ABSTRACT

Standing crib supports have been applied in underground

mining programs to resist large roof movements and sustain highloads. The strength and deformation capability of these systems 

has been documented under both laboratory and field conditions. 

The parameter that has not been examined and is not well

understood is the effect that a crib or other types of standing 

support has on the primary and secondary bolting systems. A crib 

support system with low system stiffness may allow large amounts 

of closure and deformation and cause the bolting systems to yield 

or even fail. Conversely, cribs or standing support that are too stiff 

may experience brittle or buckling failures, negating the advantage 

of the intrinsic supports previously installed. Utilizing a

combination of field measurements and 3-dimensional finite

element modeling techniques, the relationship between system

stiffness and the subsequent performance of the installed bolting 

system is evaluated. Additionally, a simple method for calculating 

the combined system stiffness for standing supports, when using 

materials with different strengths such as steel, concrete, wood, etc., 

is presented.

INTRODUCTION

Tailgate entries are often the most hazardous areas in longwall 

mining operations. Establishing and maintaining a stable tailgate 

entry is a complex challenge for the mine engineer. The entry 

provides access and a secondary travel-way for the miners and 

completes the ventilation loop that removes gases and dust from the 

working face. Higher loading pressures usually exist around the 

tailgate entry which makes them difficult to support and maintain. 

Moderate entry closure or partial blockage of the tailgate entry can 

be tolerated but the criteria between moderate and extreme can 

cause production delays and remedial work under hazardous

conditions. The majority of tailgates are controlled by standing or 

crib supports and are routinely supplemented with vertical cables, 

cable slings, cable trusses, rigid trusses, and longer high capacity 

bolts.

Crib support design requirements depend on the nature of the 

strata loading behavior. Researchers have shown that a zone 

approximately one-tenth of the overburden depth is the area that 

continually moves forward during longwall panel extraction and 

loads the intrinsic and standing support systems (1-3). For example, 

if the overburden depth is 500 ft, the zone approximately 50 ft in 

front of the longwall face would be subjected to higher stresses and 

subsequent loading. If the tailgate entry is examined as if it were a 

test frame for determining the physical property characteristics of 

rock or soil samples, two testing modes are routinely selected; 

displacement (rate) or load control. In the displacement control 

method, the rock testing machine is programmed to load the sample 

at an established closure rate. It will continue to “close” at the rate 

irrespective of the load applied to the sample. Conversely, in the 

load control setting the machine is programmed to apply an

established force to the sample, irrelevant of the strain or closure on 

the sample. If this concept is applied to increase the understanding 

of tailgate entry performance, the standing crib supports are the 

samples being tested. Assuming that tailgate entry loading is 

completely displacement or rate-controlled, the standing supports 

can’t stop the continuous closure and are compressed and

eventually failed by the constant roof to floor deformation. If the 

primary and secondary bolting system can prevent the decoupling 

of the strata layers or the creation of isolated rock sections, then 

there essentially would be no benefit to having a standing support 

system. It can be argued that a standing roof support system would, 

if anything, do more harm than good since it may puncture into the 

roof and floor and cause further instability of the rock mass. On 

the other hand, if the loading is completely load-controlled, the 

most critical design parameter would be the stiffness of the support 

because a passive standing support requires convergence of the 

mine roof and floor to compress the support and generate its load 

carrying capacity. If the support is too soft, too much deformation 

will occur which can initiate the failure of the mine roof as shown 

in figure 1. If too much defo rmation or closure occurs, the primary 

and secondary bolt systems may experience tensile failures directly 

related to combinations of excessive elongation, roof bending and 

bed slippage, which can ultimately result in shear type failures.

When both load-controlled and displacement-controlled 

behavior occurs, as is often the case, there are conflicting design 

requirements and compromises must be made to achieve the

optimum support design. Depending on the amount and timing of 

the displacement-controlled loading, the same support system may 

work fine in one application and fail in another. Displacementcontrolled loading or uncontrolled convergence can make soft 

supports perform well in areas where stiff supports fail miserably 

and vice-versa. A prime exa mple of this is the 3C1 support, which 

 

1Mention of any company name or product does not constitute endorsement 

by the National Institute for Occupational Safety and Health.

