INVESTIGATION OF THE JACKING FORCE CAPABILITY OF TUNNEL LINERS

Thomas M. Barczak Ron Smith 

 NIOSH American Commercial Inc. 

 Pittsburgh, PA Bristol, TN 

ABSTRACT

Steel liner plates, which are bolted together to form a shell,

are commonly used to provide temporary protective linings in the

tunneling industry to ensure the stability of the tunnel and safety

of the workers until more permanent ground reinforcement can be 

installed. As the ground deforms to try and close the opening, the 

liner plate is radially stressed, and by deforming redistributes

these loads back to the surrounding insitu earth. Grouting behind

the plates is done to fill any voids and ensure uniform contact to

the strata around the back of the plates. Another function of the 

liner plate in some applications is to act as a platform or structure

for the mining machine to react against in order to develop the 

necessary forward thrust to be able to excavate the face. In this 

case, the liner plates are loaded axially as a horizontal column of

bolted liner plate rings, creating a jacking force against the liner

plate assembly. In general, the buckling strength of the plate 

sections determines the jacking force capacity. Historically, the

jacking force capacity of tunnel liner plates has been theoretically

determined by some simplistic equations developed from tests of

single liner plates. Recently, full-scale tests on a complete

circular section of bolted liner plate ring segments have been 

conducted in the NIOSH Safety Structures Testing Laboratory 

using the unique Mine Roof Simulator to apply controlled loading 

in a manner that simulates the real world environment. Due to the

unique size and capacity of the Mine Roof Simulator, these are 

believed to be the first full-scale tests conducted on the jacking 

force capability of tunnel liner plate rings. These tests indicate

that the previously used theoretical equations tend to over predict 

the jacking force capacity of these tunnel liner plates.

Unanticipated deformations in the field have been observed at

loads below these theoretical design loads, therefore, a more

accurate determination of the jacking force capability was needed. 

This report summarizes the full-scale test results, examines

design parameters that impact the jacking force capability, and

develops new mathematical models to improve the prediction of

design loads. 

INTRODUCTION AND BACKGROUND INFORMATION 

Tunnel liners can be any number of materials, which are

used to prevent the ground from collapsing into the working tunnel

as it is excavated, providing protection to the workers until more

permanent reinforcement can be installed. Steel liner plate such

as that evaluated in this study (figure 1) is a very effective

material since it can be bolted together from inside the tunnel to 

form a contiguous structure of required diameter as the tunnel is

advanced. 

The research pertaining to this paper is the result of a 

request from American Commercial Inc in Bristol, VA to conduct 

full-scale testing of tunnel liner plates in the unique NIOSH Mine 

Roof Simulator. American Commercial Inc and its predecessor

company Commercial Shearing Inc have been a major

manufacturer of roof support products for tunnels for the past 75

years. Their first installation of steel liner plates dates back to 

1926 in the Moffat Tunnel in Colorado, which at the time was

declared the year’s outstanding engineering achievement (1).

Since then, American Commercial has become a world leader by

being involved in thousands of tunneling projects on six different 

continents. 

During the past year, structural damage was noticed on 

several tunnel liner plates in one tunneling operation (see figure

2). These deformations were unexpected in that the design 

calculations predicted a capacity significantly larger than the

estimated 240 tons of load in this particular application. These 

deformations prompted American Commercial Inc to study the 

problem to determine if: the theoretical calculations upon which 

the loading capabilities were based were in need of modification,

the load conditions were different than estimated, or if there were

quality control problems with the fabrication of these particular

Figure 1. Steel liner plate commonly used to provide 

temporary lining of tunnels. 

 
FUNCTION OF TUNNEL LINER PLATE 

In g ctions in

tunn

tward

bulg

liner plates. It was concluded that the best way to examine this 

problem would be to conduct full-scale testing. 

The Mine Roof Simulator is large biaxial load frame that is 

designed to simulate the weight and movement of overburden

rock masses in underground mining so that support structures 

can be analyzed in a controlled load environment (see appendix 

A). It has a load capacity of up to 3 million pounds of vertical

force through a 24-inch displacement. In addition to providing 

controlled convergence of the roof and floor, horizontal

displacements of up to 16 inches can be provided to induce

shearing type loading with capacities of up to 1.6 million pounds. 

The goal of this project was to load a series of tunnel liner plate 

rings configured in a 9-ft diameter in compression to simulate the 

jacking force induced on the liner plate assembly by the tunnel

boring machine. These dimensions are equivalent to that used in

the field application, where excessive deformation of the liner

plate sections occurred, so a meaningful comparison of the

laboratory results and field experiences could be achieved. 

Figure 3 shows one configuration of the tunnel liner section being

installed in the load frame in preparation for testing. 

This paper describes the full-scale tests that were conducted 

and the results of these tests relative to objectives of the study.

Included in the study were liner plates with and without gasket

material, which is used to prevent water infiltration into the tunnel, 

blocking material, which is normally specified to stiffen the liner 

plate for axial loading, and variations in the turning radius of the

liner plate edges. This was the first full-scale testing of liner plate

rings conducted by American Commercial Inc and it is believed to 

be the first ever tests conducted to this scale by anyone. The

tests proved to provide valuable information pertaining to the

failure mode of liner plate subjected to axial loading and the

significance of blocking material in controlling this failure mode. It

was concluded that the previous theoretical calculations of jacking 

load capacity, which were based on single liner plate testing as 

opposed to full ring segments, did not accurately predict the 

capabilities of these structures as used in the actual tunnel 

application. From the full-scale tests, new mathematical models 

would be developed to improve the load capacity prediction and

design requirements for tunnel liner plate applications. Hence, an 

improvement in the state-of-the-art of tunnel design providing 

increased safety to the workers has been achieved through these

full-scale tests.. 

