Proceedings: New Technology for

Ground Control in Multiple-seam Mining

IC 9495

INFORMATION CIRCULAR/2007

Department of Health and Human Services

Centers for Disease Control and Prevention

National Institute for Occupational Safety and Health
Information Circular 9495 

Proceedings: New Technology for Ground Control in 

Multiple-seam Mining 

Edited by Christopher Mark, Ph.D., P.E., and Robert J. Tuchman

DEPARTMENT OF HEALTH AND HUMAN SERVICES 

Centers for Disease Control and Prevention 

National Institute for Occupational Safety and Health 

Pittsburgh Research Laboratory 

Pittsburgh, PA 

May 2007
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SAFER • HEALTHIER • PEOPLETM
CONTENTS Page 

Abstract .........................................................................................................................................................................1 

Multiple-seam longwall mining in the United States: background, by C. Mark ...........................................................3 

Multiple-seam mining in the United States: design based on case histories, by C. Mark, F. E. Chase, and 

 D. M. Pappas ...........................................................................................................................................................15 

LaModel: a boundary-element program for coal mine design, by K. A. Heasley and Z. G. Agioutantis........................ 29 

The new two-dimensional LaModel program (LaM2D), by O. O. Akinkugbe and K. A. Heasley.............................35 

Multiple-seam longwall mining in the United States: lessons for ground control, by C. Mark...................................45

Extreme multiple-seam mining in the central Appalachian coalfields, by C. Mark ....................................................55 

Multiple-seam mining interactions: case histories from the Harris No. 1 mine, by F. E. Chase, P. Worley, and 

 C. Mark ....................................................................................................................................................................63 

Failure mechanics of multiple-seam mining interactions, by R. K. Zipf, Jr. ...............................................................73 
UNIT OF MEASURE ABBREVIATIONS USED IN THIS REPORT

ft foot m meter 

GPa gigapascal min minute 

hr hour mm millimeter 

in inch MPa megapascal 

ksi 1,000 psi psi pound-force per square inch 
PROCEEDINGS: NEW TECHNOLOGY FOR GROUND CONTROL 

IN MULTIPLE-SEAM MINING 

Edited by Christopher Mark, Ph.D., P.E.,1 and Robert J. Tuchman2

ABSTRACT 

Multiple-seam interactions are a major ground control hazard in many U.S. underground coal mines. In some 

U.S. coalfields, particularly in central Appalachia and the West, the majority of today’s mines are operating above 

and/or beneath previously mined seams. 

 The effects of multiple-seam interactions can include roof falls, rib spalling, and floor heave. These can seriously disrupt mining operations and threaten the safety of miners. In early 2006, a West Virginia coal miner was 

killed by rib roll that occurred in a high-stress zone beneath a remnant structure in an overlying mine. 

 Fortunately, not every multiple-seam situation results in hazardous conditions. Indeed, the vast majority do not. 

For the past several years, the National Institute for Occupational Safety and Health has been conducting research 

to develop better techniques for predicting the location and severity of multiple-seam interactions. During this investigation, more than 50 mines were visited across the U.S. coalfields. Nearly 300 case histories were collected and 

analyzed using multivariate statistical techniques. The study also employed the numerical model LaM2D to estimate 

the multiple-seam stress, the Analysis of Longwall Pillar Stability (ALPS) and the Analysis of Retreat Mining Pillar 

Stability (ARMPS) programs to determine pillar stability factors (SFs), and the Coal Mine Roof Rating to measure 

roof quality. 

 The study focused on the two most common types of multiple-seam interactions: 

• Undermining, where stress concentrations caused by previous full extraction in an overlying seam is the 

main concern; and 

• Overmining, where previous full extraction in an underlying seam can result in stress concentrations and 

rock damage from subsidence. 

The study confirmed that overmining is much more difficult than undermining, and isolated remnant pillars 

cause more problems than gob-solid boundaries. For the first time, however, it was possible to quantify these effects 

in terms of the equivalent thickness of interburden needed to compensate for them. The study also found that pillar 

design is a critical component of multiple-seam mine planning. Many of the failed cases involved pillars whose SF 

seemed inadequate once the multiple-seam stresses were accounted for. Weaker roof was also found to significantly increase the risk of multiple-seam interactions. 

 The most important result of the study is an equation that predicts the critical thickness of the interburden 

required to minimize the likelihood of a multiple-seam interaction. This equation was incorporated into a step-bystep methodology that allows mine planners to evaluate each potential interaction and take steps to reduce the risk of 

ground control failure. Such measures could include installing cable bolts or other supplemental support, increasing 

the pillar size, or changing the mine layout to avoid the remnant structure entirely. 

 These Proceedings also contain several previously published papers that cover other facets of multiple-seam 

mining research. Two papers describe the LaModel family of software developed by Professor Keith A. Heasley of 

West Virginia University. The LaModel programs were designed for calculating the stresses and displacements in 

coal mines or other thin, tabular seams in layered media. The original three-dimensional version of LaModel is 

essential for detailed analyses of complex multiple-seam scenarios. LaM2D, by contrast, implements a simplified 

two-dimensional model that is suitable for quick approximations of the multiple-seam stresses and strains. 

 Three additional papers in these Proceedings describe the extensive multiple-seam experience of the Harris 

Mine, examples of extreme multiple-seam mining from the central Appalachian coalfields, and longwall mine 

experiences in Pennsylvania, West Virginia, and Utah. The final paper reports on a numerical modeling study that 

provided some insight into the mechanics of multiple-seam mining.

 

 1Principal research mining engineer, Pittsburgh Research Laboratory, National Institute for Occupational Safety and Health, 

Pittsburgh, PA. 

 2

Technical writer-editor, Writer-Editor Services Branch, Division of Creative Services, National Center for Health Marketing, 

Coordinating Center for Health Information and Service, Centers for Disease Control and Prevention, Pittsburgh, PA. 

3

MULTIPLE-SEAM MINING IN THE UNITED STATES: BACKGROUND 

By Christopher Mark, Ph.D., P.E.1

INTRODUCTION 

 Studies have estimated that 156 billion tons of coal, 

representing two-thirds of the minable reserves in the 

United States, are subject to multiple-seam mining [Singh 

and Dunn 1981]. In some U.S. coalfields, particularly in 

central Appalachia and the West, the majority of today’s 

mines are operating above and/or beneath previously 

mined seams. 

 The effects of multiple-seam interactions can include 

roof falls, rib spalling, and floor heave, which can 

seriously disrupt mining operations and threaten the safety 

of miners. In early 2006, a West Virginia coal miner was 

killed by rib roll that occurred in a high-stress zone 

beneath a remnant structure in an overlying mine [MSHA 

2006]. 

 Fortunately, not every multiple-seam situation results 

in hazardous conditions. Indeed, the vast majority do not. 

Accurate prediction of which interactions are likely to be 

higher-risk allows mine planners to prepare for them or 

avoid them. 

 Over the years, multiple-seam mining has been the 

subject of much research, both in the United States and 

abroad. Much advice on how to mitigate the risk has been 

presented, but unfortunately it is often contradictory. For 

example, one group of researchers wrote that “stresses 

from superincumbent workings are not transferred through 

shale strata for distances of over 110 ft” [Haycocks et al. 

1982], while another group indicated that “a stress transfer 

distance of 760 ft has been recorded between longwalls” 

[Haycocks et al. 1992]. 

 For the past several years, the National Institute for 

Occupational Safety and Health has been conducting 

research to develop better techniques for predicting the 

location and severity of multiple-seam interactions. During 

this investigation, more than 50 mines were visited across 

the U.S. coalfields. The study also made extensive use of 

numerical models, particularly the LaModel family of 

software [Heasley and Agioutantis 2007]. This paper presents the background to that study. The results of the study 

are discussed by Mark et al. [2007]. 

BACKGROUND 

 Figure 1 shows the five major underground coal 

mining regions in the United States. From the standpoint 

of multiple-seam mining, by far the most significant coalfield is the central Appalachian region of southern West 

Virginia, eastern Kentucky, and southwestern Virginia. 

Currently, underground mines in this region produce 

approximately 123 million tons of coal per year, or about 

33% of the total U.S. underground production [EIA 2006]. 

Mining has been ongoing in central Appalachia for nearly 

150 years. Recent studies have indicated that perhaps 70% 

of the ultimate reserve base in the region has already been 

mined [Bate and Kvitkovich 2004]. 

 One consequence of the maturity of the central 

Appalachian coalfields is that nearly every remaining 

underground reserve has been impacted by past mining 

activity. The mountains of the central Appalachian coalfields are honeycombed with worked-out mines located 

above, below, and adjacent to today’s and tomorrow’s 

operations. Figure 2 shows a typical geologic column from 

the central Appalachian region. On this property there are 

13 seams in which mining has been or is currently being 

conducted. 

 Full extraction is also widely practiced in the central 

Appalachian coalfields. Although only 8 mines currently 

use the longwall method [Fiscor 2006], a recent survey 

indicated that approximately 315 mines, accounting for 

58% of the room-and-pillar production in the region, 

engage in pillar recovery [Mark et al. 2003]. The prevalence of full extraction adds greatly to the potential for 

multiple-seam interactions. 

 The Western United States is the next most significant 

area for multiple-seam mining. In Utah, Colorado, 

Wyoming, and New Mexico, nearly 95% of underground 

 

 1Principal research mining engineer, Pittsburgh Research 

Laboratory, National Institute for Occupational Safety and 

Health, Pittsburgh, PA. 

 Figure 1.—The five major underground coal mining 

regions in the United States. 

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin
4 

production comes from 13 longwall operations [EIA 2006; 

Fiscor 2006]. Approximately half of these are operating in 

multiple-seam configurations. In contrast to central Appalachia, in the West the same mining company is usually 

responsible for all the mining on a property. As a result, 

a greater degree of multiple-seam planning is normally 

possible. On the other hand, when combined with deep 

cover and strong roof and floor rock, multiple-seam interactions can contribute to deadly bump hazards [Peperakis 

1968; Iannacchione and Zelanko 1995]. 

 In none of the other three underground mining regions 

are multiple-seam interactions currently a major factor, 

although all three historically have had problems [Kohli 

1992; Paul and Geyer 1932; Zachar 1952], and they may 

very well have them again in the future. Factors that contribute to the relative lack of multiple-seam interactions in 

these regions include the following: 

• Most longwall production in the northern Appalachian and Alabama coalfields is from a single 

seam (the Pittsburgh and Blue Creek Seams, 

respectively), without significant mining in other 

seams above or below. 

• The depth of cover, particularly for room-andpillar mines, is relatively low in northern Appalachia and the Illinois Basin. 

• Very few room-and-pillar mines engage in fullextraction pillar recovery in the Illinois Basin. 

There is almost no room-and-pillar mining at all 

in Alabama. 

HAZARDS ASSOCIATED WITH 

MULTIPLE-SEAM MINING 

 Ground instability is usually the greatest hazard due to 

multiple-seam interaction. Interactions may be classified 

into four major categories depending on the mining 

method, mining sequence, and thickness of the interburden. Other potential hazards are associated with 

water, gas, and oxygen-deficient air. 

Undermining, the first category of interaction, occurs 

when the upper seam has been mined first and the lower 

seam is the active seam (Figure 3). In an undermining 

situation, damage is caused by load transfer from highly 

stressed remnant structures associated with full-extraction 

mining in the overlying seam. These remnant structures 

can generally be classified as either: 

 Figure 2.—Typical geologic section for Boone County, 

WV, showing coal seams. 

Overburden

Upper Seam

(Earlier Mining)

Lower Seam

(Current mining)

Interburden

UNDERMINING

Overburden

Upper Seam

(Earlier Mining)

Lower Seam

(Current mining)

Interburden

UNDERMINING

Figure 3.—Undermining interaction. 
5 

• Gob-solid boundaries, with gob on one side; or 

• Isolated remnant pillars that are surrounded by 

gob on two or more sides (Figure 4). 

 Figure 4 shows that while a gob-solid boundary 

carries a single, distributed abutment load, an isolated 

remnant pillar is subjected to two overlapping abutments. 

As a result, the stress concentration on an isolated remnant 

pillar is usually significantly larger than that on a gob-solid 

boundary, and its impact on underlying seams is proportionally greater. The interburden thickness is also 

important because the stress concentration beneath any 

upper-seam remnant structure becomes less intense the 

greater the interval between the seams. 

Overmining, the second type of interaction, occurs 

when the upper seam is extracted after mining is complete 

in the lower seam (Figure 5). Load transfer occurs in this 

situation just as it does in undermining (in other words, 

gob-solid boundaries and isolated remnant pillars cause 

stress concentrations both above and below). In addition, 

however, full extraction of the lower seam normally results 

in subsidence of the overlying beds. 

 Figure 6 is a conceptual model that illustrates the type 

of damage that can be expected within the overburden due 

to subsidence above a full-extraction panel. Five broad 

zones can be identified [Singh and Kendorski 1981; Peng 

and Chiang 1984; Kendorski 1993, 2006]: 

1. The complete caving zone, in which the roof rock 

is completely disrupted as it falls into the gob, 

normally extends two to four times the extracted 

seam height (h). 

2. The partial caving zone, in which the beds are 

completely fractured but never lose contact with 

one another, extends up to 6–10 h. 

3. The fracture zone, within which the subsidence 

strains are great enough to cause new fracturing in 

the rock and create direct hydraulic connections 

to the lower seam. The top of this zone can be as 

high as 24 h above the lower seam. 

4. The dilated zone, where the permeability is 

enhanced but little new fracturing is created, 

extends up to 60 h. 

5. The confined zone, where subsidence normally 

causes no change in strata properties other than 

occasional bed slippage. This zone extends from 

the top of the dilated zone to about 50 ft below 

the surface. 

 The dimensions of these zones vary from panel to 

panel because of differences in geology and panel geometry. The implication of this model for multiple-seam 

mining is that when the interburden thickness exceeds 

approximately 6–10 times the lower seam thickness, the 

upper seam should be largely intact, although the roof may 

be fractured or otherwise damaged. 

Dynamic interactions occur whenever active mining 

occurs above or beneath open entries that are in use. The 

most severe dynamic interactions occur when a lower 

seam is longwalled or pillared, resulting in active 

subsidence of the open overlying workings. However, 

 Figure 4.—Stress concentrations in multiple-seam 

mining: (A) gob-solid boundaries associated with a very 

large pillar and (B) remnant pillar isolated in the gob. 

Overburden

Upper Seam

(Current mining)

Lower Seam

(Earlier mining)

Interburden

OVERMINING

Overburden

Upper Seam

(Current mining)

Lower Seam

(Earlier mining)

Interburden

OVERMINING

Figure 5.—Overmining interaction. 

 Figure 6.—Overburden response to full-extraction 

mining: (A) caving zones, (B) fracture zone, (C) dilated 

zone, and (D) confined zone. 
6 

damage can also be caused by the abutment stresses 

associated with full extraction in an overlying seam or 

even, in extreme cases, by development mining above or 

below. 

 The conditions associated with dynamic interactions 

are generally far more difficult than would have been the 

case if the open workings were developed after the full 

extraction was completed. Part of the explanation is that a 

dynamic interaction subjects the preexisting works to a 

traveling wave of subsidence and/or abutment stress rather 

than the static situation where the disturbance is 

concentrated in a single area. In addition, while unmined 

ground is normally in a confined state when it is 

overmined or undermined, the presence of a mine opening 

removes the confinement. The loss of confinement greatly 

weakens the rock mass and exposes it to tensile bending 

stresses. 

Ultraclose mining is the fourth type of interaction and 

the only one in which development mining alone is 

significant. The main concern is failure of the interburden 

between the two seams. The beam of interburden can fail 

either through shear caused by pillar punching or by 

tension caused by the self-weight of the rock plus that of 

any machinery working on it (Figure 7). Ultraclose 

interactions are unlikely when the two seams are more than 

20–30 ft apart [Haycocks and Zhou 1990; Singh et al. 

2002]. Ultraclose scenarios are most likely to occur near 

where a thick seam splits or where a rider coalbed is of 

minable thickness. 

Other hazards include the potential for inundation 

from an overlying flooded mine, particularly where full 

extraction in the lower seam can create a direct pathway 

between the upper- and lower-seam gobs. Mine Safety and 

Health Administration (MSHA) regulations require that a 

permit be obtained prior to mining under a body of water. 

An evaluation of the potential hazard should consider 

[Michalek and Wu 2000]: 

• Estimates of the potential volume of water in the 

overlying mine; 

• Evaluation of the strata separating the two mines; 

• Determination of probable flow paths and 

identification of critical areas that may become 

flooded; 

• A warning system, water control plan, and evacuation plan in the event of an inundation. 

 A review of MSHA data indicates that of the 201 

inundation incidents that were reported during 1996– 

2005, only 4 resulted when caving associated with full 

extraction in a lower seam intersected water-filled overlying workings. Several other water inundations occurred 

when development in a lower seam inadvertently cut into 

uncased boreholes that were connected with an upper 

seam. No injuries were associated with any of these 

incidents. In one incident, however, the first longwall 

panel at a Kentucky mine encountered a major inflow from 

workings 150 ft above. The water posed no hazard to the 

miners, but the lack of pumping capacity in the lower seam 

resulted in major mining delays [Mark et al. 1998]. 

 Interestingly, development above gob areas has been 

associated with large, but temporary groundwater inflows 

in several instances [Stansbury 1981; Bauer et al. 1992; 

Lazer 1965]. In these cases, the fracture and dilated zones 

apparently filled with excess groundwater, which was 

drained when the entries were developed. Fractures in 

these zones can also fill with methane or oxygen-deficient 

air, resulting in inflows of methane or blackdamp when 

they are intersected by overmining. 

Overmining 

 During the hand-loading era, which lasted until about 

1950, most underground coal mines operated under 

shallow cover and emphasized complete recovery, leaving 

few remnant structures in the gob. Perhaps as a result, 

early studies of multiple-seam interactions barely 

mentioned undermining and focused almost exclusively on 

overmining. 

 One of the first comprehensive studies was reported 

by Eavenson [1923a]. He concluded that “mining an upper 

seam after a lower one can almost always be successfully 

done when the interval between the seams exceeds 19 ft,” 

although he noted several cases where some coal had to be 

abandoned with interburdens up to 120 ft. Several other 

individuals, including such notable rock mechanics pioneers as D. Bunting and G. S. Rice, took partial exception 

to Eavenson’s conclusion, pointing out the importance of 

the nature of the interburden, thickness of lower bed, and 

uniformity of extraction of the lower bed [Eavenson 

1923b]. Taking all the evidence into consideration, the 

American Institute of Mining and Metallurgical Engineers, 

Committee on Ground Movement and Subsidence, 

concluded that Eavenson’s figure of 19 ft was appropriate 

as long as the thickness of the lower seam did not exceed 

8–9 ft [AIMME 1926]. 

 Figure 7.—Ultraclose mining (after Chekan and Listak 

[1994]). 
7 

 The next significant investigations of multiple-seam 

mining were conducted at Virginia Polytechnic Institute 

and State University (VPI) in the early 1950s by C. T. 

Holland and his student D. T. Stemple. Initially, Holland 

seemed to draw optimistic conclusions similar to those of 

Eavenson. Citing 38 examples from the literature, he 

concluded that “all but two or three” were successful, 

although “success” was defined as extracting as little as 

50% of the upper seam [Holland 1951]. Holland found that 

20–25 ft of interburden was adequate to provide good 

mining conditions, although he emphasized that “no 

remnants should be left in the lower seam” because 

remnants are “certain to result in considerable disturbance 

to the overlying strata.” Holland also emphasized that at 

least 3 months, and preferably several years, should elapse 

after completion of mining in the lower seam to allow 

settlement of the gob to be complete [Holland 1951]. 

 The study by Stemple [1956] was a landmark. He 

visited 45 mines throughout the eastern coalfields and 

ultimately collected a database of 61 actual mining case 

histories. About one-third of these were overmining cases, 

and Stemple found that in nearly every one there was some 

disturbance to the overlying seam, including: 

• Cracking or horizontal parting of the roof strata; 

• Vertical displacement (subsidence); 

• Rarely, but serious when it occurred, “squeezing 

and crushing of coal, accompanied by falls of top 

or heaving of the bottom.” 

 Stemple found that the most severe damages occurred 

directly above isolated remnant pillars abandoned in the 

underlying seam. Gob-solid boundaries also caused 

trouble, but the greatest disturbance was typically observed 

“not directly above the gob edge, but rather 100–300 ft out 

over the goaf.” 

 Stemple concluded that Eavenson’s recommendation 

of 19 ft of interburden might have been adequate for handloading, but that “such conditions would probably be prohibitive to mechanized mining.” Indeed, he found that 

“damage seriously adverse to mining can be done even 

with a vertical interval greater than 300 ft.” 

 One explanation for the discrepancy between the 

conclusions of Stemple and Eavenson is that they may 

have been talking about different things. Eavenson was 

apparently concerned with mining over gob areas, which 

were usually quite extensive in hand-loading operations. 

As mechanized mining became more prevalent, remnants 

of various shapes and sizes were more often left within the 

gob areas. In addition, the new mining methods were less 

flexible, so when working an overlying seam it became 

much more difficult to avoid the underlying remnants. The 

result was that even though mining above gob areas might 

still be feasible, finding enough good mining to make 

extracting an overlying seam profitable became much 

more challenging. 

 One other factor Stemple evaluated was the effect of 

time. He concluded that mining in an overlying seam 

should not be conducted until the subsidence process is 

completed, which could require 5–10 years. 

 The next major study of multiple-seam mining was 

conducted by C. Haycocks and his students and colleagues 

at VPI. This work was conducted over a period of nearly 

2 decades, beginning in the early 1980s. Haycock’s program included the development of empirical equations 

based largely on Stemple’s data, supplemented by analytic 

work, photoelastic studies, and numerical modeling. There 

is little evidence, however, of underground in-mine data 

collection in Haycock’s work. Haycock’s research resulted 

in an extensive published literature, as well as several mine 

design computer programs. 

 To evaluate the potential for successful overmining, 

several equations were proposed [Haycocks and Zhou 

1990]: 

Ico = h/t [18.84 X – 2(Z – 50) – 1,240] (1) 

Ico = h (3.5X – 224) (2) 

Ico = h/t (15X – 973) (3) 

where Ico = critical interburden thickness (ft) for no 

 appreciable damage to upper seam; 

 h = lower-seam thickness (ft); 

 t = time since mining the first seam (years); 

 X = percent extraction in the lower seam; 

and Z = percent hard rock in the interburden. 

 Equation 2 indicates that the critical interburden 

thickness ranges from zero (for a lower-seam extraction of 

about 65%) to as much as 700 ft (for 90% extraction in an 

8-ft lower seam). 

 Application of these equations presents several problems. First, it is not made clear where they should be 

applied—above an isolated remnant pillar, a gob-solid 

boundary, or anywhere the lower seam has been mined 

out? Second, where should one determine the percent 

extraction in the lower seam? Is it an overall percent 

extraction, or does it vary from place to place? Finally, 

Equations 1 and 3 are very sensitive to the time factor, but 

does it make sense that critical interburden thickness is 

reduced by a factor of 10 if a fully subsided gob is 50 years 

old rather than 5? 

 Luo et al. [1997] looked back on the earlier VPI 

studies and concluded that “although efforts were made to 

relate the magnitude of upper-seam damage to innerburden 

thickness, mining height, time, and extraction percentage, 

the data scatter was too great to achieve this relationship.” 

Further research also showed that “upper-seam damage 

could not be correlated with subsidence strain at the upperseam elevation.” More success was reportedly achieved 

when the upper-seam roof conditions were included 

together with vertical movement in the upper seam. 
8 

 Several case histories of overmining, successful and 

unsuccessful, have been reported in the literature during 

the past 3 decades. In the Gary District of southern West 

Virginia, U.S. Steel attempted to work the Pocahontas 

No. 4 and 5 Seams about 60 ft above the worked-out No. 3 

Seam [Stansbury 1981]. Conditions were extremely difficult, particularly in the No. 4 Seam where the roof 

consisted of “3–17 ft of unconsolidated and thinly 

laminated bands of shale intermixed with thin bands of 

coal.” Particular difficulties occurred “when mining near 

or directly above gob lines or lost blocks of coal in the 

No. 3 Seam.” Ground control was reportedly achieved by 

developing 12-ft-wide entries supported by 9-ft bolts and 

trusses, both on 3-ft centers, but the roof support cost made 

mining uneconomic. Mining was significantly easier in the 

No. 5 Seam, typically just 10 ft above the No. 4, and the 

difference was attributed to a more competent shale roof. 

 In central Pennsylvania, Bethlehem Mines Corp.’s 

No. 33 Mine employed longwall methods to extract the 

B Seam and the overlying C-prime Seam [Bauer et al. 

1992]. The B Seam, averaging about 5 ft thick, was 

extracted first, and the interburden was approximately 

105 ft. The depth of cover was typically less than 600 ft, 

which allowed the upper-seam gates to be stacked above 

the lower ones. Ground conditions in the upper seam were 

generally quite good, indeed better than areas where the 

upper seam was mined over virgin B Seam. The improvement was attributed to subsidence above the lower-seam 

longwalls that apparently relieved some of the in situ 

horizontal stress. Some minor areas of poor roof were 

encountered when crossing into areas above the gob, but 

these could be handled with some additional support. 

Water inflows were a more serious impediment to upperseam mining. 

 In eastern Kentucky, Black Mountain Resources used 

room-and-pillar techniques to extract Owl Seam reserves 

located 200–235 ft above abandoned Harlan Seam 

longwall panels [Rigsby et al. 2003]. The Harlan Seam 

was 11 ft thick with up to 1,500 ft of cover. The Owl Seam 

panels were driven across the longwall stopline pillars and 

then developed over the longwall gob. Although some roof 

fractures, rib spalls, and water inflows were observed, the 

panel was developed and retreated without major incident. 

Undermining 

 Although undermining is more common than overmining and although it is the recommended mining 

sequence, it has received considerably less attention in the 

literature than overmining. The explanation may be the 

apparent simplicity of the load transfer effect. For 

significant load transfer to occur, two factors must be 

present: 

• The interburden must be relatively thin; and 

• The seams must be relatively deep. 

 Stemple [1956] included 26 cases of undermining in 

his study, and he documented interactions in about half of 

them. In those cases where interactions occurred, the depth 

of cover exceeded 500 ft and the interburden was less than 

110 ft. The disturbances all occurred beneath isolated 

remnant pillars or within 100–200 ft of a gob-solid 

boundary. Stemple also concluded that the time lag after 

the mining of the upper seam was not a factor in 

undermining. 

 Haycocks et al. [1982] emphasized the role of the 

interburden geology in determining the extent of load 

transfer. A softer overburden, either due to a large number 

of rock layers or a low modulus of the individual layers, 

results in an elongated pressure bulb that reaches deeper 

seams below. Using Stemple’s data, Haycocks et al. [1982] 

proposed two relationships for predicting the critical 

interburden thickness (Icu ) in room-and-pillar mining: 

Icu = 110 – 0.42 Z (4) 

Icu = 6.8 N + 55 (5) 

where N = the number of interbeds; 

and Z = percent hard rock in the interburden. 

 Equation 4 is illustrated in Figure 8. Elsewhere, 

Haycocks and Zhou [1990] emphasized the special role 

of isolated remnant pillars, including longwall chain pillars, in creating high-pressure zones in seams above or 

below. Pillars less than 60 ft wide were singled out as 

allowing the “abutment pressure zones from both sides to 

superimpose.” 

 In European mines, multiple-seam interactions have 

been a major concern for many years due to the deep cover 

and long history of mining. In the 1970s, the U.K. National 

Coal Board collected detailed data from 18 undermining 

case histories [Dunham and Stace 1978]. Using multivariate statistical techniques, the study concluded that the 

two most important factors affecting the condition of the 

underlying seam during longwall extraction were: 

• The type of remnant structure; and 

• The initial roadway stability (determined rock 

strength, roadway width, depth of cover, width of 

the adjacent pillar, and other factors). 

 Dunham and Stace cautioned that it is “extremely 

dangerous to dismiss interaction effects purely on the basis 

of the thickness of the interval between the seams.” In one 

case in their database, an isolated remnant pillar caused a 

disturbance 450 ft below, while in another case, a gobsolid boundary had no noticeable effect just 90 ft below. 
9 

Dynamic Interactions 

 For nearly a century, the verdict of the experts on 

dynamic interactions has been unanimous: Don’t do it! 

Some examples follow: 

• Eavenson [1923a]: “Working in an upper seam 

should not be attempted while pillar robbing is 

going on beneath it.” 

• Paul and Geyer [1932]: “Pillar recovery should 

never be commenced under advance work in the 

overlying seam.” 

• Stemple [1956]: “The greatest difficulties are 

caused when pillar falls are made in the lower 

seam beneath previously developed entry work in 

the upper seam.” 

• Lazer [1965]: “If openings are first developed in 

the upper seam and then undermined, the 

openings in the upper seam will cave totally and 

the developed pillars will be lost.” 

 Despite these warnings, this practice is still occasionally tried, with depressing results. Three relatively 

recent cases have been reported in the literature—one 

where pillars were extracted beneath previously developed 

mains 180 ft above [Su et al. 1986] and two where 

longwalls were extracted as far as 550 ft beneath open 

main entries [Ellenberger et al. 2003; Mark 2006]. In each 

case, the overlying main entries were lost or severely 

damaged. 

 Less predictable are instances in which delayed subsidence of underlying works has the same destructive 

effect on overlying entries. In one instance, a set of mains 

was developed 180 ft above pillared works, and conditions 

were excellent for 2 years [Mark 2006]. Then the roof 

began to deteriorate dramatically, and heavy supplemental 

support was required to prevent major roof collapses. In 

another instance, dewatering of 20-year-old works caused 

marginally stable support pillars to fail, causing a pillar 

collapse in an overlying seam and subsidence at the 

surface [Kohli 1992]. In yet a third case, extensive floor 

heave developed in a lower seam 2 years after it was 

developed. A year later, the 5-year-old workings in the 

upper seam were severely damaged [Matetic et al. 1987]. 

 Although subsidence associated with full extraction in 

a lower seam has the most dramatic effects, dynamic 

interactions have also been attributed to full extraction in 

an upper seam [Hill 1995]. In this situation, because the 

stress abutments are applied to the unconfined strata 

around preexisting openings in the lower seam, their 

effects are more severe than would be the case if the 

upper-seam mining was complete before the lower seam 

was developed. 

Ultraclose Mining 

 Haycocks and Zhou [1990] stated that “when the 

interburden thickness is less than two times the room 

width, interburden failure cannot be ruled out.” Typically, 

however, ultraclose interactions are only a concern when 

the interburden is less than 25 ft. Zhou and Haycocks 

[1989] determined that the minimum safe working 

thickness for a massive, unstratified sandstone was just 

6 ft, whereas for shale it was 20 ft. They also determined 

that tensile failure is unlikely when the interburden 

thickness exceeds about 4 ft, so shear failure is the main 

concern. 

 Columnization of the pillars is considered the standard 

design practice when ultraclose interactions are a concern. 

Columnization minimizes the shear stress in the 

interburden and also provides a more uniform stress on the 

pillars, minimizing the risk of pillar failure. In Indian coal 

mines, columnization is required if the interburden is less 

than 30 ft [Singh et al. 2002]. In South Africa, 

columnization is recommended where the parting distance 

is less than 0.6–0.75 times the pillar center-to-center 

distance. Barrier pillars should be columnized for interburdens up to 100 ft thick [Munsamy et al. 2004]. 