THE EFFECT OF STANDING SUPPORT STIFFNESS ON

PRIMARY AND SECONDARY BOLTING SYSTEMS

Stephen C. Tadolini, Supervisory Mining Engineer

Thomas M. Barczak, Geophysicist

National Institute for Occupational Safety and Health

Pittsburgh Research Laboratory

Pittsburgh, Pennsylvania USA

Yunqing Zhang, Research Associate

West Virginia University

Morgantown, West Virginia USA
is the predecessor to the modern Can support. As shown on 

figure 2, this particular support requires over a foot of convergence 

to provide a useful capacity for roof support. Yet the support has 

been used successfully in a longwall tailgate in the Western U.S. 

The reason it performed adequately was that the mine employed a 

yielding pillar design where both the pillar yield and large amounts 

of floor heave occurred that in effect mobilized sufficient load 

carrying capacity in the support by pre-compressing it prior to the 

passing of the longwall face. This concept can be illustrated in 

figure 3. Shown are two hypothetical ground reaction curves, one 

with little displacement-controlled activity and the second one with 

large amounts of displacement-controlled loading. As seen in the 

figure, the 3C support is much too soft to generate sufficient 

loading to achieve roof control in the first ground reaction curve, 

but does provide the necessary capacity when the uncontro lled 

convergence is large (4). In comparison, the Can support should 

perform well in both these environments, as it develops its load 

carrying capacity relatively quickly and is able to sustain this load 

carrying capacity through a large displacement range. On the 

opposite side of the spectrum, as shown in figure 4, concrete donut 

cribs and Magnum concrete supports, which are high capacity but 

very stiff supports, have failed prematurely in many western mines 

operating in similar conditions to the 3C support, yet have 

performed satisfactorily in several eastern mines. Again, different 

degrees of displacement-controlled loading were most likely the 

reason why successes turn to failures in the application of the same 

support technology.

Another related issue is the effect that the crib loading can have 

on the intrinsic bolting system installed as primary and secondary 

support. For example, in some deep cover western mines, it has 

been reported that even soft, 4-point cribs eventually are squeezed 

enough by excessive convergence, appearing to be uncontrollable, 

in these mines to cause damage to the roof beam. Again, the 

success or failure of a support will be determined by the degree of 

displacement-controlled loading behavior of the ground. How the 

stiffness of the crib supports effects the loading of the internal 

bolting support systems and how the stiffness can be examined and 

modified will be examined in detail.

Figure 1. Large roof deformations resulting from a soft crib 

system.

Figure 2. A longwall tailgate supported with 3C supports 

subjected to considerable convergence before useful support 

capacity is realized.

Figure 3. The Can support performs well in both loadcontrolled and displacement-controlled environment while 

the 3C support fails to provide roof control in the absence of 

the uncontrolled convergence.

Figure 4. Failure of stiff concrete supports in a 

displacement-controlled environment.
ä1 = 

1

1

AE

PL

THE EFFECT OF CRIB STRENGTH DESIGN ON BOLT 

AND ENTRY BEHAVIOR

Concrete cribs in the form of donuts, Magnums, and pumpables 

have been used to control the deformation and loading behaviors in 

longwall tailgates, difficult intersections, and pre -driven longwall 

recovery rooms. These base systems are routinely “capped” or 

“finished off” by applying a layer of wooden crib materials, layers 

of plywood or fibre -board, or a pumped bag or pillow systems. 

The bag or pillow systems being employed are filled with a lightweight material with low compressive strengths and stiffness

characteristics. Understanding the bag or pillow and the yield 

behavior is a critical design parameter for the accurate design of 

standing concrete type support systems. To examine the impact of 

various crib strengths, 3-dimensional finite element models were 

developed and analyzed. The purpose of the models, utilizing 3 

different crib strengths, were three-fold; to examine the closure at 

the center of the entry, calculate the expected loads exerted on the 

cribs, and finally, to evaluate the effects of crib loading on the

primary and secondary bolting systems. 

To help illustrate the concept used for the model, a crosssection of the longwall tailgate entry is shown in figure 5. Again, 

applying the rule of “one-tenth the overburden depth”, the cribs are 

not subjected to significant load before the longwall face moves 

into this calculated zone. The adjacent coal pillars, primary, and 

secondary bolting systems control most of the initial movements 

and subsequent loads. As the longwall face advances down the 

entry, the pillars, roof, and the floor are deformed under the 

forward abutment loads. 