Figure 3. Four layer section of liner plate being 

constructed in the Mine Roof Simulator for full-scale 

testing. Figure 2. Damage to liner plates in one tunneling 

application at an estimated jacking load of 240 tons. 

eneral, liner plates can provide two basic fun

eling operations. In conventional tunneling methods where

tunnels are excavated by full face, heading and bench, or multiple 

drift procedures, these tunnel liners are used to stabilize and 

prevent infiltration of surrounding material, whether it be soil or

rock, into the tunnel and in so doing provide protection to the 

workers and to secure the area for more permanent lining of the

tunnel, usually by concrete, if necessary. The steel liner plate is 

fabricated to bolt together to form a complete or partial circle of 

protection. As the ground deforms, the liner plate is actively

loaded and deflects in response to the ground pressure. 

However, as the tunnel liner deflects it then transfers the active

ground loading back into surrounding areas of the ground, which 

is now acting as a passive restraint to the tunnel liner

deformations. This concept is illustrated in figure 4. Here it is 

shown that as overburden material vertically loads the tunnel

liner, it deflects inward at the top and outward at the sides. The

ground resists the outward deflection of the sides of the liner plate

by developing a passive force equal in magnitude and opposite in

direction to the force exerted by that of the lining (figure 4). 

The ability of the surrounding ground to resist the ou

e of the liner plate ring is the key to vertical load support

induced by the weight of the rock or soil above the tunnel. With 

the ring confined to a small deflection, the thrust line induced by

the load is forced to follow the ring of liner plates. Thus, the ability

of the assembly to withstand the applied load depends upon its 

 

 
ability to transmit radial ring thrust from plate to plate around the

ring without inducing high bending stresses that would cause the

liner plate to buckle. Hence the design criteria to control the 

surrounding ground for the tunnel liner plate are clear (1). 

 Deflection Criteria -- First, it is imperative that the tunnel 

dimensions be controlled to limit the liner plate deflections 

such that the thrust line is maintained within the liner plate 

assembly. Typically, deflection should remain within 3% of 

the nominal diameter. If the tunnel is over-mined or the soil 

or rock is unstable and unable to maintain the shape of the

original opening, the tunnel is typically grouted or filled with

some form of fill material to maintain the ground presence 

needed to redistribute the active liner plate loading back into 

the ground. 

Figure 4. Function of tunnel liner to control ground 

pr g

 Minimum Stiffness – The stiffness of the liner plate assembly 

is an important parameter because it defines how much the 

liner plate will deflect under load, and as previously

described it is imperative that the liner plate be able to

maintain its structural stability to transmit active loading

through deflections elsewhere within the assembly back into

the ground. The minimum stiffness requirements depend on

the type of liner plate used. The minimum stiffness 

requirements as published by the American Association of

State Highway and Transportation Officials (AASHTO) and

the American Railroad Engineering and Maintenance of Way 

Association (AREMA) is 111 lbs/in for four-flange liner plate

 Critical Buckling Criteria – For the liner plate to function as 

designed where the thrust line is maintained within the liner 

plate assembly, each liner plate is subjected to compressive

stresses along the length of the plate. Hence, the buckling

strength of the liner plate is a critical design parameter, since 

if the buckling strength was exceeded, the liner plate would 

weaken and no longer be capable of carrying the thrust

associated with the active ground loading. The buckling 

strength of the liner plate is a function of the cross-sectional 

area of the liner plate, the moment of inertia and radius of

gyration of the plate, and the modulus of elasticity for the 

steel. The design for buckling is accomplished by limiting 

the ring compression thrust to the buckling stress, not to

exceed 28,000 psi or the yield strength of the steel,

essure by distributing loads back into the surroundin multiplied by the effective cross-section area of the plate

ground. 

Tunnel

Boring

Machine

Axial 

loading

Ground 

pressure 

restrains 

Jacking 

cylinder

Drawing not to scale 

3 

Figure 5. Diagram of axial loading of tunnel liner plate due to jacking forces associated with advancement of the miner. 
w that which

(which is generally considered as only 50% of the actual 

area) with an additional factor of safety of 2 built into the

design. Hence, the allowable thrust is well belo

should cause buckling of the tunnel liner plate. 

Joint Strength – The seam strength of the liner plates must

be sufficient to withstand the thrust from the total load 

supported by the liner plate. Tables of joint strength,

measured in kips/ft for different plate thickness, are available

from the above-mentioned specifications. A factor of safety

of 3 is typically employed in determining the joint strength,

meaning that the actu

 

al joint strength is at least 3 times that 

of expected loading. 

The second function of liner plates is to act as a reactionary 

platform or structure upon which certain types of tunnel boring

machines thrust against to develop the necessary force at the 

cutting head or to steer the cutting head in the desired direction 

during the excavation of the tunnel (figures 5 and 6). It is this

function of the liner plate that was investigated in the full-scale

labo

the tunnel

lining

 State

niversity. These results are reported in the next section. 

LINER PLATE TE YOUNGSTOWN 

transfer among the plate skin, flange, and 

stiffening members. 

ratory tests conducted as part of this study. 

Thrust is provided by a set of hydraulic rams, which exert 

force against the stationary liner plate being held in place by the

ground reactions described above. A jacking ring or solid steel 

ring is typically placed immediately behind the boring machine to

enable the jacking forces to be uniformly applied to the

subsequent rings of liner plate. It is normal to dissipate thrust 

loads due to the jacking forces within the first 50 feet of

as long as the annulus is grouted or backpacked. 