 Figure 8.—Percent sandstone in interburden versus 

interburden thickness (after Haycocks and Zhou [1990]). 
10 

NUMERICAL MODELING FOR 

MULTIPLE-SEAM MINING 

 Analysis of multiple-seam mining interactions is complex because of the many geologic and mine design 

variables, as well as the complicated three-dimensional 

(3–D) geometries that occur. This complexity makes 

empirical analysis difficult and lends attractiveness to 

numerical approaches. Two main types of numerical 

model have been employed for multiple-seam analysis: 

• Displacement-discontinuity models, including 

MULSIM and LaModel, in which only the seams 

of interest are discretized; and 

• Finite-element models, in which the entire rock 

mass must be discretized. 

 Displacement-discontinuity models provide a pseudothree-dimensional simulation of tabular deposits such as 

coal. They have undergone continuous development and 

improvement over the past 2 decades. The original 

MULSIM and MULSIM–PC codes were limited to purely 

elastic analyses [Donato 1992]. MULSIM–NL allowed 

yielding of elements within the coal seams and nonlinear 

gob elements [Zipf 1992], but the overburden was still 

simulated as one solid material. LaModel introduced a 

formulation that simulates the overburden as a stack of 

layers with frictionless interfaces, thereby providing a 

more realistic suppleness to the strata response [Heasley 

and Chekan 1999]. LaModel can also consider topographic 

relief and subsidence, and LaModel grids can be generated 

directly from AutoCAD mine maps [Heasley and 

Agioutantis 2007]. The most recent development is a 

simplified two-dimensional (2–D) version of LaModel 

called LaM2D, which is much easier to grid and which 

runs in a fraction of the time required for the full 3–D 

model [Akinkugbe and Heasley 2007]. 

 Chekan and Listak [1993, 1994] employed MULSIM–

NL in an extensive series of parametric studies evaluating 

the effects of mining sequence and orientation on multipleseam interactions. Their most significant findings were: 

• Peak multiple-seam stresses are greater when 

retreating from solid toward the gob than when 

retreating from the gob to the solid (Figure 9); 

• Stresses on the longwall face are greatest when 

the face is being retreated in a direction directly 

perpendicular to a remnant structure in the other 

seam; and 

• Orientation relative to other seam remnant 

structures is not a major factor for development 

workings. 

 Heasley and Chekan [1999] report two case histories 

in which LaModel was used to evaluate multiple-seam 

interactions. In both cases, the model results were calibrated against extensive stress mapping that was conducted 

underground. In the first instance, an undermining example 

from eastern Kentucky, a 60-ft-wide isolated remnant 

pillar in the upper seam resulted in a multiple-seam stress 

of 2,200 psi, which, when added to the 3,000-psi singleseam pillar stress, was enough to cause significant roof and 

rib failure (Figure 10). In the other case, a set of longwall 

gates encountered multiple-seam stresses of 1,300 psi 

above a barrier adjacent to high-extraction room-and-pillar 

mining. A significant feature of this study was that it was 

necessary to simulate the yielding of the lower-seam 

production pillars in order to realistically model the 

interaction. 

 Su et al. [1986] report an early example of the use of 

finite-element modeling to investigate multiple-seam interactions. Both 2–D and 3–D models were employed, and 

some allowed bedding plane slip. One significant conclusion was that caving of the lower-seam roof strata 

forced the horizontal stresses upward, potentially creating 

stress concentrations around openings in the upper seam. 

The models also showed that highly bedded rock, where 

sliding takes place along individual layers, results in a 

narrower and deeper zone of interaction. 

 Hsiung and Peng [1987a] used numerical modeling to 

develop some rules of thumb for undermining. They 

concluded that if the interburden thickness is two to three 

times the width of the upper-seam isolated remnant pillar, 

no interaction is likely to occur. On the other hand, when 

the interburden is less than 10 times the mining height of 

the upper seam, the models indicated that the lower seam 

is likely to be fractured as well as highly stressed. Hsiung 

and Peng [1987b] also indicated that it is best to retreat 

from the gob toward the solid and that the best situation 

occurs when a longwall face maintains an approach angle 

of about 30° to remnant structure. 

 Figure 9.—Influence of retreat direction on multipleseam interaction: (A) retreating from solid to gob creates 

an unfavorable “stress window,” while (B) retreating from 

gob to solid results in lower stress concentrations (after 

Chekan and Listak [1993]).
11 

 Some recent examples of finite-element modeling 

applications to multiple-seam mining include 2–D and 3–D 

analyses of pillar and roof stability in overmining cases 

from northern West Virginia [Zhang et al. 2004; Morsy 

et al. 2006]. Zipf [2007] focused on the effects of vertical 

stress, horizontal stress, stress reorientation, and bedding 

slip on failure mechanics during multiple-seam mining. 

Gale [2004] evaluated different stacked longwall chain 

pillar layouts in the Australian context and concluded (as 

have many others) that the offset arrangement is far 

superior to vertical stacking. His models also predicted that 

stress transfer might be observed up to four pillar widths 

above and below a chain pillar, which would be approximately 400 ft for a typical Australian longwall design. 

SUMMARY AND CONCLUSIONS 

 Hazards resulting from multiple-seam interactions are 

a serious issue at many U.S. coal mines, particularly in the 

central Appalachian and western mining regions. The four 

types of interaction are: 

• Undermining, where stress concentrations caused 

by previous full extraction in an overlying seam is 

the main concern; 

• Overmining, where previous full extraction in an 

underlying seam can result in stress concentrations and rock damage from subsidence; 

• Dynamic interactions, caused when full extraction takes place above or below open entries that 

are in use (the most extreme dynamic interactions 

involve mining beneath open entries in an upper 

seam); and 

• Ultraclose mining, where room-and-pillar 

development of two seams within 25–30 ft of 

each other can result in interburden failure. 

 Undermining and overmining are by far the most common types of interaction. Nearly a century of research has 

identified a number of factors that can affect the intensity 

of a multiple-seam interaction. These include: 

• Depth of cover: The deeper the overburden, the 

greater the potential stress concentration caused 

by multiple-seam mining. 

• Mining sequence: Overmining is more difficult 

than undermining because of the potential for 

rock damage caused by subsidence. Dynamic 

interactions (particularly retreating beneath open 

works) should be avoided at all costs. 

• Interburden thickness: The smaller the distance 

between the seams, the greater the intensity of the 

potential interaction. 

• Type of remnant structure: Isolated remnant pillars that are surrounded by gob cause more 

intense interactions than gob-solid boundaries. 

First workings are generally not a concern unless 

the seams are ultraclose. 

• Interburden geology: Stronger, less bedded interburden tend to distribute multiple-seam stress 

concentrations more rapidly, resulting in less 

intense interactions. 

• Immediate roof geology: Weak roof (and floor) 

are more likely to be damaged by multiple-seam 

interactions. 

• Angle of approach to remnant structure: Retreat 

mining should proceed from the gob toward the 

solid side of a gob-solid boundary, and a longwall 

should not be brought broadside into long remnant structure. 

 Figure 10.—Comparison between (A) in-mine stress 

mapping and (B) LaModel-calculated stresses for eastern 

Kentucky (after Heasley and Chekan [1999]). 

Gob

Gob

A. Stress Mapping for Mine 1

B. LAMODEL Stresses for Mine 1

Gob

Gob

Legend (psi (000))

0.9 1.7 2.6 3.4 4.3 5.1
12 

 The large number of geologic and mining variables 

involved in multiple-seam interactions has made them very 

difficult to analyze. Empirical studies have foundered 

because the databases were too small for the number of 

variables and because bivariate analyses are inappropriate 

when there are so many variables involved. Numerical 

models have been helpful, but to be most useful they have 

required site-specific calibration to underground conditions. A hybrid approach, employing multivariate statistical analysis of a large database combined with numerical 

modeling, could provide the mining community with a 

valuable tool for predicting, avoiding, or controlling 

multiple-seam hazards. 

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15

MULTIPLE-SEAM MINING IN THE UNITED STATES: 

DESIGN BASED ON CASE HISTORIES 

By Christopher Mark, Ph.D., P.E.,1

 Frank E. Chase,2 and Deno M. Pappas3

INTRODUCTION 

 Multiple-seam interactions are a major ground control 

hazard in many U.S. underground coal mines. The two 

most common types are: 

• Undermining, where stress concentrations caused by 

previous full extraction in an overlying seam is the 

main concern; and 

• Overmining, where previous full extraction in an 

underlying seam can result in stress concentrations 

and rock damage from subsidence. 

 The goal of the study described in this paper is to help 

identify the location and likely severity of these interactions. Mine planners can use this information to adjust 

the ground support, pillar design, or mine layout to minimize the hazard. 

 In conducting the study, the National Institute for 

Occupational Safety and Health (NIOSH) relied mainly on 

an empirical approach. Empirical methods in ground control start with the concept that real-world mining experience, in the form of case histories, can provide valuable 

insight into the performance of very complex rock 

mechanics systems. In recent years, statistical analysis of 

large ground control case history databases has led to the 

development of methods for longwall pillar design [Mark 

et al. 1994; Colwell et al. 1999], roof bolt selection [Mark 

et al. 2001], retreat mine pillar design [Mark and Chase 

1997], and design of rib support [Colwell and Mark 2005]. 

Although fairly uncommon in mining, modern empirical 

research methods based on quantitative data analysis using 

statistics are the foundation of econometrics, epidemiology, 

and many other scientific disciplines. 

DATA COLLECTION 

 The multiple-seam case history database was developed over the course of several years through a program of 

mine visits. The mines included in the study were identified through discussions with mining company personnel and Mine Safety and Health Administration (MSHA) 

roof control specialists in each MSHA district. The study 

focused on those mines that had experienced the most 

difficulties with multiple-seam interactions. At each operation, however, care was taken to collect both successful 

and unsuccessful case histories. In general, only those case 

histories where problems might reasonably have been 

expected were documented. The many cases where the 

seams were clearly too far apart for interaction to occur 

were ignored. 

 A total of 44 mines were investigated during the 

study, nearly all from the central Appalachian and western 

coalfields (Figure 1). Several mines were visited in the 

northern Appalachian coalfields, but none of the case 

histories collected met the criteria for inclusion in the final 

database. 

 The key goal of each mine visit was to develop a 

history of multiple-seam interactions (and noninteractions) 

for the operation. Overlay mine maps, showing both the 

active mine and past workings above and/or below, were 

reviewed with experienced mine officials who had firsthand experience of the conditions encountered. Every 

instance where the active mine had crossed a gob-solid 

boundary or a remnant pillar was discussed. The officials 

also provided their best recollection of the support used 

and other relevant information. These discussions resulted 

in a preliminary list of case histories for the operation. 

 Underground investigations were also conducted at 

nearly every mine. It was seldom possible to access more 

than a few of the historic interaction sites because many 

were in sealed or otherwise inaccessible areas. However, 

 

 1Principal research mining engineer. 

 2

Research geologist. 

 3

Research civil engineer. 

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

Central

Appalachian

Coalfields

Western

Coalfields 

Western

Coalfields 

 Figure 1.—Location of mines included in the NIOSH 

multiple-seam database. 
16 

underground observations provided a sample of the ground 

conditions associated with interactions at that mine. Mapping of the conditions was only conducted in some 

instances. The underground visits also provided raw data 

on roof geology and strength for determination of the Coal 

Mine Roof Rating (CMRR). 

 The mine officials were also asked to provide 

AutoCAD files with mine maps for all of the seams on the 

property, together with exploratory bore logs. These data 

were subsequently analyzed by NIOSH to complete the 

database. 

PARAMETERS IN THE DATABASE 

 In any empirical study, the most important parameter 

is the outcome for each case history. A nonsubjective 

measure of the outcome, one that does not rely solely on 

the opinions of different observers, is highly desirable. For 

example, in the NIOSH study that led to the development 

of the Analysis of Roof Bolt Stability (ARBS), the outcome variable was the number of MSHA-reportable roof 

falls per 10,000 ft of drivage [Mark et al. 2001]. 

 In this multiple-seam study, there was no such clear, 

unbiased measure of the outcome. Complete reliance on 

the observations of the miners themselves, however, was 

not an option. As Stemple [1956] noted while he was 

collecting his multiple-seam database: 

Bad roof conditions are present in many cases where 

there is no vertically adjacent mining, and so it is not 

always possible to state definitely that mining in 

another seam is responsible for the conditions. The 

coal miner is usually anxious to identify the cause of 

any difficulty which arises, and certainly the previous mining of a contiguous seam provides a convenient scapegoat. 

 Therefore, a combination of reported conditions and 

evidence from the mine map were used to rate the conditions for each case history. For example, where a roof fall 

occurred above or beneath a remnant structure, but the 

map showed that nearby noninteraction areas encountered 

a similar density of roof falls, the case would be eliminated 

from the database. 

 A four-level outcome rating scale was used: 

• No interaction was assigned where undermining or 

overmining had occurred, but no effect was reported 

in the target seam (i.e., the conditions were the same 

in the affected and the unaffected areas). 

• Minimal interaction was assigned where the presence of past mining was noticed underground (e.g., 

there was slightly more rib spall or an occasional 

roof crack), but this did not have any effect on 

mining operations. 

• Moderate interactions were those where mining was 

completed, but with significant difficulties attributed 

to multiple-seam interactions, including such evidence on the mine map as roof falls, entry segments 

or crosscuts that were not developed, or pillars that 

were left unmined during retreat operations. 

• Severe interactions were those where mining 

operations were abandoned. 

 In the statistical analyses, these four levels were 

collapsed into two. The “no interaction” and “minimal 

interaction” cases were combined as “successes,” while the 

“moderate” and “severe” interactions became the 

unsuccessful cases (or “failures”). A handful of cases were 

rated “borderline” where the interaction seemed to fall 

between “minimal” and “moderate.” 

 The “explanatory” or “independent” variables are 

those that are thought to possibly contribute to the outcome. Values for some of the variables were readily available, including whether the case was: 

• Overmining or undermining 

• Development or retreat 

• Longwall or room-and-pillar 

Other variable values could be obtained from the mine 

maps or drill logs: 

• Depth to the target seam 

• Thickness of the interburden 

• Seam heights for both seams 

• Time lag between the mining of the two seams 

(obtained by comparing the dates of mining for each 

seam) 

• Angle of mining at which the active section intercepted the remnant structure 

 The level of roof support was determined during the 

discussion with the mine staff. Originally, the plan was to 

use ARBS to measure the amount of support installed. 

However, it was found that in almost every instance the 

primary support consisted of 4- or 5-ft fully grouted resin 

bolts. With so little variation, it did not make sense to 

include primary support as an independent variable. Similarly, supplemental support was almost always a pattern of 

8- to 12-ft-long cable bolts or resin-assisted mechanical 

bolts. Therefore, supplemental support was included as a 

“yes/no” variable. 

 Some variables required some intermediate calculations. These included: 

• Pillar stability factors (SFs) were calculated for the 

target seam using the Analysis of Longwall Pillar 

Stability (ALPS) or the Analysis of Retreat Mining 

Pillar Stability (ARMPS) programs. In some cases, 

SFs were also determined for remnant structures in 

the previously mined seam to determine whether 

they were likely to be yielded or intact. 
17 

• Type of remnant structure (gob-solid boundary or 

isolated remnant pillar). A remnant structure was 

judged to be an isolated remnant pillar if its SF 

indicated that it was large enough to be intact, but 

small enough that it concentrated the abutment 

stress, as shown in Figure 2. Equation 1 was used to 

determine whether a pillar was considered an isolated remnant or a gob-solid boundary: 

Wp = 5 H (1) 

where Wp is the maximum allowable width (ft) for a 

remnant structure to be considered an isolated 

remnant pillar, and H is the depth of cover (ft). 

Equation 1 indicates that the maximum width of an 

isolated remnant pillar is 100 ft at 400 ft of cover 

and 200 ft at 1,600 ft of cover. Larger remnant 

structures were considered gob-solid boundaries. 

• The CMRR was normally determined underground, 

but the core logs often showed that the geology 

varied from hole to hole. Using the underground unit 

ratings for individual rock layers, CMRR values 

were calculated for each borehole. Each case history 

was then assigned the CMRR determined for the 

nearest borehole. However, since geostatistical studies have shown that immediate roof geology is 

seldom consistent between holes [Mark et al. 2004], 

an average CMRR value was also determined for 

each mine. 

• The percentage of competent rock in the interburden 

was calculated at each borehole by summing the 

total sandstone plus limestone, and then dividing by 

the total interburden thickness. 

• The number of beds in the interburden. 

 A key difficulty in multiple-seam analysis is obtaining 

estimates of the multiple-seam stresses. The complex 

three-dimensional (3–D) geometries of most multiple-seam 

situations have so far defied attempts at empirical 

estimation. In the early phases of this study, 3–D LaModel 

analyses were conducted of several case histories using 

LaModel’s capability to import pillar grids directly from 

AutoCAD mine maps [Ellenberger et al. 2003; Heasley 

and Agioutantis 2007]. However, the sheer volume of case 

histories was too great to contemplate conducting fullscale 3–D analyses on the entire database. Fortunately, 

it was possible to analyze every case using LaM2D 

[Akinkugbe and Heasley 2007]. Some simplifications were 

necessary in order to create two-dimensional (2–D) grids 

of the case histories: 

• Gob widths were defined by the least dimension of 

the gob (in plan view); 

• Entry widths in the previously mined seam were 

adjusted so that the model 2–D extraction ratio 

approximated the true 3–D extraction ratio in order 

to more accurately simulate pillar yielding; 

• Standard LaM2D defaults for the rock and coal 

moduli, lamination thickness, gob stiffness, and 

yielding properties of the previously mined seam 

were employed; 

• Coal elements in the target seam were modeled as 

elastic (nonyielding) to better estimate the loads that 

the ground was attempting to apply to the critical 

pillars; and 

• An out-of-plane extraction ratio multiplier was 

applied to the target seam to account for the crosscuts that could not be modeled in 2–D. 

 Each model was run for the development case and, 

where appropriate, the retreat mining case. Figure 3 shows 

a portion of a typical LaM2D model grid. The variable 

values that were derived from the model results included: 

• Average total vertical stress applied to the critical 

pillar (most heavily loaded pillar above or beneath 

the remnant structure); 

• Average multiple-seam stress on the critical pillar; 

• Maximum convergence in the entry adjacent to the 

critical pillar; and 

• Maximum differential convergence, defined as the 

difference between the convergence at the edge of 

the critical pillar and the convergence at the rib on 

the other side of the entry. 

 Figure 2.—Stress distribution on isolated remnant 

pillars and gob-solid boundaries. 

Gob Solid Boundary

Abutment 

Stress

Isolated Remnant 

Pillar

9.3 H 5.0 H

A B

Abutment Load

H Overburden Depth

Gob Solid Boundary

Abutment 

Stress

Isolated Remnant 

Pillar

9.3 H 5.0 H

A B

Abutment Load

H Overburden Depth
18 

 The results from the LaM2D analyses were then used 

to determine the multiple-seam stability factor (ARMPS 

SFMS) determined for the target seam using the following 

formula (using either ARMPS or ALPS, as appropriate): 

ARMPS SFMS = 

⎥

⎦

⎤ ⎢

⎣

⎡

(single seam load + multiple seam load)

single seam load ARMPS SF (2) 

 Finally, an SF rating was calculated by comparing the 

ARMPS SFMS to the recommended ARMPS SF, which 

depends on the depth of cover [Chase et al. 2002]. Those 

cases where the ARMPS SFMS exceeded the recommended 

SF were given a rating of 1; the others were rated 0. 

DESCRIPTION OF THE DATABASE 

 The final database includes 344 case histories from 

36 different coal mines. The cases include 252 development cases and 92 retreat mining cases. Since retreat mining cannot be conducted unless development mining was 

successful, every retreat mining case is also included in the 

development mining database. 

 Figure 4 shows that more than half of the cases in the 

database involved undermining during development 

(n=190). Only about 13% of these undermining development cases were judged to be failures. Retreat mining was 

later attempted in about 40% of the undermining cases and 

was successful about 65% of the time. There are about 

one-third as many overmining development cases in the 

database as undermining (n=61), and their failure rate is 

almost three times as great. Retreat mining was only 

attempted in 19 of the overmining cases in the database, 

with a 68% success rate. 

 Figure 5 shows the type of remnant structure 

encountered for the development cases only. It indicates 

that when undermining encountered a gob-solid boundary, 

the crossing was successful 90% of the time. The failure 

rate almost doubled for undermining isolated remnant 

pillars, however, from 10% to 19%. Overmining developments were successful 73% of the time when crossing gobsolid boundaries, but that rate dropped to only 59% for 

remnants. 

 Figure 6 shows the depth of cover and interburden 

thickness for the undermining cases. In about 90% of the 

cases, the depth of cover ranges between 400 and 1,200 ft. 

The interburden is less than 220 ft in all but 20 cases (and 

these are all successes). Figure 6 also shows that there are 

few cases in which the interburden falls between 90 and 

150 ft. The fact that the data fall into two main groupings 

(interburdens of 40–90 ft and 150–240 ft) is less than ideal 

from a statistical point of view. However, when the sample 

pool is dictated by geology, it is impossible to conduct perfect “random sampling.” 

 Figure 3.—Portion of a typical LaM2D model grid used 

to determine the multiple-seam stress and total vertical 

stress. 

SITE C1 SITE C1

0

20

40

60

80

100

120

140

160

180

200

Overmining -

Retreat

Overmining -

Development

Undermining -

Retreat

Undermining -

Development

Number of Cases

Failure

Success

 Figure 4.—Histogram showing the case history database by type of mining (development or retreat) and 

type of interaction (overmining or undermining). 

 Figure 5.—Histogram showing the case history 

database by type of remnant structure and type of 

interaction. 

Structure vs. Under/Over Mining for Development Mines

0

20

40

60

80

100

120

140

Overmining -

gob solid

Overmining -

remnant

Undermining

-gob solid

Undermining

- remnant

No. Cases

 Failure

Success
19 

 Depth of cover and interburden thickness for the 

81 overmining cases are shown in Figure 7. Both Figures 6 

and 7 show that the likelihood of success increases as the 

interburden thickness increases. 

 Figure 8 shows the TVS and multiple-seam stress 

determined for the critical pillar. The TVS is determined as: 

 TVS = Tributary area stress + multiple-seam stress + 

 abutment stress (if applicable) (3) 

where the tributary area and abutment stresses were calculated with ALPS or ARMPS and the multiple-seam stress 

was determined using LaM2D. 

 The range of the TVS is from approximately 1,000 to 

4,000 psi, while the multiple-seam stress varies from near 

zero to about 2,000 psi. Failures are more likely to occur at 

higher levels of both types of stress. 

 The range of CMRR values in the database is shown 

in Figure 9. None of the cases had a CMRR of less than 

44, which means that truly weak roof conditions are not 

represented in the database. 

 Figure 10 shows the ARMPS SF adjusted for the 

multiple-seam stress plotted against the depth of cover. Of 

the failed cases, about one-third had ARMPS SFs that 

were below the recommended values. These included 

nearly half of the retreat cases. In contrast, more than 87% 

of the successful cases (including two-thirds of the retreat 

successes) had ARMPS SFs that exceeded the 

recommended values. The clear implication is that many 

multiple-seam interactions could be avoided simply by 

adjusting the pillar design to account for the additional 

multiple-seam stresses. On the other hand, there are still 

many unsuccessful cases in the database that cannot be 

explained by improper pillar design. 

 Figure 6.—Plot showing depth of cover and interburden thickness for the undermining cases in 

the database. 

0

500

1000

1500

2000

0 100 200 300 400

Interburden (ft)

Depth of Cover (ft)

Failure

Success

0

500

1000

1500

2000

0 50 100 150 200 250 300

Interburden (ft)

Depth of Cover (ft)

Failure

Success

 Figure 7.—Plot showing depth of cover and interburden thickness for the overmining cases in the 

database. 

 Figure 9.—Plot showing the range of CMRR values in 

the database. 

0

200

400

600

800

1000

1200

1400

1600

1800

2000

40 50 60 70 80

CMRR

Depth of Cover (ft)

Failure

Success

 Figure 8.—Plot comparing the total vertical stress and 

multiple-seam stress in the database. 

0

0.5

1

1.5

2

2.5

012345

Total Vertical Stress (ksi)

Max Mult. Seam Stress (ksi)

Success

Failures
20 

LOGISTIC REGRESSION 

 Logistic regression is the most common multivariate 

statistical technique used when the outcome variable is 

binary (i.e., there are two possible outcomes). In the 

NIOSH multiple-seam study, the two possible outcomes 

are either “successful” or “unsuccessful.” 

 Logistic regression has much in common with linear 

regression. In both cases, the goal is to predict the outcome 

as a linear combination of the predictive variables. With 

linear regression, the method of least squares is used to 

estimate the unknown parameters (or coefficients, or 

slopes). The analogous function in logistic regression is 

called the method of maximum likelihood. According to 

Hosmer and Lemeshow [2000], “in a very general sense, 

the method of maximum likelihood yields values for the 

unknown parameters which maximize the probability of 

obtaining the observed set of data.” The method of maximum likelihood solves for the “logit” g(x) of the logistic 

regression model as: 

 g(x) = B0 + B1x1 + B2x2 + ... + Bnxn (4) 

where a “B” represents a coefficient and an “x” represents 

a value of a variable in the data set. The more positive the 

value of g(x) is for an individual case, the greater the 

likelihood of a positive (successful) outcome; the more 

negative, the greater the likelihood of a negative 

(unsuccessful) outcome. 

 The process of model-building with logistic regression 

is also similar to that with linear regression. In general, the 

goal is to obtain the model that best explains the data with 

the fewest variables. Logistic regression software provides 

standard deviations and significance levels for each individual coefficient, which allows their contribution to the 

overall outcome to be assessed. Variables with little 

significance can be eliminated from the model. 

 It is important that the effect of any continuous 

variables be “linear in the logit.” When the relationship 

between a variable and the logit is nonlinear (i.e., 

quadratic, logarithmic, binary, or something else), then the 

variables can be transformed so that the transformed variable does have a linear effect. It is, of course, essential that 

the transformed variable make scientific sense. 

 When logistic regression is used to classify cases into 

two groups, a plot of sensitivity versus specificity is used. 

Sensitivity is the probability of correctly identifying a 

positive (or successful) case, while specificity is the probability of correctly identifying a negative (or unsuccessful) 

case. Where the two curves cross, the likelihood of 

detecting a false positive (1-sensitivity) equals that of 

detecting a false negative (1-specificity). This is considered the optimal “cut-point” [Hosmer and Lemeshow 

2000]. Mathematically, the selection of the cut-point 

affects the value of the B0 term in Equation 4. 

 One disadvantage of logistic regression is that there is 

no universally accepted measure of the overall model 

goodness-of-fit analogous to the r-squared used in linear 

regression. One of the most recommended measures is the 

“area under the ROC curve.” The ROC, or receiver 

operating characteristic, is a concept borrowed from signal 

detection theory. It plots the sensitivity against 1-specificity for an entire range of possible cut-points [Hosmer 

and Lemeshow 2000]. As a general rule: 

• If ROC = 0.5, there is no discrimination (same as 

flipping a coin); 

• If 0.7 < ROC < 0.8, discrimination is considered 

acceptable; 

• If 0.8 < ROC < 0.9, discrimination is considered 

excellent; and 

• If ROC > 0.9, discrimination is outstanding. 

 All of the statistical analyses conducted for the 

NIOSH multiple-seam study were performed using the 

Stata statistical package [StataCorp 2005]. 

STATISTICAL ANALYSIS 

 With such a large database and so many variables, the 

statistical analysis was a complex iterative process. 

Numerous logistic regression models were tested and 

evaluated. As more refined models were developed with 

new combinations of variables, it was often necessary to 

test parameters that had previously been excluded to 

ensure that they were still nonsignificant. Eventually, the 

process arrived at a single design equation. 

 The first step in the analysis was to weight the cases. 

Weighting was necessary because some mines provided a 

large number of case histories, whereas others provided 

only a few. To fairly represent all of these cases without 

allowing the database to be overwhelmed by a few mines 

0.00

1.00

2.00

3.00

4.00

250 500 750 1000 1250 1500 1750 2000

Depth of Cover (ft)

Actual ARMPS SF

Successes

Failures

Suggested

ARMPS SF

 Figure 10.—ARMPS SF for the case histories compared with the suggested values of the ARMPS SF. 
21 

that contributed many cases, the following weighting equation was used: 

N m

1

case weight = (5) 

where Nm = the total number of cases from this mine. 

In other words, the more cases there were from an individual mine, the smaller the weight of each individual 

case, but the greater the weight of the mine’s total 

experience. 

 Another important issue is that of correlations 

between variables within the database. In general, it does 

not make sense to include two variables that are highly 

correlated with each other in the same analysis. In the 

NIOSH database, for example, the TVS, multiple-seam 

stress, and depth of cover are all highly correlated with 

each other. Logistic regression trials indicated that of these 

three, the TVS gave the best overall results. 

 Several continuous variables were tested for the 

linearity of their effects on the outcome. Ultimately, 

logarithmic transformations were applied to the thickness 

of the interburden and the CMRR, although linear versions 

of the variables were also retained. Using the natural log of 

the interburden thickness eliminates the possibility that in 

some low-stress undermining cases the logistic equation 

might predict a negative value for the critical interburden 

thickness. On the other hand, when the log of the interburden is used, the assessment of extreme high-stress overmining scenarios can result in unreasonably large values 

for the critical interburden thickness. 

 In the case of the CMRR, the evidence in the data was 

not clearly in favor of a logarithmic transformation, perhaps because of the relatively narrow range of CMRRs in 

the database. However, the scientific logic is that increasing the CMRR from 40 to 50 has a greater effect on stability than an increase from 80 to 90. In effect, a logarithmic 

transformation implies that it is the percent change in the 

value of a variable that matters, not the absolute value of 

the change. The transformed variable that was used in the 

multiple-seam analysis was: 

 lnCMRR20 = ln (CMRR-20) (6) 

 Another reason for using the transformed CMRR variable in the analysis is that there were so few cases where 

the CMRR was less than 45. Since the transformed variable implies that the effect of multiple-seam interactions is 

amplified when the roof is weak, using it means that the 

ultimate design equation is more conservative for the low 

range of CMRR values. 

 One other transformation that was applied was to the 

angle at which the active mining intercepted the remnant 

structure. The transformed variable was binary, with 

angles greater than 20° given a value of “perpendicular” 

while the others were “parallel.” This variable was not 

significant in the final model, however. 

 In designing the analysis, a key issue was whether the 

undermining, overmining, development, and retreat cases 

should be analyzed separately or together. There are good 

scientific arguments for separating them, since the 

mechanics of the interactions may be different in the different groups. The disadvantage is that the four databases 

would each be much smaller. As discussed in the next section, the analyses showed that all four groups could be 

combined and analyzed together. 

 When the values for both the CMRR and the 

interburden competence (expressed as the percent of 

strong rock in the interburden) were obtained from the 

borehole nearest the case history, the analysis showed 

that neither variable was statistically significant. When 

mine-wide averages were employed instead, the CMRR 

was highly significant, and it is included in the final 

design equation. 

