12 16 20 24 

TIME, min 

Figure 5.-Effect of MRS 1 unloading on MRS 2. 
ROOF BOLT LOAD CELLS 

Load cells were installed on 1.2-m (4-ft) long, point-anchor 

roof bolts at 10 locations to determine the impact of MRS setting loads on the roof bolts. These bolts were installed approximately 1 month after normal bolting had been completed. The 

roof bolts with load cells were centered between existing bolts. 

Installation tension varied from approximately 25 to 53 kN 

(5,700 to 12,000 Ibf). 

These roof bolts were installed in the center of the entry and 

located between the MRS's at the first MRS set location. 

During setting of the MRS, the load decreased from about 1.8 

to 0.44 kN (400 to 100 lbf) (figure 7). The amount of decrease 

depended on the position of the MRS's with respect to the 

position of the roof bolt, the setting load of the MRS at that location, and roof conditions. 

During unloading of MRS's 1 and 2 during the first and 

second moves, load on the roof bolts increased from approximately 1.8 to 3.6 kN (400 to 800 lbf) (figure 8). Similar results 

were observed when MRS's 3 and 4 were unloaded. This indicates that during unloading, this amount of load was being 

transferred to the roof bolts. (Note that the monitored roof bolts 

were installed after the normal bolting cycle; however, it is 

likely that similar unloading and loading conditions existed on 

the other roof bolts.) These results confirm that there should be 

optimum MRS setting loads established that will have the least 

effect on the existing support and roof strata, yet will still 

provide adequate support. 

The roof bolt load cells showed a significant increase in load 

when the last lifts were being mined at each test site, from 36 to 

48 kN (8,000 to 10,500 lbf) at the MRS site and from 34 to 

43 kN (7,500 to 9,500 Ibf) at the post site. This increase in load 

was relative to the load increase on the resin-grouted bolts and 

roof movement. 

ROOF-TO-FLOOR CLOSURE DEVICES 

(SAGMETERS) 

Roof-to-floor closure devices (sagmeters) were installed near 

the load cells. The purpose was to monitor roof movement and 

correlate roof-to-floor movement to displacement of the MRS's, 

as well as to monitor the roof after the MRS's advanced in each 

of the entries. These readings were then compared with roof 

movement in the same area when posts had been used. 

Results showed a significant amount of closure beginning 

just prior to moving MRS's 1 and 2 (figure 8) and continuing at 

a high rate throughout the mining of the pillar. Also, when 

MRS 1 was moved for the first time, the sagmeters did not show 

any significant change. This is consistent with figure 5, which 

shows that the unloading of MRS 1 did not have any effect on 

the loading of MRS 2. At MRS's 3 and 4, roof response was 

similar. At the post site, the sagmeters showed a significant rate 

of movement as the last lift in the first wing was being mined 

and when mining in the last wing was started. 

- KEY 

MRS 1 leg 2 --- MRS 2 leg 2 ---- Load cell 1 -------- Load cell 2 

250 1 I I 

0 0 

1 1.25 1.5 1.75 

TIME, hours 

Figure 7.-Loading of MRS 1, MRS 2, and roof bolts for (A) 

entire pillar and (B) during one cycle. 

MRS 1 

MRS 2 

KEY 

leg2 ---- 

leg 1 ----.--- 

Sagmeter 

Sagmeter 

I 

0 

P 

0 1 2 O b 

3 0 

TIME, hours [r 

Figure 8.-Loading of MRS 1 and MRS 2, and roof-to-floor 

closure for entire pillar. 

FULL-COLUMN INSTRUMENTED BOLTS 

Twelve strain gauges were placed on each of two 1.8-m 

(6-ft) long, full-column, resin-grouted bolts. One was installed 

in the center of each of the test site intersections [Signer et al. 

19931. The results showed that strain at both the MRS site and 

the post site increased rapidly during mining of the last lift (13) 

and the last wing (lo), respectively. Those gauges positioned 
135 cm (54 in) above the roof line had exceeded the yield 

strain. When mining was completed at the test sites, 75% of the 

strain gauges had exceeded 66% of the yield strain at the MRS 

site; at the post site, 25% of the strain gauges had exceeded 66% 

of the yield strain. 

SONIC PROBES 

Sonic probes were installed to measure roof displacement at 

increments of 0.3 m (1 ft) to a depth of 6 m (20 ft) above the 

roof line. One sonic probe was installed in the center of the intersection at each test site. significantly higher movement was 

found in the intersection where MRS's were used compared 

with the intersection where posts were used. Table 1 compares 

roof movements measured by the sonic probes immediately 

after the crew completed mining of the pillar and was moving 

out to the next pillar. There was approximately 4.4 times the 

amount of movement in the MRS test sitk, which is consistent 

with the higher loads recorded on the roof bolts in the intersection area. 

Table 1 .-Intersection roof displacement 

Movement of 

roof line with Maximum Location of 

Test site probe end maximum separation, 

as reference, mm (in) separation, 

mm (in) mm (in) . . . . . . . . . 

Mobile roof 

support . . . . 12.7 (0.500) 3.3 (0.13) 61 0 (24) 

Post . . . . . . . 2.9 (0.1 14) 0.76 (0.03) 91 4 (36) 

SURVEY TARGETS 

Survey targets were installed on seven roof bolts and monitored to determine the amount of roof sag in the intersection 

area at both test sites. From a distance of about 30 rn (100 ft), 

the targets were surveyed with a transit and monitored from the 

time mining of the pillar was started until mining was completed and miners had left the area. There was only a slight 

difference between the amount of roof sag at the two sites. The 

MRS test site showed a slightly higher amount of total roof sag. 

SllMMARY AND CONCLUSIONS 

1. The location of an MRS is most important during 

retreat pillar mining. The MRS should be moved as often as 

possible and kept as close as possible to the continuous miner 

to reduce the possibility of premature roof caving. It is especially important that the MRS be in the intersection as much as 

possible to protect miners and equipment, because the results of 

this test show that the most roof-fall-prone area is the intersection. 

In this test, the MRS's were not positioned in the intersection. 

2. Proper setting loads need to be established to match 

geologic conditions. It was shown that setting loads were appropriate for the conditions encountered during this test. This 

conclusion is based on the finding that loads on the roof bolts 

decreased as the MRS's were set, yet the bolts were still able to 

support increases in load significantly when the MRS's were 

released. Other indicators of proper setting loads were that the 

roof did not break above or around the MRS's and that caving 

patterns remained the same. 

3. There were higher loads on the roof bolts in the 

intersection area and more roof displacement when MRS's were 

used. This can be attributed to one or both of the following 

conditions: 

(a) Type and location of support: In the MRS test site, 

MRS's were not positioned in the intersection. In the post test 

site, posts were installed and left in the intersection. 

(b) Method of mining: The Christmas tree method was 

used at the MRS test site, whereas the pocket-and-wing 

method was used with posts. This may have contributed 

to some of the difference in roof support loading and 

displacement. 

As stated in item 1 above, the MRS's should be placed into 

the intersection as much as possible to safeguard miners and 

equipment. 

4. The load loss that occurred immediately after setting the 

MRS's to the roof can be caused by creep or leaks in the MRS's 

hydraulic system, soft roof and/or floor conditions, loose coal and 

rock on the floor, setting pressures that are too high (which causes 

creep in the roof and floor), or load transfer as a result of mining. 

A test of the MRS determined that creep was not a significant 

factor in the MRS system. Of the remaining possibilities, the 

greatest loss of load would likely be caused by loose coal and 

rock on he floor, because the loss of load was very rapid 

immediately after setting. To eliminate this possibility, it is important to clean the floor prior to setting an MRS. This can be 

done with a continuous miner or the dozer blades on the MRS. 

5. A major advantage to using an MRS is that mine personnel no longer need to go into hazardous areas in entries or 

intersections to install supports as they do when setting posts. 

Dial gauges are installed on all MRS's to establish setting pressures and to indicate how the MRS's are loading up. Some of this 

improvement in safety is jeopardized when MRS operators must 

get close to the MRS to read the dial gauges while they are setting 

the machine or determining how the MRS's are being loaded. 

Larger, more visible gauges should be installed to eliminate the 

necessity of approaching the MRS to read the gauges. 

6. Research is continuing to establish optimum setting loads 

for a variety of conditions, define precursors to roof caving, 

investigate possible combinations of MRS's and other types of 

supports, and improve safety when using MRS's. 
REFERENCES 

Chase FE, Mark C [1993]. Ground control design for pillar extraction. Soc International Conference on Ground Control in Mining. Morgantown, WV: 

Min Eng preprint 93-282. West Virginia University, pp. 16-23. 