The allowable jacking force is again calculated based on the

buckling strength of the bolted liner plate rings, except the loads

are now applied axially along the direction of the tunnel

excavation instead of radially through the circumference of the

liner plate ring as would occur from the ground loading.

Traditionally, computing the buckling strength of the flanges

where the liner plates are bolted together, and adding the load

carrying capacity of the skin of the liner plate in between the

flanges have calculated the jacking capacity of the liner plate

assembly. Several assumptions about the effective width of the

flange and the maximum stress in the skin portion of the plate

between the stiffeners were typically made in these computations. 

In order to clarify these variables, which are critical in the 

determination of jacking thrust capability, a series of tests on

single and three-section liner plates was conducted in 1973 by

Commercial Shearing Inc in cooperation with Youngstown

U

STS CONDUCTED AT

STATE UNIVERSITY 

Two tests to evaluate the jacking capability of steel tunnel

liner plate were conducted in this unpublished study in 1973. The

first test was conducted on a single liner plate and the second

was conducted on three plates bolted together as shown in figure

7. The liner plates were instrumented with strain gages in an

attempt to measure the stress profile the across the liner plate. A 

load of 50,000 lbs in the first test and 25,000 lbs in the second

test was applied in 5,000 lbs increments uniformly to the liner 

plate section(s) by a Forney Compression Testing machine. The

goal was not necessarily to test the plates to failure as much as to

determine the load

AXIAL LOAD 

Figure 6. Photo showing tunnel boring machine 

equipped with jacking ring to uniformly transfer jacking 

loading into the tunnel liner. 

Fig at

Youngstown State University in 1973. 

ure 7. Three-segment test conducted

 
e could actually carry since the buckling of the 

plate

y the Navy on

The results of the three-section liner plate test revealed that

the stress in the skin of the plate reached an average value of

about 10,000 to 15,000 psi and more or less remained constant. 

This indicates that the skin of the plate was buckling elastically 

and that the load was then being transferred from the plate skin to

the stiffeners and/or the plate flanges. Based on this test, the 

10,000 psi stress in the skin plate was used to calculate the 

expected jacking force capacity of the liner plate. Prior to this 

study, the full yield strength of the plate (i.e., 28,000 psi) was

erroneously used in calculating the contribution of the nonreinforced skin of the liner plate to the overall jacking force 

capacity, resulting in an estimated capacity that was considerably 

higher than the plat

was ignored. 

Another outcome of these initial tests was a revision in the

way that the effective width of the flange section was calculated.

Previous methods for determining the effective width came from

studies conducted b submarine hulls. They

recommended using 1.57⋅ r⋅ t where r is the radius of the

hull (or liner plate in this case) and t is the thickness of the plate. 

It was concluded from this test and other tests conducted by 

Commercial Shearing Inc that a more reasonable value for

Where t = thickness of liner plate, inches 

the

effective width (We) could be found by the following equation. 

 Fy = yield strength of the steel, psi 

gnores the corner radii and the effect of the holes in the

nges.

. Calculate the effective width of the bolted flange section. 

2. he strength of the bolted flange section as follows

(figure 8): 

Determine the neutral axis for the flange section. 

rea of top section (A1) = (0.2092) x (1.8305) x 2 = 0.7659 in2

rea of bottom section (A2) = (0.2092) x (1.9158) x 2 = 0.8016 in2

 . = Poisson ratio = 0.3 

 E = modulus of elasticity, 29 x 106 psi

Based on this work, the estimated jacking force of a full set

of liner plates could be theoretically calculated. Here is an

example for a non-reinforced liner plate using 5-gage plate

constructed in a 9-ft diameter assembly requiring 9 liner plate

sections to form a full circle ring segment. Note that this also

assumes the plates are manufactured with the flanges at exactly 

90° and i

fla

1

Calculate t

4.0794 inches

28 10

29 10

3 (1 0.3 )

0.2092 W 3

6

2 e = ⋅

⋅ ⋅ ⋅ − =

e E Fy

t W /

3(1 )

2 ⋅ − . = A

A

0.2092 in

We= 4.0794 inches 

1.8305 in

X X

Neutral Axis 

1.4770 in

1.9158 in

2.0204 in

0.9579 in

Top section 

Centroid

0.5434 in

Bottom

section

Centroid

0.5191 in

Figure 8. Diagram of bolted flange section at ends of liner plate sections. 

 
Moment arm to centroid of top section from X – X axis – (Y1) =

2.0204 in 

m to centroid of bottom section from X – X axis – (Y2) = 

.9579 in

Neutral axis is computed from (A1 + A2) y = A1y1 + A2y2 as follows: 

.7659) + (0.8016) = (0.7659 x 2.0204) + (0.8016 x 0.9579) 

 

ertia

is computed using the parallel axis theorem as follows: 

 of inertia

 

 A.I.S.C.

pinned end condition is assumed resulting in k= 1.0. 

Moment ar

0

(0

_

y

Moment of inertia – The moment of inertia is computed by

dividing the bolted flange section in two areas: (1) the top

two pieces of the flange which are in the same plane as the 

skin of the liner plate and (2) the two perpendicular sections 

which form the flange arrangement. The moment of in

Compute the radius of gyration (r) form the moment

using column formulas from

specifications (2): 

0.664 in

1.5675

0.6902

A A

I

r

1 2

Total = = + =

(ITotal) and the area (A) of the flange section. 