 Interburden competence was still not significant, 

however. Two related factors may have contributed to this: 

• The percent of competent rock was based entirely on 

the geologic descriptions included with the core 

logs. In many cases, the description was little more 

than the rock type (shale, sandstone, etc.). In the 

central Appalachian coalfields, however, some siltstones and even shales can be very strong [Rusnak 

and Mark 2000]. Without an actual geotechnical 

description, some weak rocks may have been labeled 

strong and vice versa. 

• Because the case histories are all from two coalfields 

where the rocks tend to be strong, there may not be 

sufficient variability in the database to capture the 

effect of interburden competence. 

 Two other variables that were expected to be significant for the overmining cases were the time lag since 

mining the bottom seam and the lower coalbed-tointerburden thickness ratio. However, the analysis did not 

find that either was significant. In the case of the time lag, 

the database contained a total of 12 overmining cases in 

which the time lag was less than 10 years. Of these, all but 

two were successes, indicating that time lag by itself is 

unlikely to be a major factor. However, one of the two 

failures proved to be a major outlier when compared with 

the rest of the database. It seems quite likely, in this 

instance at least, that the settling time was important. 

 The lack of influence of the lower coalbed-to-interburden thickness ratio may also be due to the sample size. 

There are 30 cases (21 development and 9 retreat) in which 

the interburden thickness was 7.5–10 times the lower coalbed thickness. Of these, 13, or 44%, are failures, which is a 

relatively high failure rate. However, the effect may be 

captured by other variables, particularly the interburden 
22 

thickness, which was less than 50 ft in all but one of these 

cases. It seems likely that the upper-seam mining in these 

30 cases probably took place in the fracture zone, above 

the top of the caving zone, which is normally 6–10 seam 

heights above the lower bed [Kendorski 2006]. It may be 

that once the upper seam is above the caving zone, the 

lower coalbed-to-interburden thickness ratio may not be 

significant. However, since all of these cases (except one) 

come from just two mines in Virginia, it is possible that 

more trouble might be encountered in other geologic 

environments. 

 Retreat mining was another factor that was not significant in the final analysis. However, the effect of retreat 

mining is indirectly included in the TVS variable. On 

average, the vertical stress was 20% greater in the retreat 

cases than in the development cases. 

FINAL MODEL AND DESIGN EQUATION 

 Of the 344 cases in the database, the outcomes in 9 of 

them were considered “borderline,” and these were 

excluded. An additional 26 failed cases were excluded 

because their ARMPS SFMS were less than the recommended values. It was believed that the poor mining 

conditions in these cases were likely attributable to inadequate pillar design rather than multiple-seam interaction 

per se. As a result, the final database included 309 case 

histories. 

 The final, best model is given below: 

 g(x) = –0.81*TVS + 1.79*UO + 0.0233*INT + 2.02*EX 

 – 1.80*REMPIL + 1.95*lnCMRR20 – 6.47 (7)

where TVS = total vertical stress on the critical 

 pillar (thousands of psi); 

 UO = 1 for undermining, 0 for overmining; 

 INT = interburden (ft); 

 EX = 1 for extra support, 0 for none; 

 REMPIL = 1 for isolated remnant pillar, 0 for 

 gob-solid boundary; 

and lnCMRR20 = ln (CMRR-20). 

 The logistic regression table for this model is shown 

in Table 1. It indicates that all of the parameters are 

statistically significant at well above the 99% confidence 

level. 

 Figure 11 shows that the ROC = 0.883 for this model. 

The model’s optimal cut-point is approximately p=0.86, 

as shown in Figure 12. The classification table (Table 2) 

indicates that this model correctly classifies approximately 

80% of the cases overall, including 81% of the failures, 

when this cut-point is used. 

Table 2.—Performance of the design equation 

against the NIOSH database 

Model predictions No. of 

cases 

% of total 

cases 

Successes correctly predicted ....................... 212.2 68.7 

Failures correctly predicted............................ 36.2 11.7 

Total cases correctly predicted ...................... 248.4 80.4 

Failures predicted as successes.................... 9.6 3.1 

Successes predicted as failures..................... 51.0 16.5 

Total cases incorrectly predicted.................... 60.6 19.6 

 Total cases................................................. 309 100 

Table 1.—Logistic regression table for final model (Equation 7) 

Variable Coef. Std. err. z P > z 95% confidence interval 

TVS........................ –0.80655 0.125674 –6.42 0.000 –1.05287 –0.56023 

UO......................... 1.79438 0.21375 8.39 0.000 1.37543 2.21333 

INT........................ 0.02331 0.00229 10.16 0.000 0.01881 0.02780 

EX......................... 2.02037 0.27899 7.24 0.000 1.47356 2.56719 

REMPIL.................. –1.79952 0.23288 –7.73 0.000 –2.25596 –1.34307 

InCMRR20............... 1.95214 0.47521 4.11 0.000 1.02074 2.88354 

Constant.................. –6.467325 1.805228 –3.58 0.117 –10.0055 –2.92914 0.00 0.25 0.50 0.75 1.00 Sensitivity

0.00 0.25 0.50 0.75 1.00

1 - Specificity

Area under ROC curve = 0.8827

 Figure 11.—Receiver operating characteristic (ROC) 

curve for Equation 7. 
23 

 A version of Equation 7 using the log transformation 

of the interburden was also determined. The lnINT model 

was very similar to Equation 7, but its ROC was slightly 

less at 0.873. 

 Table 3 compares the model derived from the overall 

data set (Equation 7, shown as “Linear-all” in Table 3) 

with the lnINT model (“Log-all” in Table 3) and models 

using subsets consisting of just the undermining, overmining, development, or retreat cases. It is remarkable that 

there is so little variation in the coefficients (b) among 

these different models. The influence of the interburden 

thickness, for instance, varies by no more than ±10%. The 

effects of EX, REMPIL, and lnCMRR20 are also very 

stable. The ROC values for all four of these models are 

also excellent. Overall, these results support the use of a 

single universal equation to predict the severity of 

multiple-seam interactions. 

 For use in design, Equation 7 was adjusted to correspond to the cut-point of 0.86 and transformed to predict the 

critical interburden thickness (INTcrit, ft): 

 INTcrit = 35*TVS – 77*UO – 87*EX + 77*REMPIL 

 – 83*(lnCMRR20) + 359 (8) 

 One disadvantage of Equation 8 is that in some 

extreme cases it can predict a negative value of the critical 

interburden thickness. Therefore, the lnINT model was 

used to derive a second equation for the critical interburden thickness INTcritLN : 

INTcritLN = EXP[0.35*TVS – 0.74*UO – 0.99*EX 

 + 0.74*REMPIL – 0.91*(lnCMRR20) + 7.23] (9) 

 Figure 13 shows that, in general, INTcritLN > INTcrit

when the interburden is less than about 50 ft or greater 

than about 170 ft. Since the linear interburden model fits 

the data slightly better and since it provides more conservative answers where the data are sparse (when the 

interburden falls between 90 and 150 ft), Equation 8 was 

preferred for most situations. The log interburden model is 

preferred only for the thinnest interburdens, where Equation 9 provides the most conservative value for the critical 

interburden. 

 When the actual interburden thickness exceeds the 

INTcrit determined by Equation 8, there is a very good 

likelihood that conditions will be satisfactory. The data in 

Table 2 and Figure 14 indicate that within the NIOSH 

database only 4% of the cases (9.6 of 221.8 weighted 

cases) where the interburden exceeded the critical value 

defined by Equation 8 were failures. Of the cases where 

the actual interburden was less than the critical value, the 

multiple-seam interaction resulted in unsatisfactory conditions 43% of the time. 

Table 3.—Comparison between logistic regression models 

Variable Linear-all 

coef. (b) 

Log-all 

coef. (b) 

Undermining 

coef. (b) 

Overmining 

coef. (b) 

Development 

coef. (b) 

Retreat 

coef. (b) 

TVS.............................. –0.81 –0.80 –0.61 –1.14 –1.13 –0.43 

UO............................... 1.79 1.69 NA NA 2.48 0.40 

INT.............................. 0.0233 NA 0.024 0.024 0.024 0.026 

EX............................... 2.02 2.26 1.64 2.45 1.83 2.56 

REMPIL........................ –1.80 –1.68 –1.74 –2.05 –1.82 –1.98 

lnCMRR20..................... 1.95 2.07 1.72 2.59 2.71 0.61 

Constant....................... –6.47 –14.72 –0.91 –3.82 –10.07 1.13 

No. of cases.................. 309 309 234 75 237 72 

NA Not applicable. 

0.00

0.25

0.50

0.75

1.00

0.00 0.25 0.50 0.75 1.00

Probability cutoff

Sensitivity Specificity

Sensitivity/Specificity

 Figure 12.—Plot of specificity versus sensitivity for 

Equation 7, showing that the optimal cut-point is 

approximately p=0.86. 
24 

 Equation 8 also indicates that: 

• Each additional 1,000 psi of vertical stress is equivalent to subtracting 35 ft of interburden. 

• All other factors being the same, overmining 

requires 77 more feet of interburden than undermining. 

• All other factors being the same, an isolated remnant 

pillar requires 77 more feet of interburden than a 

gob-solid boundary. 

• All else being equal, a CMRR = 45 roof requires 

approximately 50 more feet of interburden than a 

CMRR = 65 roof. 

 The analysis indicates that installing a pattern of cable 

bolts or other heavy supplemental support is equivalent to 

adding as much as 87 ft of interburden. However, while 

supplemental support may make mining possible, the 

likelihood of encountering rib spalling, floor heave, or 

hazardous roof also increases when the analysis suggests 

that supplemental support is necessary. 

 In Figure 15 the case histories are plotted again, but 

this time each point is plotted with its suggested overburden for the “no extra support” (EX=0) condition. Three 

regions are defined on the graph. The uppermost region, 

where the actual interburden exceeds the critical interburden when EX=0, is labeled “Predicted Successes.” 

Within this region, 97% of the case histories that maintained an adequate pillar SF were successful. In the middle 

region, success is predicted only if a pattern of supplemental support is installed. In this region, 93% of the cases 

that did install supplemental support were successful, 

whereas only 63% of those that did not were successful. In 

the bottom region, where failure is predicted, only 52% of 

the cases were successful. 

 Figures 16–19 are “design charts” that illustrate the 

critical interburden thickness for CMRR = 45 and 

CMRR = 65 situations, for undermining and overmining, 

with and without extra support. 

 Figure 13.—Comparison of Equations 8 and 9, showing that the logarithmic version (Equation 9) can yield 

the most conservative value of the critical interburden 

when the calculated interburden is less than about 50 ft. 

0

50

100

150

200

250

1234567

Vertical Stress (ksi)

Critical Interburden (ft)

Linear Equation (8)

Log Equation (9)

 Figure 14.—Performance of the design equation, comparing the suggested critical interburdens with the actual 

interburdens in the case histories. 

0

50

100

150

200

250

300

0 50 100 150 200 250 300

Suggested Interburden (ft)

Actual interburden (ft)

Successes

Failures

Fails: SF

Predicted

Successes

Predicted

Failures

 Figure 16.—Suggested values of the critical interburden for undermining a gob-solid boundary. 

0

40

80

120

160

200

1 1.5 2 2.5 3 3.5 4 4.5 5

Vertical Stress (ksi)

Critical Interburden (ft) .

CMRR=45 ExSupp=0

CMRR=45 ExSupp=1

CMRR=65 ExSupp=0

CMRR=65 ExSupp=1

Check for Ultra-Close Interactions when I<30 ft

Failures: No Extra Support Successes: No Extra Support

Success: With Extra Support Failures: With Extra Support

 Figure 15.—Suggested critical interburden values setting 

EX=0.
25 

LIMITATIONS OF THE STUDY 

 Every research study has limitations that need to be 

kept in mind when the findings are applied to the real 

world. One of the advantages of statistical modeling is that 

the process is relatively transparent, and the limitations are 

not hidden within the “black box.” Although the NIOSH 

multiple-seam database is the largest of its kind ever 

collected, it is subject to at least three major limitations: 

• Data quality: Many of the parameters included in 

the multiple-seam database, such as the depth of 

cover and the interburden thickness, can be determined very reliably for each individual case history. 

However, there is some uncertainty about critical 

parameters such as the CMRR, multiple-seam stress, 

interburden geology, and even (in some cases) the 

outcome (success or failure). The reasons for these 

uncertainties have been discussed earlier. Fortunately, the larger the database, the less influence that 

possible measurement errors will have on the overall 

trends. 

• Number of parameters: The size of the database also 

limits the number of parameters that can be included 

in the model. The NIOSH database seems to be large 

enough to estimate the effects of the main parameters, as reflected by the stability of the model when 

different portions of the database are analyzed (see 

Table 3). However, it may not have been large 

enough to capture some second-order effects, such 

as the orientation of the panel relative to the remnant 

structure. The database was also not large enough to 

analyze interactions between the variables. For 

example, some analyses indicated that extra support 

is more beneficial to low CMRR roof than to strong 

roof, but this trend was not statistically significant. 

• Nonrandom sampling: The NIOSH database 

includes a large percentage of the recent multipleseam interactions that have occurred in U.S. 

underground mines, but even if it included every 

single one it would still suffer from sampling 

limitations. This is because the actual mines 

represent a tiny fraction of all the combinations of 

parameters that could exist. One example that has 

already been discussed is the lack of case histories 

with interburdens between 90 and 150 ft. There are 

also many combinations of parameters that are not 

represented, particularly among the smaller group of 

failed cases. 

 The most important limitation to remember, however, 

is the range of the data. Empirical models should always 

be used with caution outside the range of the data that were 

employed in their development. In this instance, it is 

possible that new failure mechanisms may occur when the 

CMRR is less than 45, the interburden is less than 30 ft 

 Figure 17.—Suggested values of the critical interburden for undermining an isolated remnant pillar. 

0

50

100

150

200

250

300

1 1.5 2 2.5 3 3.5 4 4.5 5

Vertical Stress (ksi)

Critical Interburden (ft) .

CMRR=45 ExSupp=0

CMRR=45 ExSupp=1

CMRR=65 ExSupp=0

CMRR=65 ExSupp=1

Check for Ultra-Close Interactions when I<30 ft

 Figure 18.—Suggested values of the critical interburden for overmining a gob-solid boundary. 

0

50

100

150

200

250

300

1 1.5 2 2.5 3 3.5 4 4.5 5

Vertical Stress (ksi)

Critical Interburden (ft) .

CMRR=45 ExSupp=0

CMRR=45 ExSupp=1

CMRR=65 ExSupp=0

Check for Caving Zone Interactions when I<10 h

 Figure 19.—Suggested values of the critical interburden for overmining an isolated remnant pillar. 

0

50

100

150

200

250

300

350

1 1.5 2 2.5 3 3.5 4 4.5 5

Vertical Stress (ksi)

Critical Interburden (ft) .

CMRR=45 ExSupp=0

CMRR=45 ExSupp=1

CMRR=65 ExSupp=0

Check for Caving Zone Interactions when I<10 h
26 

thick, or an upper seam is developed within the caving 

zone of a previously mined lower seam. Checking this 

range limitation is therefore an important step in the 

suggested procedures for multiple-seam analysis that are 

discussed below. 

PROCEDURES FOR MULTIPLE-SEAM 

MINE DESIGN 

 The results of this study can be used to evaluate the 

potential for multiple-seam interactions and provide 

guidance for pillar sizing, supplemental support, and other 

aspects of mine design. The suggested step-by-step process 

follows: 

1. Identify critical remnant structures on the maps of 

mining in seams above and below the target seam. 

Every remnant structure that may be crossed by 

active mine workings should be evaluated. 

2. For each potential remnant structure crossing, 

determine the— 

• Depth of cover to the target seam 

• Interburden thickness 

• Seam heights (both seams) 

• Age of the older workings 

• CMRR for the roof of the target seam 

3. Check that the parameters of the case being considered fall within the limits of the NIOSH 

multiple-seam database. If the roof is very weak 

(CMRR < 45) or the stress is very high (>5,000 

psi), then the equations should be used with 

caution. The same is true if the case involves 

overmining and the lower coalbed thickness-tointerburden ratio is less than 10. If the interburden 

thickness is less than 30 ft in either undermining or 

overmining, then potential for an ultraclose interaction should be given primary consideration. 

4. Determine whether the remnant structure is a gobsolid boundary or an isolated remnant pillar. Equation 1 or Figure 20 may be used if the remnant 

structure is a pillar. If the remnant pillar is so thin 

that it may have failed completely, it may be helpful to determine its ARMPS SF. 

5. Determine the ARMPS SF or ALPS SF (single 

seam) for the proposed section in the target seam. If 

retreat mining is planned (either pillar recovery or 

longwall mining), determine the maximum abutment stress applied to the critical pillar. 

6. Conduct a LaM2D analysis of the remnant structure 

crossing to determine the multiple-seam stress 

applied to the critical pillar in the target seam. Also 

determine the TVS using Equation 3. 

7. Determine the ARMPS or ALPS multiple-seam 

stability factor (SFMS) for the target seam using 

Equation 2 and compare it to the recommended 

ARMPS or ALPS SF. If the calculated SFMS is 

lower than the recommended value, then the pillar 

size should be increased. 

8. Use Equation 8 to determine the critical interburden 

thickness, both with and without supplemental 

support (EX=1 and EX=0). Note that if the INTcrit

determined from Equation 8 is less than 60 ft, then 

Equation 9 should be used if it provides a more 

conservative value. 

9. Compare the actual interburden to the two INTcrit

values determined in step 8. Three cases are 

possible: 

• If INTcrit without supplemental support is significantly less than the actual interburden, then 

a major multiple-seam interaction can be considered unlikely. 

• If the actual interburden is less than INTcrit

without supplemental support but greater than 

INTcrit with supplemental support, then adding 

a pattern of cable bolts or other equivalent supplemental support could greatly reduce the 

probability of a major interaction. 

• If INTcrit even with supplemental support is 

greater than the actual interburden thickness, 

then a major interaction should be considered 

likely with this design. 

10. If desired, the mine design can be adjusted by 

changing the pillar size or the entry width (both of 

which affect the TVS) and then repeating steps 5–9. 

CONCLUSIONS 

 To conduct this study, NIOSH collected the largest 

database of multiple-seam case histories ever assembled. 

These data were analyzed with the multivariate statistical 

technique of logistic regression. The study also employed 

LaM2D to estimate the multiple-seam stress, ALPS and 

ARMPS to determine pillar SFs, and the CMRR to measure roof quality. 

400

600

800

1000

1200

1400

1600

0 20 40 60 80 100 120 140 160 180 200

Pillar Width (ft.)

Depth of Cover (ft.)

Maximum width of 

isolated remnant

Minimum width range 

of isolated remnant

 Figure 20.—Upper and approximate lower limits for 

determining whether a remnant structure is an isolated 

remnant pillar or a gob-solid boundary. 
27 

 Several of the study’s findings confirm the conventional wisdom about multiple-seam interactions. Overmining was found to be much more difficult than 

undermining, and isolated remnant pillars caused more 

problems than gob-solid boundaries. For the first time, 

however, it was possible to quantify these effects in terms 

of the equivalent thickness of interburden needed to compensate for them. 

 The study also found that pillar design is a critical 

component of multiple-seam mine planning. Many of the 

failed cases involved pillars whose SF seemed inadequate 

once the multiple-seam stresses were accounted for. 

Weaker roof was also found to significantly increase the 

risk of multiple-seam interactions. Some factors that were 

not found to be statistically significant included the 

interburden competence, the time lag between mining the 

two seams, the lower coalbed-to-interburden thickness 

ratio, and the angle between the active mining and the 

remnant structure. 

 The most important result of the study is an equation 

that predicts the critical thickness of the interburden 

required to minimize the likelihood of a multiple-seam 

interaction. This equation has been incorporated into a 

step-by-step methodology that allows mine planners to 

evaluate each potential interaction and take steps to reduce 

the risk of ground control failure. Such measures could 

include installing cable bolts or other supplemental support, increasing the pillar size, or avoiding the remnant 

structure entirely. 

REFERENCES 

 Akinkugbe OO, Heasley KA [2007]. The new twodimensional LaModel program (LaM2D). Paper in these 

proceedings. 

 Chase FE, Mark C, Heasley KA [2002]. Deep cover 

pillar extraction in the U.S. coalfields. In: Peng SS, Mark C, 

Khair AW, Heasley KA, eds. Proceedings of the 21st 

International Conference on Ground Control in Mining. 

Morgantown, WV: West Virginia University, pp. 68-80. 

 Colwell M, Mark C [2005]. Analysis and design of rib 

support (ADRS): a rib support design methodology for 

Australian collieries. In: Peng SS, Mark C, Tadolini SC, 

Finfinger GL, Khair AW, Heasley KA, eds. Proceedings of 

the 24th International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 12–22. 

 Colwell M, Frith RC, Mark C [1999]. Calibration of 

the Analysis of Longwall Pillar Stability (ALPS) chain 

pillar design methodology for Australian conditions. In: 

Peng SS, Mark C, eds. Proceedings of the 18th International Conference on Ground Control in Mining. 

Morgantown, WV: West Virginia University, pp. 282–290. 

 Ellenberger JL, Chase FE, Mark C, Heasley KA, 

Marshall JK [2003]. Using site case histories of multipleseam coal mining to advance mine design. In: Peng SS, 

Mark C, Khair AW, Heasley KA, eds. Proceedings of the 

22nd International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 59–64. 

 Heasley KA, Agioutantis ZG [2007]. LaModel: 

a boundary-element program for coal mine design. Paper 

in these proceedings. 

 Hosmer DW, Lemeshow S [2000]. Applied logistic 

regression. 2nd ed. New York: John Wiley and Sons. 

 Kendorski FS [2006]. Effect of full-extraction underground mining on ground and surface waters: a 25-year 

retrospective. In: Peng SS, Mark C, Finfinger GL, Tadolini 

SC, Khair AW, Heasley KA, Luo Y, eds. Proceedings 

of the 25th International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 425-429. 

 Mark C, Chase FE [1997]. Analysis of retreat mining 

pillar stability (ARMPS). In: Mark C, Tuchman RJ, eds. 

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 and Prevention, National Institute for 

Occupational Safety and Health, IC 9446, pp. 17–34. 

 Mark C, Chase FE, Molinda GM [1994]. Design of 

longwall gate entry systems using roof classification. In: 

Mark C, Tuchman RJ, Repsher RC, Simon CL, eds. 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. 5–17. 

 Mark C, McWilliams LJ, Pappas DM, Rusnak JA 

[2004]. Spatial trends in rock strength: can they be 

determined from coreholes? In: Peng SS, Mark C, 

Finfinger GL, Tadolini SC, Heasley KA, Khair AW, eds. 

Proceedings of the 23rd International Conference on 

Ground Control in Mining. Morgantown, WV: West 

Virginia University, pp. 177–182. 

 Mark C, Molinda GM, Dolinar DR [2001]. Analysis 

of roof bolt systems. In: Peng SS, Mark C, Khair AW, eds. 

Proceedings of the 20th International Conference on 

Ground Control in Mining. Morgantown, WV: West Virginia University, pp. 218–225. 

 Rusnak J, Mark C [2000]. Using the point load test to 

determine the uniaxial compressive strength of coal 

measure rock. In: Peng SS, Mark C, eds. Proceedings of 

the 19th International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 362–371. 

 StataCorp [2005]. Stata statistical software: release 9. 

College Station, TX: StataCorp LP. 

 Stemple DT [1956]. A study of problems encountered 

in multiple-seam mining in the eastern United States. Bull 

Va. Polytech Inst 49(5):65. 

29

LAMODEL: A BOUNDARY-ELEMENT PROGRAM FOR COAL MINE DESIGN 

By Keith A. Heasley, Ph.D.,1 and Zacharias G. Agioutantis, Ph.D.2

ABSTRACT 

 LaModel 2.1 is a PC-based boundary-element program for calculating the stresses and displacements in coal 

mines or other thin, tabular seams or veins. This type of 

mine modeling software can be used by mine design engineers in the industry to investigate and optimize the pillar 

sizes and pillar layouts in relation to pillar stress, multiseam stress, and/or bump potential (energy release). This 

paper introduces the LaModel program and highlights the 

Windows-based pre- and postprocessors. 

INTRODUCTION 

 Mine planners have a variety of modeling methods, 

either empirical or numerical, available for analyzing pillar 

stresses and determining safe pillar sizes for various mine 

geometries and geologic structures. Empirical methods 

emphasize the collection and interpretation of case histories of pillar performance. These empirical methods are 

closely linked to reality, and for many “typical” mining 

geometries, they work extremely well. 

 However, it is difficult to apply these empirical 

methods to mining situations beyond the scope of the 

original empirical database. Therefore, when complicated 

stress conditions arise from complex single- or multipleseam mining geometries, numerical modeling techniques 

such as finite-element, boundary-element, discreteelement, or finite-difference are usually applied. In 

general, the numerical, or analytical, design methods are 

derived from the fundamental laws of force, stress, and 

elasticity. Their main advantage is that they are very 

flexible and can quickly analyze the effect of numerous 

geometric and geologic variables on mine design. Their 

main disadvantage is that they require difficult-to-obtain 

and/or unknown information about material properties, 

failure criteria, and postfailure mechanics. 

 In order to analyze the displacements and stresses 

associated with the extraction of large tabular deposits 

such as coal, potash, and other thin, vein-type deposits, 

the displacement-discontinuity variation 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. Using this 

technique, only the planar area of the seam needs to be 

discretized, or gridded, in order to obtain the stress and 

displacement solution on the seam. Often, this limited 

analysis is sufficient, since in many practical applications 

only the distributions of stress and convergence on the 

seam horizon are of interest. Also, by limiting the detailed 

analysis to only the seam, the displacement-discontinuity 

method provides considerable computational savings over 

other techniques, which discretize the entire body (such as 

finite-element, discrete-element, or finite-difference). It is 

a direct result of this computational efficiency that the 

displacement-discontinuity method is able to handle large 

areas of tabular excavations as needed in many practical 

coal mining problems. 

 A displacement-discontinuity program incorporating a 

laminated media was originally developed by the National 

Institute for Occupational Safety and Health [Heasley 

and Chekan 1999]. This program, called LaModel, was 

designed for calculating the stresses and displacements in 

coal mines or other thin, tabular seams in layered media. In 

order to facilitate the transfer of the LaModel technology 

to the mining industry for improving overall mine design, 

an intuitive, easy-to-use preprocessor, LAMPRE, and postprocessor, LAMPLT, were also developed. The LAMPRE 

program handles all of the numerical parameter input and 

allows the mine plan to be graphically entered into the 

program. The new postprocessor, LAMPLT, also uses an 

intuitive graphical interface and allows the user to quickly 

and easily plot and analyze the output data from the 

numerical calculation phase of LaModel. 

LAMODEL

 Traditional displacement-discontinuity programs use a 

homogeneous isotropic elastic formulation that simulates 

the overburden as one solid material. In contrast, the 

LaModel program simulates the geologic overburden 

stratifications as a stack of layers with frictionless interfaces. Specifically, each layer is homogeneous isotropic 

elastic and has the identical elastic modulus, Poisson’s 

ratio, and thickness. This “homogeneous layering” formulation does not require specifying the material properties 

for each individual layer, yet it still provides a realistic 

suppleness to the mining overburden that is not possible 

with the classic homogeneous, isotropic elastic overburden 

model. From our experience, this suppleness provides 

 

 1Associate professor, West Virginia University, Department of

Mining Engineering, Morgantown, WV. 

 2

Professor, Department of Mineral Resources Engineering,

Technical University of Crete, Greece. 
30

more accurate strata response than a homogeneous overburden for modeling local deformations, interseam interactions, and/or surface subsidence [Heasley and Barton 

1999; Heasley and Chekan 1999; Heasley and Salamon 

1996]. 

 To utilize the LaModel program, the user must input 

the seam geometry, seam orientation, stress field, and 

geologic material properties. Up to 26 different in-seam 

materials from 5 different material models (elastic, elasticplastic, strain-softening, bilinear strain-hardening, and 

exponential strain-hardening) can be used in a model, and 

the grid size can be a maximum of 1,000 by 1,000. Once 

the input is developed, the LaModel program calculates the 

stresses and displacements at the seam level and at 

requested locations in the overburden or at the surface. The 

program also has the ability to analyze (1) the interseam 

stresses resulting from multiple-seam mining, (2) effects of 

topographic relief on pillar stress and gob loading, 

(3) stress changes during mining through multiple mining 

steps, and (4) surface subsidence. 

LAMPRE: THE PREPROCESSOR 

 In order to facilitate the transfer of the LaModel 

technology to the mining industry for improving overall 

mine design, the requirements on the occasional user for 

data input need to be simplified as much as possible. To 

this end, an intuitive, easy-to-use preprocessor, LAMPRE, 

that allows the user to quickly and easily enter both the 

required material data and the necessary model geometric 

data was initially developed. LAMPRE runs in the 

Windows environment using an intuitive graphical user 

interface with ample context-sensitive online help. The 

preprocessor handles all of the numerical parameter input 

and also allows graphical input of the material codes for 

the seam grids or automatic grid generation from an 

AutoCAD mine map. Also, a set of “material wizards” for 

automatically calculating reasonable coal and gob 

properties from typical values was added to LAMPRE. 

These wizards greatly reduce the burden on the novice user 

of specifying the essential material properties for the seam 

and gob. 

 A typical LAMPRE input form for parameter values is 

shown in Figure 1 (in this case, the form for the “General 

Model Information”). In this form, the input parameters 

are entered using either a text edit, slider, or radio button 

object, whichever is most convenient. If help on the 

parameter is needed, simply pressing the “F1” function 

key, while the parameter is active, will open the 

appropriate help topic. Also, the program checks the value 

of all of the input parameters before they are saved to 

make sure that they fall within reasonable prescribed 

ranges. Using a number of these input forms, all of the 

geometric and geologic parameters for the LaModel 

program are easily entered. 

 Once the numerical parameters are entered, the user 

then needs to define the geometry and material of the 

pillars and openings in the seam grid. In LaModel, the 

different materials are represented by alphabetical 

“material codes” (A to Z), and openings are represented by 

the character “1”. LAMPRE provides a graphical editor for 

creating and modifying the various seam grids, as shown 

in Figure 2. The grid editor essentially works like a 

spreadsheet. Material codes for individual cells may be set 

by clicking on a cell and then typing and entering an 

appropriate letter from A to Z. Blocks of cells may be 

changed by clicking and dragging the mouse over a range 

and then typing and entering the material code, or entire 

rows/columns may be selected and changed using the row/

column number. Also, a block of cells may be copied and 

pasted (greatly simplifying the creation of numerous 

pillars). Finally, the grid editor allows for automatic yield 

zone generation based on materials that were automatically 

defined using the coal property wizard. 

 The grid editor in LAMPRE greatly speeds the 

creation of “regular” rectangular pillar plans aligned with 

the grid. However, even with the graphical grid generation, 

if the mining plan is complex or the grid does not align 

with the pillars, then it can take hours or even days to 

meticulously determine and specify the material code for 

every grid element, especially since the maximum grid size 

has increased to 1,000 by 1,000, or 1 million elements. 

 In response to the difficulty of creating complex mine 

grids, one of the most recent developments in the 

evolutionary history of LaModel is the “automatic” grid 

generator, which can take an AutoCAD map of the mine 

pillars and automatically create the material grid for the 

LaModel input file. Also, a program for automatically 

generating the topographic input grid from an AutoCAD 

file was recently developed. These new automatic grid 

generators greatly reduce the effort and shorten the time 

Figure 1.—Typical parameter input form for LAMPRE. 
31

required to produce a practical model. In most cases, the 

time required to create the mine model can be reduced 

from days and hours to minutes using the automatic grid 

generators. 