The load applied to the crib is dependent on the level of

abutment pressure, the stiffness of the crib and pillars, the behavior 

of the roof and floor strata, and the bolt support system during the 

initial development cycle. For example, a stiff crib system will 

attract a large portion of the abutment load, possibly causing the 

cribs to fail before the longwall can safely mine past the area. 

Conversely, a soft crib system will attract less abutment load but 

may provide insufficient support to the immediate and main roofs, 

potentially causing the primary and secondary bolting systems to 

fail and the roof to fall before the longwall can safely move past the 

crib supports.

THE CONCEPT OF THE COMBINED MODULUS AND 

STIFFNESS FOR CRIB SYSTEMS

When the concrete crib base is used in conjunction with a 

pillow or bag, which is softer than the crib, to support the recovery 

room, the combined modulus or stiffness of the bag and crib can be 

determined. It is important to distinguish the difference between 

combined stiffness, which is a structural property, and combined 

modulus, which is a material property that is independent of the 

size or shape of the support structure.

Figure 6 illustrates a diagram that represents the stiffness of the 

base, KCB, and the bag or pillow, KB. The determined value is the 

support system stiffness, KT, of the crib and bag, estimated as 

follows:

 (1)

Where: KB = stiffness of the bag, lb/inch

KCB = stiffness of the crib base, lb/inch, and 

KT = combined crib and bag or system stiffness, 

lb/inch.

Because stiffness is a structural property that is directly related 

to the geometry of the structure being considered, it must be 

converted to a form useable in numerical techniques. Since 

stiffness is related to the modulus of the material, a physical

property required for numerical analysis, an equation is developed 

to combine the modulus of different materials. This is termed 

combined modulus.

The Combined Modulus Equation

The combined modulus is a variable that is usually required in 

most numerical models. The value is used primarily to input a 

single physical property parameter for multiple layers of different 

materials. An equation, which makes the combination of modulus 

values for the crib material and the bag material, is presented. 

Assume a load P is applied to the bag and crib, as shown in figure 6. 

The vertical compression of the bag created by the load P is: 

 (2)

while the vertical compression of the crib created by the load P is:

 (3)

Figure 5. The abutment load distribution on a longwall 

tailgate entry.

B CB

B CB

T K K

K K

K

+

×

=

Figure 6. Concept of combined stiffness.

ä2 = 

2

2

AE

PL
where: L1 = height of the bag, inches,

L2 = height of the crib, inches,

E1 = Young’s modulus of the bag, psi,

E2 = Young’s modulus of the crib, psi

A = cross-section area of the bag and crib, in 2, and

 ä1, ä2 = displacement for bag and crib, respectively, inches

The overall compression of the bag and crib can also be 

calculated by the combined Young’s modulus as:

 (4)

where: Ec = combined Young’s modulus, psi, and

äT = total compression of both the bag and crib base, 

inches.

The compression of the bag and crib calculated by the

combined Young’s modulus in equation (4) is equal to the

summation of the compression calculated separately in both

equation (2) and (3). 

Therefore: 

 (5)

The combined Young’s modulus can be derived from equation (5) 

as:

 (6)

The combined Young’s modulus of the bag and crib is a value 

between the modulus of the crib and bag. 

The relationship between the modulus and stiffness can be 

derived because stiffness is a function of the area and the length. 

Using figure 6:

 

 and (7) 

KT can be derived which is the same as equation (1): 

 

 (8)

The Application of the Combined Stiffness Equation

Even if the crib base is very stiff, a soft bag will always result 

in a soft combined system stiffness. Therefore, the stiffness of both 

the bag and crib base can be adjusted by design to provide an 

appropriate support system with a combined stiffness capable of 

controlling the closure of the entry and maintaining a stable roof 

condition. 

A sensitivity analysis conducted to examine the impact of the 

height or thickness of the bag or pillow placed on top and also, 

varying the material modulus for the grout used in the bag or the 

combined stiffness of this support system is presented. As

observed during underground investigations, the irregularities in 

the mine floor and roof can create different final heights for the 

bags or pillows. To complete the analysis, the material modulus for 

the bag grout defined from physical property testing was utilized as 

a base value. This modulus value was varied 50 percent higher and 

lower to provide a range for calculated values and subsequent 

analysis of the results. The bag or pillow heights were also varied 

from 2 to 15 inches thick, heights routinely encountered during 

field investigations. Table 1 summarizes the parameters evaluated.

Table 1. Values for parametric examination of combined crib 

modulus and stiffness.