Assume the flanges act as columns and calculate the 

allowable stress

A

Since ( )

r

k ⋅ L < Cc, the following equation from the A.I.S.C. is 

used to compute the allowable stress (F ) in the flange section 

:

Compute the load carried by the skin of the plate section 

using the assumption that the stress is 10,000 psi. 

 

data to refine the methodology for theoretical 

calcu

w

p

placement at a rate of 0.5 inches per 

minute by the load frame. Load-displacement profiles for all tests 

are provided in Appendix B. 

a

(2)

( )

( )

15,988 psi

143

23.4663

8

1

143

23.4663

8

3

3

5

143

23.4663 2

1

28,000 1

C

r k L

8

1

C

r k L

8

3

3

5

c

C

r k L2

1

F 1

a

F

3

3

2

2

3

C

3

C

2

y 2

=

+ ⋅ − ⋅















 ⋅ −

=



















 ⋅ − 













 ⋅ +



















 ⋅ ⋅ −

=

Neutral Axis 1.4770 inches _y = =

( ) ( ) 2

2 2 2

2 I Total = I1 + A 1d1 + I + A d

4

3 3

1

0.0028 in

1.8305 0.2092 2 12 b h 1

12 I 1

=

= ⋅ ⋅ = ⋅ ⋅ ⋅

4

3 3 2

0.2452 in

0.2092 1.9158 2 12 b h 1

12 I 1

=

= ⋅ ⋅ = ⋅ ⋅ ⋅

d 0.5434 in 1 =

d2 =0.5191 in

2 4

2

Total

(0.2452 0.8016 0.5191 ) 0.6902 in

I (0.0028 0.7659 0.5434 )

+ ⋅ =

= + ⋅ +

12.53 tons

2000

1.5675 15988

Flange Load Area of flange stress

= ⋅ =

= ⋅

( )

0.2092 10000 70,307 lbs 35.2 tons

Plate Load Area Stress 37.6875 4.0794

⋅ ⋅ = =

= ⋅ = −

The total load for a nine plate ring segment would be: 

Skin Load ( ) 9 12.53 (9 35.2) 429 tons

Jacking Load Capacity Flange Load

+ = ⋅ + ⋅ =

=

FULL-SCALE TESTS CONDUCTED AT NIOSH 

The first full-scale tests of an assembled tunnel liner from

bolted plate sections were conducted at the NIOSH Mine Roof

Simulator (MRS) facility in Pittsburgh, PA in December of 2000

and January of 2001. The goals of the test program were: (1) to

compare the full ring assembly tests with previous test results 

conducted on only single plate sections, (2) to evaluate the impact

and failure mode of liner plate with and without reinforcement, and

(3) to use this

lation of expected jacking load capacity for various liner plate

constructions. 

The test program consisted of testing variations of a 9-ft

diameter assembly of 5-gage liner plate with plate sections 

measuring 37.6875 inches in length and 16 inches in width. A

total of six tests were conducted, three tests on a single ring of

liner plate forming a 9-ft diameter assembly (see figure 9), and

three tests conducted on a stack of four, 9-ft-diameter ring

segments as shown in figure 10. The parameters evaluated in 

the study included: (1) reinforcement of the plate ranging from no

reinforcement to wood timbers to 3-inch steel channel, (2) sealing 

material around the perimeter of the plate including either gasket 

material or no sealing material, and (3) the flange radius which

as either standard or a tight radius. The test arrangements and

eak loading observed for each configuration are summarized in

table 1. Axial loading of the tunnel liner assembly was provided

through a controlled axial dis

( ) ( ) 143.0

28,000

2 3.1416 29 10

F

2 E

c

C

2 6

y

2

= ⋅ ⋅ ⋅ = ⋅ ⋅ = π

( ) ( ) 23.4663

0.664

1.0 16 0.2092 0.2092

r

k L = − − = ⋅

 
Table 1. Test configurations and measured peak loading for full-scale tests in the MRS 

Test Identification Perimeter 

Material 

Ring 

Sections 

Reinforcement

Material 

Flange 

Radius

Peak Load, tons

Tunnel Liner #1 Gasket 4 None Standard 287 

Tunnel Liner #2 Gasket 4 3x6x16-in oak 

timbers 

Standard 325

Tunnel Liner #3 Plate 4 3x6x16-in oak 

timbers 

Standard 360

Tunnel Liner #4 Plate 1 None Standard 332 

Tunnel Liner #5 Plate 1 3-inch steel 

channel 

Standard 765

Tunnel Liner #6 Plate 1 None Tight 360 

 

The conclusions drawn from this full-scale testing are 

summarized as follows: 

1. The previous methodology based on single liner plate testing 

significantly over predicts the jacking load capacity of steel

liner plate fabrications. Referring to the previous section, a

capacity of 429 tons was computed for a non-reinforced liner 

plate assembly using the assumption that the full skin of the 

liner plate carries a stress of 10,000 psi. The full-scale tests 

show an axial load capacity ranging from 287 to 360 tons. 

This compares favorably from observations made in the

field, where jacking forces of approximately 240 tons on a 

non-reinforced liner plate assembly caused deformations 

indicating that the capacity of the liner plate had been

exceeded in this application. 

2. Full-scale testing is necessary to properly evaluate the load

interaction between the rings or layers of liner plates, which

are present in the actual construction in a tunnel application.

As shown in figure 11, the middle two rings of the four-ring

liner plate test structure observed deformations that were not

found in the rings that contacted the roof and floor of the

load frame or in the single ring test arrangement. The joint 

or boundary between the layers of rings controls the 

deformation to a large degree as seen in figure 11.