 Once the input file is created, the LaModel numerical 

analysis is run in batch mode for calculating the stress and 

displacements at the seam level. Model runs can take several minutes to several days depending on the computer 

speed and model complexity (which includes such factors 

as number of steps, number of seams, and grid size). The 

output from the calculation phase is stored in a data file 

for subsequent analysis by the postprocessing program, 

LAMPLT. 

LAMPLT: THE POSTPROCESSOR 

 The postprocessing program for LaModel, called 

LAMPLT, allows the user to quickly and easily plot and 

analyze the tremendous amounts of output data from the 

calculation phase of LaModel. With LAMPLT, the user 

can interactively examine the seam convergence and the 

vertical seam stress, and the individual components of the 

seam stress from topography, multiple seams, and/or 

surface effects. These output values can be displayed using 

either a pseudo-three-dimensional (3–D) “colored-square” 

plot (Figure 3) or a two-dimensional (2–D) cross-section. 

The program also allows interactive selection of plot types, 

step numbers, scaling, and color. For additional output, 

LAMPLT allows cut-and-paste of the graphics for input to 

other reporting programs. 

EXAMPLE CASE STUDY 

 The example case study site is a longwall mine in 

Greene County, PA, operating in the Sewickley Coal 

Seam. This mine is underlain by an abandoned room-andpillar operation in the Pittsburgh Coal Seam. The primary 

problem at this site was the transfer of multiple-seam stress 

from the lower mine. Yielding of smaller pillars and the 

subsequent transfer of their load to larger pillars in the 

lower seam apparently caused increases in vertical stress in 

the upper seam, which were noticed during development of 

the headgate entries (Figure 4). Severe pillar spalling and 

poor roof conditions were experienced when mining the 

headgate over these large pillars in the lower seam. Mine 

management was concerned these underlying abutment 

pillar stresses would continue to be a problem further inby 

in the headgate and also in the longwall panel because 

there were several areas in the lower seam where similar 

pillar conditions seemed to exist. 

 In order to optimally use LaModel for accurate stress 

prediction at a given mine, the program should first be 

calibrated to the site-specific geomechanics based on 

previously observed stress conditions at that mine. At this 

site, the model properties were calibrated using underground stress mapping [Heasley and Chekan 1999]. Then, 

with the calibrated material properties, LaModel analyses 

were created and run in order to predict areas of potential 

problems within the remaining headgate and the future 

longwall panel. 

 Figure 4 shows two areas of the headgate and 

longwall panel that were modeled using optimized 

properties from the calibration process. These colored 

plots show the interseam stress, which is the additional 

stress on the upper mine due to the lower-seam mining. In 

this figure, Zone 1 covers the upper (inby) part of the 

headgate panel and the first 365 m (1,200 ft) of the 

Figure 2.—Grid editor in LAMPRE. 

Figure 3.—Colored square output in LAMPLT. 
32

longwall panel. Zone 2 covers the lower part of the 

headgate (where the stress problems were first noticed) 

and the last (outby) 330 m (1,100 ft) of the longwall panel. 

In these two zones, the lower mine pillar conditions and 

the overburden depths appear similar; therefore, the poor 

pillar conditions encountered in Zone 2 were expected in 

Zone 1. 

 However, when comparing the interseam stress 

between these two zones, as shown in Figure 4, it is 

obvious that the stress is considerably greater in Zone 2 

than in Zone 1. Closer investigation reveals two main 

reasons for this difference. First, the maximum depth over 

the gate roads and panel in Zone 2 is over 280 m (920 ft), 

whereas in Zone 2, the maximum depth is just over 250 m 

(820 ft). Second, when examining the model output for the 

lower mine, there seems to be less pillar yielding in Zone 1 

than in Zone 2. During headgate development in Zone 1, 

no pillar problems were encountered. Thus, the calibrated 

model successfully predicted the reduced stress conditions 

in the headgate of Zone 1. 

 The mine management was also concerned about the 

multiple-seam stresses adversely affecting the retreating 

longwall panel. In particular, a large, irregularly shaped 

barrier pillar in the lower mine is superimposed under the 

center line of the initial half of the longwall panel in 

Zone 1 (Figure 4). However, the interseam stress calculated by the model from this barrier pillar only reaches 

about 3 MPa (450 psi). When the panel was mined, this 

slightly increased face stress posed very little problem. 

 However, in the lower part of the panel near the 

headgate location where poor ground conditions were first 

encountered (see Zone 2, Figure 4), an area of interseam 

stress up to 9 MPa (1,300 psi) is evident in the panel. 

Because of the underlying barrier pillar, the mine 

anticipated difficult face conditions in this area. Indeed, 

when the longwall face reached this area, ground control 

problems, which included severe face spalling and poor 

roof conditions in the headgate entries, were encountered. 

In fact, the stress interaction with the lower seam was 

severe enough to stop the longwall face about 15 m (50 ft) 

short of the longwall recovery chute and make recovery of 

the supports difficult. 

SUMMARY 

 In this paper, the unique laminated formulation of 

LaModel is detailed and the realistic suppleness this 

provides to the mining overburden is advocated. Also, the 

program’s ability to analyze interseam stresses, topographic stresses, and surface subsidence is promoted. Next, 

the preprocessor, LAMPRE, is described to handle all of 

the numerical parameter input and to allow graphical input 

of the material codes for the seam grids in LaModel. Then, 

the postprocessor, LAMPLT, is presented, and its 2–D and 

3–D plots are illustrated. 

 Finally, in the case study, the numerous features of 

LaModel (including laminated overburden, multiseam 

simulations, strain-softening seam materials, and topographic effects) were used advantageously in simulating an 

actual mining scenario. Once realistic pillar strengths and 

load distributions were established by calibration, the 

mechanics-based overburden behavior in LaModel effectively analyzed the complicated stresses and displacements 

associated with the complex multiple-seam mining scenarios and successfully predicted upcoming high-stress 

conditions in advance of mining for preventive action by 

mine management. 

Figure 4.—Interseam stress for Zone 1 and Zone 2. 

Stress

(MPa)

9

1

3

5

7

-1

200

Scale (m)

0

Zone 2

Zone 1

19

220

0

280

190

2

5

0

220

190

220

250

Headgate
33

REFERENCES 

 Heasley KA, Barton TM [1999]. Coal mine subsidence prediction using a boundary-element program. SME 

preprint 99–181. Littleton, CO: Society for Mining, Metallurgy, and Exploration, Inc. 

 Heasley KA, Chekan GJ [1999]. Practical boundaryelement modeling for mine planning. In: Mark C, Heasley KA, Iannacchione AT, Tuchman RJ, eds. Proceedings 

of the Second International Workshop on Coal Pillar 

Mechanics and Design. Pittsburgh, PA: U.S. Department

of Health and Human Services, Public Health Service, 

Centers for Disease Control and Prevention, National Institute for Occupational Safety and Health, DHHS (NIOSH) 

Publication No. 99–114, IC 9448, pp. 73–87. 

Heasley KA, Salamon MDG [1996]. New laminated 

displacement-discontinuity program. In: Aubertin M, 

Hassani F, Mitri H, eds. Proceedings of the Second North 

American Rock Mechanics Symposium: NARMS ’96. 

(Montreal, Quebec, Canada, June 19–21, 1996), pp. 1879–

1886. 

This paper was previously published as: 

Heasley KA, Agioutantis ZG [2001]. LAMODEL: a boundary-element program for coal mine design. In: 

Proceedings of the 10th International Conference on Computer Methods and Advances in Geomechanics, 

Tucson, AZ, January 7–12, 2001), pp. 1679–1682.

35 

THE NEW TWO-DIMENSIONAL LAMODEL PROGRAM (LAM2D) 

By Olayemi O. Akinkugbe1

 and Keith A. Heasley, Ph.D.2

ABSTRACT 

 The safe, productive exploitation of coal in multipleseam situations requires specific design technology. At 

present, there are two general design approaches that the 

mining engineer can use to analyze the multiple-seam 

problem in question: simple empirical relationships or 

complex numerical models. Often, the empirical 

approaches are too general for the specific problem, and 

the numerical models require too much time and expertise 

for the practicing engineer. In this paper, a new program 

designed for quickly and easily calculating the stresses, 

strains, and safety factors associated with basic multipleseam mining situations is presented. The program 

implements a simplified two-dimensional (2–D) boundaryelement method in order to model the complex multipleseam stress and displacement interactions. The program 

incorporates automatic coal and gob property generation to 

simplify the input and inherently calculates pillar safety 

factors to enhance the output. 

INTRODUCTION 

 Numerical modeling as a tool for engineering design 

has been used for decades. During this time, the awareness 

of its relevance has greatly increased and, at the same time, 

considerable developments have been made at improving 

its accuracy. Presently, there are several general numerical 

techniques for solving engineering problems; one may 

choose to use finite-element, boundary-element, discreteelement, finite difference, and/or a hybrid combination of 

the above to solve the problem at hand. Based on the 

mathematical approximations involved, numerical methods 

can be classified into two categories: domain methods and 

boundary methods. The domain methods are characterized 

by area and volume discretization of the problem domain. 

They use relatively large numbers of elements, resulting in 

large systems of equations that usually consume large 

amounts of computer resources and require lengthy times 

for solution. The boundary-element method, on the other 

hand, requires only the discretization of the boundaries of 

the domain of interest. Relative to subsurface mineral 

extraction, only the planar area of the seam is discretized 

in order to obtain a solution for the desired problem. The 

boundary-element method usually leads to smaller systems 

of equations, faster computing times, and a reduced need 

for computing resources. 

 The displacement-discontinuity method, a version of 

the boundary-element method promoted early by Salamon 

[1963], is usually the method of choice for analyzing 

stresses and displacement distribution in slitlike or thinseam openings. In this approach, the relative movement 

between the roof and floor of a mine is treated mathematically as a discontinuity in the displacement field of the 

surrounding media, in other words, a crack/slit in a continuum. The displacement-discontinuity method has been 

used to create several well-known computer packages for 

solving geomechanical problems in underground mines. 

 The first implementation of the displacementdiscontinuity method into a personal computer code for 

calculating stress and displacement from multiple seams 

was done by the U.S. Bureau of Mines. The program was 

called MULSIM/BM. In 1992, MULSIM/BM was 

upgraded to MULSIM/NL [Zipf 1992a,b] by the addition 

of nonlinear seam material models, calculations for determining the mechanical strain energy changes associated 

with the seam materials, and a capability for “stepping” 

though a series of mining sequences. A new displacementdiscontinuity program called LaModel was later introduced by the National Institute for Occupational Safety 

and Health [Heasley and Chekan 1999]. LaModel implemented the same nonlinear seam material models in 

MULSIM/NL, but introduced a laminated overburden 

model. This model assumes the overburden to be a stack of 

horizontally lying strata, which naturally increases the 

flexibility of the overburden. As a result, LaModel estimation of the stresses, displacements, and surface subsidence 

is considered more accurate compared to the homogeneous 

overburden model implemented in MULSIM/NL. 

 In this paper, a new 2–D boundary-element program, 

LaM2D, is introduced. This program implements the laminated overburden model [Heasley 1998; Salamon 1991, 

1962] used in the three-dimensional (3–D) LaModel program into a greatly simplified 2–D program. This 2–D 

implementation of the laminated boundary-element program was specifically developed in order to provide a fast, 

simple, multiple-seam analysis program for the practicing 

engineer. LaM2D only requires that the user input the 

minimum geometric parameters from the multiple-seam 

mining situation. The program automatically incorporates 

default overburden, coal, and gob properties to simplify 

the input and inherently calculates pillar safety factors for 

the program output. 

 

 1Research assistant. 

 2

Associate professor. 

 West Virginia University, Department of Mining Engineering,

Morgantown, WV. 
36 

2–D LAMINATED OVERBURDEN MODEL 

 In the 3–D MULSIM and LaModel programs, the 

horizontal plan of the seam was used to determine a grid of 

elements from which the vertical (out-of-plane) stresses 

and displacements were determined. For example, given 

the mine plan in Figure 1A, a square grid (with an element 

size optimized to best fit the entry and pillar dimensions 

and large enough to cover the intended modeling area) was 

superimposed on the mining plan, as shown in Figure 1B. 

Each element of the grid was then designated as a particular material using a letter code (Figure 1C). (In this case, 

the pillar material is “A”, the open entries are “1”, and the 

gob material is “E”.) This grid of materials was then used 

as the representation of the mine for which the vertical 

stresses and displacements were determined. 

 In the 2–D laminated model, a cross-section of the 

pillars in the seam is used to determine a linear array of 

elements from which the vertical stresses and displacements are determined. For example, given the crosssection of pillars in Figure 2A, a linear array of elements 

(with an element size optimized to best fit the entry and 

pillar dimensions and large enough to cover the intended 

modeling area) is superimposed on the mining plan, as 

shown in Figure 2B. Each element of the material array is 

then designated as a particular material using a letter code 

(Figure 2C). (In this case, the pillar material is “A”, and 

the open entries are “1”.) This array of material properties 

is then used as the representation of the mine for which the 

vertical stresses and displacements are determined. 

Advantages 

 The big advantages of the LaM2D program over the 

previous 3–D programs are that the mine grid is much 

simpler and faster to input (minutes instead of hours) and 

the solution times are much quicker, in fact, almost instantaneous (seconds instead of minutes or hours). To further 

speed the input process, LaM2D (in default mode) only 

requires that the user input the minimum geometric 

Gob Area

50% extraction area

Area of

interest

Mine Plan

Model Grid Overlay

Material Codes

A

B

C

Figure 1.—Schematic of LaModel mine grid. 

Cross-Section of Mine Plan

Model Grid Overlay

Material Codes

A

B

C

Figure 2.—Schematic of LaM2D mine grid. 
37 

parameters to define the seam geometries. Default properties for overburden, coal, and gob are automatically generated. (In the advanced mode, the user has complete control 

over all of the input parameters.) This ease of input and 

instantaneous solution enables the user to quickly analyze 

and compare many different design variations or perform 

parametric studies. 

Disadvantages 

 The main disadvantage of LaM2D is that it is only a 

2–D plane strain analysis. Therefore, the variation in the 

mine plan in the out-of-plane direction is ignored, and only 

mining plans that can be reasonably represented in 2–D 

can be accurately simulated. The plane strain approximation means that the pillars and entries in the cross-section 

are essentially assumed to be infinitely long in the out-ofplane direction. Therefore, mining plans that change 

dramatically in the out-of-plane direction will not be 

modeled properly. However, LaM2D does have a capability for inputting an out-of-plane extraction ratio, which 

can be used to accurately simulate the crosscuts in the 

“infinitely” long pillars (see below). With the out-of-plane 

extraction ratio, not only infinitely long pillars, but also 

mining plans that are reasonably continuous and symmetric across the plane of the cross-section can be accurately simulated. 

Stability Analysis in LaM2D 

 In addition to its capability to quickly analyze basic 

multiple-seam interactions, LaM2D also provides several 

practical stability indices for the affected mine and pillars. 

The presently available computer programs for analyzing 

seam interactions in single- or multiple-seam mining situations typically produce only seam displacement and stress 

values as output. The new LaM2D program goes several 

steps further by using the seam displacement and convergence values to determine pillar safety factors, multipleseam subsidence, and multiple-seam horizontal strain 

values for output. 

Safety Factors 

 To ensure overall mine stability, the general approach 

is to size the pillars large enough, and therefore strong 

enough, to support the applied loads. In classical engineering, this generally requires a pillar safety factor greater 

than 1.0, with some reasonable margin for error. In situations where seam interactions exist as a result of a superjacent or a subjacent seam, simple empirical methods for 

calculating pillar safety factors are not generally applicable. Also, computer programs based on the single-seam 

empirical methods are not recommended for multiple-seam 

interaction analysis; rather, more complex numerical methods are often required. However, most practicing mining 

engineers are accustomed to working with pillar safety factors, and the stresses and displacements generated by the 

complex numerical modeling programs are not generally 

as useful or as comparable. In order to provide the 

customary safety factors and enhance the practical 

functionality of the LaM2D program, an automatic calculation of pillar safety factors has been incorporated into 

the program. 

 The primary safety factor implemented in the LaM2D 

program is stress-based and uses the typical safety factor 

calculation: 

B

S

SF p = (1) 

where SF = the pillar safety factor; 

 Sp = the pillar strength; 

and B = the pillar load. 

In order to incorporate realistic pillar strengths into the 

program, the Mark-Bieniawski pillar strength formula 

[Mark and Chase 1997] was used to determine pillar 

strengths because of its strong empirical basis and long 

successful history of application in the U.S. coalfields: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ⎟ −

⎠

⎞ ⎜

⎝

⎛ = +

lh

w

h

w S S p i

2

0.64 0.54 0.18 (2) 

where Si = the in situ coal strength; 

 w = the pillar width; 

 h = the pillar height; 

and l = the pillar length. 

 In the program, the height of the pillar is determined 

from the input seam height, and the width of the pillar is 

automatically determined from the number of elements 

across the pillar. For calculating the pillar load, LaM2D 

averages the values of all of the calculated element stresses 

across the pillar. The safety factor is then determined by 

dividing the Mark-Bieniawski pillar strength by the 

average pillar stress. Ultimately, the pillar safety factors 

can be displayed in a colored-square or line plot (see the 

subsequent case study). 

Multiple-seam Subsidence and Strain 

 When a seam is undermined, it experiences multipleseam stress transfer similar to an overmining situation. 

However, a seam that is undermined by total extraction 

mining can also be subjected to subsidence and the 

associated horizontal tensile and compressive strains. The 

multiple-seam subsidence effects are an important ground 

control consideration and can result in serious roof and 

floor instability and damages in the overlying seam. For 
38 

instance, field studies of room-and-pillar operations 

[Chekan and Listak 1994] showed that severe roof conditions are typically experienced within the tension and compression zones of a subsidence trough. 

 In previous multiple-seam analysis programs, the 

upper-seam subsidence and strain were never explicitly 

determined, although these calculations can be critical to 

analyzing the overlying mine stability. In order to provide 

the expected multiple-seam subsidence and strain values to 

the practicing engineer and increase the information available for design, the LaM2D program specifically calculates these values and provides them to the user. In the 

program, the upper-seam subsidence is calculated using 

the lower-seam convergence distribution and a laminated 

influence function. Once the seam subsidence is known, 

the resulting horizontal strains are determined using the 

method promoted by the National Coal Board in the 

Subsidence Engineers Handbook [National Coal Board 

1975]. A demonstration of these strain calculations is presented later in the case study. 

Out-of-plane Extraction Ratio 

 Because of the 2–D plane-strain formulation of 

LaM2D, the pillars and entries in the cross-sectional mine 

plan are essentially assumed to be infinitely long in the 

out-of-plane direction. This assumption is typically 

unrealistic for many pillar geometries in practical applications, where the pillars have a finite length in the out-ofplane direction. In many 2–D cross-sections, such as 

across a longwall gate road or through a room-and-pillar 

section, the pillar plan in the out-of-seam direction is fairly 

continuous (although not infinite). In these situations, an 

“out-of-plane extraction ratio” can be used to get a more 

accurate calculation of the true 3–D stress and displacements in the modeled mine plan. 

 In LaM2D, the in-plane extraction ratio is inherently 

provided through the specification of pillar material and 

openings in the input material properties (see Figure 2C). 

To specify the out-of-plane extraction ratio, the functional 

extraction ratio for the noninfinitely long pillars needs to 

be input. For example, consider a room-and-pillar section 

with 40- by 80-ft pillars and 20-ft-wide entries. The 2–D 

cross-section in LaM2D goes through the 80-ft length of 

the pillar; therefore, the input mine plan shows 80 ft of 

pillar material coupled with 20 ft of opening (an in-plane 

extraction ratio of 20%). In the out-of-plane direction for 

this situation, there is 40 ft of pillar coupled with 20 ft of 

opening. This results in an out-of-plane extraction ratio 

of 33%. 

 If the user inputs an out-of-seam extraction ratio, the 

stress from the infinitely long simulation will be adjusted 

to account for the limited pillar length. Essentially, the 

material stress and convergence values calculated for the 

infinitely long pillar are increased by a factor of: 

1− e

1 (3) 

where e = the out-of-plane extraction ratio. Thus, the practical result of using the out-of-plane extraction ratio is that 

the calculated stress values will realistically represent the 

average stress on the finite pillar, while the convergence 

values will represent an out-of-plane average across the 

finite pillar and the adjacent opening. Therefore, the 2–D 

simulation can give realistic 3–D pillar stress when the 

out-of-plane extraction ratio is used. 

 However, using the out-of-plane extraction ratio does 

not provide realistic 3–D safety factors. Without using the 

out-of-plane extraction ratio, the program calculates accurate pillar stresses, pillar strengths, and safety factors for 

the plane strain (infinitely long pillar) situation. When the 

out-of-plane extraction ratio is used, the program accurately adjusts the pillar stress values up to account for the 

limited length of the pillar; however, the pillar strength is 

still calculated based on an infinite-length pillar (for lack 

of sufficient information on the true pillar size). Therefore, the safety factor is realistically decreased from the 

2–D situation, but is still not an accurate safety factor for 

the 3–D situation implied by the out-of-plane extraction 

ratio. 

LAM2D INPUT 

 Compared to most numerical modeling programs, the 

input and output for the LaM2D program were simplified 

as much as possible to make it very user-friendly. Typical 

numerical modeling programs involve parameter input, 

solution of the model based on the input parameters, and 

an output or plot of the solution values. In many cases, 

three separate programs are used to perform the parameter 

input, solution, and the output (e.g., LaModel); however, 

the full-featured LaM2D program incorporates all of the 

numerical modeling steps into a single program. 

 Furthermore, the many material property input parameters necessary for modeling often seem daunting for practicing engineers who do not have the time to gather and 

verify specific material properties for both the seam and its 

surrounding rocks. In the LaM2D program, realistic 

average coal seam and overburden parameters are automatically defined. In fact, if the program is run in the 

standard mode, the only parameters required from the user 

are: the general model information, which includes model 

title, number of seams, and the system of units to be used 

(Figure 3); the seam geometry information, which includes 

element size, grid size, seam depths, and thicknesses 

(Figure 4); and the seam material grids (Figure 5). In 

standard mode, all of the seam and overburden properties 

are automatically entered with average values. However, if 

the user desires and the program is run in advanced mode, 
39 

the program allows the user the total flexibility of adjusting any or all parameters in the program, including 

the rock mass parameters (Figure 6), in-seam element 

parameters (Figure 7), boundary conditions (Figure 8), and 

program control values (Figure 9). 

Figure 3.—General model information form. 

Figure 4.—Seam geometry form. 

Figure 5.—Grid editor form.
40 

Grid Editor 

 The grid editor is used for inputting the seam material 

codes. It contains a simulated cross-section for each of the 

modeled mining seams (Figure 5) where the user can enter 

the appropriate letter codes into the proper grid location in 

order to represent the 2–D mine layout. The letter codes 

are entered using a mouse-driven interface similar to an 

Excel spreadsheet. The user only needs to input the primary coal letter code for each seam, and the program automatically generates appropriate yield zones based on a 

Bieniawski pillar strength. Also, the grid editor is where 

the user inputs the out-of-plane extraction ratio. 

Running the Model 

 Once all of the input has been completed, the LaM2D 

program typically runs in just a few seconds. Then, the 

output can be viewed using a number of plotting routines. 

The values output from the program include: seam convergence, seam stress, multiple-seam stress, multiple-seam 

displacement, and horizontal strain. Any of these values 

can be plotted separately using the traditional “coloredsquare” plot or line graph. Also, a new “comparison plot” 

can be used to visualize, compare, and analyze the 

interacting effect of different values from one seam to 

the other. 

CASE STUDY 

 In order to test the accuracy and utility of using 

the new, simplified multiple-seam program, LaM2D, 

a comparative case study was performed using both 

LaModel and LaM2D. The case study was taken from the 

literature [Ellenberger et al. 2003] and documents a situation where a longwall panel undermined an active roomand-pillar mainline development. The upper mine operates 

in the 9-ft-thick Coalburg Seam and has driven a sevenentry mainline with 40- by 80-ft pillars. The lower mine 

operates in the 5.6-ft-thick No. 2 Gas Seam and was 

 Figure 6.—Overburden and rock mass parameters 

form. 

Figure 7.—In-seam material form. 

 Figure 8.—Seam geometry and boundary condition 

form. 

Figure 9.—Program control form. 
41 

retreating a 1,000-ft-wide longwall panel directly under the 

overlying mains. The interburden at this site averaged 

560 ft, and the maximum overburden over the upper mine 

was about 400 ft. The idealized mine map for this site, 

as used in the comparative study, is shown in Figure 10. 

 For the idealized comparative study in this paper, the 

mains are modeled as perfectly aligned with the middle of 

the longwall panel, and all of the pillars in the mains and 

the gate roads are modeled with consistent sizes. In reality, 

the mains were somewhat offset from the center of the 

longwall panel and slightly skewed to the longwall 

advance. Also, the pillars in both the mains and the longwall gate roads have some variations in dimension due to 

practical and operational considerations [Ellenberger et al. 

2003]. 

 In the field, when the upper seam was undermined by 

the longwall, tension cracks began to develop in the roof 

when the underlying face approached to within 70 ft. Roof 

and rib control problems intensified as the longwall moved 

under and continued beyond the upper seam. The sandstone roof in the upper mine was extensively fractured, 

with some apertures measured at 4 in. Several roof falls 

occurred and severe rib spalling was experienced. As the 

dynamic subsidence settled and reached equilibrium, the 

roof fractures mostly closed and conditions significantly 

improved. 

 For the mine plan in Figure 10, a 3–D LaModel model 

was developed. This model used 10-ft elements on a 250-

by-250 grid to represent the mine plan shown. Entering the 

material parameters and the mine grid in order to build this 

simplified model took about 1 hr (for the real mine plan 

with variable size pillars and off-angle pillar, more than a 

day was used to build the model). The program then took 

another hour to solve the model for the seam stresses and 

displacements. Overall, LaModel took about 2 hr to build 

and solve the model. 

 Using LaM2D, a 2–D cross-section from the critical 

area (section A–A′) in Figure 10 was modeled. This crosssection essentially consists of the upper-seam pillars and 

the lower-seam longwall, as shown in Figure 11. With the 

simplified 2–D LaM2D, a linear array of 250 10-ft elements was used to represent each seam. The minimum 

parameter input for the model was used, as shown in 

Figures 3 and 4, and only took a minute or two to enter. 

Also, the material grids were entered into the program 

using the grid editor (Figure 5) and took another few minutes. After all of the parameters and grids were entered, the 

program solved for the seam stresses, displacements, 

safety factors, and strains in a few seconds. Overall, the 

total parameter input and solution required less than 

10 min using LaM2D. 

 Figure 12 compares the stresses calculated with the 

3–D LaModel program with those calculated with the simplified LaM2D program. For all practical purposes, the 

vertical stresses calculated by the two models are the same. 

Both models show the stress abutment in front of the 

advancing longwall face and the reduced stress in the gob 

behind the longwall. Further, the LaM2D program can be 

used to determine the pillar safety factors and the 

subsidence-induced strain. For instance, the pillar safety 

factors for the case study are shown in Figure 13. It can be 

seen in this figure that the safety factors are generally high 

and pillar failure would not be expected. This is consistent 

with observations in the field, which reported pillar spalling but no pillar failure. The horizontal strain was also 

calculated for the case study, as shown in Figure 14. The 

tensile stress in front of the moving longwall and the 

compressive stress behind the longwall on the order of 

250 microstrains are clearly evident. Based on the National 

Coal Board experience with surface strains [National Coal 

Board 1975], the 250 microstrains should only cause 

“slight damage.” However, the correlation between strain 

and underground roof damage is not very well defined, and 

250 microstrains could be consistent with the observed 

roof cracking and roof falls. 

 Figure 10.—Idealized mine plan for the case study 

analysis. 

Gob

Gob

Upper Mine

Lower Mine

A A'

Figure 11.—Cross-section used in the LaM2D program. 
42 

SUMMARY AND CONCLUSIONS 

 A simplified 2–D boundary-element method 

designed to model the complex multiple-seam stress and 

displacement interactions has been implemented into the 

LaM2D program. This new program allows the user to 

quickly and easily calculate the stresses, displacements, 

safety factors, and strains associated with basic multipleseam mining situations that can be accurately represented 

in two dimensions. The input to LaM2D has been greatly 

simplified by incorporating automatic overburden, interburden, coal, and gob property generation. Also, the program is the first multiple-seam analysis package to 

inherently calculate pillar safety factors and multipleseam-induced subsidence and strain. In a practical comparison with the full-featured multiple-seam program 

LaModel, the simplified LaM2D provided comparable 

stress and displacement calculations in minutes instead of 

hours. Further, the unique safety factor and strain calculations of the LaM2D program provided practical accurate 

assessments of the observed pillar and roof behavior in the 

modeled mine. 

REFERENCES 

 Chekan GJ, Listak JM [1994]. Design practices for 

multiple-seam room-and-pillar mines. Pittsburgh, PA: U.S. 

Department of the Interior, Bureau of Mines, IC 9403. 

 Ellenberger JL, Chase FE, Mark C, Heasley KA, 

Marshall JK [2003]. Using site case histories of multipleseam coal mining to advance mine design. In: Peng SS, 

Mark C, Khair AW, Heasley KA, eds. Proceedings of the 

22nd International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 59–64. 

 Heasley KA [1998]. Numerical modeling of coal 

mines with a laminated displacement-discontinuity code 

[Dissertation]. Golden, CO: Colorado School of Mines, 

Department of Mining and Earth Systems Engineering. 

 Heasley KA, Chekan GJ [1999]. Practical boundaryelement modeling for mine planning. In: Mark C, Heasley KA, Iannacchione AT, Tuchman RJ, eds. Proceedings 

of the Second International Workshop on Coal Pillar 

Mechanics and Design. Pittsburgh, PA: U.S. Department 

of Health and Human Services, Public Health Service, 

Centers for Disease Control and Prevention, National Institute for Occupational Safety and Health, DHHS (NIOSH) 

Publication No. 99–114, IC 9448, pp. 73–87. 

 Mark C, Chase FE [1997]. Analysis of retreat mining 

pillar stability (ARMPS). In: Mark C, Tuchman RJ, eds. 

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 and Prevention, National Institute for 

Occupational Safety and Health, IC 9446, pp. 17–34. 

 National Coal Board [1975]. Subsidence engineers 

handbook. London: National Coal Board, Mining 

Department. 

 Salamon MDG [1962]. The influence of strata movement and control on mining development and design 

[Dissertation]. Durham, U.K.: University of Durham, 

King’s College, Department of Mining. 

 Figure 12.—Comparison of LaModel and 

LaM2D stresses. 

0

200

400

600

800

1000

1200

0 500 1000 1500 2000 2500

Distance Along Cross Section (ft)

Vertical Stress (psi)

LaModel Output

LaM2D Output

Figure 13.—LaM2D safety factor output. 

Figure 14.—LaM2D horizontal strain output. 
43 

 Salamon MDG [1963]. Elastic analysis of displacements and stresses induced by the mining of seam or reef 

deposits, part 1. J S Afr Inst Min Metall 64(4):128–149. 