Parameter Description Value

Height of the concrete base Fixed height = 96 inches

Concrete base modulus E = 3,000,000 psi

Thickness of bag or pillow Varied from 2 to 15 inches

Bag or pillow modulus E = 5,000, 10,000, and 15,000 psi

Crib and pillow area A = 1,810 in2

If the combined stiffness, KT , is examined using the parametric 

values and a crib and bag area of 1,810 in2, reductions in stiffness 

also occurs as the bag thickness increases. The actual reductions 

are shown in figure 7. If the material in the bag has a modulus of 

15,000 psi and the thickness is increased from 2 to 15 inches the 

combined system stiffness is reduced from approximately

10,950,000 to 1,750,000 lb/inch. Even if the crib base has an 

extremely high stiffness value, a bag with a low stiffness can 

impact the stiffness of the entire system. Therefore, the material 

comprising both the bag and crib can be designed to provide an 

appropriate system with the combined strength and relative

stiffness capable of sustaining the desired load within a designed 

amount of convergence. The effects of reduced system stiffness, 

with respect to entry behavior, and the intrinsic primary and 

secondary support loading should be considered. 

Effect of Crib Stiffness on Roof Support System Behavior

To complete the numerical analysis on the effects of crib 

stiffness, with respect to the roof-to-floor convergence and support 

äT =

AEc

P(L L ) 1 + 2

AEc

P L L

AE

PL

AE

PL ( ) 1 2

2

2

1

1 +

+ =

1 2 2 1

1 2 1 2

( )

L E L E

E E L L

Ec

+

+

=

1

1

L

AE

KB =

2

2

L

AE

KCB =

B CB

c B CB

T

K K

K K

L

AE

L

AE

L

AE

L

AE

L E L E

AE E

L L

AE K

+

×

=

+

×

=

+

=

+

=

2

2

1

1

2

2

1

1

1 2 2 1

1 2

1 2

0.0E+00

2.0E+06

4.0E+06

6.0E+06

8.0E+06

1.0E+07

1.2E+07

0 2 4 6 8 10 12 14 16

Bag height, inches

Combined stiffness, lb/inch

15,000 psi

10,000 psi

5,000 psi

Figure 7. Combined support stiffness with various bag 

thicknesses and material modulus of elasticity.
system behavior, 3-dimensional finite element modeling was

completed using ABAQUS (5). The sub-model that was extracted 

from a calibrated global and gob model was utilized to provide the 

detail required for both the primary and secondary cable bolts load 

comparisons (6-7). The immediate roof, in the calibrated submodel, consists of a 6-ft thick laminated shale overlain by gray 

shale of various thicknesses. A weak layer of claystone in the 

immediate floor is 1.5 ft thick with a competent limestone material 

underneath which is about 5 ft thick. In general, the roof in the 

calibrated model is considered weak with a calculated Coal Mine 

Roof Rating (CMRR) values that range from 40 to 45 (8).

A modeling matrix was designed to examine the tailgate entry 

as the longwall face approached the area. The entry was supported 

with the primary support system, resin -assisted, 7/8-inch diameter 

tensioned bolts and 15-ft long secondary cable bolts. A single row 

of concrete cribs on 10-ft row spacings was selected. Figure 8 

shows the finite element mesh for a cross section of the longwall 

tailgate entry with the primary bolts and cribs installed. A plan 

view of the recovery room and the placement of the concrete cribs 

is shown in figure 9. Figure 9 also shows a cross-section of the 

bolting system with the specific installation lengths, diameters, and 

installed loads. 

The crib and support loading was determined when the state of 

the longwall panel was computed at 100 and 10% of the intact coal 

yield state. This occurred when the longwall face was

approximately 35 ft outby and 10 ft outby, respectively. This 

longwall panel state occurs after a large amount of yielding has 

taken place and the coal can only maintain 10% of its original 

strength. The numerical model and field results predicts that this 

would occur when the longwall face was approximately 10 ft away 

from the panel corner. It is difficult to determine or estimate how 

much residual strength a pillar will have after completely yielding. 

In this modeling application, the panel material is being completely 

removed by the longwall shearer so no residual strength can be 

realized from the extracted coal. As the longwall begins to remove 

the material in front of the longwall face, it will eventually collapse 

and lose load carrying capacity. A value of 10% of the original 

coal strength was used and deemed appropriate for this high 

loading situation. The pillar adjacent to the longwall panel was 

developed with a final crosscut width of 50 ft that resulted in a 

solid coal block of 32 ft. The numerical model predicted that the 

pillar would experience a yield zone of about 4 ft as the longwall 

face was mined to within appro ximately 16 ft. 