Essentially, the joint arrangement in the four-ring 

configuration provides a degree of instability that is not 

present in the single-ring configuration due to manufacturing 

tolerances and the fastener locations, etc. The axial 

stiffness of the structure is also significantly different if only a

single ring of liner plate is tested (see figure 12). 

Figure 9. Single layer test of 9-ft diameter ring of tunnel 

liner plate (test conducted at NIOSH Mine Roof Simulator 

facility). 

3. Gasket material does not have a significant effect on the 

jacking capacity of the tunnel liner plate assembly. The

gasket material is relatively soft and deforms quickly under

load. One test showed a slight reduction in jacking capacity 

when the gasket material was used, but it is not certain that

the gasket material was fully responsible for this drop (about

10 pct) in capacity. 

4. The tight radius on the liner plate does make a small 

improvement in jacking force capacity. It probably reduces 

the eccentricity in the load interaction between the liner plate

rings. 

5. The failure mode is sensitive to the reinforcement used in 

the liner plate construction. With no reinforcement, the

vertical flange where the ends of the plates bolt together is 

by far the stiffest area, and axial loading is concentrated in

this section before and after the remaining skin of the plate 

buckles. The vertical flange loading damages the adjacent 

ring as seen in figure 13. This damage probably causes a 

progression of failure in the skin that eventually causes

buckling of the skin and subsequent rotations creating 

moments about the horizontal joints between ring segments 

where the plates bolt together. The vertical flanges for the

most maintain their structural integrity in this configuration

(figures 10 and 12). 

Figure 10. Full-scale test of four-ring assembly of tunnel 

liner plate in the Mine Roof Simulator. 

 
Figure 13. Vertical flange loading causes damage to the 

flange of the adjacent ring. 

The use of timber planking changes this failure mode. The

timber is stronger than the vertical flange and transfers

loading across the skin of the liner plate into the vertical 

flange section in the plate segments above and below. The

load is focused on the flange causing the flange to buckle as 

its strength is exceeded, while there is no apparent damage 

to the timber planking (see figure 14). 

Figure 11. Middle two rings controlled the failure of the 

four-layer configuration. 

N IO SH SAFET Y ST RUCT URS T EST IN G LAB O R AT O R Y

AM ER IC AN CO M M ERCIAL INC -- T UNNEL LIN ER PLAT E

0

100

200

300

400

500

600

700

0.00 1.00 2.00 3.00 4.00 5.00 6.00

DISPLACEM ENT, inche s

SUPPORT LOAD, kips

Tunnel Liner Plate -- Standard Des ign (5 gage) -- No reinforcem ent -- 9 ft dia -- 4 lay er sec tion

Tunnel Liner Plate -- Standard Des ign (5 gage) -- No reinforcem ent -- 9 ft dia -- Single lay er

Figure 14. Timber planking transfer loading and cause

damage to the flange area below with no apparent 

damage to the timber planking. 

The steel channel reinforcement was not tested in the fourring configuration, but it is likely that the failure mode would

be similar to that of the timber-planking configuration 

previously described. The steel channel, since it is welded

to the liner plate skin, should stiffen the plate and increase

its load capacity, so the steel channel reinforcement is likely

to have a higher jacking load capacity than that of timber

planking. The single-ring test produces an artificially high

axial loading capability because the channel load is not

being transferred into the flanges of the adjacent ring 

segments. In the single-ring test, the steel channel actually 

buckles (see figure 15), but again this is not expected to

occur in the multi-ring arrangement since the channel

Figure 12. The axial stiffness of the structure is

different if only a single ring of liner plate is tested. 

 
 

buckling strength is greater than the vertical flange strength

and the channel will more than likely transfer load to the 

adjacent flange without failing itself. It is also seen from the

tests that localized failure is occurring in the region near the 

middle bolthole of the flange (see figure 16), which is 

reducing the capacity of the flange in this load environment. 

This type of failure was also observed with the four-ring

timber planking test configuration. 

RECOMMENDATIONS FOR IMPROVING THE 

THEORETICAL PREDICTION OF JACKING FORCE 

It is clear from the full-full-scale tests that the simple model 

developed as a result of the 1973 testing significantly over

predicts the jacking capacity as the predicted load was 429 tons 

while the measured load for non-reinforced liner plate during the

full-scale testing was only 252 tons. Most of the error in this

model is likely from over-estimating the contribution of the skin of

the liner plate. 

The load transfer though a series of liner plate sections 

resulting from the jacking force of the mining machine is quite

complex, more than likely involving complex inelastic buckling and 

interactions between the various structural sections of which the

complete assembly is made and the boundary conditions which 

control the load development. None-the-less, improvements in

the approximation of the jacking capacity can still be made by 

improving some of the basic assumptions made in the elastic 

buckling model. 

 Non-Reinforced Liner Plate – The non-reinforced liner plate

is modeled in three sections (see figure 17): (1) the bolted 

flange section with its load capacity determined by its 

buckling strength, (2) a stiffened area adjacent to the flange

with an assumed average stress of 10,000 psi based on the 

previous 1973 study, and (3) the center section which is 

assumed to buckle elastically.

Single plate section 

Figure 15. Buckling of steel channel in single ring test.

Non-reinforced

skin 

 Bolted Flange Section – The capacity of the bolted flange

section is computed as shown in the previous section using

the A.I.S.C. column buckling equation to compute the 

buckling stress (Fa). Using this equation and the area of the 

flange section, it was determined that the flange is capable 

of carrying 12.53 tons of load before it buckles.