Follington IL, Trueman R, Medhurst TP, Hutchinson IN [1992]. Continuous Thompson RT, Frederick JR [1986]. Design and field testing of a mobile 

monitoring of mechanized breaker line supporn to investigate roof and pillar roof suppon for retreat mining. In: Proceedings of the 5th International Conbehavior. In: Proceedings of the I lth International Conference on Ground Con- ference on Ground Control in Mining, Morgantown, WV: West Virginia 

trol in Mining. Wollongong, New South Wales, Australia: University of university, pp. 73-79, 

Wollongong, pp. 345-357. 

Signer SP, Mark C, Franklin G, Hendon G [1993]. Comparisons of active 

versus passive bolting in bedded mine roof. In; Proceedings of the 12th 
APPENDIX.-LOAD RATE MONITORING SYSTEM 

Two pressure gauges mounted on each MRS were monitored 

continually by mine personnel to determine loads and loading 

rates on the machines. This loading information is used by 

MRS operators to determine when to cease mining operations 

andlor remove miners and equipment from an area before a 

dangerous fall occurs. The dial gauges could be difficult to 

read, requiring the MRS operator to approach the MRS to 

monitor the gauges. Spokane Research Center personnel and 

equipment manufacturers are currentiy cooperating in the 

development of a load rate monitoring (warning) system for the 

MRS that can be easily seen by all miners in the vicinity of the 

MRS. The system monitors dynamic loading on the MRS and 

activates green, amber, and red warning lights on the canopy 

near the dial gauges. Each light represents a different loading 

rate on the machine. 

The system uses a dedicated embedded processor to monitor 

pressure inside the four hydraulic jacks associated with the 

MRS. Loading is proportional to the internal pressure and the 

surface area of the piston head of the hydraulic cylinder and is 

determined by the formula 

F=AxP, (A-1) 

where F = force in pounds, 

A = area, in2, 

and P = pressure, psi. 

The embedded processor reads changes in cylinder pressure 

through four multiplexed data acquisition channels and converts 

these pressure changes to load rates, which are then displayed 

by the three load rate indicator lights. 

External components of the load rate monitoring system are 

shown in the back row in figure A-1. From left to right, the 

components are pressure transducers and connecting cables for 

monitoring the internal pressure in the hydraulic cylinders; an 

explosion-proof, MSHA-approved container that houses the 

internal components of the system; and three colored fluorescent load rate indicator lights with magnetic mounting bases. 

The internal components of the explosion-proof container 

are shown in the foreground of figure A-1. From left to right, 

these components are the power supply for the internal components, the embedded processor and associated control program, the solid-state control relays for the lights, and intrinsically safe power supplies and barriers for the external 

components. 

As an integrated package, these components can be retrofitted into existing MRS equipment. Further underground studies of MRS loading histories will be performed to determine 

critical load rate settings for the system and to evaluate its 

performance. 

Figure A-1 .-Load rate monitoring system components. 
FULL-SCALE PERFORMANCE EVALUATION OF MOBILE ROOF SUPPORTS 

By Thomas M. Barczak' and David F. Gearharf 

ABSTRACT 

Two mobile roof supports (MRS's), one manufactured by J. H. Fletcher and Co. and one manufactured by 

Voest-Alpine Mining and Tunneling, were evaluated under controlled load conditions in the Strategic 

Structures Testing Laboratory at the Pittsburgh Research Center. A unique load frame, called the mine roof 

simulator, provided a realistic simulation of mining conditions by inducing vertical, horizontal, and lateral 

loading on the support. The purpose of these tests was to determine the performance capabilities and 

limitations of the supports and to investigate factors that influence the measurement of loading and loading 

rate. An evaluation of the support design and load conditions that can cause support failure or loss of support 

capacity is presented relative to the laboratory tests. In general, lateral loading perpendicular to the 

longitudinal axis of the canopy is most severe, although horizontal loading in the direction of the longitudinal 

axis of the canopy can also produce critical loading is some cases. The tests indicate that both setting force 

and leg pressure measurement are influenced by the staging of the leg cylinders. The implications of these 

factors on load rate measurement are evaluated. Differences in design philosophy between the two supports 

are identified and related to support performance. The difference in leg design, two- versus three-stage, had 

the most impact on support performance. Safety issues pertaining to support operation and maintenance are 

also discussed. Lastly, MRS capacity and stiffness characteristics are compared with those of conventional 

timber supports. 

'~esearch physicist, Pittsburgh Research Center, National Institute for Occupational Safety and Health, Pittsburgh, PA. 

'Project engineer, SSI Services, Inc., Pittsburgh, PA. 
INTRODUCTION 

Mobile roof supports (h4RSrs) have improved the safety of 

pillar extraction during secondary mining by providing superior roof control and significantly reducing the materials 

handling associated with timber posts in pillaring operations. 

The superior capabilities of the MRS's have promoted pillar 

extractions in conditions such as weak roof and floor geologies prone to unpredictable caving that were previously too 

dangerous when using timber posts. 

Since the introduction of MRS's in the United States in 

i988, MRS technology has matured, with installations in more 

than 40 coal mines. Overall, MRS's have experienced widespread success. Few failures have been reported; these are 

typically attributed to lack of operating experience or severe 

conditions, such as those associated with the first or large 

areas of caving strata. Only one fatality has occurred on a mobile section; however, the fatality was not attributed to failure 

of the mobile supports. Nevertheless, some questions arose 

during the fatality investigation in 1995 regarding the performance capabilities of MRS's and the capability to assess 

ground instabilities from MRS loading. 

In an effort to evaluate their support design, but unrelated 

to the fatality investigation, the two manufacturers of MRS's, 

J. H. Fletcher and Co., Huntington, WV, and Voest-Alpine 

Mining and Tunneling (VAMT), Pittsburgh, PA, made arrangements to have their supports tested at the Pittsburgh 

Research Center. One support from each manufacturer was 

evaluated at the Center's Strategic Structures Testing Laboratory through full-scale testing in the unique mine roof simulator load frame. Although the supports that were tested are 

similar in operating range and capacity, the Fletcher support 

utilized three-stage leg cylinders, whereas the VAMT support 

utilized two-stage leg cylinders. This difference in leg design 

should be considered when making comparisons of support 

performance, particularly the stiffness of the support. 

A series of tests was conducted under controlled load conditions, which provides a better understanding of the performance capabilities and limitations of MRS's and factors 

that influence the measurement of loading and loading rate. 

This paper presents the results of these laboratory studies and 

compares differences in design philosophies and evaluates 

their impact on support performance. The supporting capabilities of MRS's is compared with those of conventional timber 

posts and cribs. Safety issues relative to support maintenance 

and operation are also discussed. 

STRATEGIC STRUCTURES TESTING LABORATORY 

The Srrategic Structures Testing Laboratory is a unique 

laboratory where full-scale mining equipment and roof support structures can be tested in a controlled environment. 

Figure 1 shows an MRS in the laboratory's mine roof simulator load frame. This unique load frame is designed to simulate the loading induced on support structures due to the behavior of rock masses during mining. The load frame can 

provide controlled roof and floor movements to simulate the 

closure of the mine opening while generating up to 13.334 kN 

(3 million lb) of vertical force and 7,117 kN (1.6 million lb) 

of horizontal (shear) force. 

The test procedure for the MRS evaluation was as follows: 

1. The MRS was positioned in the proper orientation to 

allow the load frame to induce vertical, horizontal, or lateral 

loading on the support. 

2. The MRS was actively set against the load frame 

platens using the internal hydraulic power to establish the 

initial load condition. 

3. Subsequent loading was applied by controlled displacement of the load frame's lower platen to simulate closure 

of the mine entry. Three different load vectors were evaluated 

through applied vertical, horizontal, and lateral displacements, 

as depicted in figure 2. 

4. The support response to the applied loading was 

measured through strain gauges and pressure transducers 

Figure 1 .-Full-scale testing of a mobile roof support In the 

unique mine roof simulator load frame. 
VERTICAL LOAD 

NEGATIVE HORIZONTAL LOAD POSITIVE HORIZONTAL LOAD 

NEGATIVE LATERAL LOAD POSITIVE LATERAL LOAD 

Figure 2.-Vertical, horizontal, and lateral loading applied to the mobile roof supports by the mine roof simulator. 
installed on the various MRS components. A typical instrumentation arrangement is shown in figure 3. 

Parameters investigated included (1) setting pressure, 

(2) support height, (3) load vector (direction of loading), and 

(4) canopy contact configuration. Additionally, a variety of 

eccentric crawler frame contact configurations were evaluated 

with the Fletcher support. The testing effort focused on the 

following studies: 

1. Rated support capacity: Determination of maximum 

support capacity in relation to the support's rated design 

capacity. 

2. Stifiess characteristics: Measurement of support resistance and component responses to roof movements in the 

vertical, horizontal, and lateral directions. 