MODELING RESULTS FOR DIFFERENT CRIB STIFFNESS 

VALUES

The effects on individual bolting systems, roof-to-floor closure, 

and crib loading were analyzed using the 3-dimensional sub-model 

extracted from the global model. To complete the analysis, 

materials were selected and used for the crib base and the bag or 

pillow top and compared with the results from the calibrated case. 

The stiffness values used for the soft, normal, and stiff cribs are 

shown in table 2. The numerical results obtained for the roof-tofloor closure, determined at the center of the entry, and the 

maximum crib loads for both distances are shown in table 3 and 

graphically in figure 10. As the cribs sustain or carry more of the 

applied loads, the roof-to-floor closure is reduced. This does not 

imply that if enough stiff cribs are used closure could be

completely eliminated. A certain amount or degree of closure will 

always be unavoidable or uncontrollable, even with the highest 

capacity cribs on dense patterns.

Table 2. Crib stiffness values used in numerical model.

Crib stiffness (lb/inch)

Soft crib 5.66 x 104

Normal crib 5.66 x 105

Stiff crib 5.66 x 106

Table 3. Roof-to-floor closure and crib loading results.

Longwall face 

distance, >35 ft outby

Longwall face 

Crib distance, 10 ft outby

Stiffness Closure, 

inch

Crib load, 

lbs

Closure, 

inch

Crib load, 

lbs

No cribs 1.65 2.54

Soft crib 

system

1.61 80,698 2.45 131,089

Normal 

crib system 1.50 626,417 2.32 1,124,371

Stiff crib 

system

1.44 1,130,400 1.82 1,632,800

Figure 8. Cross section of the finite element mesh of the 

longwall tailgate sub-model section.

Figure 9. Plan view of the sub-model and cross-section of 

the installed bolting systems.
As the loads carried by the yielding longwall panel block and 

adjacent chain pillar are reduced, the entry roof is subjected to 

additional loading and subsequent closure; requiring the cribs to 

carry more of the abutment loads. When the face is 35 ft outby the 

selected crib location, a normal crib allows 1.50-inches of roof-tofloor closure and is subjected to approximately 626,000 lbs of load. 

If no cribs are present, this closure is increased to 1.65 inches. A 

soft crib allows 1.61-inches of entry closure and is subjected to 

smaller crib loads. During the critical loading phase, when the face 

is 10 ft away, the entry would be subjected to 2.54-inches of 

closure if no cribs are installed. A soft crib would prevent only 

0.09-inches of closure and support 131,000 lbs of load. A stiff crib 

would allow only 1.82-inches of roof-to-floor closure but be 

subjected to 1,632,800 lbs of load. This high degree of loading is 

approaching the ultimate capacity of these types of support systems 

which could result in brittle or ultimate crib failure. The results 

indicate that as the crib system becomes softer, the roof-to-floor 

closure increases and the load carried by the cribs reduced. This 

relationship can be intuitive once mine loading and stresses are 

well understood. What is not easily understood is the effect that a 

softer or stronger crib support system can have on the intrinsic 

primary and secondary bolting support system or the effect that a 

roof bolting system can have on minimizing crib loading and room 

closure.

THE LOADING BEHAVIOR OF BOLTING SYSTEMS 

WHEN USED IN CONJUNCTION WITH CRIB SYSTEMS 

OF VARYING STIFFNESSES

To examine the behavior of the primary and secondary roof 

support systems, the model was used in conjunction with the 

various crib systems previously described. The specific loads 

applied to the bolts were analyzed with no cribs and the three 

defined crib support systems when the longwall face was 35 and 

10 ft inby. 

Tensioned Bolts Installed in Row 1

The specific individual loads on the first row of bolts, presented 

in table 4 and graphically shown in figure 11, were determined with 

no cribs installed, and cribs that were soft, normal, and stiff when 

the longwall face was 35 and 10 ft inby. If no cribs are installed, 

the mine roof entry sag is the highest in the middle when the face is 

35 ft outby, loading the middle entry bolts, no. 2 and 3, to 16,605 

and 15,309 lbs, respectively. As the face continues to advance and 

the panel reaches critical yielding, the loads on the bolts nearest the 

panel, no.’s 1 and 2, load to 45,378 and 34,619 lbs, respectively. 