 Salamon MDG [1991]. Deformation of stratified rock 

masses: a laminated model. J S Afr Inst Min Metall 91(1):

9–26. 

 Zipf RK Jr. [1992a]. MULSIM/NL application and 

practitioner’s manual. Pittsburgh, PA: U.S. Department of 

the Interior, Bureau of Mines, IC 9322. 

 Zipf RK Jr. [1992b]. MULSIM/NL theoretical and 

programmer’s manual. Pittsburgh, PA: U.S. Department of 

the Interior, Bureau of Mines, IC 9321. 

This paper was previously published as: 

Heasley KA, Akinkugbe OO [2004]. A simple program for estimating multiple-seam interactions. SME preprint 04–137. Littleton, CO: Society for Mining, Metallurgy, and Exploration, Inc. 

45

MULTIPLE-SEAM LONGWALL MINING IN THE UNITED STATES: 

LESSONS FOR GROUND CONTROL 

By Christopher Mark, Ph.D., P.E.1

ABSTRACT 

 Relatively few longwall mines in the United States 

operate under multiple-seam conditions where the two 

seams are less than 200 ft apart. This paper describes the 

experience of six that do. These operations are located in 

Pennsylvania, West Virginia, and Utah and include examples of both undermining and overmining. Some operate 

above or beneath their own workings; others are in historic 

mining districts and must contend with abandoned mines 

that are decades old. The lessons learned by these mines 

cover a broad range of topics, including: 

• Whether to stack gate roads or place them under 

old gob areas; 

• How to size pillars and select artificial support to 

cross longwall stop lines; 

• How to use yield pillars to minimize multipleseam stresses and coal bump potential; and 

• When to anticipate the creation of pathways for 

gas, water, or oxygen between current and abandoned gobs. 

INTRODUCTION 

 Multiple-seam interactions can cause roof falls, rib 

spalling, and floor heave, disrupting mining operations and 

threatening the safety of miners. In early 2006, a West 

Virginia coal miner was killed by rib roll that occurred in a 

high-stress zone beneath a remnant pillar structure in an 

overlying mine [MSHA 2006]. Longwalls may be 

uniquely vulnerable to multiple-seam interactions because 

they tend to operate under deep cover, generate large 

abutment stresses, and have little flexibility to avoid localized zones of difficult conditions. 

 For the past several years, the National Institute for 

Occupational Safety and Health has been conducting 

research to develop better techniques for predicting the 

location and severity of multiple-seam interactions. During 

this research, more than 40 mines were visited across the 

U.S. coalfields, including a number of longwalls. The 

study also drew upon past studies conducted by the author 

and an extensive literature review. The purpose of this 

paper is to discuss the potential impacts of multiple-seam 

mining on longwalls based on an evaluation of the different geologic environments in which longwall mining is 

conducted, combined with analyses of past case histories. 

OVERVIEW 

 Figure 1 shows the five major underground coal mining regions in the United States. Also shown are the numbers of operating longwall faces in each region [Fiscor 

2006]. 

 The central Appalachian region of southern West Virginia, eastern Kentucky, and southwestern Virginia is the 

most significant coalfield from the standpoint of multipleseam mining. Mining has been ongoing in central Appalachia for nearly 150 years. Recent studies have indicated 

that perhaps 70% of the ultimate reserve base in the region 

has already been mined [Bate and Kvitkovich 2004]. One 

consequence of the maturity of the central Appalachian 

coalfields is that most underground reserves have been 

impacted by past mining activity. The mountains of the 

central Appalachian coalfields are honeycombed with 

worked-out mines located above, below, and adjacent to 

today’s and tomorrow’s operations. 

 Although longwall mining has a long history in central 

Appalachia, including the first mechanized faces in the 

 Figure 1.—The five major underground coal mining 

regions in the United States and the number of longwall 

faces in each [Fiscor 2006]. 

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin

13

3

7

17

10

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin

13

3

7

17

10

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin

13

3

7

17

10

Central

Appalachian

Western 

Alabama

Northern

Appalachian

Illinois

Basin

13

3

7

17

10

 

 1Principal research mining engineer, Pittsburgh Research

Laboratory, National Institute for Occupational Safety and

Health, Pittsburgh, PA. 
46 

United States [Barczak 1992], the 10 longwalls active 

there today account for less than 20% of the region’s 

underground production [EIA 2006; Fiscor 2006]. Of 

these, three are extracting the very deepest minable seam 

in the area—the Pocahontas No. 3—without any significant mining above or below. The others are operating in 

multiple-seam environments, and two case histories are 

discussed in some detail later in this paper. 

 The Western United States is the next most important 

area for multiple-seam mining. In Utah, Colorado, Wyoming, and New Mexico, nearly 95% of underground production comes from the 13 longwall operations [EIA 2006; 

Fiscor 2006]. Approximately half of these are operating 

in multiple-seam configurations, including several mines 

in the Price, UT, area and three others in the North Fork 

Valley near Paonia, CO. 

 In contrast to central Appalachia, in the West the same 

mining company is usually responsible for all the mining 

on a property. As a result, a greater degree of multipleseam planning is normally possible. Recently, however, 

one North Fork Valley mine was forced to abandon a panel 

prematurely when an interaction associated with an 

abandoned mine resulted in roof falls in the headgate 

[Buchsbaum 2006]. 

 Multiple-seam interactions in western longwalls can 

also contribute to the deadly bump hazard because of the 

deep cover and strong roof and floor rock [Peperakis 

1968]. To minimize the hazard, many western longwalls 

employ yield pillar gate entry designs. Some of the lessons 

that have been learned from their experience are discussed 

later in this paper. 

 The northern Appalachian region is home to nearly 

one-third of the nation’s longwalls—17 faces producing 

87 million annual tons [EIA 2006; Fiscor 2006]. Currently, 

all but one of these longwalls is extracting the Pittsburgh 

Seam. The Freeport Seam, the nearest minable coalbed 

beneath the Pittsburgh, is over 600 ft below, and it has 

never been extracted beneath any Pittsburgh longwall. 

There has been some mining in the Sewickley Seam that 

overlies the Pittsburgh. In the distant past it caused some 

disruption in the Pittsburgh [Zachar 1952], but it has not 

been known to impact the modern longwall mines. 

 Although none are operating today, in the recent past 

longwalls have operated in the Sewickley Seam and other 

seams in central Pennsylvania. These operations provide 

two examples of longwall mining above abandoned works 

that are discussed below. 

 More than 90% of the underground coal produced in 

the Illinois Basin comes from two seams: the Springfield 5/

Kentucky 9 and the Herrin 6/Kentucky 11 [EIA 2006]. 

These seams are often less than 80 ft apart, so there is 

significant potential for multiple-seam interaction. However, today there are just two longwall mines operating in 

the Illinois Basin. One of these does work both seams, but 

the mining has been planned so that the workings in one 

seam have never crossed previously mined gob areas in the 

other seam. 

 Seven longwalls are currently active in the Alabama 

coalfields. Six of these are extracting the Blue Creek 

Seam, which is again the deepest minable seam in the 

section, lying hundreds of feet below the nearest minable 

overlying seam. The only non-Blue Creek longwall is also 

free of multiple-seam interactions. 

HAZARDS ASSOCIATED WITH 

MULTIPLE-SEAM MINING 

 Ground instability is usually the greatest hazard due to 

multiple-seam interaction. Interactions may be classified 

into four major categories depending on the mining 

method, mining sequence, and thickness of the interburden. Other potential hazards are associated with water, 

gas, and oxygen-deficient air. 

Undermining, the first category of interaction, occurs 

when the upper seam has been mined first and the lower 

seam is the active seam (Figure 2). In an undermining situation, damage is caused by load transfer from highly 

stressed structures associated with full-extraction mining 

in the overlying seam. 

Overmining, the second type of interaction, occurs 

when the upper seam is extracted after mining is complete 

in the lower seam (Figure 3). Load transfer occurs in this 

situation just as it does in undermining (in other words, 

gob-solid boundaries and remnant pillars cause stress concentrations both above and below). In addition, however, 

full extraction of the lower seam normally results in subsidence of the overlying beds, potentially damaging 

the roof. 

Dynamic interactions occur whenever active mining 

occurs above or beneath open entries that are in use. The 

most severe dynamic interactions occur when a lower 

seam is longwalled or pillared, resulting in subsidence of 

the overlying workings. Although there have been several 

relatively recent cases in which longwalls were extracted 

as far as 550 ft beneath open main entries, causing 

Overburden

Upper Seam

(Earlier Mining)

Lower Seam

(Current mining)

Interburden

Overburden

Upper Seam

(Earlier Mining)

Lower Seam

(Current mining)

Interburden

Figure 2.—Undermining interaction. 
47 

extensive damage to the upper mine [Ellenberger et al. 

2003; Mark 2006], there do not seem to be any instances in 

the United States where a working longwall was the victim 

of a dynamic interaction. 

Ultraclose mining is the fourth type of interaction and 

the only one that applies to development mining. The main 

concern is failure of the interburden between the two 

seams. Ultraclose interactions are unlikely when the two 

seams are more than 20–30 ft apart. 

 Other hazards include the potential for inundation 

from an overlying flooded mine, particularly where full 

extraction in the lower seam can create a direct pathway 

between the upper- and lower-seam gobs. In one example, 

the first longwall panel at a Kentucky mine encountered a 

major inflow from workings 150 ft above. In this instance, 

the water posed no hazard to the miners, but the lack of 

pumping capacity in the lower seam resulted in major 

mining delays [Mark et al. 1998]. Fractures in rock above 

subsided gob areas can also fill with methane or oxygendeficient air, resulting in inflows of gas or blackdamp 

when they are intersected by overmining. 

OVERMINING INTERACTIONS: 

TWO CASE HISTORIES FROM PENNSYLVANIA 

 Two Pennsylvania longwalls, both of which are now 

closed, provide examples of the potential for overmining 

previously extracted areas. The first example was a successful one, the second less so. 

 In Cambria County, PA, Bethlehem Mines Corp.’s 

No. 33 Mine employed longwall methods to extract the 

B Seam (Middle Kittanning) and the overlying C-prime 

Seam (Upper Kittanning). The mining was sequenced so 

that extraction was completed in the B Seam first. More 

than 25 longwall panels were later recovered in the upper 

seam over almost 2 decades [Bauer et al. 1992]. 

 The B Seam averaged about 5 ft in thickness, and its 

depth of cover was 500–800 ft. High horizontal stresses 

were typical of the lower seam, where they caused extensive damage to the gate entries during both development 

and longwall retreat [Mark and Mucho 1994]. 

 The C-prime Seam was typically 3.5 ft thick, with an 

additional 1 ft of drawrock normally extracted. The immediate roof consisted of interbedded sandstone and shale, 

with a Coal Mine Roof Rating (CMRR) of about 50. The 

interburden between the seams averaged 105 ft and consisted mainly of shales and sandy shales. Typically, at least 

10 years was allowed to elapse between the completion of 

mining in the lower seam and development in the upper. 

Contrary to usual multiseam practice, the gate entries in 

the C-prime were superpositioned above the gates in the 

lower seam (Figure 4). In general, conditions were quite 

good both on development and longwall retreat. Six-ftlong fully grouted resin bolts, supplemented in the tailgate 

by two rows of wooden posts, provided adequate roof support. In only one instance, when an attempt was made to 

develop a gate road slightly over the gob, were poor 

ground conditions attributed to multiple-seam interactions. 

Conditions along the longwall face were also benign and 

were not noticeably different near the panel edges than 

they were in the center of the subsidence trough. 

 Interestingly, ground conditions in the upper seam 

actually seemed to be worse above virgin, undisturbed 

B Seam areas than above areas where the B Seam was 

mined out. Above the virgin areas, C-prime workings 

encountered severe ground conditions and methane emissions. It seemed that perhaps the subsidence associated 

with the extraction of the B Seam relieved the horizontal 

stresses and removed the methane that would otherwise 

have been present. 

 The second example of overmining is provided by 

New Warwick Mining Co.’s Warwick Mine in Greene 

County, PA. The Warwick Mine worked the Sewickley 

Seam beneath 500–900 ft of cover. The Sewickley roof 

was a dark sandy shale with a CMRR of about 40. 

 The Pittsburgh Seam was approximately 100 ft below 

the Sewickley, and underneath one panel it had been 

Overburden

Upper Seam

(Current mining)

Lower Seam

(Earlier mining)

Interburden

Overburden

Upper Seam

(Current mining)

Lower Seam

(Earlier mining)

Interburden

Figure 3.—Overmining interaction. 

 Figure 4.—Stacked gate road layout at Bethlehem 

Mines Corp.’s No. 33 Mine [Bauer et al. 1992]. 
48 

worked by high-extraction room-and-pillar methods (Figure 5). Analysis indicated that many of the small fenders 

that had been left in the Pittsburgh Seam had probably 

yielded, transferring much of their load to a few large, isolated remnant pillars [Heasley and Chekan 1999]. Severe 

pillar spalling and poor roof conditions were encountered 

when developing gate entries above these remnants. 

 The inby portion of the Sewickley longwall panel, 

where the cover was shallowest, was developed and longwalled with little difficulty. However, stress mapping and 

LaModel analysis indicated that the most severe interactions would occur beneath the deepest cover, near the 

mouth of the panel [Heasley and Chekan 1999]. When the 

longwall face reached this area, ground control problems 

included face spalling and headgate roof instability. The 

conditions were so severe that the longwall face had to be 

stopped 50 ft short of the recovery chute, which greatly 

complicated recovery of the longwall. 

UNDERMINING INTERACTIONS: 

TWO CASE HISTORIES FROM 

CENTRAL APPALACHIA 

 Two mines in southern West Virginia provide examples of longwalls undermining full-extraction workings in 

an upper seam. The first, the Harris No. 1 Mine, is 

operated by Peabody Energy. During the past 30 years, 

Harris has mined more than 60 longwall panels in the 

Eagle Coal Seam. The depth of cover exceeds 1,400 ft 

beneath the highest ridges. The roof is fairly typical of 

central Appalachia and consists of strong shale and 

sandstone, with CMRR values normally in the 45–60 

range. 

 The Eagle is located 180–200 ft beneath the No. 2 Gas 

Seam. Figure 6 is a generalized stratigraphic column of the 

interburden that shows it contains a high percentage of 

massive sandstone. The No. 2 Gas has also been extensively mined. The longwall panels in the upper seam were 

much smaller than the recent Harris panels, and they lie in 

several different orientations. As a result, in many cases it 

has not been possible to plan the lower-seam workings to 

minimize the potential for interaction. 

 Chase et al. [2007] evaluated 17 locations where gate 

entries at Harris crossed beneath remnant pillar structures 

in the upper seam. In about 60% of these cases, development and retreat mining were successful, although it was 

often necessary to install cable bolts or other secondary 

support to control the ground conditions. In seven cases, 

however, the interactions were serious enough that the 

headgate or tailgate squeezed shut during panel recovery. 

 Figure 5.—Map of the Warwick Mine in the Sewickley 

Seam and the underlying Pittsburgh Seam workings 

[Heasley and Chekan 1999]. 

520

820

720

620

720

620

620

720

820

920

620

Scale (ft)

0 650

520

820

720

620

720

620

620

720

820

920

620

520

820

720

620

720

620

620

720

820

920

620

Scale (ft)

0 650

 Figure 6.—Stratigraphic column from the Harris Mine, 

Boone County, WV. 
49 

Most of these cases occurred beneath chain pillars where 

the depth of cover exceeded 1,000 ft (Figure 7). 

 There have also been several instances when poor 

ground conditions on the face were attributed to overlying 

structures. Most often the problem was severe floor heave, 

although in two instances roof control was lost at the face. 

All but one of these incidents occurred more than a decade 

ago, when the longwall hydraulic face supports were much 

less capable than they are today. 

 Based on past experiences, mine planners at Harris try 

to adhere to the following rules of thumb: (1) the long axis 

of the panel to be mined should parallel that of the upperseam panel; (2) the future headgate should be positioned 

under and as close to the center of the gob as possible; and 

(3) avoid advancing the longwall face under a gob/solid 

boundary [Hsiung and Peng 1987a,b]. 

 The engineers at Harris also use the LaModel program 

to identify high vertical stress areas that are caused by 

deep cover, abutment loads, and/or multiple-seam stress 

transfer [Heasley and Agioutantis 2007]. Once multipleseam stress transfer magnitudes are obtained, they can be 

incorporated into the Analysis of Longwall Pillar Stability 

(ALPS) or the Analysis of Retreat Mining Pillar Stability 

(ARMPS) programs to obtain more realistic pillar stability 

factors. 

 At West Virginia Mine B, longwall mining was 

initially conducted in the Lower Cedar Grove Seam before 

moving to the underlying Alma Seam. The interburden 

between the two seams was only 60–90 ft and consisted 

largely of sandstone. In the highly mountainous terrain 

the cover varied from 300 to 1,100 ft [Vandergrift et al. 

2000]. The lower-seam roof is typically a strong siltstone or fine-grained sandstone, with a CMRR in excess 

of 55. 

 Following an extensive LaModel study, the lowerseam layout was planned to minimize the potential for 

interactions. Particular care was taken to ensure that no 

development crossed beneath the heavily loaded upperseam chain pillars. The lower-seam mains were developed 

beneath the overlying ones, but the longwalls were laid out 

so that the gates were offset by approximately one-quarter of 

a panel width (Figure 8). Because the lower-seam panels 

were longer than the overlying ones, the gates crossed both 

the upper-seam start and stop line barrier pillars. At both 

crossings, the entry widths were reduced to 18 ft and the 

crosscut spacing was increased. Extra support, including 

truss bolts and steel props, were installed in the belt entry. 

 The first panels were retreated in the lower seam (see 

Figure 8). No serious gate entry stability problems were 

encountered, although the high stress zones beneath the 

stop and start lines were clearly visible, particularly when 

the depth of cover exceeded 800 ft. Although the upperseam chain pillars were located above the longwall face, 

they did not seriously impact ground conditions either. As 

one miner put it: “Beneath the chain pillars, the coal was 

so broken that the shearer didn’t have to cut anything; all it 

had to do was load.” Several of these faces were also 

successfully retreated across the upper-seam start pillars. 

 The next two panels were not as successful, however. 

They were oriented perpendicular to the initial panels and 

crossed a 110-ft-wide remnant barrier pillar located under 

about 800 ft of cover between two pillared areas in the 

upper seam. It was difficult to develop the gates across this 

structure, but when the longwall face approached it, the 

conditions on the face became so severe that the panel was 

abandoned early. To prevent a similar event, the second 

 Figure 7.—Harris Mine longwalls in the underlying 

Eagle Seam (light gray) crossing beneath upper-seam 

chain pillars (dark gray). 

 Figure 8.—Multiseam longwall layout employed in the 

Lower Cedar Grove Seam [Vandergrift et al. 2000]. 

Scale, ft

0 1,000 2,000

Upper Seam Gob
50 

panel was also recovered before it reached the remnant 

barrier. 

 Subsequently, eight additional longwall panels have 

been extracted in the lower seam. The most serious difficulties have been associated with an unsuccessful attempt 

to stack the gate roads. The roof in this area consisted of an 

unusually thick shale, and the conditions became so severe 

once the depth of cover exceeded 800 ft that the development was abandoned. 

MULTIPLE-SEAM LONGWALL MINING 

WITH YIELD PILLARS IN UTAH 

 Longwall mines in Utah work in a region characterized by rugged canyon and mesa topography, where cover 

depths can reach 3,000 ft. The roof and floor strata tend to 

be relatively strong (15,000–25,000 psi), and the overburden consists largely of massive sandstone units that can 

be as much as 500 ft thick. Many of the mining properties 

also contain two or more thick coal seams in close proximity [Barron et al. 1994]. 

 Unique among U.S. coal operations, Utah longwalls 

rely almost exclusively on yielding pillar gate entry 

designs. According to DeMarco [2000], yielding pillar gate 

systems were developed to— 

1. Mitigate the severe coal bumps (dynamic coal 

seam failures), most commonly experienced in the 

tailgate entries; and 

2. Abate the high-stress concentrations associated 

with remnant gate pillars during close-proximity 

multiseam mining. 

 Based on his comprehensive study of yielding pillar 

case histories, DeMarco [2000] concluded that successful 

gate entry systems— 

• Are largely limited to settings where the immediate roof is strong enough to withstand large 

deformations (the CMRR in every successful case 

history exceeded 50); 

• Employ chain pillar width-to-height ratios of less 

than 5 to ensure that timely, nonviolent yielding 

occurs during first panel mining; and 

• Normally require that artificial roof support plays 

a major role. 

 At Utah Mine A, approximately 20 lower-seam longwalls have been extracted beneath upper-seam panels. The 

interburden has ranged from 60 to 90 ft, with maximum 

depths of cover from 1,400 to 2,100 ft. Upper-seam mining 

was typically completed 3–5 years before development 

began beneath it. Both seams employ two-entry gate 

systems with 30-ft-wide yield pillars. The roof in the lower 

seam is variable, but typically a strong sandstone either 

lies directly on the coal or can be reached with roof bolts, 

so the CMRR is usually near 70. 

 In the very first multiple-seam application at this 

mine, the gate entries were stacked. A 160-ft massive roof 

fall ensued, and the gate had to be redriven. Since then, all 

gate entries have been driven beneath the upper-seam gob, 

normally offset at least 100 ft from the upper gate. Lowerseam panels have usually been designed so that they begin 

and end under the upper-seam gob. However, Figure 9 

shows a multipanel layout where several of the longwall 

faces did cross the stop lines of the overlying panels. The 

depth of cover at the crossings was less than 1,400 ft in 

every case. 

 No special support is normally required to develop the 

gates underneath the upper-seam stop lines. However, 

when the cover exceeds about 1,600 ft, bumps are more 

likely to occur if the gate crosses the stop line near the 

middle of the overlying panel rather than near the edge 

where the upper-seam stop line barrier pillar has been 

slightly “softened” by the presence of the upper-seam 

gates. 

 One of the big advantages of a two-seam longwall 

mine is that, other than the chain pillars, there should be no 

isolated remnant pillars in the overlying gob. One incident 

at Mine A illustrates just how important this advantage is. 

A “barrier pillar” was created in the upper seam between 

the longwall start line and some bleeder entries. When the 

lower-seam longwall had mined beneath the bleeders and 

partially beneath the barrier, a major bump occurred along 

approximately 80% of the face (Figure 10). The depth of 

cover at this location was in excess of 2,000 ft. Subsequently, the setup rooms in the lower seam were always 

placed under the gob if the cover exceeded 1,400 ft. 

 In general, face conditions have been excellent in the 

lower seam. Although the imprint of the overlying yield 

pillars could clearly be seen in the roof and coal face, 

 Figure 9.—Layout of yield pillar longwall panels at 

Utah Mine A. 

Lower seam

Upper seam
51 

it caused no problems. Under deep cover, the panels start 

outby the upper setup rooms and stop inby the upper-seam 

stop line, so nearly all the longwall mining has been 

beneath gob. 

 With a 10-ft extraction height in the lower seam and 

as little as 60 ft of interburden, it seems likely that subsidence fractures can create hydraulic connections between 

the gobs of the two seams. Nonetheless, there has been just 

one instance where water from an insufficiently dewatered 

area entered the lower-seam mine in sufficient quantities to 

obstruct production. 

 At Utah Mine B, the two seams are 50–75 ft apart. 

The cover is lighter than at Mine A, reaching a maximum 

of about 1,100 ft, but the roof is weaker, with a CMRR 

averaging about 50. Bumps have not been a problem at this 

operation, but roof control is a significant concern. 

 The upper-seam panels at Mine B have all employed 

three-entry gates with 30-ft-wide yield pillars. A total of 

nine lower-seam panels have been extracted. As was the 

case in Utah Mine A, it is possible to “see” the upper-seam 

pillars on the face and in the setup rooms of the lower 

seam, but they have never caused problems. 

 The lower-seam panels at Mine B have been planned 

so that they are set up, mined, and stopped all under the 

gob of the upper seam. The critical zone of interaction has 

been where the gates cross the upper-seam stop line barriers. These have been troublesome and require substantial 

artificial support. Currently, cable bolts are installed in 

these areas as primary support on development in a pattern 

of four 12- to 16-ft-long cables per row, with rows on 4-ft 

centers. Over time, Mine B has also found that conditions 

are better if larger pillars (currently 100-ft centers) are 

used to cross the barriers (Figure 11). 

 Another design issue is the location of the lower-seam 

setup room. In the first lower-seam panels, the setup was 

developed just 40 ft inside the gob from the edge of the 

upper-seam setup (Figure 12). Conditions were less than 

favorable. They improved considerably when the offset 

was increased to 100 ft. 

 A mine planner at Mine B observed that one 

undesirable characteristic of multiple-seam interactions is 

that “they just keep coming.” In one instance, a long-term 

mains crossed beneath an upper-seam gob line. Over 

4 years, more than 4 ft of floor heave developed, and the 

roof was cable-bolted three times but still eventually 

failed. 

 Spontaneous combustion is another safety concern at 

Mine B. In one of the panels where the setup was developed with a 40-ft offset, the caving behind the longwall 

created a connection that allowed oxygen to reach the 

upper-seam bleeder system. A heating developed, which 

was fortunately detected and quickly extinguished. There 

have also been several instances of large water inflows to 

the lower seam despite numerous holes drilled from the 

lower seam into low areas of the upper-seam gob to drain 

water pools from the upper-seam workings before 

undermining. 

60' interburden

200'

Key

Coal

Gob

 Figure 10.—Cross-section showing panel geometry 

at the time of a large face bump at Utah Mine A. 

 Figure 12.—Setup room offset at Mine B: (A) 40-ft 

offset; (B) 100-ft offset. 

40' offset

60' interburden

60' interburden

100' offset

A

B

Key

Coal

Gob

40' offset

60' interburden

60' interburden

100' offset

A

B

Key

Coal

Gob

40' offset

60' interburden

60' interburden

100' offset

A

B

Key

Coal

Gob

 Figure 11.—Layout of stop line crossings employed 

at Utah Mine B. 

gob

Upper seam

Lower seam
52 

CONCLUSIONS 

 Although few U.S. longwalls face serious multiseam 

challenges, some valuable lessons can be learned from 

those that do. These include the following: 

• Gate entry configuration: The first question that 

often faces mine planners is whether to stack 

the gate roads or offset them. Most of the 

U.S. longwall mines have preferred to offset their 

gate roads at least 100 ft from the gates in the 

previous seam. Some have had bad experiences 

when they tried stacking. However, the No. 33 

Mine successfully employed a stacked gate configuration for many years, and the Harris Mine 

has used stacked pillars on occasion. 

• If the offset design is employed, then the gate 

entries must be developed across the gob-solid 

boundary at the stop line of the previous seam 

longwall. The experience of the mines described 

in this paper is that this crossing may require a 

special pillar design or extra roof support, but 

it has not usually been a major concern. 

• Panel layout: The preferred layout is one in which 

the panel start positions are set up outby the 

overlying or underlying setup room and stop 

before they reach the stop line of the previous 

panel. This way all longwall mining is beneath 

the old gob area (except for the coal beneath the 

chain pillars). However, there are a number of 

instances where these mines have successfully 

brought their faces across old stop lines or setup 

rooms, at least occasionally. 

• Longwall face conditions have rarely been a problem at any of these mines, whether they are overmining, mining underneath large chain pillars, 

or crossing gob-solid boundaries. The most difficult interactions have been associated with large 

remnant pillars that were left in the previously 

mined seam, particularly when the remnant barrier was oriented parallel with the longwall face. 

• Roof strength: The experience of these mines is 

that roof strength is a major factor determining 

the success of multiple-seam mining. For example, the weaker roof at Utah Mine B has caused it 

to generally have more trouble than Utah Mine A, 

even though Mine A’s cover is significantly 

greater. Similarly, the Warwick Mine has weaker 

roof than the No. 33 Mine and also encountered 

greater difficulties. 

• Planning: The ability to preplan the multipleseam mining on a property is a major advantage. 

Mines that overmine or undermine their own 

longwall panels can more easily overlap the 

panels, offset the gates, and longwall almost 

entirely under gob areas. Most importantly, they 

can ensure that isolated remnant pillars are not 

left in the gob. The experience of the Utah mines 

also shows that yield pillars can also be a very 

useful planning option as long as the roof strength 

is adequate. 

• Mines like Harris and Warwick that do not have 

the luxury of preplanning all the mining in 

advance have been forced to deal with remnant 

structures and gob areas at odd angles. For 

instance, it seems that perhaps the extra 100 ft of 

interburden at Harris compared to West Virginia 

Mine B may have compensated for Harris’ more 

difficult remnant structures. 

 As the demand for coal continues to expand, mining 

companies are considering a number of new multiseam 

longwall mines in such areas as central Illinois and the 

Hunter Valley in Australia. Hopefully, these future 

operations will be safer because they have benefited from 

the experience of the operations described in these case 

studies. 

REFERENCES 

 Barczak TM [1992]. The history and future of longwall mining in the United States. Pittsburgh, PA: U.S. 

Department of the Interior, Bureau of Mines, IC 9316. 

 Barron LR, DeMarco MJ, Kneisley RO [1994]. Longwall gate road stability in four deep western U.S. coal 

mines. Denver, CO: U.S. Department of the Interior, 

Bureau of Mines, IC 9406. 

 Bate RL, Kvitkovich JF [2004]. Quantifying the coal 

reserve dilemma in the central Appalachian mining region. 

SME preprint 04–108. Littleton, CO: Society for Mining, 

Metallurgy, and Exploration, Inc., 8 pp. 

 Bauer ER, Chekan GJ, Sames GP [1992]. Influence of 

subjacent gob on longwall development mining in the 

upper Kittanning coalbed of south-central Pennsylvania. 

Pittsburgh, PA: U.S. Department of the Interior, Bureau of 

Mines, RI 9403. 

 Buchsbaum L [2006]. Longwall mining thrives in 

Colorado’s North Fork Valley. Coal Age 111(8):28–34. 

 Chase FE, Worley P, Mark C [2007]. Multiple-seam 

mining interactions: case histories from the Harris No. 1 

mine. Paper in these proceedings. 

 DeMarco M [2000]. Review of current yielding gate 

road design approaches and applications in U.S. longwall 

operations. Published as Appendix 11 to Volume 1 of University of New South Wales/SCT Collaborative ACARP 

Project C9018 – Final Report, Australian Coal Association 

Research Program (2005). 

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Mark C, Khair AW, Heasley KA, eds. Proceedings of the 

22nd International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

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of the Second International Workshop on Coal Pillar 

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Publication No. 99–114, IC 9448, pp. 73–87. 

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a boundary-element program for coal mine design. Paper 

in these proceedings. 

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multiple-seam mining, part I. Coal Min 24(9):42–46. 

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central Appalachian coalfields. SME preprint 06–060. 

Littleton, CO: Society for Mining, Metallurgy, and Exploration, Inc., 7 pp. 

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This paper was previously published as: 

Mark C [2007]. Multiple-seam longwall mining in the United States: lessons for ground control. SME preprint 

07–104. Littleton, CO: Society for Mining, Metallurgy, and Exploration, Inc. 

55

EXTREME MULTIPLE-SEAM MINING IN THE CENTRAL APPALACHIAN COALFIELDS

By Christopher Mark, Ph.D., P.E.1

ABSTRACT 

 Coal has been mined in the central Appalachian 

coalfields of southern West Virginia, western Virginia, and 

eastern Kentucky for more than a century. The dwindling 

reserve base consists in large part of coal that would have 

been considered unminable by earlier generations. Nearly 

every current operation is working on a property where 

coal has been extracted in the past from seams either 

above, below, or both. 