 Stiffened Area Adjacent to Flange – It is clear from the tests

that the area adjacent to bolted flange section is capable of 

carrying a higher load than the center section of the plate.

The effective width of this stiffened area is assumed to be

equal to 1.5 times the effective area of the flange section. 

Hence, for the 5-gage liner plate, the effective area adjacent 

to each flange (see previous example) is computed as

4.0794 x 1.5 = 6.12 inches. If it assumed that the average

stress across this section is 10,000 psi, then each of the two

stiffened sections per liner plate is capable of carrying a load 

of 6.4 tons.

 Center Section of Plate – The remaining portion of the plate

is assumed to buckle. The buckling load is computed using 

A.I.S.C. specifications. We begin by analyzing a 1-inch wide

section of plate. The calculations necessary to compute the

buckling load are as follows:

Flange

section 

Flange

section

Stiffened areas

(10,000 psi)

Figure 17. Non-reinforced liner plate is modeled in three 

sections (flange section, stiffened area, and nonreinforce skin).

Figure 16. Localized failure occurs near middle bolthole 

of flange section. 
ince

0.0604 in

0.2092

0.000763

A

I Radiusof gyration(r)

=

= =

S Cc r

k L = ⋅ the following equation is used to compute the 

buckling stress (2):

Total Load -- The total load of a 9-ring segment would be the 

Timber Planking Reinforced Liner Plate

sum of these three components equating to 273 tons. 

4

3

3

0.000763 in

12

(1) (0.2092)

b h 12 Momentof Inertia 1

= ⋅

= ⋅ ⋅ =

2 Area = 1 ⋅ 0.2092 = 0.2092 in

 – Timber is often

sed to reinforce the liner plate. The timber is cut to length 

 

u

to fit snugly in between the horizontal flanges of the liner 

plate and is wedged in place as the liner plate ring is 

constructed. Since the timber is not fixed to the plate, it 

does not alter the properties of the plate and functions 

primarily to transfer the vertical flange loading from the plate

above or in front across the liner plate skin and into the 

flange in the ring below or behind. Hence, the timber is 

modeled by assuming a load capacity equivalent to that of

the steel flange, which was previously calculated to equal to 

12.53 tons. However, the test data shows that the flange in

this application tended to have localized buckling at the 

middle bolthole, so only 75% of the flange capacity is used

in this computation. Hence, the jacking load capacity for the

timber planking reinforced liner plate in a 9-ft diameter

section with 5-gage plate is 329.5 tons, computed as the 

sum of the vertical flange capacity, the capacity of the

stiffened area adjacent to the flange, the capacity of the

unstiffened area of the plate, and the timber which has an 

assumed load equal to that of the vertical flange. 

Flange Capacity – As previously determined, the flange 

capacity is computed to be 12.53 tons, which is r

( ) 258.0126

0.0604

1 16 0.2092 0.2092

r

k L = ⋅ − − = ⋅

143.0

28000

2 3.1416 29 10

F

2 π E C

2 6

y

2

c

= ⋅ ⋅ ⋅

= ⋅ ⋅ = educed by

 

25% due to premature buckling at the bolthole. 

Stiffened Section of Plate – As shown above, the stiffened

area is assumed to equal 6.12 inches of plate adjacent to 

 

each of the two vertical flanges, with each of these areas 

providing a load capacity of 6.4 tons. 

Unstiffened Section of Plate - As shown above, the

remaining 21.37 inches of liner plate is expected to buckle at

 

tons.

Stee late

a load of 5.02 tons. 

Total Load – This provides a total load from the sum of these

components of 329.5

l-Channel Reinforced Liner P – In order to stiffen the plate

nd improve its buckling strength, a section of channel can be

ent was tested. The capacity of this arrangement is 

com

in the tests conducted at Youn

study.

a

welded to the interior skin of the liner plate, which was shown in

figure 14. 

Only a single ring of liner plates with 3-inch channel

reinforcem

puted as above for the non-reinforced plate except the

channel strength is now added to the strength of the plate and the

entire plate is assumed to be stiffened by the addition of the steel

channel. Since the channel is welded to the plate, it is also

assumed that the stiffening of the plate by the channel allows the

stress to increase to 15,000 psi, which is the maximum measured

gstown State University in the 1973

 Flange Capacity – As previously determined, the flange 

capacity is computed to be 12.53 tons. 

2,243 psi 23 258.0126

12 3.1416 29 10

r

k L 23

12 π E F

2

2 6

2

2

a

= ⋅

⋅ ⋅ ⋅

=





 



 ⋅ ⋅

⋅ ⋅ =

Plate Load (center section) = Area ⋅ stress

( ( ))

5.02 tons

2,000

2,243 0.2092

37 .6875 4 .0794 4 .0794 2 1 .5

Plate Load (center section)

⋅ ⋅ =

− − ⋅ ⋅

=

( ) ( )

( ) 9 5.02 (0.75 9 12.53) 329.5 tons

Total Load 0.75 9 12.53 9 2 6.4

⋅ + ⋅ ⋅ =

= ⋅ ⋅ + ⋅ ⋅ +

 Stiffened Section of Plate – The stiffened length of the plate

is the section between the flange and the 3-inch-wide 

 

channel. With an overall plate length of 37.6875 inches and 

an effective flange width of 4.0794 inches, the stiffened 

length of plate is computed as 30.6081 inches. Using the 

plate thickness of 0.2092 inches, the area of this section of

plate is 6.4 in2

, which translates into a load capacity of 48

tons assuming a stress of 15,000 psi. 