This loading profile is a function of the effective roof span that 

increases toward the panel side as the coal yields and looses load 

carrying capacity. On the opposite side, the chain pillar yields but 

the depth is only about 4 ft which allows the remaining pillar to 

carry load and prevent load transfer onto the bolting system. The 

bolt behavior changes when the cribs are placed in the entry. The 

combination of a yielding panel and chain pillar is resisted by the 

crib, creating a shorter roof span that results in the bolt nearest the 

panel edge to carry the highest amount of load. As shown, the 

standing supports are not able to eliminate all the closure. The 

respective loads are a function of both the closure and crib stiffness. 

Clearly, the standing support has an immediate impact on the 

subsequent loading of the primary tensioned bolting system,

especially when the cribs are stiff.

Table 4. Individual bolt loads when the longwall is 35 and 10 ft inby.

No cribs Soft Normal Stiff Roof Bolt No.

35 ft 10 ft 35 ft 10 ft 35 ft 10 ft 35 ft 10 ft

1 9,187 45,378 9,048 40,626 7,760 35,793 6,650 16,680

2 16,605 34,619 16,469 32,962 15,639 26,325 14,799 20,199

3 15,309 22,914 15,244 22,426 14,810 21,136 14,271 18,932

4 7,665 10,605 7,650 9,889 7,543 9,646 7,461 8,425

Figure 10. The roof-to-floor closure and the individual crib 

loads when the longwall face is 35 and 10 ft outby.

0.0E+00

2.0E+05

4.0E+05

6.0E+05

8.0E+05

1.0E+06

1.2E+06

1.4E+06

1.6E+06

1.8E+06

no cribs soft normal stiff no cribs soft normal stiff

 35 ft 10 ft

Crib load, lbs

1.2

1.4

1.6

1.8

2.0

2.2

2.4

2.6

2.8

Roof-to-floor convergence, inch

load 35 ft

load 10 ft

closure 35 ft

closure10 ft

1

2

3

4

Stiff

Normal

Soft

No cribs

0

5,000

10,000

15,000

20,000

25,000

30,000

35,000

40,000

45,000

50,000

Bolt load,lbs

Bolt no.

Stiff

Normal

Soft

No cribs

Panel

Chain Pillar

Figure 11. Individual loads on the bolts in row 1 when the 

longwall panel is 10 ft outby with no, soft, normal, and stiff 

cribs.
Table 5. Individual bolt loads when the longwall is 35 and 10 ft inby.

Cable No cribs Soft Normal Stiff

Bolt No. 35 ft 10 ft 35 ft 10 ft 35 ft 10 ft 35 ft 10 ft

1 6,222 13,914 6,162 12,354 5,743 12,284 5,266 8,293

2 3,301 7,846 3,102 7,037 2,900 3,894 2,745 2,866

3 2,989 3,966 2,814 3,775 2,740 3,494 2,635 3,226

4 2,335 3,632 2,178 3,425 2,121 3,075 2,009 2,904

5 2,264 3,840 2,259 3,650 2,208 3,232 2,110 3,023

Cable Bolts Installed in Row 2

To continue this brief analysis, the next bolt row analyzed 

consisted of 15 ft long cable supports. The cables are simply 

thrust tight against the roof and 1 ton (2,000 lbs) of tension is 

applied. The two outside cables were angled over the longwall 

panel and the pillar to help maintain the roof after undermining 

takes place on the panel side and help minimize the damage of 

the high stress concentrations and subsequent plastic roof failure

on the chain pillar side. The specific cable loads for the 4 support 

conditions are presented in Table 5 and graphically in figure 12. 

The cable row examined is 2 ft closer to the crib and the 

stiffness of cable supports are considerably lower than the 7/8-

inch diameter resin assisted bolts. The cable subjected to the 

highest loading in all cases is the unit installed at an angle over 

the yielding longwall panel edge. As previously explained, this is 

the result of the effective roof span increasing and additionally, a 

slight rotation and movement of the roof onto the yielding 

longwall panel block. The loading condition on the cables, when 

no cribs were installed and the longwall face was 10 ft inby, 

indicate a maximum load on cable no. 1 of only 13,914 lbs, 

which is significantly below the design capacity strength of

58,600 lbs. The largest load changes take place on cable no. 1 

and no. 2. It is interesting to note that the loads on cable no. 2 

can be reduced by increasing the stiffness of the crib from a soft 

system to a normal and stiff unit. The cables installed at an angle 

over the chain pillar are subjected to only small amounts of load 

because the yield zone of 4 ft does not create high shear strains in 

the immediate and main roofs. If these forces are increased the 

cables loads would increase substantially.