 The National Institute for Occupational Safety and 

Health (NIOSH) is conducting research aimed at helping 

mine planners prevent hazardous conditions due to 

multiple-seam interactions. To date, nearly 300 case histories have been collected from underground mines, mainly 

from central Appalachia. This paper focuses on several of 

the more challenging situations that have been encountered, including: 

• Room-and-pillar development 20 ft (6 m) beneath 

full-extraction workings at a depth of 1,000 ft 

(300 m) of cover (Virginia); 

• Pillar recovery 45 ft (14 m) above full-extraction 

workings at 900 ft (270 m) of cover (Virginia); 

• Near-simultaneous room-and-pillar mining with pillar recovery, with 40 ft (12 m) of interburden and 

1,500–2,000 ft (450–600 m) of cover (Kentucky); and 

• Longwall mining directly beneath main entries in 

overlying seams (West Virginia). 

 Some of these operations have been highly successful 

in overcoming the challenges; others less so. The lessons 

learned from their experience will help ensure that these 

and similar difficult reserves can be mined safely. 

INTRODUCTION 

 The central Appalachian region of eastern Kentucky, 

western Virginia, and southern West Virginia has produced more than 17 billion tons of coal since mining began 

there nearly 150 years ago. Production peaked in the late 

1990s at approximately 275 million tons/yr (250 million 

tonnes/yr) and has since dropped to about 240 million 

tons/yr (215 million tonnes/yr). Recent studies have indicated that perhaps 70% of the ultimate reserve base in the 

region has already been mined [Bate and Kvitkovich 2004]. 

 One consequence of the maturity of the central Appalachian coalfields is that nearly every remaining underground reserve has been impacted by past mining activity. 

The mountains of the central Appalachian coalfields are 

honeycombed with worked-out mines located above, 

below, and adjacent to today’s and tomorrow’s operations. 

 In making future mine projections, mine planners must 

evaluate the potential impacts of these multiple-seam 

interactions. When undermining (mining beneath old 

workings), the new developments may be subjected to 

excessive load transfer (Figure 1). In the overmining 

situation (Figure 2), load transfer can also occur and, 

in addition, the ground may have been damaged by subsidence. In some cases, multiple-seam interactions can be so 

severe and hazardous that mining is impossible. In others, 

it may be possible to develop pillars, but not recover them. 

Figure 2.—Overmining. 

Figure 1.—Undermining. 

 

 1Principal research mining engineer, Pittsburgh Research

Laboratory, National Institute for Occupational Safety and

Health, Pittsburgh, PA. 
56

In many cases, however, the interaction may be barely 

noticeable.2

 Some rules of thumb are available to aid in planning 

for multiple-seam interactions. Westman et al. [1997] cite 

traditional reserve estimation criteria that state that when 

the interval to mining above or below is less than 40 ft 

(12 m) the coal is considered to be sterilized, but otherwise 

accessible. Haycocks and Zhou [1990] found that load 

transfer interactions were unlikely when the interburden 

between the seams exceeded 110 ft (33 m), but that some 

factors (such as strong sandstone or a limited number of 

interbeds) could reduce this to as little as 60 ft (18 m). 

With thinner interburdens, interactions could be expected. 

Haycocks and Zhou [1990] also state that columnization of 

pillars “is considered the traditional approach to multiseam 

mining, especially when the interburden is less than 50 ft 

(15 m).” 

 For overmining situations, Luo et al. [1997] developed 

a technique for calculating a damage rating based on the 

lower-seam extraction ratio. However, Lazer [1965] 

reported that if the lower seam has been completely 

extracted, the upper seam can often be easily mined. The 

overburden mechanics model developed by Kendorski 

[1993] indicates that mining might be difficult within 

the “caving zone” (where the interburden-to-seamthickness ratio (I/t) is less than 6–10) or the “fracture 

zone” (I/t < 24). 

 The time lapse between mining has also been cited as 

an important factor. Intervals of at least 2 years have been 

suggested to allow the gob to fully consolidate [Haycocks 

and Zhou 1990]. Some studies have indicated that the 

longer the time lapse, the better the conditions that are 

anticipated. 

 For the past several years, NIOSH has been studying 

multiple-seam interactions with the goal of providing mine 

planners with more precise guidelines than are currently 

available. Nearly 40 mines have been visited, and a total of 

nearly 300 individual case histories have been documented. Approximately 80% of these case histories are 

from central Appalachian mines, with the remainder from 

northern Appalachia and the western coalfields. 

 Each case history in the database has been classified 

according to the severity of the observed interaction. There 

are four levels: 

• No interaction where conditions appear to be no 

different from those in areas where no past mining has 

been conducted; 

• Minor interactions where minimal pillar spall or roof 

cracks indicate that there are some changes that can 

 

 2Other hazards may arise when mining under abandoned works 

that are flooded or contain bad air [Michalek and Wu 2000], but 

these are not discussed in this paper. 

be attributed to past mining, but they had no significant effect on mining; 

• Difficult interactions where conditions were severe 

enough to require supplemental support, design 

changes, or (on retreat) abandonment of a few pillars; and 

• Severe interactions where the area was abandoned 

and judged unminable. 

 In the course of collecting the case histories, NIOSH 

found that mining is being conducted under many 

“extreme” situations, where previous mine workings are 

close by the target seam. Currently available rules of 

thumb imply that mining should be severely restricted at 

these operations but, in many cases, most of the target 

reserves are being mined with some success. Severe interactions have been encountered in some areas of nearly all 

these mines, however. The goal of this paper is to focus on 

a few of these extreme situations, identify those factors 

that have contributed to severe interactions, and discuss the 

control techniques that have proved to be successful. 

CASE NO. 1: UNDERMINING WITH 20 FT (6 m) 

OF INTERBURDEN 

 In southwestern Virginia, NIOSH researchers visited 

two mines that have exploited seams lying just 20 ft (6 m) 

beneath previously worked seams. In one instance, the 

Marker Seam is being mined beneath the Taggart Seam; 

in the other, the Tiller is being mined beneath the Jawbone. 

The interburden geology is similar in both situations, 

consisting mainly of competent sandstone and siltstone. 

Lower-seam Coal Mine Roof Rating (CMRR) values are 

typically in the mid-60s. The depth of cover (H) ranges 

from 600 to 1,000 ft (180 to 300 m). 

 One of the mines is a two-seam operation, with both 

seams being worked by the same operator. All mining has 

been development, with no pillar recovery. The pillars 

have been stacked directly above one another. More than 

800 pillars have been developed in this fashion, reportedly 

without serious incident. 

 At the other operation, mining in the upper seam was 

completed approximately 30 years ago and included large 

areas where pillars were recovered. Due to the variety of 

upper-seam pillar sizes and uncertainty about the surveying, there has been no attempt to columnize the pillars. 

Nevertheless, lower-seam mining has been largely successful beneath upper-seam first workings. In some areas, 

it has even been possible to partially extract pillars in the 

lower seam. Roof support beneath first workings consists 

of 4-ft, No. 5, full-column resin bolts supplemented by 6-ft 

“superbolts” in the intersections. 

 Problems have been encountered when attempting to 

cross upper-seam gob lines. On at least two occasions, 

mining had to be abandoned despite the use of longer 

bolts, cribs, and steel posts. Conditions were particularly 
57

difficult above an 80-ft (24-m) wide barrier pillar separating two gob areas. Where the gob line was successfully 

crossed, the pillar sizes were increased in addition to 

installing extensive supplemental support (Figure 3). 

CASE NO. 2: OVERMINING WITH MINIMAL 

INTERBURDEN 

 Three mining operations, also in Virginia, are extracting seams that lie less than 45 ft (14 m) above previous 

workings. In two instances, the Jawbone is being mined 

above Tiller Seam workings; in the third, the target seam 

is the Splashdam above Upper Banner workings. The 

Jawbone-Tiller interburden again consists of strong, 

competent rock, while the Splashdam-Upper Banner 

interval is somewhat weaker (CMRR=45 for the immediate roof). The cover is mainly in the 800–1,000 ft (240–

300 m) range for the Jawbone-Tiller mines and 500–600 ft 

(150–180 m) at the other operation. 

 At one operation, the interburden is just 20 ft (12 m). 

Here both seams are being mined by the same operator, 

and pillars have been columnized. There has been no 

second mining in the lower seam, but pillars have been 

fully extracted in four upper-seam panels with abandoned 

first workings directly underneath. No problems were 

reported. 

 At the other two operations, the interburden is 35–45 ft 

(11–14 m). Extensive second mining was conducted in the 

lower seams, but was completed at least 10 years ago. 

Recent upper-seam mining has been largely interactionfree above first workings or over gob areas. These mines 

typically use 5–6 ft (1.5–1.8 m) resin bolts for roof support, supplemented by “superbolts” or cables when crossing lower-seam structures (isolated remnant pillars or gobsolid boundaries). 

 The lower seams are 4–5 ft (1.2–1.5 m) thick, so the 

I/t ratio is approximately 9. Above lower-seam structures, 

the roof can be severely cracked or even “pulverized” into 

small pieces that fall out upon mining and require short 

cuts. Several types of lower-seam structures have been 

encountered: 

• Gob-solid boundaries, where the “solid” can be either 

unmined coal or development pillars, and is at least 

150 ft (45 m) wide; 

• Isolated remnant pillars and narrow barriers, approximately 50–100 ft (15–30 m) wide with gob on at least 

two sides; and 

• Sandstone channels that were left between gob areas. 

Depending on the width of the channel, these cases 

were classified into one of the previous two 

categories. 

Figure 3.—Interactions resulting from undermining remnant structures with just 20 ft (6 m) of interburden. 
58

Figure 4 illustrates these different kinds of structures. 

 At the two mines, a total of 23 crossings or attempted 

crossings were analyzed. In 11 of these cases, the panels 

were subsequently pillared. Such panels were analyzed 

twice, first as successful development cases and then as 

retreat mining cases. 

 Tables 1 and 2 show the results. During development, 

difficult or severe interactions were encountered above 

only 2 of the 16 gob-solid boundaries (13%). In contrast, 

of the seven isolated remnants that were overmined, four 

caused interactions that were so severe that they stopped 

mining completely, and a fifth resulted in very difficult 

conditions on advance (71% total). The two relatively 

successful crossings were both at relatively shallow cover. 

 Nine of the successful gob-solid crossings were 

subsequently pillared, about half with minimal problems. 

When pillars were retreated above the two successful 

isolated remnant crossings, one resulted in difficult 

conditions. 

Table 1.—Conditions encountered crossing lower remnant 

structures: development (case history No. 2) 

Condition 

Structure type Minor/None Difficult Severe 

Gob-solid boundary 9 5 2 

Isolated remnant pillar 2 1 4 

Table 2.—Conditions encountered crossing lower remnant 

structures: retreat (case history No. 2) 

Condition 

Structure type Minor/None Difficult Severe 

Gob-solid boundary 4 5 0 

Isolated remnant pillar 1 1 0 

 Clearly, isolated remnants pose much more significant 

hazards at these mines than gob-solid boundaries. The 

obvious explanation is that isolated remnants can result in 

much greater stress concentrations in the adjacent seams. 

First, they normally carry two stress abutments (one from 

each gob), while gob solid boundaries usually only carry 

one. Perhaps just as important is the load distribution that 

develops within the remnant. As Chase et al. [2007] 

point out, three different kinds of pillars may be defined 

based on their load distributions: 

• Small yielded structures that carry relatively small 

loads (Figure 5A); 

• Wide pillars or gob-solid boundaries that have 

localized high-stress zones, but distribute the load 

(Figure 5B); and 

• Isolated remnants that are highly stressed throughout 

(Figure 5C). 

 Wide pillars may carry the same (or even greater) load 

as a remnant, but because their load is distributed over a 

much larger area, their “footprint” is less noticeable in 

seams above or below. 

CASE NO. 3: NEARLY SIMULTANEOUS MINING 

 A mining complex in Kentucky is extracting the 

Kellioka and Darby Seams, which are separated by 40–

70 ft (12–21 m) of interburden. The interburden consists 

largely of sandstone. The depth of cover reaches 2,000 ft 

(600 m), so the H/t ratio can be as high as 50. 

 Both mines are room-and-pillar with full-pillar extraction. Mining is sequenced from the top down. During the 

15 years since mining began on the property, a number of 

 

 Figure 5.—Stress distributions within three types of remnant 

structures.

A B C

 Figure 4.—Different types of remnant structures in the 

 underlying workings. 
59

lessons have been learned and incorporated into mine 

planning. 

 Figure 6 shows three early attempts to develop production panels beneath fully extracted Darby Seam works. In 

each case, severe roof conditions above thin barrier pillars 

isolated between two gob areas. The problems were 

encountered despite modifications to the pillar size and 

supplemental roof support. 

 Subsequently, Kellioka workings have been laid out to 

parallel the overlying workings. The width of the Kellioka 

retreat panels, including slab cuts, exactly matches that of 

the Darby gobs. With this panel stacking design, most of 

the lower-seam panel development and pillar recovery 

takes place under the Darby gob. The potential difficulty 

with panel stacking is that the development must cross a 

gob-solid boundary in order to access the reserve beneath 

the gob. At the time NIOSH visited the complex, seven 

lower-seam panels had been successfully extracted using 

the stacking design. Although some difficult conditions 

were encountered at gob-solid boundaries, they were much 

less severe than those associated with the thin isolated 

barrier pillars. 

 In the early planning, the mine tried to wait at least 

6 months after completion of the upper-seam retreat mining before developing the lower-seam works. However, 

experience showed that “settling time” did not have a large 

effect on the conditions encountered. Recently, some 

lower-seam developments have begun as early as 1 month 

after the overlying panel was extracted. 

CASE NO. 4: UNDERMINING PREEXISTING 

WORKINGS 

 In southern West Virginia, mining on several properties has been conducted in as many as 10 seams. Longwalls have mined large portions of the Powellton Seam 

and are currently working near the bottom of the geologic 

column in the Eagle and No. 2 Gas Seams. 

 NIOSH studies found numerous instances of successful mining above previously longwalled areas. In most of 

these cases, the interburden between the target seam and 

the longwall gob is at least 180 ft (55 m). 

Figure 6.—Interactions caused when trying to undermine thin isolated barrier pillars in an overlying seam. 
60

 There have been several instances in which longwalls 

undermined open entries, usually mains, in overlying 

mines. The results have almost always been unsatisfactory. 

In one instance, a mine was maintaining main entries in the 

9-ft (2.7-m) Coalburg Seam, 560 ft (170 m) above the 6-ft 

(1.8-m) No. 2 Gas. The I/t ratio in this case was nearly 

100, and the overburden-interburden ratio was less than 

1.0. In addition, over 50% of the overburden was sandstone, and the immediate roof consisted of competent 

sandstone (CMRR=70). Finally, 16-ft (4.9-m) vertical 

cable bolts and cable straps were installed together with 

standing supports (steel props). 

 The longwall directly undermined the mains, as shown 

in Figure 7. Within days, the Coalburg Seam subsidence 

measured 36–42 in (0.9–1.1 m). The immediate roof was 

severely fractured, with some open apertures of 4 in 

(100 mm). Numerous large roof falls resulted (Figure 8). 

 This example, and at least four others in the database, 

show that the normal subsidence prediction rules are 

completely inapplicable when open entries are involved. 

The reasons are not hard to understand. Referring again to 

Kendorski’s overburden mechanics model, the Coalburg 

Seam would normally have been safely within the “confined zone” within which the ground subsides but no new 

fracturing takes place. However, the entries removed the 

compressive confining pressure, so the rock around the 

mine openings was subjected to severe tensile stress. If the 

mains had been developed after the longwall had been 

extracted, there might have been no obvious evidence of 

its passage. 

 Another curious case in the database involved a mining complex in Kentucky. A room-and-pillar panel was 

retreated in the Pond Creek Seam, and approximately 

2 years later a set of main entries was developed in the 

Cedar Grove Seam 180 ft (55 m) above. The I/t ratio was 

about 5, and conditions were initially excellent with just 

4-ft (1.2-m) fully grouted bolts. After about 2 years, however, the roof began to deteriorate dramatically. Extensive 

supplemental support, including full cable bolting, wood 

cribs, and polyurethane injection, eventually had to be 

installed. The most likely explanation is that ground 

between the two seams had not fully subsided when the 

upper-seam entries were developed. When it did subside 

later, it apparently caused the same kind of damage as in 

the longwall case described previously. 

CONCLUSIONS 

 The case histories presented in this paper, and others 

contained in the NIOSH database, clearly show that the 

existing multiple-seam guidelines should be refined. One 

general point is that it is seldom possible to evaluate the 

minability of an entire reserve with broad criteria based on 

factors like the extraction ratio or interburden thickness. 

Multiple-seam interactions are highly localized, so it is 

necessary to evaluate the interaction potential from each 

structure left in the previously mined seam. 

 Experience seems to show that where the previous 

mining has been limited to development, it may have little 

impact on reserves separated by interburdens of as little as 

20 ft (6 m). In one instance described in this paper, 

columnization was not even necessary. Similarly, mining 

beneath gob areas, where the ground has been largely 

destressed, seldom presents serious problems. 

 When mining above gob areas, some roof fracturing 

can be expected up to a distance of perhaps 24 times the 

lower-seam height. However, the mines described here 

have encountered few difficulties even just 40 ft (12 m) 

above gob areas. 

 Difficult ground conditions are often encountered 

when crossing from the solid into the gob (or vice versa). 

However, by using control techniques including longer 

 Figure 8.—Damage caused when open entries were 

 subsided. 

 Figure 7.—Longwall mining that subsided open entries 

560 ft (170 m) above. 
61

pillars, narrower entries, and additional roof support, the 

mines described in this paper have been able to cross most 

of these structures. 

 The most serious interactions occur above or beneath 

isolated remnant pillars, normally 40–100 ft wide, located 

between two gob areas. These types of structures are 

apparently too wide to have yielded, but too narrow to 

effectively distribute the load. The high stresses associated 

with these types of structures have often stopped mining 

completely. A reserve area that contains many such isolated remnants may indeed be unminable. 

 Severe interactions are also likely if an open entry is 

undermined by longwall or pillar extraction. A large interburden thickness, a low depth of cover, and even strong 

roof may be no protection from the damage caused when 

an open entry is subsided. 

 The case histories seem to indicate that the necessary 

time lag between mining the two seams may not be fully 

understood. In one instance, as little as 1 month seemed to 

be an adequate “settling time”; in another case, 4 years 

may not have been enough. The important factor may not 

be the elapsed time, but whether the subsidence is complete. Observations from longwall mining indicate that 

subsidence at the surface is often complete within weeks, 

while some abandoned mines have collapsed decades after 

mining [Iannacchione and Mark 1990]. 

 NIOSH is continuing its evaluation of the entire 

multiple-seam case history database. The results will be 

used to develop suggested guidelines for analyzing potential multiple-seam interactions. It is hoped that these guidelines will help mine operators to more safely extract the 

increasingly difficult reserves in central Appalachia and 

elsewhere. 

REFERENCES 

 Bate RL, Kvitkovich JF [2004]. Quantifying the coal 

reserve dilemma in the central Appalachian mining region. 

SME preprint 04–108. Littleton, CO: Society for Mining, 

Metallurgy, and Exploration, Inc., 8 pp. 

 Chase FE, Worley P, Mark C [2007]. Multiple-seam 

mining interactions: case histories from the Harris No. 1 

mine. Paper in these proceedings. 

 Haycocks C, Zhou YX [1990]. Multiseam mining: 

a state-of-the-art review. In: Proceedings of the Ninth 

International Conference on Ground Control in Mining. 

Morgantown, WV: West Virginia University, pp. 1–11. 

 Iannacchione AT, Mark C [1990]. Possible mechanism for surface vibrations near Maxwell Hill, West Virginia. Bull Assoc Eng Geol 27(3):341–353. 

 Kendorski FS [1993]. Effect of high-extraction coal 

mining on surface and ground waters. In: Proceedings of 

the 12th International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 412–425. 

 Lazer B [1965]. Mining seams above mined-out lower 

seams. Min Eng Sep:75–77. 

 Luo J, Haycocks C, Karmis M [1997]. Gate road 

design in overlying multiple-seam mines. SME preprint 

97–107. Littleton, CO: Society for Mining, Metallurgy, 

and Exploration, Inc. 

Michalek SJ, Wu KK [2000]. Potential problems 

related to mining under or adjacent to flooded workings. 

In: Peng SS, Mark C, eds. Proceedings of the 19th International Conference on Ground Control in Mining. 

Morgantown, WV: West Virginia University, pp. 199–203. 

 Westman E, Luo J, Haycocks C, Karmis M [1997]. 

Ground control criteria for coal reserve estimation in 

multiple-seam mines. In: Peng SS, ed. Proceedings of the 

16th International Conference on Ground Control in 

Mining. Morgantown, WV: West Virginia University, 

pp. 311–315. 

This paper was previously published as: 

Mark C [2006]. Extreme multiple-seam mining in the central Appalachian coalfields. SME preprint 06–060. 

Littleton, CO: Society for Mining, Metallurgy, and Exploration, Inc. 

63

MULTIPLE-SEAM MINING INTERACTIONS: 

CASE HISTORIES FROM THE HARRIS NO. 1 MINE 

By Frank E. Chase,1

 Phyllip Worley,2 and Christopher Mark, Ph.D., P.E.3

ABSTRACT 

The Harris No. 1 Mine, located in Boone County, 

WV, has been longwalling the Eagle Coalbed for over 

30 years. Harris has experienced numerous interactions 

associated with the extensive room-and-pillar and longwall 

mining operations that have been conducted in the overlying No. 2 Gas Coalbed. The problems have included roof 

falls, excessive rib sloughage, and gate road and bleeder 

entry closure. A detailed evaluation of the multiple-seam 

experiences at Harris No. 1 Mine was conducted as part of 

the National Institute for Occupational Safety and Health 

(NIOSH) nationwide multiple-seam mining case history 

database. One observation from the Harris gate road case 

histories was that smaller, critically loaded, upper-seam 

pillars seemed to cause more severe ground conditions 

than wider pillars. The LaModel program was used to 

investigate this supposition. The results confirmed that 

“critical” sized pillars do transmit the highest amounts of 

stress to adjacent seams. In addition, the data suggest that 

the probability of a major multiple-seam mining interaction increases when the depth of cover is 1,000 ft or 

greater and when the Eagle Seam pillars have an Analysis 

of Longwall Pillar stability factor less than 1.50. 

INTRODUCTION 

 NIOSH recently completed a comprehensive nationwide database of multiple-seam mining case histories. To 

collect the case histories, underground geotechnical evaluations were conducted at more than 45 U.S. coal mines. 

The data are currently being analyzed to ascertain the 

relative importance of the various contributory mining and 

geologic parameters responsible for multiple-seam mining 

interactions. The ultimate goal is to provide the mining 

community with a design methodology for multiple-seam 

mining that will aid in determining the likelihood of 

adverse interactions so that corrective measures can be 

taken to prevent injuries and fatalities. 

 During the study, 22 multiple-seam case histories 

were collected from the Harris No. 1 Mine, more than at 

any other mine site. An area was deemed to be a case history if a multiple-seam interaction occurred or should have 

been anticipated. The accumulation of such a significant 

number of cases over a relatively small geographic area 

presented an excellent opportunity to conduct a study that 

would evaluate the current state of the art in multiple-seam 

design. In other words, can the criteria that engineers 

employ to predict whether or not a multiple-seam interaction will occur be used to explain Harris’ experiences? 

 The Harris No. 1 Mine is operated by Eastern Associated Coal Corp., a subsidiary of Peabody Energy. Harris is 

located in Wharton, Boone County, WV, (Figure 1) and 

began operations in 1966. Since then, Harris has driven 

and retreat mined more than 60 longwall panels in the 

Eagle Coalbed. The No. 2 Gas Coalbed is situated approximately 200 ft above the Harris Mine workings. Both 

longwall and room-and-pillar retreat mining have been 

conducted in the No. 2 Gas. In many cases, remnant structures such as barrier pillars, isolated gate roads (gate roads 

that are bordered by gob on both sides), etc., that were left 

in the 2 Gas have caused difficult ground conditions in 

Harris due to downward load transfer. In other instances, 

upper-seam structures have not noticeably impacted mining. From the mine planning perspective, the paramount 

question is: When will multiple-seam problems occur, and 

how severe will the interaction be? The purpose of this 

investigation was to shed some light on these questions by 

conducting detailed analyses of Harris’ experiences. 

Wharton

Charleston

Figure 1.—Location of Harris No. 1 Mine in Wharton, WV.

 

 1Research geologist, Pittsburgh Research Laboratory, National

Institute for Occupational Safety and Health, Pittsburgh, PA.

 2Senior manager of engineering, Peabody Energy, Wharton, 

WV. 

 3

Principal research mining engineer, Pittsburgh Research

Laboratory, National Institute for Occupational Safety and

Health, Pittsburgh, PA. 
64

GEOLOGIC SETTING

 The topography above Harris No. 1 Mine is fairly 

rugged. The valleys are narrow and V-shaped, and ridges 

are steep and prominent. These physiographical features 

can cause rapid changes in cover over relatively short 

horizontal distances. The overburden at Harris ranges from 

100 ft at the drift to slightly over 1,400 ft under the highest 

ridges. As is the case with most central Appalachian coal 

mines, the overburden is relatively competent. 

 Previous researchers [Holland 1947; Stemple 1956; 

Haycocks et al. 1982] have determined correlations 

between multiple-seam interactions and the interburden 

competency, thickness, and number of interbeds (number 

of distinct rock units within the interburden). Therefore, 

considerable emphasis was placed on obtaining corehole 

information as close to the case history sites as possible. 

The information on interburden characteristics is listed in 

the appendix to this paper. As indicated in the appendix, 

the interburden between the Eagle and No. 2 Gas ranges in 

thickness from 176 to 213 ft. 

 Figure 2 is a generalized stratigraphic column of the 

interburden between the No. 2 Gas and Eagle Coalbeds. 

It should be noted that the major sandstone and shale units 

shown in Figure 2 vary in thickness. For example, in a few 

of the coreholes the upper two sandstone units merge into 

a 100-ft-thick unit. The same can be said for the lower two 

sandstone units. These rock unit thickness variations suggest ancient stream channel activity. Usually, the interburden contains six distinct rock units; however, the actual 

number varies from four to seven. In general, the interburden is rather competent, with the percentage of sandstone, sandy shale, and limestone ranging from 59% to 

80%. The coalbeds between the Eagle and No. 2 Gas 

shown in Figure 2 have not been mined above Harris. 

 Another factor identified in determining the magnitude of the interaction is the immediate roof rock competency [Luo et al. 1997]. The shale unit shown in Figure 2 

directly above the Eagle Coalbed varies in thickness from 

0 to 10 ft. In areas of Harris, this shale unit can be either 

laminated, sandy, or nonexistent (replaced by a sandstone 

scour). These fluctuations explain the range in Coal Mine 

Roof Rating (CMRR) values [Molinda and Mark 1994] 

from 44 to 71. These values indicate that the immediate 

roof rock is moderately strong to strong. 

 

GATE ROAD DESIGN AND SUPPORT 

 Harris began longwall operations with a 300-ft-wide 

plow face and 40-ton walking frames in 1966. Since then, 

numerous technological innovations have led to improvements in the longwall systems and gate road supplemental 

supports used. Currently, Harris is mining 3.2 million 

clean tons of coal per year. Gate road pillar design and 

supplemental support selection have also gone through an 

evolutionary process at Harris based on the performance of 

past longwall faces and gate roads. In fact, 12 different 

gate road designs that incorporated various elements of a 

three-entry, four-entry, and yield pillar designs have been 

tried at Harris. The gate road system design was progressively refined and calibrated through the back analyses 

of previous successful and not so successful mining 

attempts. 

 The engineers at Harris use the novel approach of 

integrating the multiple-seam stress transfer values 

obtained from the LaModel program [Heasley 1998] into 

the Analysis of Longwall Pillar Stability (ALPS) program 

[Mark 1990] in order to obtain a more realistic stability 

factor (SF). This methodology is described later in the 

“Discussion” section of this paper. For the past 5 years, 

Harris has been using a three-entry gate road system with 

entries on 90-ft centers and crosscuts on 140-ft centers. 

This system has worked well, and no gate road blockages 

have occurred since its usage began. Based on past experiences, during mine design Harris’ engineers adhere to the 

following rules of thumb as much as possible: (1) the long 

axis of the panel to be mined should be parallel to that of 

the upper-seam panel; (2) the future headgate should be 

positioned under and as close to the center of the gob as 

possible; and (3) avoid advancing the longwall face under 

a gob/solid boundary [Hsiung and Peng 1987a,b]. 

 Harris uses 5-ft full-column resin bolts on 4-ft centers 

in the headgate entry. In the remaining gates and bleeders, 

4-ft full-column resin bolts on 4-ft centers are standard. 

The roof control plan also stipulates that a minimum of 

two crib equivalents be installed every 12 ft in the tailgate. 

No. 2 G as

Lower No. 2 G as

Pow ellton "A"

Pow ellton

M atewan

U pper Eagle

Eagle

Legend

 Coal

 Shale

 Sandstone

0

50'

100'

150'

200'

 Figure 2.—Generalized interburden stratigraphy.
65

Floor heave has always been a major concern at Harris. 

Because conventional cribs (both four- and nine-point) are 

inclined to roll out when subjected to heave, Harris began 

using 30-in engineered timber supports. These supports 

have performed well in that the floor tends to heave up 

around the supports. 

 The engineers at Harris also use the LaModel program 

to identify high vertical stress areas that are caused by 

deep cover, abutment loads, and/or multiple-seam stress 

transfer [Heasley and Agioutantis 2007]. In highly stressed 

areas, either two or four 12-ft-long cable bolts are installed 

in between each row of primary supports. Sometimes 

additional engineered timber supports are warranted in tailgate locations. The spacing of these supports is dependent upon the expected level of stress. 

CASE HISTORY ANALYSES

 A detailed examination of both the No. 2 Gas and 

Harris No. 1 workings (Figure 3) revealed 22 case histories 

where multiple-seam interactions happened or might have 

been anticipated. In each case history, gate roads were 

driven and panels were extracted under various upperseam structures, and the outcomes are listed in the 

Figure 3.—No. 2 Gas workings superimposed on Harris No. 1 Mine.
66

appendix. Overburden depth, interburden thickness and 

composition, and additional consequential mining parameters that are believed to determine whether or not 

interactions will occur [Holland 1947; Stemple 1956; 

Haycocks et al. 1982] are also listed in the appendix. Prior 

to the analyses, the database was separated into two categories—gate entry workings (17 cases) and longwall face 

stability (5 cases)—because of the major differences 

between the two. A rating system from 1 to 6 (see the 

appendix for details) was developed to numerically evaluate the conditions or degree of interaction for each case. 

For the purpose of analyses, conditions 1 and 2 were 

combined and categorized as a “minor” interaction because 

the interactions were “barely negligible” to “minor.” 

Conditions 3 through 6 were combined and designated as a 

“major” interaction because the interactions were 

“troublesome” to “major” and warranted that special measures be taken. 