Channel Capacity -- For the tests conducted in this study, a 

C3” x 6# channel section (3 inch wid

Section Load Center Section Load

Total Load Flange Load Stiffened

+

= +

( ) ( )

( ) 9 5.02 273 tons

Total Load 9 12.53 9 2 6.4

+ ⋅ =

= ⋅ + ⋅ ⋅

e channel weighing 6

lbs per linear foot) of Grade 50 steel was welded to center

 
 

span of the liner plate. The area of the channel is 1.76 in2, 

the moment of inertia is 0.305 in 4

, a column slender ratio 

limit (Cc ) is 107, and the least radius of gyration is 0.416 in. 

The buckling stress (Fa) is then computed from the A.I.S.C.

column buckling equation to be 26,184 psi, which translates 

into a load capacity for the 3-inch channel section equaling 

23.0 tons. 

Total Load – This provides a total load from the sum of these

components of 751.5 tons. 

ents where rings of liner plates 

re butted together, the weakest member will control the failure in 

the s

SUMMARY ASSESSMENT 

Again, there are many assumptions used in these relatively 

crude models, luated through

testin

COMMENDATIONS 

The Mine Roof Simulator facility at NIOSH provides the

capabi scale

testin

at the blocking or

stren

One of the goals of this study was to improve the theoretical 

mputation of jacking force capacity. Previous computations 

were

ing the required jacking force 

capa

)

However, in multiple segm

a

ystem. It is more likely that the channel will serve the same

basic function as the timber planking, acting to simply transfer the

loading to the adjacent ring. Hence, a more appropriate model

may be that used for the timber planking where the channel is 

simply modeled with a capacity equal to that of the flange. The

only difference would be that the channel stiffens the plate to

produce a higher loading than would be expected with the timber

planking. The stiffening assumed in this model would be an area

adjacent to the flange, just as was done in the timber model, as

well as an area adjacent to the channel using the same 1.5 

multiplier. Hence the stiffened areas are summarized as follows:

(1) 6.12 inches adjacent to the flange providing a load capacity of

6.4 tons and (2) 12 inches in the center section of the plate where

the channel is located providing a capacity of 12.6 tons. The 

remaining 6.37 inches of plate is assumed to be non-reinforced

and is modeled with a buckling load capacity of 1.5 tons. In this 

model, the capacity of the steel channel arrangement would be 

410.9 tons. 

which have not been fully eva

g. A more detailed numerical analysis of the plate stress by 

finite element modeling may produce more accurate models than 

these presented here, but as the comparison in figure 18 shows, 

these crude models do provide a reasonable prediction of the

tunnel liner jacking capacity and are an improvement over

previous theoretical computations. 

CONCLUSIONS AND RE

lity to conduct full-scale testing of tunnel liners. Fullg of liner plate can provide valuable insight into the capacity 

and failure mode of these structures, as was observed in the few 

tests conducted as part of this study. The loads observed in the 

full-scale tests were consistent with damage and loading that has 

been observed in the field, hence the tests are believed to provide

a valid examination of jacking force capability.

In terms of the jacking force capacity of tunnel liner plate, it

is clear from the full-scale laboratory tests th

gthening material can make a large difference in both the

capacity and mode of failure for these structures. It seems 

prudent that some form of block material should always be used

to enhance the strength of the plate relative to axial loading when

the plate is used as a reaction fixture for advancement of the

tunnel boring machine. This too is confirmed by field experience.

In the one application referenced as the stimulus for conducting 

this full-scale laboratory study, an unblocked tunnel liner plate 

suffered deformation when other field installations using the same

plate with reinforcement performed well without any signs of

failure. Timber planking is an easy and effective form of blocking.

The major problem with timber planking is the inconsistency in

spacing of these on site and possible variations in length and

retention of the timber planks under fluctuating loads when

thrusting. Reinforcement by welding additional steel members

eliminates these problems and is generally a more effective

solution. 

( ) (

( ) 9 23.0 751.5 tons

Total Load 9 12.5 9 48

⋅ =

= ⋅ + ⋅ + NIOSH SAFETY STRUCTURES TESTING

LABORATORY

Tunnel Liner Axial Load Tests -- 5-gage plate in 9-ft

diameter section

0

100

200

300

400

500

600

700

800

900

NonreinforcedTimber

Planking

Reinforced

Steel

Channel

Reinforced

(Single

Steel

Channel

Reinforced

(multiple

JACKING CAPACITY, tons

co

largely based on a study completed in 1973 in which only 

single liner plate sections were tested. These computations 

tended to over predict the jacking capacity of tunnel liner plate

assemblies, most likely by due to poor assumptions made in 

regard to the buckling of the plate under the loading application. 

Although the models developed in the current study are still fairly

crude with several assumptions made regarding the stress 

transfer though the liner plate, they are able to reasonable predict 

the jacking capacity as measured in the full-scale tests conducted

under controlled loading in the laboratory using the Mine Roof

Simulator load frame. Additional testing is needed to confirm that 

the models would work equally well with different plate gages,

diameters, and fabrications. A finite element analysis of the

stress development in a liner plate may also provide valuable

information into this issue and should be explored as part of a

continuing study of this subject. 

From a design perspective, it is clear that the vertical flange

is a critical member in provid

city. It is also seen from the tests that the boltholes, which 

layer) layer)

Figure 18. Comparison of test results and predictive models of 

jacking force capacity. 