3. Setting force: Evaluation of setting force as a function 

of leg staging and hydraulic pump pressure. 

4. Load and load rate measurement: Evaluation of factors 

that affect the measurement of roof load and loading rate. 

5. Conditions that reduce support capacity: Identification 

of load conditions that reduce support capacity. 

6. Critical load conditions: Identification of load conditions that maximize component loading and those that produce 

critical loading where the structural integrity of the supports 

could be jeopardized. 

rCanopy Left Center Gauge 

Gauge 

PIOW Legs PT 

Left Tilt Cylinder PT 

Figure 3.-Instrumentation installed on the VAMT support to assess support performance. PT = pressure transducer. 
MOBILE ROOF SUPPORT DESCRIPTION AND BASIC DESIGN PHILOSOPHIES 

The VAMT support tested during this study was a model 

185/380-540. The Fletcher support was a model MRS-13 with 

1.45 m (57 in) to 3.71 m (146 in) operating height. These support designs are representative of the support philosophies of 

the two manufacturers, although both manufacturers offer a 

variety of machines and designs to operate in mining heights 

ranging from 1.17 m (46 in) to 3.96 m (13 ft). 

The similarities in support design for the two supports 

evaluated in this study are as follows: 

Both supports were rated at 5,338 kN (600 tons) of support capacity and designed to operate in high seams at heights 

up to 3.81 m (12.5 ft). The maximum capacity is controlled 

by hydraulic yielding of the leg cylinders. 

The canopy is connected to the base frame by four 

hydraulic leg cylinders and a lemniscate assembly. The hydraulic cylinders provide the (vertical) support capacity or 

resistance to roof-to-floor convergence. The lemniscate assembly acts to minimize horizontal canopy movement during 

raising and lowering of the support and provides resistance to 

horizontal and lateral loading. 

The connection of the lemniscate assembly to the canopy 

is articulated to permit pitch and roll rotations of the canopy 

independent of the lemniscate assembly to allow full contact 

in uneven roof and floor conditions. 

An internal hydraulic power supply provides active 

setting of the support against the mine roof and floor with independent control of the front and rear legs. 

Ground contact is established through the crawlers with 

the crawler frame designed to support the full 5,338 kN (600 

tons) of roof load. 

There are four significant differences in design philosophy 

between the two supports tested: (1) a flat-plate canopy construction versus a sloped-edge canopy construction, (2) a tiltframe lemniscate assembly versus a rigid link lemniscate assembly, (3) internal versus exposed lemniscate assembly, and 

(4) a two- versus a three-stage leg cylinder design. 

The Fletcher support utilized a canopy construction that is 

sloped at the edges, whereas the VAMT support utilized a flatplate canopy design. The rationale for Fletcher's sloped-edge 

design is to accommodate edge and point loading with reduced deflection and stress at full load. The sloped edges are 

also intended to facilitate moving the support in uneven or 

jagged roof strata. A result of this design is increased canopy 

stiffness as the edge plates reduce the size and significantly 

stiffen the top canopy plate. The flexibility of the flat-plate 

canopy utilized in the VAMT support is illustrated in figure 4, 

where deflections as great as 7.6 cm (3 in) were observed over 

the length of the VAMT canopy when the support was loaded 

with a single contact placed near the canopy tip. The flat-plate 

structure. However, the greater roof contact will not necessarily translate into larger support loads, since the roof typically behaves as some sort of beam with support loading 

controlled by roof displacements and not the dead weight of 

the rock mass. The larger canopy can result in higher stress 

developments due to greater bending moments when the canopy is not uniformly loaded. 

Another major design difference pertained to the lemniscate assembly. The VAMT support utilized a lemniscate 

assembly connected to a tilt frame that permits single degreeof-freedom rotation of the lemniscate assembly due to lateral 

loading (see figure 5). The rotation is controlled by hydraulic 

cylinders called tilt cylinders. This design minimizes stress 

development in the lemniscate assembly due to lateral loading, 

but allows lateral translation of the canopy once the yield pressure of the tilt cylinders is reached. The Fletcher MRS as tested did not incorporate a tilt frame for the lemniscate assembly 

and relies on the strength of the lemniscate structure and 

connecting joints to resist lateral loading. The consequence of 

this design is significantly larger stress development in the 

lemniscate assembly due to lateral loading; however, the 

lateral translation of the canopy as a function of applied load 

is less than that of the VAMT tilt-frame design, particularly 

when the yield pressure of the tilt cylinders is reached. 

The VAMT support also utilized a hydraulic cylinder, 

called an aligning cylinder (see figure 6), for the top lemniscate link, versus a rigid steel link in the Fletcher support. 

The aligning cylinder limits horizontal load development, 

thereby minimizing stress development in the lemniscate assembly due to horizontal loading. When yield pressure is 

reached, the aligning cylinder yields through a 60-mm (2.4-in) 

stroke, permitting an equivalent horizontal displacement of the 

canopy relative to the base. When the rear legs are retracted, 

the aligning cylinder returns to its initial stroke and restores 

the canopy to its initial horizontal position. 

concept typically provides greater roof coverage due to roof Figure .?.-Deflection of the VAMT canopy under partial 

contact across the full width and length of the canopy contact loading. 
Breather port 

\ 

Pump pressure 

PIet 

--I t60 mm stroke 

Figure 6.-Aligning cylinder designed to control horizontal 

loading on the VAMT support. 

The position of the lemniscate assembly also differed for 

the two supports tested. The caving shield protruded beyond 

the rear of the canopy in the Fletcher support, whereas the 

Model 1401320-540 entire lemniscate assembly was internal (within the confines 

of the canopy) on the VAMT support. The VAMT support 

utilized a chain curtain to resist gob flushing into the support. 

Fletcher contends that the external position of the caving 

shield provides increased protection to machine components 

from gob material and can act as a wedge to help to push the 

support from heavily caved areas. The exposed caving shield 

can also cause additional loading on the lemniscate assembly 

due to gob loading. 

The Fletcher support that was tested utilized three-stage leg 

cylinders, as opposed to two-stage leg cylinders in the VAMT 

support. The rationale for the three-stage design is to enhance 

( 

operating range. A consequence of the three-stage design is 

13'2u_1 larger diameter leg cylinders, which impacts the operating 

pressure and several performance parameters, as described Figure 5.-Tilt-frame lemniscate assembly utilized on the 

VAMT support. later in this paper. 

ASSESSMENT OF LABORATORY TEST RESULTS 

MAXIMUM SUPPORT CAPACITY 

The maximum support capacity is controlled by hydraulic 

yielding of the leg cylinders, with a yield valve controlling the 

maximum pressure in the bottom stage of the leg cylinders. 

Normally, the left and right legs in the front and rear set are 

hydraulically connected together. As a result, the yield valve 

with the lowest operating pressure will control both legs in the 

set. The yield pressure required to provide a designated support capacity is a function of the effective area of the bottom 

stage of the leg cylinder. For example, the required yield 

pressure to produce 5,338 kN (600 tons) of support capacity 

was 26.3 MPa (3,820 psi) for the Fletcher support and 36.3 

MPa (5,263 psi) for the VAMT support. This difference is 

due to the difference in leg diameters: 25.4 cm (10 in) for the 

Fletcher and 21.8 cm (8.6 in) for the VAMT support. The 

measured yield settings were approximately 38.4 MPa (5,575 

psi) for the VAMT support, providing a maximum support 

capacity of 5,649 kN (635 tons), and 27.6 MPa (4,000 psi) for 

the Fletcher support, providing a maximum support capacity 

of 5,604 kN (630 tons). 

STIFFNESS CHARACTERISTICS 

Stiffness is a measure of how much roof movement is required to produce load resistance in the support. The stiffness 

characteristics of the support are evaluated for vertical, horizontal, and lateral displacements of the canopy relative to the 

base. Horizontal and lateral displacements are imposed in 

both a positive and negative direction (see figure 2). A comparison of the vertical, horizontal, and lateral stiffness at a 

2.4-m (96-in) operating height is presented in table 1. As seen 

in the table, both supports are much stiffer vertically than horizontally or laterally; this means that much more roof movement is required to produce equivalent support resistance to 

the applied displacement in the horizontal or lateral direction 

than for roof-to-floor convergence. It should also be noted 

that the initial horizontal and lateral stiffness of the support is 
sensitive to translational freedom in the various joints of the 

lemniscate assembly and the gear train of the crawler drive 

assembly. The stiffnesses shown in table 1 represent the support 

response once this translational freedom has been removed. 

Table 1 .--Cornparison of support stiffness at a 2.4-rn (96-In) 

operating height 

Vertical displacement . . . . 3,002 (! ,714) 2,140 (1,222) 

Horizontal displacement . . . '271 (155); '137 (78) 315 (180) 

Lateral disolacement . . . . 791 (52) '137 (78): '39 (22) 

'Stiffness measured when no leg stage is fully extended. 

qwo-stage leg cylinder support design. 