CONCLUSION

The design and subsequent performance of a crib system can 

have a direct effect on the bolting system used for primary and 

secondary supports. As the forward abutment forces are

increased during subsequent longwall extraction, the residual

strength of the longwall coal block is minimized and eventually is 

zero. When this occurs the intrinsic bolts are required to carry 

more loads. If the cribs systems are stiff, the loads on the cribs 

are higher and the subsequent bolt loads are less. Conversely, 

when the system stiffness or strength of the cribs was reduced, 

the cribs carried less abutment load and the closure necessitated 

that the bolting systems provided more of the resistance to 

maintain the entry. As shown, using the design equations 

presented, the desired system stiffness for a crib can be created 

by adjusting heights and stiffness values of individual crib

components.

As evidenced by the 3-dimensional finite element modeling 

and subsequent analysis, the bolting systems not directly

supported near a crib may be subjected to higher loads and were 

required to perform more of the work. During the tailgate 

support design process, the bolting system used must consider

two distinct loading conditions; the immediate roof loads realized 

during the initial gate road development and the heavy abutment 

forces applied to the secondary supports used to maintain an 

adequate tailgate during the longwall panel extraction process. 

As clearly shown, if the crib system is soft, the bolts carry more 

loads and roof-to-floor closure is increased. If the cribs are too 

stiff and experience brittle failure before the longwall panel can 

safely mine past the area, the bolting support system may play a 

vital role in maintaining the immediate roof and assist in delaying 

main roof caving. Longwall tailgate behavior is complex, which 

necessitates the correct combinations of intrinsic bolting systems 

and/or roof-to-floor supports. The ultimate goals of these

research efforts are to provide support systems that ensure the 

stability of the entries during the longwall extraction process, 

which directly improves the health and safety for the

underground workforce. 

REFERENCES

1. Haramy, K.Y. and Kneisley, R.O. Yield Pillars for Stress 

Control in Longwall Mines- Case Study. International

Journal of Mining and Geological Engineering, 1990, Vol. 8, 

pp. 287-304.

2. Haramy, K.Y. and McDonnell, J. P. Causes and Control of 

Coal Mine Bumps. U.S. Bureau of Mines Report of

Investigation, RI 9225, 1988, 35 pp.

1

2

3

4

5

stiff

normal

soft

no crib

0

2,000

4,000

6,000

8,000

10,000

12,000

14,000

Cable load, lbs

Cable no.

stiff

normal

soft

no crib

Panel

Chain Pillar

Figure 12. Individual loads on the cables in row 2 when the 

longwall panel is 10 ft outby with no, soft, normal, and stiff 

cribs.
3. Tadolini, S.C. and Haramy, K.Y. Gateroads with Yield 

Pillars for Stress Control. Proceedings, Fourth Conference on 

Ground Control for Midwestern U. S. Coal Mines, SIU,

Collinsville, IL, November 2-4, 1992, pp. 1-16.

4. Barczak, T.M. and Gearhart, D.F. Assessment of Wood and 

Alternative Materials for Supplemental Roof Support

Construction. Proceedings, 13th International Conference on 

Ground Control in Mining, WV Univ., Morgantown, WV, 

August 5-7, 1994, pp. 190-200.

5. ABAQUS, Standard User’s Manual, Version 5.6, 2000.

6. Morsy, K.M. and Peng, S.S. Evaluation of a Mine Panel 

Failure Using the Local Mine Stiffness Criteria - A Case 

Study. Transactions, Vol. 312, 2002, pp. 8-19.

7. Zhang, Y. and Peng, S.S. Design Considerations for

Tensioned Bolts. Proceedings, 21st International Conference 

on Ground Control in Mining, WV Univ., Morgantown, WV, 

August 6-8, 2002, pp. 131-140.

8. Molinda, G.M. and Mark, C. Coal Mine Roof Rating

(CMRR): A Practical Rock Mass Classification for Coal 

Mines. U.S. Bureau of Mines Information Circular, IC 9387, 

1994, 83 pp.