( ) ( ) ( )

( ) 9 12.6 (0.75 9 12.5) 410.9 tons

Total Load 0.75 9 12.5 9 2 6.4 9 1.5

⋅ + ⋅ ⋅ =

= ⋅ ⋅ + ⋅ ⋅ + ⋅ +

 
allow the plates to be bolted together

, weaken the vertical flange. The flange strength could be enhanced by eliminating the middle 

bolthole or enhancing the section modulus in this area to prevent

failure in this region and improve the overall capacity of the liner

plate assembly in terms of jacking force. The horizontal flange is 

also important. Strengthening this flange by adding gussets 

would help to reduce the bending of the flange and reduce the

rotation of this joint that contributes to buckling of the body of the 

plate. 

REFERENCES 

American Institute of Steel Construction, Ninth Edition, 1989, 

“Manual of Steel Construction: Allowable Stress Design,”

Proc

,” Youngstown Printing Company, Youngstown,

APPENDIX A -- DESCRIPTION OF THE 

MINE ROOF SIMULATOR LOAD FRAME 

The trolled

h aulic press custom built by MTS Systems Corporation to U.S.

Bure

stics. The 

size

 up to 3 million 

poun

ed on both the left and right

side

(y-direction). The load frame has a reactive

capa

pp5-42. 

tor, R. V. and White, T. L., 1968, “Rock Tunneling with Steel

Supports

OH, 291 pp. 

Mine Roof Simulator (MRS) is a servo-con

ydr

au of Mines (USBM) specifications. It is designed specifically

for longwall shield testing, and is the only active load frame in the

United States that can accommodate full-size shields. 

A functional diagram of the load frame is shown in figure A1. The load frame has several distinctive characteri

of the platens is 20-ft x 20 ft. The upper platen can be

moved up or down and hydraulically clamped into a fixed position

on the directional columns to establish a height for testing. With a

maximum vertical opening between the upper and lower platen of

16-ft, the load frame can accommodate the largest shields

currently in use. Controlled movement of the lower platen

provides load application. The load frame is a biaxial frame,

capable of applying both vertical and horizontal loads. Load

actuators are equipped with special hydrostatic slip bearings to

permit simultaneous load and travel. This allows vertical and

horizontal loads to be applied simultaneously. The capability to

provide controlled loading simultaneously in two orthogonal

directions is unique at this scale. 

A set of four actuators, one on each of the corners of the

lower platen provides vertical loading. Loads of

ds can be applied in the vertical direction by upward

movement of the lower platen. Each actuator is capable of

applying the full 3 million pounds of force, so that the specimen

can be placed anywhere on the platen surface and the full 3 

million capacity can be provided. The vertical (upward) range of

motion of the lower platen is 24 in. 

Horizontal loading is also provided by a total of four

actuators, with two actuators locat

of the load frame just below the floor level. These actuators 

act in pairs to provide horizontal displacement of the lower platen 

in either a positive or negative (x) direction, reacting off the corner

columns of the load frame. The horizontal range of motion of the

lower platen is 16 in. 

There is no programmable control of the lower platen in the 

lateral horizontal axis

city of 1.6 million pounds in this direction, but loads cannot

be applied in the lateral direction. The range of motion of the

lower platen in this direction is + 0.5 in. 

The unstressed reference frame that provides feedback on

platen displacements and rotations to the closed-loop control

syste

displ

d application. The 

rate

m provides Six degrees of freedom control of the lower

platen. Pitch, yaw, and roll of the lower platen are controlled to

keep the lower and upper platens parallel during load application. 

A shock absorber actuator is positioned on the left and right 

side of the lower platen. These shock absorbers will control the

acement of the lower platen to less than 0.1 inch in the event

of sudden failure of the support specimen. The shock absorber

action absorbs energy stored in the load frame so that it is not

unintentionally released to the test specimen. 

Two hydraulic pumps provide up to 3,000 psi of pressure to 

the vertical and horizontal actuators during loa

of movement of the lower platen is limited by the 140 gpm

capacity of the hydraulic pumps. The maximum platen velocity 

assuming simultaneous vertical and horizontal displacement is 

5.0 in per minute.

 
APPENDIX B – FULL-SCALE LOAD-DISPLACMENT TEST DATA 

NIOSH SAFETY STRUCTURS TESTING LABORATORY

AMERICAN COMMERCIAL INC -- TUNNEL LINER PLATE -- 18 JAN 01

0

200

400

600

800

1,000

1,200

1,400

1,600

0.00 0.20 0.40 0.60 0.80 1.00 1.20

DISPLACEMENT, inches

SUPPORT LOAD, kips

Tunnel Liner Plate (Test #4)-- Single ring test (9-ft diameter) -- 5 gage plate -- no reinforcement

Tunnel Liner Plate (Test #5) -- Single ring test (9-ft diameter) -- 5 gage plate -- tight flange radius

Tunnel Liner Plate (Test #6)-- Single ring test (9-ft diameter) -- 5 gage plate -- 3 in channel reinforcement

Test #4

Test #5

Test #6

 

NIOSH SAFETY STRUCTURES TESTING LABORATORY

AMERICAN COMMERCIALINC -- 5-GAGE TUNNEL LINER PLATE -- 6-DEC-2000

0

100

200

300

400

500

600

700

800

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0

DISPLACEMENT, inches

LOAD, kips

Tunnel Liner Plate (Test #1) -- Four-ring test (9-f t diameter) -- gasket edging -- no reinf orcement

Tunnel Liner Plate (Test #2) -- Four ring test (9-f t diameter) -- gasket edging -- timber blocking

Tunnel Liner Plate (Test #3) -- Four ring test (9-f t diameter) -- gasket edging -- timber blocking

Test #1

Test #2

Test #3