Three-stage leg cylinder support design. 

4Horizontal stiffness shown for horizontal displacement toward the plow 

of the support. 

'Initial stiffness prior to pressure development in the aligning cylinder. 

'Stiffness after aligning cylinder begins to develop pressure. 

71nitial stiffness prior to yield of tilt cylinders. Load applications that 

would produce yielding of the tilt cylinders were not evaluated. 

'Initial stiffness during first 1.3 cm (0.5 in) of lateral movement. 

'Stiffness beyond initial 1.3 cm (0.5 in) of lateral movement. 

Vertical Stiffness 

Vertical stiffness is a measure of support resistance to roofto-floor convergence. It is controlled almost entirely by the 

stiffness of the hydraulic leg cylinders. Vertical stiffness depends on the height of the support and decreases with increasing 

height (figures 7A and 7B). Therefore, the supporting force at 

a high operating height will be less than at a lower operating 

height for the same roof-to-floor convergence. Using the 

VAMT support as an example, the supporting force at a height 

of 3.8 m (148 in) is only 38% of the supporting force at a height 

of 2.4 m (96 in) for the same roof-to-floor convergence. 

When none of the leg stages are fully extended, the support 

stiffness is constant from set to yield, and the setting pressure 

does not have a significant effect on the support stiffness. When 

the support is set with the bottom stage fully extended, the support capacity as a function of displacement is bilinear. The 

initial stiffness is high, since the effective column length is reduced to that of the upper stage of the leg cylinders, and decreases once the upper stage force exceeds that of the lower 

stage setting force. An example for the VAMT support with 

two-stage leg cylinders is shown in figure 8A, where the stiffness decreases at about 4,226 kN (950 kips) of loading, which 

is where the bottom stage is dislodged from its mechanical stop 

when set at 28.96 MPa (4,200 psi) setting pressure. Figure 8B 

shows an example of the change in stiffness for a Fletcher 

support with three-stage leg cylinders when both the bottom and 

middle stages were fully extended. The initial stiffness was 

reduced when the top stage force exceeded the setting force of 

2,829 kN (636 kips) developed in the second stage. 

As expected, the VAMT support was stiffer in response 

tovertical loading than the Fletcher support (see table 1). This 

is primarily due to the two-stage leg cylinder design in the 

VAMT support, compared with the three-stage leg cylinder 

design in the Fletcher support. All other things being equal, a 

three-stage leg cylinder will always be less stiff than a two-stage 

leg cylinder, because the stages act in series with the equivalent 

stiffness reduced as the number of stages increases, as shown in 

equation 1 for a three-stage leg cylinder. Equation 1 also indicates that the equivalent stiffness is never greater than the least 

stiff member. Tne stiffness of individual stages is governed primarily by the area and suoke of the cylinder, decreasing in 

stiffness as the area decreases or the stroke increases. Thus, the 

stage with the smallest diameter will be the least stiff stage and 

is likely to control the equivalent stiffness of the entire leg: 

where K, = equivalent stiffness of the leg cylinder, 

K, = stiffness of stage 1, 

K, = stiffness of stage 2, 

and K, = stiffness of stage 3. 

When both supports are set at the same leg pressure, they 

will reach yield load at nearly the same displacement. For example, with a setting pressure of 17.3 MPa (2,500 psi) at the 

2.4-m (96-in) operating height, the VAMT support will reach 

yield load (5,338 kN (1,200 kips)) after 0.94 cm (0.37 in) of 

roof-to-floor convergence, compared with 0.86 cm (0.34 in) for 

the Fletcher support (see figure 9). However, when set to the 

same setting force, the Fletcher support will require 40% more 

displacement to reach yield load (see figure 10). 

Horizontal Stiffness 

Horizontal stiffness is a measure of support resistance to 

forward or rearward displacements of the canopy relative to the 

base. The action of the lemniscate assembly primarily controls 

the horizontal stiffness of MRS's, since the leg cylinders are 

nearly vertical and do not provide much resistance to horizontal 

loading. Horizontal stiffness is at least an order of magnitude 

less than the vertical support stiffness. 

As previously indicated, the initial horizontal stiffness is 

controlled by translational freedom in the connecting joints of 

the le~nniscate assembly and gear train of the drive motors. For 

example, up to 1.91 cm (0.75 in) of horizontal displacement of 

the canopy relative to the base was required in the VAMT 

support before any significant load resistance was generated. 

The horizontal stiffness of the support is also height-dependent, 

decreasing at increasing heights, as shown in the example in 

figure 11 for the Fletcher support. 
FIETCHER MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

- 96 IN HEIGHT 

120 M HEIGHT 

-.-. 140 IN HEIGHT 

-- -- 

. . 1 ...-- --a 

.....- 

. . . . - _.-- .-- . - 

. . . - - . - - - _/.-- . - . - . - 

_.C-- 

- . C 

_#.*. 

1 

- - *.-. __c- 

.---.- 

I 

0 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 

VERTICAL DISPLACEMENT, inches 

B VAMT MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

1,200 

0 

0 0. I 0.2 0.3 0.4 0.5 0.6 0.7 0.8 

VERTICAL DISPLACEMENT, inches 

% I 

- 

- % M HEIGHT 

113 MHEIGHT - I.-*... -.-- 148 IN HEIGHT 

.- - 

. . 

Figure 7.-Effect of support height on vertical stiffness. A, Fletcher support; B, VAMT support. 
VAMT BREAKER LINE SUPPORT - 148 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 

VERTICAL DISPLACEMENT, inches 

FLETCHER MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

0 0. I 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 

VERTICAL DISPLACEMENT, inches 

Figure &-A, Decrease in vertical stiffness on the VAMT support when bottom stage is fully extended at 0.4 in of displacement. 

B, Reduction in Fletcher support stiffness when both bottom and middle stages are fully extended at 0.15 in of displacement. 
STRATEGIC STRUCTURES TESTING LABORATORY 

96 IN HEIGHT - 2,500 PSI SETTING PRESSURE 

. . 

FLETCHER SUPPORT 

VERTICAL DISPLACEMENT, inches 

Figure 9.-Both the VAMT and Fletcher supports reach yield load at nearly the same displacement when set to the 

same leg pressure. 

STRATEGIC STRUCTURES TESTING LABORATORY 

FLETCHER SUPPORT - 96 IN HEIGHT - 1,814 PSI SET 

VMT SUPPORT - 96 IN HEIGHT - 2,500 PSI SET 

- VAMT SUPPORT 

...... FLETCHER SUPPORT 

.' 

/ - 

VERTICAL DISPLACEMENT, inchFigure 10.-Roof-to-floor convergence required to produce yield load in Fletcher and VAMT supports when both 

set to the same setting force. 
Unlike the Fletcher support, which utilizes rigid lemniscate displacement. The horizontal stiffness at a particular height is 

links, the horizontal force in the VAMT support, which reduced by as much as 50% when the aligning cylinder begins 

utilizes a hydraulic aligning cylinder to limit the maximum to develop load (see figure 12). For the example shown in 

horizontal loading, is a bilinear function of horizontal figure 12 (horizontal displacement of the canopy toward the 

FLETCHER MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

HORIZONTAL DISPLACEMENT, inches 

Figure 11 .-Horizontal stiffness increases at decreasing support heights. 

VAMT BREAKER LINE SUPPORT - 96 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

--v 

-0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0 

HORIZONTAL DISPLACEMENT, inches 

Figure 12.-Reduction in horizontal stiffness when pressure develops in the aligning cylinder. 
caving shield), the initial horizontal stiffness is 233 kNlcm in the VAMT support. As shown in table 1 , the Fletcher sup- 

(133 kipslin), followed by a stiffness of 116 kN1cm (66 port is stiffer than the VAMT support initially, whereas the 

kipdin) after the aligning cylinder pressure began to increase. VAMT support is stiffer than the Fletcher support when the 

The horizontal stiffness of the Fletcher support is 2.3 times lateral movement exceeds 1.3 cm (0.5 in). 

that of the VAMT support when the aligning cylinder is con- The lateral stiffness is less than the horizontal stiffness by 

trolling horizontal load development. a factor of 3 for the VAMT support and a factor of 2.3 (initial 

stiffness) or a factor of 8 (final stiffness) for the Fletcher sup- Lateral Stiffness port at a 2.4-m (96-in) operating height. 

Lateral stiffness is a measure of support resistance to applied left or right displacements of the canopy relative to the 

base. Thus, the direction of loading is across the width of the 

canopy versus along its length in horizontal stiffness evaluations. Lateral stiffness, as shown in figure 13, is also heightdependent. 