 A series of XY scatter plots were generated to examine the various mining and geologic parameters for correlations. Figure 4 shows that six out of seven of the major 

interaction gate road workings cases occurred when 

Harris’ depth of cover was 1,000 ft or greater and the 

ALPS SF was less than 1.50. Further, Figure 5 shows that 

five out of seven of the major interaction cases occurred 

when the No. 2 Gas ALPS SF was less than 1.0 and the 

depth of cover was 1,000 ft or more in Harris. Finally, 

Figure 6 illustrates a weak correlation between problematic 

cases and a No. 2 Gas overburden/interburden ratio of 3.9 

or greater. As for the five longwall face stability cases, the 

only parallels that could be drawn were that the depth of 

cover was primarily 1,000 ft or greater and the immediate 

roof rock was generally relatively weak. Upper-seam pillar 

design did not seem to be an issue; however, both it and 

the findings mentioned in this section warrant additional 

examination and discussion. 

UPPER-SEAM PILLAR DESIGN 

 As indicated in the previous section, most of the 

multiple-seam interaction problems in Harris’ gate entries 

occurred when the upper-seam ALPS SFs were less than 

1.00. At first, it might seem counterintuitive that smaller 

upper-seam pillars would cause more severe stress conditions in an underlying seam than wider pillars. However, 

a consideration of the load distribution in the upper-seam 

pillars provides an explanation. Essentially, three load distributions are possible: 

• Figure 7A illustrates a small, yielded pillar that 

carries a relatively small load; 

• Figure 7B illustrates a wide pillar with localized 

high-stress zones near the ribs, but a lightly 

loaded core; and 

• Figure 7C illustrates the load distribution of a 

critical pillar, with a highly loaded core. 

 Figure 4.—Relationship between degree of interaction 

and the Harris No. 1 ALPS stability factors and overburden.

0

200

400

600

800

1000

1200

1400

012345

ALPS Stability Factors

Overburden (ft)

MINOR INTERACTION

MAJOR INTERACTION

 Figure 5.—Relationship between degree of interaction 

and the No. 2 Gas ALPS stability factors and Harris No. 1 

overburden.

0

200

400

600

800

1000

1200

1400

01234

No. 2 Gas Stability Factors

Harris No. 1 Overburden (ft)

"Minor Interaction"

Major Interaction

0

200

400

600

800

1000

1200

1400

01234567

No. 2 Gas Overburden/Interburden Ratio

Harris No. 1 Overburden (ft)

MINOR INTERACTION

MAJOR INTERACTION

 Figure 6.—Relationship between degree of interaction and 

the No. 2 Gas overburden/interburden ratio and Harris No. 1 

overburden.
67

 The critical pillar would result in the most severe 

“footprint” on the lower seam, because it produces an 

intensified downward “point load” type of stress transfer to 

the underlying workings. The wide pillar may carry a 

larger total load, but because that load is distributed over a 

much larger area, its effect on the lower seam is less 

noticeable. A good analogy would be the imprints that a 

petite woman in high heels might make in wet sand compared with those made by a sizable football player wearing 

tennis shoes. 

 LaModel, a displacement-discontinuity boundaryelement program, was used to evaluate the hypothesis 

described above. The models were run using standard 

default parameters and yield zones. Figure 8 displays the 

basic layout of the two mine designs that were modeled. In 

the Harris design case, a three-entry longwall gate entry 

development section (oriented from top to bottom in 

Figure 8) was driven on 120-ft-entry and crosscut centers 

in a 6-ft-high reserve. The pillars had an ALPS SF of 3.07, 

and the depth of cover was 1,200 ft. A three-entry isolated 

gate road system (oriented from left to right in Figure 8) 

was then situated 200 ft above Harris. The crosscut center 

spacing in the No. 2 Gas remained constant at 140 ft. The 

entry centers were varied from 30 to 180 ft in 10-ft 

increments for each LaModel run, and the mining height 

was 6 ft. As shown in Figure 8, the No. 2 Gas and Harris 

workings are situated perpendicular to one another so that 

four pillars were stacked in the center of the LaModel grid. 

Figure 8 also displays the LaModel analysis results for a 

No. 2 Gas gate road system with 60-ft-wide pillars. 

Figure 8 clearly shows that the multiple-seam stress transfer magnitudes in Harris are the highest beneath the isolated gate roads. Conversely, the destressing effects of the 

overlying gob are also evident in Figure 8. 

 Figure 9 displays the peak multiple-seam stress transfer value and the ALPS SF for each pillar width modeled. 

Figure 9 shows the wide range in multiple-seam peak 

stress transfer values, which are dependent on the width of 

Multiple Seam Stress SCALE

 -500

 -385

 -270

 -160

 -50

 70

 185

 300

 410

 525

 635

 ( psi )

 Figure 8.—LaModel output for 60-ft-wide No. 2 Gas

pillars. 

0

100

200

300

400

500

600

700

0 20 40 60 80 100 120 140 160 180

No. 2 GAS PILLAR WIDTH (ft)

MULTIPLE SEAM STRESS (PSI) .

0

0.5

1

1.5

2

2.5

3

3.5

ALPS STABILITY FACTOR .

MULTIPLE SEAM STRESS

ALPS SAFETY FACTOR

 WIDE

PILLAR

ZONE

CRITICAL

PILLAR

ZONE

YIELDED

 PILLAR

 ZONE

Figure 9.—Peak multiple-seam stress transfer values and the ALPS SFs for modeled pillar widths.

A B C

 Figure 7.—Pillar load distribution diagrams: (A) yielded,

(B) wide, (C) critical. 
68

the pillar. The multiple-seam stress transfer curve in 

Figure 9 seems to have three distinct regions that correspond to the three upper-seam load distributions shown in 

Figure 7. The peak or “critical” multiple-seam stress transfer values occur when the chain pillars in the upper seam 

are in the 50- to 90-ft range. The models indicated that the 

cores of these pillars were all heavily loaded. On the left 

side of the critical pillar region, the models showed that the 

stresses in the cores of smaller, upper-seam pillars were 

much lower than for the critical pillars. The smaller the 

pillar, the lower the peak stress and the less the multipleseam stress experienced in the lower seam. On the right 

side of the critical pillar region, as the upper-seam pillars 

get wider, they distribute their load more evenly. The 

result is a steady decreasing trend in downward stress 

transfer as the pillar width is increased up to around 130 ft. 

Once the pillar reaches a certain width, there is essentially 

no interaction between the two high-stress zones at the 

ribs, and the peak stress transfer levels out at approximately 350 psi. 

DISCUSSION 

 For lack of a better adjective, the term “critical” was 

used to describe the pillars whose size transferred the 

highest multiple-seam stress values. Obviously, the word 

“critical” conjures up different meanings depending on 

whether you are designing deep-cover gate road yield 

pillars or mining in bump-prone ground conditions. However, from a multiple-seam aspect, the LaModel analyses 

indicate that critically sized upper-seam pillars can 

increase the lower-seam pillar stresses substantially. In this 

study, the LaModel results were used to calculate the 

average stress increase in a Harris tailgate pillar system 

caused by isolated No. 2 Gas gate roads on 80-ft-wide 

entry centers. The calculated average multiple-seam pillar 

stress was 396 psi, which is approximately equivalent to 

increasing the depth of cover by 360 ft. Therefore, a Harris 

tailgate system that was initially designed for 1,200 ft of 

overburden with a conservative ALPS SF of 1.23 was, 

in actuality, being subjected to cover loads equivalent to 

1,560 ft of overburden, which effectively reduces the 

ALPS SF to 0.88. This example emphasizes the 

importance of both estimating and incorporating multipleseam stress transfer into the pillar design process. It 

implies that wider pillars with higher ALPS SFs should be 

used; however, gate road developmental constraints also 

need to be considered. The engineers at Harris are currently using this methodology to design gate road pillar 

systems and, based on past experiences, an ALPS SF in the 

1.0–1.2 range (taking into account the additional multipleseam stress) has been determined to provide satisfactory 

results. It should be noted that the stress transfer values 

and critical pillar dimension widths previously mentioned 

are case-specific and will vary depending on the input 

parameters. 

 As stated in the “Case History Analyses” section, six 

out of the seven major interactions occurred when the 

Harris depth of cover was 1,000 ft or greater and the ALPS 

SF was less than 1.50 (Figure 4). The cover relationship is 

noteworthy in that most operators maintain that there is a 

correspondence between multiple-seam interaction difficulties and overburden. Typically, operators state that 

problems generally begin occurring at roughly 800 ft of 

cover. Essentially, it takes a certain amount of cover load 

to cause downward load transfer problems. One possible 

explanation for the higher cover value at Harris may be 

interburden competency. It is conceivable that the three 

sandstone units, which comprise 59%–80% of the interburden, are bridging and therefore dampening the downward load transfer. As for the Harris ALPS SFs, Figure 4 

suggests that the probability of a major interaction occurring decrease as the stability factor increases. The same 

can be said for the No. 2 Gas ALPS SFs. As shown in 

Figure 5, five out of seven (71%) of the major interaction 

cases occurred when the No. 2 Gas ALPS SF was less than 

1.0 and the depth of cover was 1,000 ft or more in Harris. 

Based on the abovementioned findings, a certain amount 

of concern and supplemental support are probably warranted when dealing with deep cover and lower upper- and 

lower-seam ALPS SFs. As the old longwall adage goes, 

“it is better to be safe than be shut down.” (It should be 

noted that multiple-seam stress transfer values were not

taken into account when determining the ALPS SFs listed 

in the appendix or shown in the figures.) 

 Data analyses also indicated that there was no 

relationship between the degree of interaction and the 

percentage of competent interburden. The same can be said 

for the interburden thickness/number of beds ratio. Conversely, there was a weak correlation with immediate roof 

rock competency. Generally, the CMRR was higher for the 

minor interaction cases. Another weak association previously indicated was the overburden/interburden thickness 

ratio value of 3.9. As a rule of thumb, problems generally 

do not occur until this ratio reaches 7 or 8. However, critically sized pillars may be an overriding factor in this particular situation. 

CONCLUSIONS 

 The most significant finding of this investigation was 

that the size of the remnant upper-seam structure can 

influence the extent of the multiple-seam interaction. More 

specifically, this study suggests that smaller, critically 

loaded upper-seam pillars are more likely to cause lowerseam ground control problems than wider pillars. The 

LaModel program was used to examine this supposition, 

and the results verified this premise. 

 This investigation also demonstrated the effectiveness 

of LaModel in determining multiple-seam stress transfer 

magnitudes. Once this value is obtained, it can be incorporated into the ALPS or the Analysis of Retreat Mining 
69

Pillar Stability (ARMPS) programs to obtain a more realistic stability factor. 

 The back analyses of 17 gate road case histories at 

Harris No. 1 indicate that the probability of a major 

multiple-seam mining interaction occurring increases when 

(1) the depth of cover is 1,000 ft or greater, (2) the upperseam pillars are critically loaded, and (3) the Eagle Seam 

pillars have a nonadjusted ALPS SF (excludes multipleseam load transfer) less than 1.50. In areas where these 

criteria are met, Harris engineers have mitigated problems 

through pillar design modifications and the installation of 

supplemental support. Based on past experiences, the engineers at Harris have determined that an adjusted ALPS SF 

in the 1.0–1.2 range provides satisfactory results. 

 Finally, the analyses also identified a weak correlation 

between the degree of multiple-seam interaction and 

the immediate roof rock competency (CMRR) and the 

overburden/interburden thickness ratio. However, no 

relationship between the degree of interaction and the 

percentage of competent interburden or the interburden 

thickness/number of beds ratio was evident. This may be 

attributable to the lack of variability in this site-specific 

database. Possibly, the conclusions drawn from the analyses of the nationwide multiple-seam mining database will 

concur with previous researchers’ findings. 

ACKNOWLEDGMENTS 

 The authors express their gratitude to Keith A. 

Heasley, Ph.D., Associate Professor of Mining Engineering, West Virginia University, for his technical advice 

during the data analysis, interpretation, and review phases 

of this project. We also acknowledge the contributions of 

John K. Marshall, Electronics Technician, NIOSH Pittsburgh Research Laboratory. In addition, special thanks go 

to Eastern Associated Coal Corp. and Peabody Energy for 

providing the mining data and for allowing their valuable 

experiences to be published. 

REFERENCES 

 Haycocks C, Karmis M, Ehgartner B [1982]. 

Multiple-seam mine design. In: Proceedings of the International Symposium on Ground Control in Longwall Coal 

Mining and Mining Subsidence: State of the Art. Littleton, 

CO: American Institute of Mining, Metallurgical, and 

Petroleum Engineers, pp. 59–65. 

 Heasley KA [1998]. Numerical modeling of coal 

mines with a laminated displacement-discontinuity code 

[Dissertation]. Golden, CO: Colorado School of Mines, 

Department of Mining and Earth Systems Engineering. 

 Heasley KA, Agioutantis ZG [2007]. LaModel: 

a boundary-element program for coal mine design. Paper 

in these proceedings. 

 Holland CT [1947]. Effects of unmined seams of coal 

by mining seams above and below. In: Proceedings of the 

West Virginia Academy of Science. Huntington, WV: 

Marshall University, pp. 113–132. 

Hsiung SM, Peng SS [1987a]. Design guidelines for 

multiple-seam mining, part I. Coal Min 24(9):42–46. 

 Hsiung SM, Peng SS [1987b]. Design guidelines for 

multiple-seam mining, part II. Coal Min 24(10):48–50. 

 Luo J, Haycocks C, Karmis M [1997]. Gate road 

design in overlying multiple-seam mines. SME preprint 

97–107. Littleton, CO: Society for Mining, Metallurgy, 

and Exploration, Inc. 

 Mark C [1990]. Pillar design methods for longwall 

mining. Pittsburgh, PA: U.S. Department of the Interior, 

Bureau of Mines, IC 9247. NTIS No. PB 90–222449. 

 Molinda GM, Mark C [1994]. Coal mine roof rating 

(CMRR): a practical rock mass classification for coal 

mines. Pittsburgh, PA: U.S. Department of the Interior, 

Bureau of Mines, IC 9387. 

 Stemple DT [1956]. A study of problems encountered 

in multiple-seam coal mining in the eastern United States 

[Thesis]. Blacksburg, VA: Virginia Polytechnic Institute 

and State University. 

This paper was previously published as: 

Chase FE, Worley P, Mark C [2005]. Multiple-seam mining interactions: case histories from the Harris No. 1 

mine. In: Peng SS, Mark C, Tadolini SC, Finfinger GL, Khair AW, Heasley KA, eds. Proceedings of the 24th 

International Conference on Ground Control in Mining. Morgantown, WV: West Virginia University, pp. 79–86. 
70

APPENDIX.—HARRIS NO. 1 CASE HISTORY DATABASE 

Case H h LC SF INT COMP 

INT % 

No. 2 

Gas h LC SF H/INT

# 

Beds Angle CMRR Rating and comments 

1 750 6.6 TGL 1.34 180 62 6.0 ISO 1.69 3.2 25.7 33 54 

(1) No problems were encountered 

while crossing under isolated gate 

roads. 

2 925 6.3 TGL 1.01 180 62 6.0 ISO 0.69 4.1 25.7 33 54 

(1) No problems were encountered 

while crossing under isolated 

gate roads. 

3 469 6.6 HGL 4.03 180 62 6.0 BL 3.49 1.6 25.7 33 54 

(1) Headgate was driven under 

bleeder entries without any problems. 

Panel was recovered without 

gate road cribbing. 

4 875 6.7 HGL 1.86 176 59 6.0 HGL 1.15 4 25.1 32 63 

(1) No problems were encountered 

during gate road advance or panel 

retreat under gate road pillars. 

5 725 6.8 TGL 1.37 199 72 6.5 BL 2.90 2.6 33.2 0 47 

(1) Gate roads were successfully 

driven under longwall bleeder entries 

the entire length of the panel. 

6 800 6.8 TGL 1.18 180 62 6.0 ISO 1.50 3.3 25.7 33 54 

(2) Additional gate entry cribbing was 

required while crossing under 

isolated gate roads. 

7 1,000 6.1 TGL 0.90 193 74 6.0 ISO 0.93 4.2 32.2 33 56 

(2) Additional gate entry cribbing was 

required while crossing under 

isolated gate roads. 

8 1,100 6.2 DEV 2.29 193 74 6.0 ISO 0.75 4.7 32.2 33 56 

(2) Poor ground conditions required 

cable bolting on development while 

crossing under isolated gate roads. 

9 800 6.8 TGL 1.20 193 74 6.0 ISO 0.87 3.1 32.2 33 56 

(2) Additional gate entry cribbing was 

required while crossing under 

isolated gate roads. 

10 1,050 7.2 DEV 2.16 193 74 6.0 ISO 0.85 4.4 32.2 33 56 

(2) Poor ground conditions required 

cable bolting under isolated 

gate roads. 

11 1,000 7.1 TGL 1.07 201 80 6.0 ISO 0.60 3.9 50.3 25 44 

(3) Tailgate entries located below 

isolated gate roads experienced 

several roof falls. Numerous 

tensioned cable bolts were installed 

on 4-ft centers. 

12 1,200 7.3 TGL 0.52 201 80 6.0 ISO 0.34 5 50.3 25 44 

(3) During face recovery, tailgate 

entries situated below isolated 

gate roads experienced excessive 

floor heave and roof falls. 

13 1,200 6.9 HGL 1.18 176 59 6.0 BL 0.95 5.8 25.1 32 63 (4) During panel recovery, 500 ft of 

tailgate closed. 

14 800 6.1 TGL 1.30 180 62 6.0 BL 1.88 3.4 25.7 33 54 (4) During panel recovery, 1,200 ft of 

the headgate entry heaved closed. 

15 1,000 6.3 HGL 1.49 199 72 6.5 ISO 0.78 4 33.2 0 47 

(5) During panel recovery, the tailgate squeezed closed under a 

headgate. 

16 1,000 6.2 BL 1.46 178 71 6.0 BL 1.42 4.6 35.6 58 71 

(5) During panel recovery, 750 ft of a 

four-entry bleeder system squeezed 

shut. 
71

APPENDIX.—HARRIS NO. 1 CASE HISTORY DATABASE–Continued 

Case H h LC SF INT COMP 

INT % 

No. 2 

Gas h LC SF H/INT

# 

Beds Angle CMRR Rating and comments 

17 1,200 5.8 HGL 1.44 192 66 6.0 BL 0.86 5.3 32 76 44 

(5) The headgate squeezed closed 

beneath bleeder entries after panel 

extraction. 

18 1,181 6.8 LW 

face NA 192 66 6.0 ISO 0.73 5.2 32 14 44 (2) 2 ft of face heave occurred while mining under isolated gate roads. 

19 1,000 6.7 LW 

face NA 199 72 6.5 ISO 0.69 4 33.2 0 47 (2) 2 ft of face heave occurred while 

mining under isolated gate roads. 

20 675 7.6 LW 

face NA 199 72 6.5 ISO 1.77 2.4 33.2 0 47 (2) 2 ft of face heave occurred while mining under isolated gate roads. 

21 1,200 5.7 LW 

face NA 178 79 5.1 LC2 15.86 5.7 44.5 63 62 (5) Longwall face went on squeeze under a gob/barrier pillar boundary. 

22 1,200 6.6 LW 

face NA 213 71 6.5 BL 1.22 4.6 35.5 90 44 

(5) Roof falls and weight on the face 

halted recovery under bleeder/gob 

boundary. 

 Appendix abbreviations Rating scale 

Angle Intersection angle (degrees) 1

BL Bleeder loading 

Panel was developed and retreat mined with little or no evidence of 

multiple-seam interactions. 

CMRR Coal Mine Roof Rating 

COMP Competent 2

DEV Development loading 

H Mining height (ft) 

Panel was developed and retreat mined with minor to moderate floor 

heave (less than 2 ft) and/or rib sloughage (less than 4 ft). Infrequent 

roof falls may also have occurred. 

h Overburden (ft) 

HGL Headgate loading 3

INT Interburden thickness (ft) 

Panel was developed with minor difficulties. On retreat, pillars were 

occasionally abandoned due to roof falls and/or heavy pillar loading. 

ISO Isolated loading 

LC Loading condition 4

LC2 Loading condition 2 (ARMPS) 

Panel was developed with greater difficulties, and several pillars 

were lost on retreat due to adverse conditions. 

LW Longwall 

NA Not applicable 5

SF Stability factor 

Panel was extremely difficult to advance and could not be retreat 

mined. 

TGL Tailgate loading 

# Beds Number of beds in interburden 6

% Percentage 

Ground conditions necessitated that the panel be abandoned on 

development, or deteriorating conditions over time closed the section. 

73 

FAILURE MECHANICS OF MULTIPLE-SEAM MINING INTERACTIONS 

By R. Karl Zipf, Jr., Ph.D., P.E.1

ABSTRACT 

 Multiple-seam mining interactions caused by fullextraction mining, whether due to undermining or overmining, frequently involve tensile failure of the affected 

mine roof. The adverse ground control conditions may prevent mining for both safety and economic reasons. Prior 

research has identified the geometric, geologic, and mining 

factors controlling multiple-seam mining interactions. This 

numerical study examines the mechanics of these interactions using a modeling procedure that (1) incorporates 

the essential constitutive behavior of the rock, such as 

strain-softening of the intact rock and shear and tensile 

failure along bedding planes, and (2) captures the geologic 

variability of the rock, especially the layering of weak and 

strong rocks and weak bedding planes. 

 Specifically, the numerical study considered the effect 

of vertical stress, interburden thickness, and the immediate 

roof quality of the affected seam in both undermining 

and overmining situations. The models show that for 

overburden-to-interburden (OB/IB) thickness ratios of less 

than 5, interactions do not occur and that for OB/IB more 

than 50, extreme interaction is a certainty. In between, the 

possibility of an interaction was found to depend on gob 

width-to-interburden thickness ratio, site-specific geology, 

and horizontal stress to rock strength ratio, in addition to 

the OB/IB ratio. The models also showed that horizontal 

stress was profoundly altered well above or below a fullextraction area and that these changes are likely to 

influence the success or failure of multiple-seam mining. 

The role of horizontal stress in multiple-seam mining 

interactions has received little attention in prior 

investigations. 

 Four factors control the mechanics of multiple-seam 

mining interactions: 

1. Vertical stress concentration; 

2. Horizontal stress concentration; 

3. Stress redirection; and 

4. Bedding plane slip bands. 

 A combination of vertical and horizontal stress 

increase and high-stress gradients in the vicinity of fullextraction areas reorients principal stresses into a very 

unfavorable direction. This seemingly small stress reorientation has a profoundly adverse effect on bedded rock. 

INTRODUCTION 

 Most underground coal mines face multiple-seam 

mining situations with the potential for interactions that 

can pose challenging ground control conditions. Knowing 

the location of prior mining, planning engineers may seek 

to access and mine new reserves above or below old workings. Two common questions arise: 

1. Will workings above or below cause excessive 

stresses in the proposed workings that lead to 

difficult ground control conditions? 

2. Will subsidence from workings below cause ruinous ground control conditions in the upper seam? 

 In addition to ground control issues, multiple-seam 

mining interactions can create other safety issues. For 

example, an interaction can produce pathways for air, gas, 

or water migration that can cause spontaneous combustion 

and inundation hazards. 

 Whether an adverse multiple-seam mining interaction 

occurs or not depends on numerous factors that are well 

documented by many ground control experts [Chekan and 

Listak 1993, 1994; Haycocks and Zhou 1990; Hill 1995; 

Hladysz 1985; Hsiung and Peng 1987a,b]. These include: 

1. Mining geometry – overburden depth, interburden 

thickness, and seam thicknesses; 

2. Mine design – layout, sequence, and percent 

extraction; and 

3. Geology – immediate roof rock quality and interburden rock strength. 

Combination of these factors may lead to various degrees 

of multiple-seam mining interaction ranging from none to 

additional ground support required to abandonment of 

an area. 

 The ground control research program at the National 

Institute for Occupational Safety and Health (NIOSH) is 

seeking to reduce ground control failures resulting from 

multiple-seam mining interactions through the development of design-based control technology. This study 

reexamines the failure mechanics of multiple-seam mining interactions using a new numerical modeling approach 

under development at NIOSH. In this approach, numerical 

models are created with sufficient geologic detail and 

proper constitutive behavior. With these two conditions 

met, numerical models can predict the behavior of the rock 

mass and indicate whether an adverse multiple-seam mining interaction might occur. 

 

 1Senior research mining engineer, Pittsburgh Research

Laboratory, National Institute for Occupational Safety and

Health, Pittsburgh, PA. 
74 

 This study examines three fundamental types of 

multiple-seam interaction: 

1. Undermining (Figure 1A): This situation represents classic top-down multiple-seam mining. The 

upper seam is mined first and abandoned prior to 

mining the lower seam. Gob-solid boundaries, 

barrier pillars, pillar remnants, or other structures 

left over from full-extraction mining in the upper 

seam may cause adverse stress concentrations in 

the lower seam. 

2. Overmining (Figure 1B): This situation represents 

classic bottom-up multiple-seam mining. Fullextraction mining in the lower seam causes the 

upper seam to fully subside prior to its development. In addition to stress concentrations due to 

pillar remnants or gob-solid boundaries in the 

lower seam, the upper seam and surrounding rock 

may suffer damage from subsidence-induced displacement and fracture. 

3. Simultaneous mining (Figure 1C): This situation 

implies that both seams are active simultaneously. 

In the worst case, workings are developed in the 

upper seam followed by full-extraction mining in 

the lower seam. Subsidence of the upper seam 

occurs after the lower-seam workings are 

extracted. 

NIOSH INPUT PARAMETERS FOR 

NUMERICAL MODELING 

 Ground control researchers at NIOSH follow a 

philosophy developed by Gale [Gale and Tarrant 1997; 

Gale 2004; Gale et al. 2004] of “letting the rocks tell us 

their behavior.” Numerical models that are constructed 

with sufficient geologic detail and proper constitutive 

behavior can predict response of the rock mass, including 

deformation, stress redistribution, failure modes, and support requirements. For general modeling of rock behavior 

in coal mine ground control, Itasca’s FLAC program 

[Itasca Consulting Group 2000] contains many useful features, in particular, the SU constitutive model. SU stands 

for the strain-softening, ubiquitous joint model and is ideal 

for simulating laminated coal measure rocks. In essence, 

this constitutive model allows for strain-softening behavior 

of the rock matrix and/or failure along a predefined weakness plane (in this case bedding planes). Failure through 

the rock matrix or along a bedding plane can occur via 

shear or tension, and the dominant failure mode can 

change at any time. Conveniently, the “state” variable 

within FLAC tracks the failure mode in each model element as either shear or tensile failure through the rock 

matrix or along a bedding plane. 

 The SU constitutive model requires four major input 

parameters, namely, cohesion, friction angle, dilation 

angle, and tensile strength for both the rock matrix and the 

 Figure 1B.—In overmining, the lower seam is mined 

first and abandoned, followed by upper-seam mining. 

The upper seam is fully subsided prior to its mining.

 Figure 1A.—In undermining, the upper seam is mined first

and abandoned, followed by lower-seam mining. 

 Figure 1C.—In simultaneous mining, the upper seam 

is developed first, followed by lower-seam mining. Fullextraction mining of the lower seam subsides existing 

upper-seam workings. 
75 

bedding planes. Each of these parameters begins at some 

peak value and decreases to a residual value as postfailure 

strain increases. It is this decrease in parameter value with 

postfailure strain that gives rise to strain-softening behavior of both the rock matrix and the weakness planes. FLAC 

permits an infinite combination of these requisite input 

parameters; however, in order to facilitate rational numerical modeling, NIOSH researchers created an organized 

suite of material input parameters. Figure 2 summarizes 

the names for this suite of “numerical rocks” and the 

corresponding values for the unconfined compressive 

strength (UCS) of the rock matrix and the strength of the 

bedding planes. 

 The strength values shown in Figure 2 are laboratoryscale values determined from standard UCS tests. Alternatively, the point load test provides excellent, economic 

estimates of the UCS. These UCS values require scaling to 

reduce the laboratory values to the field values needed by 

the numerical model. Following the lead of Gale and 

Tarrant [1997], laboratory values of UCS are reduced by 

0.56 universally. 

 The material suite shown in Figure 2 includes very 

weak soils and claylike materials with a UCS of 0.02 MPa 

and weak, medium, and finally strong rocks with a UCS of 

about 150 MPa. Also included is coal, which ranges from 

the most friable with a UCS of 2 MPa to a strong coal with 

a UCS of 12 MPa. The soil material models are isotropic, 

that is, the soil matrix properties are the same as those for 

the horizontal weakness plane. However, the rock models 

exhibit anisotropy since the strength along bedding planes 

is less than the UCS of the rock matrix. Following results 

of point load tests by Molinda and Mark [1996], weak 

rocks are the most anisotropic with the strength along bedding planes about 50% of the rock matrix UCS, while 

stronger rocks have less anisotropy with the strength along 

bedding planes about 90% of the rock matrix. The coal 

models have a similar trend in strength anisotropy, with 

the stronger coal less anisotropic than the weaker coal. 

 Note that in proposing this suite of numerical rocks, 

the UCS of the rock matrix is independent from the 

strength of the bedding planes. In the absence of specific 

data, the user will usually specify the rock matrix and bedding plane strength as a pair, with strength ratio similar to 

that noted by Molinda and Mark [1996] for an extensive 

database of axial and diametral point load tests. However, 

the strength values for the rock matrix and bedding planes 

are independent in the material property suite, and the user 

can specify any value for the bedding plane strength up to 

that of the rock matrix UCS. 

 Also note that the material model suite has a relation 

to the unit ratings in the Coal Mine Roof Rating (CMRR) 

system. Mark et al. [2002] proposed that the CMRR unit 

rating for a coal measure rock layer is comprised of a UCS 

rating for the rock matrix strength and a discontinuity rating for the bedding plane strength. The UCS rating ranges 

from 5 to 30 points for a rock matrix strength ranging from 

0 to 138 MPa as determined from axial point load tests. 

Similarly, the discontinuity rating ranges from 25 to 60 for 

a bedding plane strength ranging from about 6 to 52 MPa 

as determined from diametral point load tests. The proposed material property suite correlates to CMRR unit ratings from 30 to 90 and represents the range from the 

weakest to the strongest coal measure rocks. Figure 3 

shows these relations. Given CMRR unit ratings from core 

logging, the relations shown in Figure 3 provide 

approximate material choices for input to numerical 

models. 

 The material model suite and UCS values shown in 

Figure 2 imply a range of cohesion and friction angle 

values for the rock matrix and bedding planes. Based on a 

Mohr-Coulomb strength model, the UCS of a rock depends 

on cohesion and friction angle as— 

0.01

0.10

1.00

10.00

100.00

1000.00

Soil_0

Soil_1

Soil_2

Soil_3

Soil_4

Soil_5

Rock_A

Rock_B

Rock_C

Rock_D

Rock_E

Rock_F

Rock_G

Rock_H

Rock_I

Rock_J

Coal_1

Coal_2

Coal_3

Coal_4

Strength - MPa .