For supports equipped with a tilt-frame lemniscate assembly such as the tested VAMT support, lateral stiffness is 

controlled primarily by the tilt cylinders, which control rotation of the lemniscate tilt assembly. The lateral stiffness of 

the Fletcher support tended to be bilinear with a high initial 

stiffness during the first 1.3 cm (0.5 in) of lateral movement, 

followed by a reduced stiffness for lateral movements beyond 

this, as shown in figure 14. The decrease in stiffness was 

greatest at the 2.4-m (96-in) operating height, with a 70% 

reduction in stiffness when the lateral movement exceeded 1.3 

cm (0.5 in). The bilinear nature of the lateral stiffness is probably due to the interaction of the leg cylinders and the 

lemniscate assembly. This bilinear behavior was not observed 

ASSESSMENT OF SETTING FORCE 

Setting force is defined as the force exerted against the mine 

roof and floor by actively setting the support using the internal 

hydraulic power. The setting force is determined by the effective leg area times the hydraulic pressure with the total 

setting force equal to the sum of all four leg cylinder forces. 

The effective leg area depends on the staging of the leg cylinders. Figure 15 compares the setting force as a function of 

hydraulic leg pressure with no stages fully extended for the 

Fletcher and VAMT supports. Because the VAMT support has 

smaller diameter leg cylinders-21.8 cm (8.6 in) compared with 

25.4 cm (1 0 in) for the Fletcher support--greater pressures are 

required to produce equivalent setting forces. For example, 

approximately 17.4 MPa (2,530 psi) of pressure is required to 

produce 3,558 kN (800 kips) of setting force with the Fletcher 

support, whereas 24.1 MPa (3,500 psi) would be required to 

produce an equal setting force with the VAMT support. 

VAMT BREAKER LINE SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

LATERAL DISPLACEMENT, iocha 

Figure 13.-Effect of height on lateral support stiffness. Left-to-right lateral displacement of the canopy. 
FLETCHER MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

LATERAL DISPLACEMENT, inches 

Figure 14.-Bilinear stiffness response to lateral loading. 

The VAMT support utilized a two-stage leg cylinder, whereas the Fletcher support utilized a three-stage leg cylinder. Table 

2 shows the reduction in setting force due to leg staging for the 

VAMT and Fletcher supports. As shown in the table, setting 

force can be reduced by as much as 70% for three-stage leg 

cylinders when the bottom and middle stages are fully extended. 

Because variances can also exist in each leg of the support with 

regard to staging, setting forces between the values shown in 

table 2 are possible. Thus, a wide range of setting forces can be 

provided for both supports even if the hydraulic setting pressures remain constant from set to set. An example of this is 

shown in figure 16. 

Table 2.-Reductions in setting force due to leg staging 

The effect of leg staging on setting force development can be 

explained by examining the operation of the leg during setting 

and the associated leg mechanics, as depicted in figure 17 for a 

three-stage leg cylinder. Operationally, when the support is 

raised, the bottom stage is designed to extend to full extension 

fust, followed by the middle and top stages. Likewise, when the 

support is lowered, the bottom stage retracts first, followed by 

the middle and top stages. The setting force will always equal 

the force developed in the stage with the largest diameter that is 

r7nt fully extended, equaling the pump pressure times the area of 

that stage. 

When the support is initially raised from a collapsed position 

to a height greater than the bottom leg extension, the setting 

force is diminished in proportion to the area reduction of the 

next stage. as de~icted in table 2. On subseauent setting events, u. " Reduction in setting force, % 

Leg stage condition Fletcher VAMT the setting force depends on whether full extension of leg stages 

sup~ort ' suDood 2 is required due to changes in operating height. Once a support 

No stages fully extended. . . . . . 0 0 is extended to an operational height with a diminished setting 

Bottom stage fully extended . . . 45 42 force due to the bottom or middle stage being fullv extended, .+ - - 

Bottom and middle stage 

NAp the setting force will be restored to its maximum capability if the 

fullv extended . . . . . . . . . . . . . 70 

NAp Not applicable. support is reset at any lower height, provided the bottom stage 

'Three-stage leg cylinder design. has not been fully retracted, and the setting force again will be - -. 

Two-stage leg cylinder design. diminished if the support is reset at an equal or greater height. 
MOBILE ROOF SUPPORT TESTS 

STRATEGIC STRUCTURES TESTING LABORATORY 

0 500 1,000 1,500 2,000 2,500 3,000 3,500 4,000 4.500 

HYDRAULIC PRESSURE, psi 

Figure 15.4omparison of setting forces as a function of leg pressure with no stages fully extended for Fletcher 

and VAMT supports. 
EQUILIBRIUM REQUIREMENTS 

ff I= 02 =a3= PUMP PRESSURE 

EQUII IBRIUM REQUREMFNTS 

01 = 02 = PUMP PRESSURE 

03- ff2 

TOP STAGE 

03 A3 - 03 A3=0 

MlDDlf STAGF 

ff3A3+(ff2A2 -ff3A3)-gZA2=0 

BOTTOM STAGE 

ff2AZ-(ffzAz -~3A3) 

+(ffiA~ -ff3A3)-grAl=O 

jXUILIBRIUM REOUIREMENTS 

a,= 8.02 

m= 

rn = PUMP PRESSURE 

TOP STAGF 

ff3A3- 03 A3=0 

MIDDLF STAGE 

ff3A3- 02A2=0 

03 A3- me A2=O 

BOTTOM STAGE 

rn A1 - 01 A1 =o 

TOP STAGE 

03 A3 - 03 A3=0 

PUMP 

PRESSURE 

Figure 17.-Leg mechanics. A, bottom and middle stages fully extended; B, bottom stage fully extended; C, no stage fully extended. 
An example is shown in figure 18 for a three-stage leg cylinder. Behavior of a two-stage leg cylinder can be deduced by 

elimination of the middle stage. Initially, the support is set at 

a height (H,) that causes the bottom stage to be fully 

extended, providing a diminished setting'force. In preparation 

for the next cycle, the support is lowered, during which the 

bottom stage is partially retracted while the upper stages 

remain extended. When the support is reset (second cycle) at 

a lower operating height, full extension of the bottom stage is 

not required since the upper stages remain extended from the 

previous cycle. As a result, the setting force is restored to its 

inaximlim capability, equaling the setting pressure times the 

area of the bottom stage. Two scenarios are examined for the 

third cycle. In both cases, the support is reset at a higher 

operating height than the second cycle. In the first case, the 

support is raised to a height greater than the initial height. In 

this case, the bottom stage is fully extended once again and the 

setting force is once again diminished. However, if the 

support is raised to a height on the third cycle that is less than 

the initial height, full extension of the bottom stage is not 

required and full setting capacity is maintained. 

In summary, during underground operation, the setting 

force will always be reduced on the mining cycle that establishes a new maximum operating height after an initial operating height that causes full extension of the bottom stage. 

All other cycles should provide full setting capability because 

extension of the bottom stage will not be required. Operationally, the probability of achieving maximum setting forces 

can be enhanced by establishing a maximum operating height 

as soon as possible. Ideally, when the support is initially taken 

underground, it can be brought to a location that is higher than 

where it will be placed into operation during pillar extraction, 

and fully extended. This will ensure full setting forces for all 

load cycles, provided the support is not lowered to the point 

where the bottom stage is fully collapsed, which would then 

cause retraction of the upper stages. In this case, a new maximum operating height would have to be estabiished to prevent 

reductions in setting force. 

FACTORS AFFECTING LOAD AND LOAD 

RATEMEASUREMENTS 

Since the dial gauges on the support measure pressure in 

only the bottom stage of the leg cylinder, an assessment of 

load and loading rates can only be determined through the full 

load cycle when none of the stages are fully extended. If the 

bottom stage or bottom and middle stages (three-stage cylinder design) are fully extended, the dial gauges will not 

record changes in pressure until the setting forces in the extended stages are overcome by additional load development in 

the upper stages. When this condition occurs, roof loading 

during a beginning portion of the loading cycle will go undetected by the dial pressure gauges. The period of undetected 

roof loading depends on the setting pressure and will increase 

with increasing setting pressure in a particular support. 

Using the VAMT support as an example, if the support is 

set with 29.0 MPa (4,200 psi) of hydraulic pressure with the 

bottom stage fully extended, a force of approximately 4,226 

kN (950 kips) is generated in the bottom stage against the 

mechanical stops and 2,558 kN (575 kips) is generated in the 

upper stage acting on the mine roof. Because the bottom stage 

is fully extended, the dial gauges will remain inactive until the 

roof load acting on the support increases by i,668 kN (375 

kips) to cause the force in the upper stage to exceed 4,226 kN 

(950 kips) and cause the bottom stage to be moved off of its 

mechanical stops, resulting in an increase in pressure. 

Figure 19 shows the magnitude of roof loading that is not 

recorded by the dial pressure gauges when one or more leg 

stages are fully extended as a function of the setting pressure. 

As seen in the figure, the unrecorded roof loads increase 

linearly with increasing setting pressure. As expected, the 

magnitude of unrecorded roof loading is much greater for the 

Fletcher three-stage leg cylinders than for VAMT two-stage 

leg cylinders because the bottom stage area is 35% larger in 

the Fletcher support, creating a higher setting force in the 

bottom stage compared with the VAMT support at the same 

hydraulic setting pressure. Additionally, when the bottom and 

middle stages are fully extended, the load difference between 

the top and bottom stages governs the unrecorded roof load. 

As shown in figure 19, unrecorded roof load ranged from 

approximately 445 kN (100 kips) at 6.9 MPa (1,000 psi) of 

setting pressure to as high as 1,690 kN (380 kips) at full pump 

pressure for the VAMT support and 609 kN (137 kips) at 6.9 

MPa (1,000 psi) of setting pressure to 2,202 kN (495 kips) at 

full pump pressure when the bottom stage of the Fletcher 

support is fully extended. When both the bottom and middle 

stages are fully extended, 3,509 L"J (789 kips) of roof loading 

can go undetected by the dial gauges when the Fletcher 

support is set to full pump pressure. 

Therefore, a false sense of loading and loading rate can be 

interpreted from the pressure gauges when the bottom leg 

stage is fully extended. This can result in unreliable information for operators that utilize support loading to assess roof 

stability and impending roof caving. Full extension of the 

bottom stage can occur at heights greater than 50% of the 

operating range for a two-stage leg cylinder and at heights 

greater than 33% of the operating range for a three-stage leg 

cylinder (assuming equal stroke of the leg stages). However, 

the inaccurate information occurs only when a new maximum 

operating height is attained; therefore, the probability of 

inaccurate information depends on the mining conditions. 
1 

INITIAL SET SECOND CYCLE THIRD CYCLE (A) THIRD CYCLE (0) 

F set = A, (LOW) F set =q A, (HIGH) F set = q A, (LOW) F set = q A, (HIGH) 

HEIGHT = Hint HEIGHT = Hint- AH=H2 HEIGHT = HZ+ (AH+HAz) HEIGHT = Hz+ @ti-HA,) 

Figure 18.--Conditions that produce diminished setting force and unrecorded roof loads. 

MOBILE ROOF SUPPORT TESTS 

STRATEGIC STRUCTURES TESTING LABORATORY 

- VAMT TWO-STAGE LEGS WITH BOlTOM STAGE FULLY EXTENDED 

--- FLETCHER THREE-STAGE LEGS WITH BOTTOM STAGE FULLY EXTENED 

...... FLETCHER THREE-STAGE LEGS WITH BOTTOM AND MIDDLE STAGES FULLY EXTENDED 

800 - 

700 YI 

P 

Z 6 

a 

3 500~ 

8 s 400.- 

s 

g 300 

8 

2 

1 

-- 

600--~.- 

100 

. . . . 

_..' . , 44- . . 

/==----- 

_..---I 

. - .' 

--- 

C 

. . . - 

0 1,000 - 1.400 Tipp 1,800 2,200 2,600 3.000 3.400 3,800 4,200 

HYDRAULIC SE'ITING PRESSURE, psi 

Figure 19.-Unmeasured roof load when one or more of the stages are fully extended as a function of setting 

pressure in the leg cylinders. 

-_--. - - - 

----- - 

. 

. . , - A- 

- 

- - --- 

_--- ' , - - -, 

. 

- - - - - - 

-- , 

1 _--/- 

/ 

200-'---.-~---- 
CONDITIONS THAT REDUCE SLIPPORT CAPACITY 

One cause of reduced support capacity is the bleed-off of 

hydraulic pressure from the leg cylinders under static loading 

conditions. Bleed-off rates of 69 kPa (10 psi) to 138 kPa (20 

psi) per minute were common to both supports tested. As 

shown in figure 20, approximately 356 kN (80 kips) of load 

resistance was lost in 30 min because of loss of leg pressure 

under static loading for the VAMT support. 

Horizontal loading can either increase or decrea-e support 

capacity depending on the change in leg pressures between the 

front and rear set of legs and the reaction of lemniscate assembly. 

Leg cylinders that are inclined toward the direction of the 

horizontal displacement will generally increase in pressure; those 

inclined away from the direction of the horizontal displacement 

will generally lose pressure. The net pressure change between the 

front and rear set of legs will generally determine whether the 

support capacity will be reduced or increased. However, the 

reaction of the lemniscate assembly must also be considered. For 

horizontal displacement of the roof acting to push the canopy 

toward the caving shield, the lemniscate assembly develops an 

upward reaction at the canopy connection, which increases 

support capacity. Likewise, when the horizontal displacement is 

toward the plow, a downward reaction is developed at the canopy 

connection, which reduces support capacity. 

For the two supports tested, horizontal displacement produced the most change in support capacity at the lower heights 

because of the greater leg inclination. Figure 21 depicts the 

effect of horizontal loading on support capacity for the VAMT 

and Fletcher supports at a 2.4-m (96-in) operating height. As 

shown in the figure, support capacity was reduced for horizontal roof displacement toward the caving shield end of the 

canopy, and support capacity was increased when the horizontal canopy displacement was toward the plow. A maximum 

reduction in support capacity of 334 kN (75 kips) was observed for the VAIviT support as a result of 2.0 cm (0.78 in) of 

horizontal roof displacement toward the rear of the canopy. 

Figure 22 is an example of an increase in VAMT support 

capacity despite a reduction in leg pressures on both the front 

and rear set due to the reaction of the lemniscate assembly. 

Lateral displacements of the canopy in both directions 

tended to produce a loss of leg pressure that resulted in loss of 

support capacity. An example is shown in figure 23 for the 

Fletcher support. Support capacity was reduced by 378 kN 

(85 kips) on the VAMT support for left-to-right lateral displacement of the canopy at a 2.4-m (96-in) operating height 

with no significant loss of leg pressure (see figure 24), which 

suggests a negative reaction by the lemniscate assembly. 

Figure 25 compares the effects of horizontal and lateral 

loading on support capacity at a 2.4-m (96-in) operating 

height for the Fletcher and VAIviT supports. As shown in the 

figure, reductions in support capacity were greater for 

horizontal loading than lateral loading for both supports. 

VAMT BREAKER LINE SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

1.280 

TIME, minutes 

Figure 20.-Reduction in support capacity due to bleed-off of leg pressures after leg yielding. 
VAMT AND FLETCHER MOBILE ROOF SUPPORTS - 96 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

HORZ DISP TOWARD REAR 1 HORZ DISP TOWARD PLOW ----------> 

1,040 - I- 

- VAMT SUPPORT 

1.020 -- 

-FLETCHER SUPPORT 

1,000 - 

EL 'j 980 - 

w' 8 

960 - / 

e 

-1 < 940 - 

U iw - 

b 3 920 .--- 

> 

900 

-0.9 -0.6 -0.3 0 0.3 0.6 0.9 

HORlZONTAL DISPLACEMENT, inches 

Figure 21 .-Increase in support capacity tor horizontal displacement toward the plow and decrease in 

support capacity for horizontal displacements toward the rear. 

VAMT BREAKER LINE SUPPORT - 137 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

-2 -1.8 -1.6 -1.4 -1.2 - 1 -0.8 -0.6 -0.4 -0.2 0 

HORIZONTAL DISPLACEMENT, inches 

Figure 22.-Increase in support capacity despite a reduction in leg pressures. 
FlETCHER MOBILE ROOF SUPPORT - 120 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

LATERAL DISPLACEMENT, inches 

Figure 23.-Reduction in support capacity as a function of direction of lateral loading. 

LATERAL DISPLACEMENT, inches 

Figure 24.-Reduction in support capacity due to lateral loading with no significant loss of leg pressure. 
A VAMT BREAKER LINE SUPPORT - 96 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

980 

960 

B 

940 J 

8 

920 

-1 

5 

900 

s 

880 

860 

840 

-2 -1.8 -1.6 -1.4 -1.2 - 1 -0.8 -0.6 -0.4 -0.2 0 

HORIZONTAL OR LATERAL DISPLACEMENT, inches 

B FLETCHER MOBILE ROOF SUPPORT - 96 IN HEIGHT- STIFFNESS TEST 

STRATEGIC STRUCTURES TESTING LABORATORY 

910 

905 

900 

895 

3 

J 890 

885 

J 

< 

880 

er 

Y 

> 875 

870 

865 

860 

-2 -1.8 -1.6 -1.4 -1.2 -I -0.8 -0.6 -0.4 -0.2 0 

HORIZONTAL OR LATERAL DISPLACEMENT, inches 

Figure 25.--Comparison of the effects of horizontal and lateral loading on reductions in support capacity. A, 

VAMT support; 8, Fletcher support. 
CRITICAL LOAD CONDITIONS 

In general, the worst-case load condition for MRS's is 

lateral loading that causes lateral displacement of the canopy 

relative to the crawler frame. All of the rotational joints within the support structure are designed with a single rotational 

degree of freedom. Because lateral loading produces rotations 

along axes perpendicular to this rotational degree of freedom, 

it is the most severe load condition. 