Laboratory strength of soil

Laboratory strength of rock matrix

Laboratory strength of rock bedding planes

Laboratory strength of coal matrix

Laboratory strength of coal bedding planes

Figure 2.—Laboratory-scale strength of matrix and bedding planes for suite of “numerical rocks.” 
76 

(1) 

where c is the cohesion and φ is the friction angle. Friction angle for the different materials in the suite is assumed 

to vary, as shown in Figure 4. Soil and claylike materials 

have friction angles of 21°, while progressively stronger 

rocks have friction angles up to 36°. This assumption for 

friction angle along with Equation 1 implies the values for 

peak cohesion shown in Figure 5. Thus, the UCS of the 

rock matrix and the bedding plane strength provide two of 

the four major input parameters to the SU constitutive 

model in FLAC. 

 Other major assumptions within this material model 

suite are as follows: 

1. Moduli for the materials range from 1 to 20 GPa, 

as shown in Figure 6. Weaker materials have a 

lower modulus, while stronger materials have a 

higher modulus. The ratio of modulus to UCS of 

the rock matrix varies from about 10,000 for the 

weakest to about 100 for the strongest materials. 

2. Cohesion decreases from its peak value given in 

Figure 5 to a residual value of 10% of peak over 

5 millistrain of postfailure strain. 

3. Friction angle remains constant at the values 

shown in Figure 4, even in the postfailure regime. 

4. Tensile strength is equal to cohesion for the soil 

materials and decreases to 0 over 1 millistrain of 

postfailure strain. 

5. Tensile strength ranges from about 10% of UCS 

for the weakest rocks to about 2% of UCS for the 

φ

φ

1 sin

2 cos

− = c UCS

 Figure 3.—Correlation of “numerical rocks” to unit ratings of the Coal Mine Roof Rating (CMRR) system. Weak rocks 

have a CMRR unit rating from 30 to 45; moderate rocks, from 45 to 60; and strong rocks, from 60 to 85. 

0

10

20

30

40

50

60

70

80

90

100

Rock_A Rock_B Rock_C Rock_D Rock_E Rock_F Rock_G Rock_H Rock_I Rock_J

Rating .

Matrix Strength Rating

Discontinuity Rating

CMRR Unit Rating

0

5

10

15

20

25

30

35

40

Soil_0

Soil_1

Soil_2

Soil_3

Soil_4

Soil_5

Rock_A

Rock_B

Rock_C

Rock_D

Rock_E

Rock_F

Rock_G

Rock_H

Rock_I

Rock_J

Coal_1

Coal_2

Coal_3

Coal_4

Friction angle (degrees) .

Friction angle of soil

Friction angle of rock matrix

Friction angle of rock bedding planes

Friction angle of coal matrix

Friction angle of coal bedding planes

Figure 4.—Friction angle for matrix and bedding planes in suite of “numerical rocks.” 
77 

strongest rocks. It also decreases to 0 over 1 millistrain of postfailure strain. 

6. Dilation angle is initially 10° and decreases to 0° 

over 5 millistrain of postfailure strain. 

 This suite of material models provides a convenient 

method for the modeler to go from a geologic log to a 

numerical model in a rational, systematic, and efficient 

manner. Table 1 illustrates this process and shows the level 

of detail needed for geologic logging. On the left is a typical core log with geologic description. Geologic features at 

a scale as small as 50 mm are typically recorded for this 

log. Of particular importance to note are soft clay layers or 

major bedding planes with weak infilling. In the middle of 

0.01

0.10

1.00

10.00

100.00

Soil_0

Soil_1

Soil_2

Soil_3

Soil_4

Soil_5

Rock_A

Rock_B

Rock_C

Rock_D

Rock_E

Rock_F

Rock_G

Rock_H

Rock_I

Rock_J

Coal_1

Coal_2

Coal_3

Coal_4

Cohesion (MPa) .

Cohesion of soil

Cohesion of rock matrix

Cohesion of rock bedding planes

Cohesion of coal matrix

Cohesion of coal bedding planes

Figure 5.—Cohesion of matrix and bedding planes in suite of “numerical rocks.” 

0

5

10

15

20

25

Soil_0

Soil_1

Soil_2

Soil_3

Soil_4

Soil_5

Rock_A

Rock_B

Rock_C

Rock_D

Rock_E

Rock_F

Rock_G

Rock_H

Rock_I

Rock_J

Coal_1

Coal_2

Coal_3

Coal_4

Modulus (GPa) .

Modulus of soil (GPa)

Modulus of rock (GPa)

Modulus of coal (GPa)

Figure 6.—Modulus of matrix in suite of “numerical rocks.” 

Table 1.—Going from core log to numerical model input parameters 

Stratigraphic column 

Rock type Thickness (m) 

UCS from axial point 

load tests (MPa) 

UCS from diametral point 

load tests (MPa) 

Rock matrix 

strength code 

Bedding plane 

strength code 

 Strong sandstone 2.4 90 80 Rock_I Rock_I 

 Siltstone (stackrock) 1.2 35 5 Rock_E Rock_B 

 Black shale 1.9 10 5 Rock_B Rock_B 

 Soft clay 0.05 0.2 0.2 Soil_3 Soil_3 

 Gray shale 1.8 25 10 Rock_D Rock_C 
78 

Table 1 are UCS estimates for each unit from axial and 

diametral point load tests. Finally, on the right of the table 

are property codes for generating input parameters to 

numerical models using the material model suite presented 

herein. The modeling approach described has been verified 

against detailed monitoring of a coal mine entry done by 

Oyler et al. [2004]. 

MULTIPLE-SEAM MINING INTERACTION MODEL 

CONSTRUCTION 

 Using the material input parameters described above, 

models were created for the three interaction types, as 

shown in Figure 7. All models examine mining either 

above or below a gob-solid boundary, which is representative of most interactions, including mining above or below 

pillar remnants or barrier pillars. In undermining, a longwall is mined first in the upper seam followed by roomand-pillar development in the lower seam. In overmining, 

a longwall is mined first in the lower seam followed by 

room-and-pillar development in the upper seam. In 

simultaneous mining, room-and-pillar development is done 

in the upper seam followed by longwall mining in the 

lower seam. These model types enable detailed examination of the failure mechanics of coal mine entries subject 

to multiple-seam mining interactions, with the focus on the 

transition zone either above or below the gob-solid 

boundary. 

 The models consider a slice of the rock mass 160 m 

wide and up to 75 m high. Each model contains 450 elements along the width and up to 515 elements along the 

height. Coal seam thickness is 2 m, entry width is 6 m, and 

entry height is also 2 m. The thickness of each geologic 

layer is about 0.15 m on average and ranges from 0.05 to 

0.25 m. Most of the rock mass layers are assigned properties corresponding to CMRR unit ratings of 45 to 60, 

which is equivalent to a moderate strength rock mass. 

However, as discussed later, the immediate roof of the 

affected seam is considered separately. The stratigraphy 

used in the model is extracted from a detailed core log 

similar to that shown in Table 1. Weak, moderate, and 

strong sections from this log were assembled as needed. 

While the models are artificial, they have a basis on a real 

geologic log. 

 Following recent work on horizontal stress by Dolinar 

[2003], an average horizontal stress of 10 MPa is applied 

to the model via the equivalent horizontal strain. Thus, 

horizontal stress varies according to the relative stiffness 

of each geologic layer. Stiffer layers have higher horizontal stress than softer layers. Vertical stress is applied at 

the top of the model to simulate cover load. 

 The full-extraction area is approximated as a strainhardening material using the DY, or double yield, constitutive model that can simulate irreversible compaction. Thus, 

a gob layer replaces the mined coal seam over a height of 

three times the coal seam thickness. This approximation 

leads to subsidence over the full-extraction area of about 

50% of the seam thickness. 

 As shown in Figure 8, each model type examined the 

effect of three factors on the mechanics of multiple-seam 

mining interaction. Each variable in the matrix has just two 

values. Vertical stress is either 3 MPa or 9 MPa, which 

implies an overburden depth of 120 m (shallow) or 360 m 

(deep). Interburden thickness is either close (7 m) or intermediate (24 m). Finally, the immediate roof quality is 

either weak (CMRR unit rating of 30–45) or strong 

(CMRR unit rating of 60–80). 

 Figure 7.—Multiple-seam mining interaction models for

three interaction types. In undermining, a longwall is mined in

the upper seam first, followed by room-and-pillar mining in

the lower seam. In overmining, a longwall is mined in the

lower seam first and the upper seam subsides fully, followed 

by room-and-pillar mining in the upper seam. In simultaneous

mining, the upper seam is developed first and then a longwall

mines the lower seam, causing subsidence of the existing

workings in the upper seam. Dark shading indicates intact

rock; lighter shading indicates matrix failure or slip along bedding planes. 

Undermining

Overmining

Simultaneous mining

Longwall first

Room-and-pillar second

Longwall first

Longwall second

Room-and-pillar second

Room-and-pillar first

Interburden thickness

Immediate roof rock quality

Vertical stress

Deep 

(9 MPa) 

Shallow 

(3 MPa) 

Weak 

(30 - 45) 

Strong 

(60 - 80) 

Intermediate 

(24 m) 

Ultra-close 

(7 m) 

 Figure 8.—Three variables for numerical modeling matrix of

multiple-seam mining interactions. 
79 

 Preliminary analysis of case history data presented by 

Ellenberger et al. [2003] suggests that multiple-seam mining interactions are possible when the OB/IB thickness 

ratio exceeds 7 for both undermining and overmining 

cases. By implication, this modeling matrix considers three 

OB/IB ratio values, namely, 5 where no interaction is 

expected, about 15–17 where an interaction is possible, 

and 51 where an interaction is likely. 

Undermining OB/IB ratio = 5; vertical stress = 3 MPa Overmining 

 Undermining OB/IB ratio = 17; vertical stress = 6 MPa Overmining 

 Undermining OB/IB ratio = 51; vertical stress = 9 MPa Overmining 

 

 Figure 9.—Vertical stress comparison. Dark shading indicates high stress; light shading indicates 

low stress. Applied horizontal stress is 10 MPa. Hatching indicates longwall. 
80 

Undermining OB/IB ratio = 5; vertical stress = 3 MPa Overmining 

 Undermining OB/IB ratio = 17; vertical stress = 6 MPa Overmining 

 Undermining OB/IB ratio = 51; vertical stress = 9 MPa Overmining 

 

 Figure 10.—Horizontal stress comparison. Dark shading indicates high stress; light shading indicates 

low stress. Applied horizontal stress is 10 MPa. Hatching indicates longwall. 
81 

Undermining OB/IB ratio = 5; vertical stress = 3 MPa Overmining 

 Undermining OB/IB ratio = 17; vertical stress = 6 MPa Overmining 

 Undermining OB/IB ratio = 51; vertical stress = 9 MPa Overmining 

 Figure 11.—Failure state comparison. Dark shading indicates intact rock. Lighter shading indicates matrix failure 

or slip along bedding planes. Applied horizontal stress is 10 MPa. Hatching indicates longwall. 
82 

COMPARISON OF MULTIPLE-SEAM MINING 

INTERACTION MODELS 

 Figures 9–11 compare vertical stress, horizontal stress, 

,and failure state images, respectively, for the undermining 

and overmining type of multiple-seam mining interactions. 

The top image in each figure is for an OB/IB ratio of 5, 

where no multiple-seam mining interaction is expected. 

Vertical stress is low (3 MPa), and immediate roof rock 

quality is weak. These models lie near the lower left corner 

of the matrix shown in Figure 8. In the middle image, the 

OB/IB ratio is 17, so a multiple-seam mining interaction is 

possible. Vertical stress is medium (6 MPa), and immediate roof rock quality is medium. These models lie at the 

center of the matrix. Finally, the lower image is for an 

OB/IB ratio of 51, where an adverse multiple-seam mining 

interaction is highly likely. Vertical stress is high (9 MPa), 

and immediate roof rock quality is weak. These models lie 

at the upper left corner of the matrix shown in Figure 8. 

 The following general observations are noted. 

1. Vertical stress concentrations occur in a narrow 

band above and below the gob-solid boundary. 

This band is inclined about 20° toward the gob 

both above and below the full-extraction seam. 

A vertical stress shadow occurs above and below 

the gob, and it diminishes slowly about 50 m 

from the gob-solid boundary where the gob has 

fully reconsolidated. 

2. Full-extraction mining produces horizontal stress 

changes several tens of meters above and below 

the mined area. The horizontal stress changes 

occur much farther above and below the mined 

area than do the associated vertical stress changes 

laterally away from that mined area. The 

horizontal stress concentrations may in turn 

induce rock failure in select geologic layers well 

above or below the mined area that can further 

amplify horizontal stress concentrations in nearby 

layers. 

3. Bedding plane slip and tensile failure through the 

rock matrix occur in a narrow band directly above 

and below the gob-solid boundary. This band 

is more extensive above the extracted seam; 

however, it also extends a considerable distance 

below it. 

4. For coal mine entries in moderate strength immediate roof rock (CMRR unit rating of 45– 60), the 

extent of rock failure through bedding plane shear 

or tensile failure of the rock matrix is about 

1 times the entry width. 

 With respect to undermining, the following additional 

observations are noted. 

1. A zone of vertical stress relief occurs under the 

full-extraction mining area beginning past the 

gob-solid boundary and extending several tens of 

meters under the gob. This zone is well understood and correlates well with the best practice of 

offsetting gate roads under the gob for optimal 

stability in multiple-seam mining. 

2. There is increased bedding plane slip in entries 

close to directly below a gob-solid boundary. The 

additional failure is slight and should not 

correspond to significant additional support 

requirements. 

3. There is a small increase in the amount of pillar 

failure in the zone below the gob-solid boundary. 

This increase might correspond to additional pillar spalling and nothing more. As before, the 

additional pillar failure is slight and not indicative 

of severe ground control conditions. 

 With respect to overmining, the following additional 

observations are noted. 

1. A significant increase in bedding plane slip and 

tensile failure occurs in the interburden and 

immediate roof rock along with pillar failure in 

the upper coal seam within a narrow band above 

the gob-solid boundary. This observation correlates with known decreases in entry and pillar stability in the transition zone above a gob-solid 

boundary [Rigsby et al. 2003]. 

2. There is no apparent increase in bedding plane 

slip or additional tensile failure in the immediate 

roof above coal mine entries developed in a coal 

seam that has subsided due to prior mining below. 

This numerical observation also correlates well 

with the good stability generally observed in 

entries driven in fully subsided coal seams. The 

failure zone induced by entry development in 

subsided ground is not substantially different 

from that of an entry in completely undisturbed 

ground. 

 With respect to simultaneous mining, the following 

additional observations are noted without showing the 

associated models. 

1. The horizontal and vertical stress distribution is 

virtually identical to that shown for simple 

overmining. 

2. The failure mode situation is completely different. 

When the longwall is created in the lower seam, 

calculations show a wave of tensile failure that 

propagates upward through the rock mass and 

completely envelopes the developed entries 

within this failure zone. Thus, these entries are 

likely to experience deteriorating ground control 
83 

conditions. This situation is completely different 

from the prior situation of entry development in a 

previously subsided coal seam. 

 Tables 2 and 3 compare notes about the undermining 

and overmining models within the modeling matrix 

described by Figure 8. These tables compare conditions 

around an entry not subject to any multiple-seam mining 

interaction to an entry subject to full interaction in the area 

directly above or below a gob-solid boundary. The comparisons examine relative changes on vertical stress, horizontal stress, and failure mode as the OB/IB ratio increases 

from 5 to 51. Figures 9–11 help illustrate this semiquantitative comparison of changes in relative stress and 

failure zone size. As expected, the tables show interesting 

trends as the OB/IB ratio increases from 5, where no interaction is expected, to over 50, where a serious multipleseam mining interaction is expected. 

Vertical Stress Comparison 

 In Tables 2 and 3, the vertical stress comparison 

documents stress changes in pillars above or below a gobsolid boundary compared to a pillar far from the 

interaction area. As indicated in Table 2, when 

undermining with a low OB/IB ratio of 5 (Figure 9, top), 

there is little change in vertical stress within pillars near 

the gob-solid boundary compared to pillars far away, no 

matter whether the immediate roof rock is weak or strong. 

As the OB/IB ratio increases to 17 or 51 (Figure 9, middle 

and bottom), the relative vertical stress concentration 

increases significantly. Again, the strength of the 

immediate roof rock makes little difference in this 

increase. 

Table 2.—Comparison of undermining models 

OB/IB 

ratio 

Weak immediate 

roof rock 

Strong immediate 

roof rock 

Vertical stress comparison 

 5 Little change Little change 

 15–17 Increases about 25% to 

 50% 

 Increases about 50% 

 51 Increase more than 50% Increase more than 50% 

Horizontal stress comparison 

 5 Increases less than 10% Increases about 25% 

 15–17 Decreases due to failure Increases about 50% 

 51 Decreases due to failure Increases more than 50% 

Failure size comparison 

(remote entry and under gob-solid boundary) 

 5 Slight increase in size 0.1 × entry width and 

 no increase in size 

 15–17 0.5 to 1 × entry width to 1 

 to 1.5 × entry width 

 0.1 to 1 × entry width to 

 0.25 to 1 × entry width 

 51 1 × entry width to more 

 than 2 × entry width 

 1 × entry width to 

 1.5 × entry width 

Table 3.—Comparison of overmining models 

OB/IB 

ratio 

Weak immediate 

roof rock 

Strong immediate 

roof rock 

Vertical stress comparison 

 5 Increases about 20% Increases about 20% 

 15–17 Increases about 25% to 

 50% 

 Increases about 25% to 50% 

 51 Increase more than 50% Increase more than 50% 

Horizontal stress comparison 

 5 Increases about 10% Increases about 50% 

 15–17 Small increases and 

 decreases due to failure 

 Increases about 50% 

 51 Decreases due to failure Increases more than 50% 

 and decreases due to failure 

Failure size comparison 

(remote entry and under gob-solid boundary) 

 5 0.3 × entry width to 

 1 × entry width 

 0.1 × entry width and 

 no increase in size 

 15–17 0.5 to 1 × entry width to 

 1 to 1.5 × entry width 

 0.1 to 1 × entry width to 

 0.1 to 1.5 × entry width 

 51 1.5 × entry width to 

 more than 2 × entry width 

 1 × entry width to 

 1.5 × entry width 

 When overmining, it seems there is more upward 

transmission of vertical stress concentration. As indicated 

in Table 3 and shown in Figure 9, at low OB/IB ratio, 

a significant relative increase in vertical stress does occur. 

At higher OB/IB ratios of 17 and 51, the relative vertical 

stress concentration increases, but it is not substantially 

different from that seen with undermining. Subsidence and 

the extent of the broken gob above the seam horizon may 

account for the difference at low OB/IB ratio. As with 

undermining, the strength of the immediate roof rock 

makes little difference with regard to the magnitude of the 

relative vertical stress changes in the pillars. 

 The overburden and interburden rock in all models is 

medium strength, with CMRR ranging from 45 to 60. 

Changing the physical nature of the interburden rock will 

change the vertical stress distribution; however, it will not 

change the relative vertical stress changes as noted in this 

comparison. 

Horizontal Stress Comparison 

 The horizontal stress comparison documents stress 

changes in the immediate roof rock of an entry above or 

below a gob-solid boundary compared to an entry far from 

the expected interaction area. As shown in Tables 2 and 3 

and Figure 10, at a low OB/IB ratio of 5, where no 

interaction is expected, horizontal stress does increase 

slightly above the background level and that increase 

depends on the strength of the immediate roof rock. 

A weak immediate roof rock sees a small increase, while 

strong immediate roof rock sees a much larger increase 

in horizontal stress. The relative increase in horizontal 

stress is more pronounced in overmining compared to 
84 

undermining for reasons of subsidence and gob formation 

noted earlier. 

 As the OB/IB ratio increases, the change in horizontal 

stress from background depends on the strength of the 

immediate roof rock. With weak roof, horizontal stress 

decreases in the interaction area when the immediate roof 

rock fails and stress is distributed elsewhere. With stronger 

roof, horizontal stress in the interaction area can increase 

dramatically over background (50% or more). However, at 

sufficiently high OB/IB ratio, even strong immediate roof 

rock can be made to fail in the interaction zone, with a 

resulting decrease in horizontal stress. Table 3 for overmining shows this possibility. 

 The role of horizontal stress is crucial for further 

understanding of multiple-seam mining interaction. As 

seen in Figure 10, a full-extraction area induces horizontal 

stress changes many tens of meters above and below the 

mined seam. The magnitude of these changes depends on 

several factors, namely, OB/IB ratio, site-specific geology, 

the ratio of extraction area width to interburden thickness, 

and the ratio of horizontal stress to immediate roof rock 

strength. 

 The OB/IB ratio affects the geometry and the vertical 

stress level of the particular situation. Closer proximity of 

the affected seam to undermining and overmining has the 

expected effect on horizontal stress magnitude. With 

respect to geology, the major variable is the percentage of 

strong rock in the interburden and where that strong rock is 

located relative to the affected roof. A suitably placed 

strong bed can shield the immediate roof rock of a seam 

from adverse multiple-seam mining interaction. The ratio 

of extraction area to interburden thickness is another 

geometry factor that controls how far above or below a 

full-extraction area the horizontal stress might change. 

There are limits on this ratio that depend on whether the 

full-extraction area exceeds the “critical width” at which 

maximum subsidence is achieved and vertical stress in the 

middle of the full-extraction area returns to in situ value. 

Finally, the ratio of applied horizontal stress (in situ plus 

induced) to the strength of the immediate roof rock controls the degree of multiple-seam mining interaction. 

A higher ratio due to either high horizontal stress or low 

immediate roof rock strength increases the chance of an 

adverse interaction. Horizontal stress has been found to be 

a major factor in many ground control problems, especially in the Eastern United States [Mark and Mucho 

[1994]. 

Failure Size Comparison 

 The failure size comparison documents the extent of 

rock matrix or bedding plane failure in either shear or 

tension within the immediate roof of the affected seam. 

The entry itself induces a failure zone in the immediate 

roof whose extent depends on overburden depth and 

immediate roof rock quality. This comparison notes how 

much additional failure occurs due to nearby multipleseam mining. Failure extent is gauged relative to the entry 

width. 

 When undermining at low OB/IB ratio (Table 2), the 

overlying gob-solid boundary is far away, and it induces 

little additional failure about an entry in the potential 

interaction area (Figure 9). For high OB/IB ratio, the overlying gob-solid boundary is close, and the size of the failure zone grows by more than a factor of 2 in weak rock, 

as shown in Figure 11. Strong rocks show a similar trend 

as indicated in Table 2. At low OB/IB ratio, the interaction 

is negligible, while at high OB/IB ratio, the added interaction is severe even with strong immediate roof rock. 

 At intermediate values of OB/IB ratio, induced 

changes in the failure zone extent due to multiple-seam 

interaction can vary greatly. For one case in weak roof, 

it changes from 1⁄2 to 1 times the entry width, while in 

another case, it changes from 1 to 11⁄2 times the entry 

width. With stronger rocks, the variability is even more 

pronounced. In one case, failure zone size changes from 

one-tenth to one-fourth times the entry width, and in 

another it remained the same size at 1 times entry width. 

 Failure zone size in the overmining models (Table 3) 

showed a similar trend. At low OB/IB ratio in weak rock, 

failure size is somewhat larger initially and grows more 

than in the undermining models, while in stronger rock 

there is no difference. At high OB/IB ratio, failure size 

grows significantly in both weak and strong rock, which 

is indicative of a substantial multiple-seam mining 

interaction. At intermediate OB/IB ratio, failure zone 

extent and its changes vary greatly as in the undermining 

models. 

 The observed changes in failure zone size reflect 

similar trends as seen with horizontal stress. For an OB/IB 

ratio less than 5, the chance of a multiple-seam mining 

interaction is very low, even under a weak immediate roof 

rock. An adverse interaction is expected for a high OB/IB 

ratio greater than 50, even under strong immediate roof 

rock. For an intermediate OB/IB ratio of around 17, the 

chance of an adverse interaction depends on the vagaries 

of the interburden rock, in particular site-specific geology 

and the ratio of horizontal stress to rock strength, and geometric factors such as the ratio of full-extraction width to 

interburden thickness. 

MECHANICAL FACTORS IN MULTIPLE-SEAM 

MINING INTERACTIONS 

 The simple modeling matrix reproduces successfully 

many practical observations of multiple-seam mining 

interactions, lending credibility to the numerical model and 

the NIOSH input parameters. Close inspection of the 

models considered here suggests four underlying factors 

controlling the failure mechanics of multiple-seam mining 

interactions: 
85 

1. Vertical stress concentration; 

2. Horizontal stress concentration; 

3. Stress redirection; and 

4. Bedding plane slip bands. 

 Vertical stress concentrations (Figure 12) occur in the 

vicinity of gob-solid boundaries, pillar remnants, and similar structures as vertical stress is diverted around fullextraction areas. The lateral extent of these increases is 

indicated in Figure 12 along with stress relief areas. The 

degree of vertical stress concentration decreases quickly 

with lateral distance from this boundary. The extent of 

vertical stress relief above and below the full-extraction 

area depends on the width of that area. There is also an 

associated vertical stress gradient near a gob-solid 

boundary. 

 Horizontal stress concentrations (Figure 13) also 

develop around full-extraction areas; however, their behavior is much more complex. Horizontal stress concentration 

depends on both distance above and below the fullextraction area and the relative stiffness of the geologic 

layers. Furthermore, horizontal stress concentrations can 

vertical stress 

concentration 

vertical stress 

relief area 

Overmining

vertical stress 

concentration 

vertical stress 

relief area 

Undermining

 Figure 12.—Vertical stress concentration above and 

below gob-solid boundaries. Dark shading indicates high 

stress; light shading indicates low stress. Hatching indicates longwall. 

 Figure 13.—Horizontal stress concentration above and 

below full-extraction areas. Dark shading indicates high 

stress; light shading indicates low stress. Hatching indicates longwall. 

horizontal 

stress 

concentrations 

horizontal 

stress 

relief 

Overmining

Undermining

horizontal 

stress 

relief 

horizontal 

stress 

concentrations 
86 

be expected much farther above or below a full-extraction 

area than vertical stress concentrations can be expected left 

or right of that area. Horizontal stress redistribution is seen 

much farther away, and it may induce failure in certain 

weaker layers, leading to even more horizontal stress 

redistribution. It seems that the effect of horizontal stress 

on multiple-seam mining interactions has not been 

explored in any prior studies. The extent of horizontal 

stress concentration and associated failure of select layers 

may explain certain cases of successful and unsuccessful 

multiple-seam mining in otherwise similar conditions. 

 The combination of vertical and horizontal stress 

increases in the vicinity of a full-extraction area and, 

in particular, stress gradients will reorient the principal 

stresses, as illustrated in Figure 14. This seemingly small 

stress reorientation has a profound effect on bedded rock. 

In the absence of mining, principal stresses are usually 

oriented parallel and perpendicular to geologic strata 

(Figure 14, top), which is a more favorable orientation for 

strength. Full-extraction mining reorients principal stresses 

to the weaker orientation shown in the bottom of Figure 14. Coal mine entries developed in nearby seams in 

this rotated stress field are much more likely to experience 

unfavorable ground control conditions. It is also noted 

without illustration that reorientation of the principal stress 

occurs in a fairly narrow vertical band adjacent to the gobsolid interface. 

 The rotated stress field also leads to bedding plane slip 

in narrow, subvertical zones above and below gob-solid 

boundaries, as seen in the failure state plots in Figures 9– 

11. These zones of bedding shear are also more likely 

areas for unfavorable ground control conditions. 

TOWARD DESIGN GUIDELINES FOR 

MULTIPLE-SEAM MINING 

 This research seeks to provide design guidelines that 

enable mine planning engineers to correctly assess the 

safety risk of an adverse multiple-seam mining interaction 

based on mine geometry factors, mine layout factors, and 

site-specific geologic conditions. Preliminary analysis of 

case studies by Ellenberger et al. [2003] suggested that for 

both undermining and overmining, when the OB/IB ratio 

was less than 7, there was little risk of adverse interaction. 

For an OB/IB ratio above 16, there was a possibility of 

extreme interaction; however, it became evident that other 

factors in addition to OB/IB also became important. 

 Numerical studies conducted during this research 

examined the effect of OB/IB ratio and the immediate roof 

rock quality of the affected seam on the degree of 

multiple-seam mining interaction. The numerical models 

utilized contain great geologic detail and the proper constitutive behavior and are able to capture the essential failure 

modes of the rock mass, in particular, shear or tensile failure through the rock matrix or along bedding planes. 

 The numerical models confirm aspects of the initial 

multiple-seam mining interaction guidelines above. When 

the OB/IB ratio is less than 5, the models clearly show 

little, if any, interaction between mining in nearby seams. 

When the OB/IB ratio exceeds 50, the models clearly show 

an extreme interaction, even with strong roof conditions in 

the affected seam. For the intermediate OB/IB ratios considered (15–17), the models show that an adverse interaction is possible, and they provide some insight into the 

controlling factors. 

 Numerical models show how vertical stresses divert 

around a full-extraction area in a seam above or below an 

active mining seam. The lateral extent of vertical stress 

increase is relatively narrow compared to the width of the 

full-extraction area. In addition, a zone of vertical stress 

relief occurs above and below the full-extraction area. 

It will extend to a distance up to the “critical width” of the 

extraction area. Horizontal stresses also divert around the 

full-extraction area; however, the distance that such 

stresses increase above or below the seam is much larger 

than the lateral extent of vertical stress increase. This 

distance may be approximately equal to the minimum 

width of the full-extraction area up to the “critical width.” 

Thus, the size of this zone of vertical stress relief in 

conjunction with horizontal stress increase defines the 

BEDDING

FH

FV

 Figure 14.—Reorientation of principal stresses leads to

failure due to multiple-seam mining interaction. In single-seam 

mining, far-field principal stresses are generally aligned parallel

to the bedding planes in which a test specimen is relatively

strong. If the far-field principal stresses are rotated due to

nearby multiple-seam mining, the bedding planes are oriented

in an unfavorable direction in which a test specimen is relatively

weak. 
87 

extent to which adverse multiple-seam mining interaction 

could occur. 

 Numerical modeling suggests four factors that control 

multiple-seam mining interactions and should be considered explicitly in design guidelines: 

1. OB/IB thickness ratio; 

2. Gob width-to-interburden thickness ratio; 

3. Site-specific geology; and 

4. Horizontal stress to rock strength ratio. 

 As mentioned earlier, the OB/IB ratio affects the 

geometry and the vertical stress level of the particular 

situation. Greater depth or closer proximity of the affected 

seam to undermining and overmining are both known to 

increase the chance of an interaction. The minimum gob 

width of the extraction area relative to the interburden 

thickness is another geometric factor that controls how far 

above or below a full-extraction area the horizontal stress 

might change. There are limits on this ratio that depend on 

whether the full-extraction area exceeds the “critical 

width” at which maximum subsidence is achieved and the 

vertical stress in the middle of the full-extraction area 

returns to its original in situ value. With respect to geology, the major variable is the percentage of strong rock in 

the interburden and where that strong rock is located relative to the affected roof. A suitably placed strong bed can 

shield the immediate roof rock of a seam from adverse 

multiple-seam mining interaction. Finally, the ratio of 

applied horizontal stress (in situ plus induced) to the 

strength of the immediate roof rock strength controls the 

degree of multiple-seam mining interaction. A higher ratio 

due to either high horizontal stress or low immediate roof 

rock strength increases the chance of an adverse 

interaction. Although horizontal stress has been found to 

be a major factor in many ground control problems, 

especially in the Eastern United States [Mark and Mucho 

1994], the role of horizontal stress in multiple-seam 

mining interactions has received little attention in prior 

investigations. 

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