Depending on the stiffness of the lemniscate assembly, 

horizontal loading can also produce critical loads in the 

lemniscate assembly components. VAMT uses a hydraulic 

cylinder in lieu of a rigid lemniscate link to limit stress development in the lemniscate assembly due to horizontal loading. 

The worst-case load conditions for canopy and base structures are partial contact configurations that induce bending. 

The associated stress development will be a function of the 

stiffness of the structure in relation to the applied loading. 

STRUCTURAL INTEGRITY OF THE VAMT 

AND FLETCHER SUPORTS 

Obviously, the effects of the above critical load conditions 

will be specific to a particular support design. A summary 

evaluation of the structural integrity of the VAMT and 

Fletcher supports based on measured component strains follows. However, it should be noted that the strain gauges were 

intended to assess load transfer through the various support 

components and were not necessarily positioned to measure 

maximum loading in any one component. All components 

were evaluated on both supports, except the crawler frame on 

the VAMT support. 

VAMT Support 

Highly loaded components on the VAMT support were the 

aligning cylinder and the canopy. 

The amount of horizontal force acting on the support required to produce pressure development in the aligning 

cylinder varied from 178 to 467 kN (40 to 105 kips) for 

support heights ranging from 2.4 to 3.8 m (96 to 148 in). 

Once pressure development begins, only another 67 to 89 kN 

(15 to 20 kips) is required to produce a yield pressure of 40 

MPa (5,800 psi) in the aligning cylinder. An example is 

shown in figure 26. In this case, 245 kN (55 kips) of horizontal loading acting to displace the canopy toward the rear of 

the support was required to produce pressure development in 

the aligning cylinder, and approximately 89 kN (20 kips) of 

additional horizontal loading produced a pressure of 40 MPa 

(5,800 psi). In this example, the displacement required to 

initiate pressure development in the aligning cylinder was 

1.14 cm (0.45 in), with 0.76 cm (0.3 in) of additional displacement required to produce a maximum pressure of 

40 MPa (5,800 psi) in the aligning cylinder (see figure 27). 

A malfunction of the aligning cylinder occurred during a 

test in which the cylinder was yielded in compression under 

the application of horizontal displacement of the canopy 

toward the plow. At the completion of the test when the pump 

pressure was applied to the cylinder during the retraction of 

the rear legs, hydraulic fluid under considerable p ure blew out 

of the breather port on the base of the cylinder, indicating that 

the lower piston seals had been damaged. Strain data were 

recorded during the test from two strain gauges located on the 

clevis that connects the cylinder to the tilt-frame assembly. 

The strain responses are displayed in figure 28. An 

examination of the strain data suggests that the damage 

occurred at approximately 13 cm (5.1 in) of horizontal 

displacement of the canopy relative to the base. The sharp 

increase in strain that occurred just prior to this suggests that 

the cylinder was fully stroked. However, an analysis of the 

lemniscate geometry indicates that approximately 23 cm (9 in) 

of horizontal canopy movement is required to compress the 

aligning cylinder through its full 60 rnrn (2.4 in) of stroke. An 

examination of the damaged cylinder by VAMT revealed that 

the cylinder was radially deformed (ballooned), suggesting 

that the failure was caused by excessive hydraulic pressure. 

However, the strain data indicate that there were not sufficient 

forces acting to generate hydraulic pressure that would darnage the cylinder. Therefore, the cause of the failure has not 

been satisfactorily determined. A new aligning cylinder was 

installed, and testing resumed. Subsequent tests at less-thanyield pressure were successfully conducted with no malfunctions of the aligning cylinder. However, at the discretion of 

VAMT, the new aligning cylinder was not tested under conditions that caused full compression or extension of the 

cylinder. 

The worst load case for the canopy was concentrated loading at the center or at one end of the canopy. However, it is 

important to note that the strain gauges were located midway 

between the front and rear leg connections, which is where the 

maximum bending moment is for the "contact at center" and 

"contact at both ends" configurations, but not for the other 

contact configurations. An assessment of stress at full support 

capacity can be made by extrapolating the canopy strains 

shown in figure 29 to 5,338 kN (1,200 kips) of support 

loading utilizing a modulus of elasticity of 206,850 MPa 

(30 x lo6 psi) for steel. The "contact at center" configuration 

produces a stress of 625 MPa (90,600 psi) at 5,338 kN (1,200 

kips) of support capacity. Assuming a yield strength of 690 

MPa (100,000 psi) for the steel, this configuration is close to 

producing permanent deformation in the canopy. A contact 

located 15.2 cm (6 in) from the canopy tip is projected to 

produce a stress of 393 MPa (57,000 psi) at the measured 

strain locations at full support capacity. However, the maximum bending moment is located farther back toward the rear 

leg in this loading condition, and the maximum stress is 

known to be greater than that measured in this test. 
VAMT BREAKER LINE SUPPORT - 96 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

-80 -70 -60 -50 -40 -30 -20 -10 0 

HORIZONTAL FORCE, kips 

Figure 26.-Horizontal force required to initiate load development and yield pressure in the aligning cylinder. 

VAMT BREAKER LINE SUPPORT - 96 IN HEIGHT 

STRATEGIC STRUCTURES TESTING LABORATORY 

HORIZONTAL DISPLACEMENT, inches 

Figure 27.-Horizontal displacement required to initiate load development and yield pressure in the aligning 

cylinder. 
VAMT MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

50 

R - RIGHT SIDE 

0 

-50 

-100 

-150 

-200 

-250 

HORIZONTAL DISPLACEMENT, inches 

Figure 20.-Malfunction of aligning cylinder during horizontal displacement toward the plow. 

VERTICAL FORCE, kips 

Figure 29.-Extrapolation of measured canopy strains to evaluate stress development at maximum support capacity. 

VAMT BREAKER LINE SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

105,000 

75,000 

45,000 

- 

m 

15,000 

VJ 

15,000 

451m 

75.000 

3,500 

2.500 . 

a .- 

2 

0 200 400 600 800 1000 1200 

5 1,500 L 

.- 

E 

I 

CANOPY CONTACT AT 

PLOW LEG CONNECTION 

I 

U 

6 IN FROM PLOW END 

-1,500 CONTACT AT BOTH& 

ENDS OF CANOPY ' _LT 

-2,500 

CONTACT AT 

CENTER OF CANOPY I 

I 

A 

I 

I 

I 

I 

I 
Fletcher Support 

The most highly stressed components in the Fletcher support were the bottom lemniscate link and sections of the base 

(crawler) frame. An objective of the testing was to determine loading limitations for these components. The following limitations are based on extrapolation of test data, 

where a margin of safety was maintained during load application. No failures of any component were observed 

under the test conditions. 

Lateral loading of 267 kN (60 kips) produced a stress of 

207 MPa (30,000 psi) in the bottom lemniscate link. Assuming a 690 MPa (100,000 psi) yield strength, extrapolation of 

the test data indicates that permanent deformation of the link 

would occur if the lateral load exceeded 556 kN (125 kips). 

Horizontal loading of 400 kN (90 kips) produced stresses as 

high as 3 10 MPa (45,000 psi) in the base cross frame member 

at a 3.1-m (120-in) support operating height. Extrapolation of 

these data suggests that the maximum horizontal loading capability for the base cross frame member at the 3.1 -m (1 20-in) 

operating height is approximately 934 kN (210 kips), assuming 

a 690-MPa (100,000-psi) yield strength. At the 3.6-m (140-in) 

height, horizontal loading of 445 kN (100 kips) produced 

stresses as high as 393 MPa (57,000 psi). Extrapolation of these 

data suggests that the maximum horizontal loading capability 

for the cross frame member at the 3.6-m (140-in) operating 

height is approximately 778 kN (175 kips). 

This analysis is conducted for full canopy and base contact. Eccentric load conditions on the crawler frame or canopy 

did not dramatically increase measured component strains. 

OTHER OPERATIONAL AND MAINTENANCE SAFETY CONSIDERATIONS 

Any MRS will become unstable if any of the lemniscate horizontal movement. Unrestrained movement of the canopy 

pins fail. Since critical stresses can be developed within the can result in serious injury or death. 

range of possible horizontal and lateral loading, these pins Caution should be used when working around the support 

should be periodically inspected. Additionally, before any of while it is pressurized. Oil leaking at these pressures can 

the lemniscate pins are removed, the canopy and lemniscate cause serious bodily damage. Likewise, pressure should be 

assembly should be supported to prevent both vertical and relieved before any hydraulic component is removed. 

COMPARISON OF MOBILE ROOF SUPPORTS WITH TIMBER POSTS 

The most obvious difference between MRS's and conventional timber posts is their size and effective roof coverage. 

Roof coverage depends on the manufacturer and support model, ranging from 3.3 to 7.9 m2 (35 to 85 ft2). In comparison, 

a wood post will provide less than 0.1 m2 (1 ft2) of roof coverage; thus, several timber posts are required to replace a single 

MRS. 

MRS's can provide an active load of up to 4,448 kN (500 

tons) to the mine roof; wood posts are strictly passive supports. The load-bearing capacity of one MRS is about the 

same as six 20-cm (8-in) diameter hardwood posts, as shown 

in figure 30. The stiffness of an MRS varies by support design and is height-dependent for a specific support. In general, an MRS operating at less than 75% of its maximum 

height is stiffer than a single 20-cm (8-in) diameter post with 

no headboard or two 20-cm (8-in) diameter posts with headboards. Figure 3 1 compares the stiffness of the Fletcher and 

VAMT supports with that of conventional timber posts and 

wood cribs. Comparisons with smaller diameter posts can be 

made by reducing the stiffness of the post in proportion to the 

reduction in cross-sectional area. 

Another significant advantage of an MRS is that it will 

continue to provide close to its full rated capacity after reaching yield load and can maintain this load capacity until the full 

leg stroke is exhausted. Thus, whereas MRS's can provide 

support through a meter or more of closure, timber posts can 

fail at less than 2.5 cm (1 in) of convergence and have no 

residual strength after failure. 

MRS's are also much better suited than timber posts to 

handle eccentric load conditions caused by horizontal and lateral roof or floor movements, gob loading, and rib rolls, which 

are common during pillar extraction and often kick out 

breaker and turn posts. In general, timber posts suffer reduced 

stability for anything but pure axial (vertical) loads. 

CONCLUSIONS 

Full-scale testing of MRS's at the Strategic Structures Test- The basic design of the VAMT breaker line support and 

ing Laboratory provided a wealth of information pertaining to the Fletcher MRS tested in this study is similar. Design differtheir performance capabilities and limitations. The tests were ences that impacted support performance included the lemnisconducted in the unique mine roof simulator load frame under cate assembly, the canopy construction, and the leg cylinder 

controlled conditions that simulate in-service load conditions. design. 
The VAMT support incorporated a tilt frame with hydraulic 

cylinders to control horizontal and lateral loading; the Fletcher 

support utilized rigid lemniscate links to resist horizontal and 

lateral loading. The tilt concept limits stress development in the 

support structure, but permits greater translation of the canopy 

relative to the base, thereby allowing greater roof movements to 

occur, particularly when the hydraulic tilt cylinders have 

yielded. The advantages and disadvantages of these designs 

from a ground control perspective have not been evaluated. 

Differences in the leg cylinder design caused most of the 

differences in support performance. The Fletcher support utilized a three-stage leg cylinder; the VAMT support, a twostage leg cylinder. Consequences of the three-stage leg design 

were (1) reduced support stiffness, (2) greater reductions in 

setting force when both the bottom and middle stages are fully 

extended, and (3) larger unrecorded roof movements, particularly when both stages are fully extended. The advantage of 

the three-stage leg design is greater operating range, providing a lower support profile for transporting and tramming 

underground. 

A critical issue pertaining to the measurement of support 

loading and loading rate is the effect of the staging of the 

leg cylinders. When the bottom stage of the leg cylinders is 

fully extended, the dial pressure gauges do not respond to 

increases in support load until the setting force established in 

the bottom stage is overcome by pressure development in the 

upper stages. The unrecorded roof load is greater at high 

setting pressures and is minimized at low setting pressures. 

Operationally, the bottom stage will be fully extended when 

the support is first raised to a height that exceeds the bottom 

stage stroke and, on subsequent cycles, whenever a new maximum operating height is established. Therefore, when possible, it is recommended that the support be taken initially to 

a location with a height greater than the expected operating 

height during pillar extraction, and fully raised. This will 

eliminate the problem of unrecorded roof loading. However, 

if this practice is followed, the support should be lowered as 

little as possible when moving the support to the section and 

during cycle changes. If the support is lowered sufficiently 

to cause the bottom stage to fully collapse, the probability of 

unrecorded roof loading will increase. 

Setting forces also greatly depend on leg cylinder staging 

and are diminished by as much as 70% for the Fletcher support with three-stage leg cylinders when the bottom and 

middle stages are fully extended. Setting pressure as measured by the dial gauges will not always reflect the true setting 

force. The same circumstances that cause unrecorded roof 

loading also cause diminished setting forces. It is desirable to 

avoid diminished setting forces because the effectiveness of 

the support to act as a breaker line for roof caving may be 

reduced for low setting forces. When comparing supports of 

different. design, it is important to remember that the smaller 

diameter leg cylinder will provide less setting force for the 

same hydraulic pressure than supports with larger diameter leg 

cylinders. This is one reason that the VAMT support operates 

at higher pump pressure than the Fletcher support. 

MOBILE ROOF SUPPORT TESTS 

STRATEGIC STRUCTURES TESTING LABORATORY 

DISPLACEMENT, inches 

Figure 30.-One VAMT support provides about the same capacity as six high-quality timber posts. 

,...> ..I 

... 

/' 

, 

, 

, 

/ 

, 

_._. 

, 

*. . . , 

-.. ...---- 

...... SIX TIMBER POSTS (8 in diam) WITH HEADBOARDS 

- VAMT MRS SET AT 4.200 PSI SElTING PRESSURE 

7 

..-- _.__ --. 

-. 
FLETCHER MOBILE ROOF SUPPORT 

STRATEGIC STRUCTURES TESTING LABORATORY 

0 0.2 0.4 0.6 0.6 1 1.2 1.4 1.6 1.8 2 

VERTICAL DISPLACEMENT, inchea 

B VAMT BREAKER LINE SUPPORT 

STRATEGIC STRUCTURES TESTING-LABORATORY 

600 

1 13 IN HEIGHT 148 IN HElGHT 

-8 in diam POST -- No headerboard . 

- - - 4 PT RED OAK CRIB 

- - -TWO 8 in diam posts with headerboard - 

\ 

\ 

VERTICAL DISPLACEMENT, inches 

Figure 31 .-Comparison of support stiffness with that of conventional timber posts. A, Fletcher support; B, VAMT 

support. 
Both the Fletcher and VAMT supports were found to be 

structurally sound for typical load conditions. The canopy is 

likely to be the most highly stressed component on the VAMT 

support for most load conditions. Partial contact can cause 

stresses as high as 690 MPa (100,000 psi). 

When horizontal or lateral loading is present, the 

lemniscate assembly and cross frame between the base crawler 

frames are likely to be the most highly stressed parts of the 

Fletcher support. Lateral loads in excess of 556 kN (125 kips) 

can cause damage to the bottom lemniscate link, and horizontal loads in excess of 778 kN (175 kips) can cause damage 

to the cross frame member. Unfortunately, there is no way to 

assess the magnitude of horizontal and lateral loads underground without installing additional instrumentation on the 

support. 

The aligning cylinder on the VAMT support was damaged 

when it was yielded in compression by approximately 13 cm 

(5 in) of horizontal displacement of the canopy relative to the 

base. The probability of such large horizontal displacements 

during underground use is not known, but it is likely that this 

is an extreme load condition that will not occur during normal 

mining cycles. The cause of the failure was not satisfactorily 

determined. The damaged cylinder was replaced, and subsequent tests at less-than-yield pressure were conducted without 

any failures. 

Because any support is unstable if the lemniscate link pins 

fail, all supports should be periodically inspected for damage 

or excessive deformation in the pin clevises. Furthermore, the 

canopy should be supported to prevent vertical and horizontal 

movement prior to removal or repair of an any lemniscate pin, 

regardless of the support manufacturer. 

MRS's provide superior supporting capabilities compared 

with conventional timber posts. Each mobile support has a 

load-bearing capacity of approximately six timber posts and 

equivalent stiffness to two or more posts. MRS's provide significantly greater roof coverage and are much more stable for 

the types of eccentric loading that is common during pillar 

extraction. Furthermore, the active loading capability provides a more effective breaker line by minimizing initial roof 

movements that can lead to roof instability or caving inby the 

supports. 

QU.S. GOVERNMENT PRlNnNG OFFICE: 190750&248