design of underground plugs.pdf

International Journal o f Mining Engineering, 1983, 1, 189-228

Design of underground plugs F .A . AULD Cementation Mining Ltd, Bentley House, P.O. Box 22, Doncaster DN5 0BT, South Yorkshire, UK

Received 24 December 1982

Summary

The design of underground plugs is well documented for the gold mines of South Africa where reasonably hard rock and relatively high water pressures are experienced at deep levels. However, very little new information has been forthcoming in the last two decades, and published design data concerning other situations in softer rocks and with lower imposed hydrostatic pressures is virtually non-existent. This paper therefore sets out to review underground plug design with the object of bringing the subject to prominence and more up to date. An attempt has also been made to rationalize the design process in relation to current practice.

Six sections are included in the paper, the various types of plug being described at the beginning. The factors to be considered in plug design form the basis for discussion in the second section and design calculations are detailed in the third. Construction aspects follow while plug sealing and resistance to leakage are the topics included as the fifth section.

To elucidate the contents of the previous five sections more fully, the last section comprises three current case studies of actual plugs. Based on the overall concepts contained in the paper, conclusions and recommendations for plug design are formulated at the end.

Key words: Mining; mine water; water inrush; underground dams; concrete

Introduction

To sink mine shafts and drive-inclined drifts, underground roadways or tunnels successfully, experience and skill is needed to maintain excavation stability and to deal with and control ground water: The presence of the latter is possibly the most serious threat to working in the underground environment and the miner must always operate with care when approaching known zones of water-bearing strata.

During development work in shafts and tunnels, techniques are available whereby strata water can be controlled temporari ly prior to installing a water-tight lining. Such methods are: pumping, where the amount is not excessive; pre-grouting of the strata for reducing water makes to within the available pumping capacity; and freezing, if excessive amounts are expected.

Before commencing development work hydrogeological boreholes are normally drilled from the surface to locate the water-bearing zones approximately. Pressure recovery tests are carried out within the boreholes to provide data for estimating water inflow quantities

0263-4546/83 $03.00 + .12 1983 Chapman and Hall Ltd.

MaOvielogo ice

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which could be expected during excavation. Subsequently forward-probe drilling is carried out prior to each section of excavation to locate the water exactly.

Such procedures allow development to take place safely irrespective of the presence of water. However, it is not always possible, or economic, to provide fully water-tight linings for shafts and tunnels and, throughout the life of the underground system, ground relax- ation and stress readjustment may allow further ingress of ground water.

Accidental inrushes of large quantities of water are also a potential hazard if mining takes place too close to undetected sources and ground instability occurs or if drilling interconnects with unexpected water-bearing zones. Therefore it can be seen that, in many cases, water will be prevalent in underground workings, whether it is expected or unexpected, and the means must be provided for sealing off areas of the workings either for temporary water control while pumping to disposal or on a permanent basis. Plugs of concrete with a designed specific length and which fill the shaft or tunnel cross-section are used for this type of sealing.

The design of underground plugs is well documented for the gold mines of South Africa where reasonably hard rock and relatively high water pressures are experienced at deep levels (Garrett and Campbell Pitt, 1958, 1961; Lancaster, 1964). However, very little new information has been forthcoming since 1964 and published design data concerning other situations in softer rocks and with lower imposed hydrostatic pressures is virtually non- existent.

This paper therefore sets out to review underground plug design with the object of bringing the subject to prominence and more up to date. An attempt has also been made to rationalize the design process in relation to current practice. Six sections are included in the paper. The following five sections are: a description of various types of plug; a discussion of the factors to be considered in plug design; detailed design calculations; construction aspects; and plug sealing and resistance to leakage.

To elucidate the contents of these five sections more fully, the sixth section comprises three current case studies of actual plugs. Based on the overall concepts contained in the paper, conclusions and recommendations for plug design are formulated at the end of the paper.

Types of plug

Four different categories of underground plugs can be defined: (1) Precautionary plugs; (2) control plugs; (3) emergency plugs; and (4) temporary or consolidation plugs.

Basic descriptions follow outlining the functions of each type.

Precautionary plugs

These plugs are normally constructed in underground roadways to limit the area of flooding should water inrushes occur. Watertight doors are built into them which can be shut when any danger of flooding arises. Precautionary plugs are installed as a safety

Design: of underground plugs 191

measure prior to development in areas known to be potential water-bearing zones and such plugs are designed to withstand full hydrostatic pressure from surface level.

Control plugs

Sealing off or controlling the inflow of water from abandoned mining areas involves the introduction of control plugs. Plugs constructed in boundary pillars between adjacent mines also fall into this category. They are referred to as boundary plugs and serve to prevent water flowing from abandoned areas of one mine into the workings of an adjacent mine.

No means of access to the sealed off areas is provided through control plugs but normally drain pipes, with valves, are cast into them. These plugs are designed to resist full hydrostatic pressure from surface level or the pressure imposed by the head of water to the highest overflow point.

Emergency plugs

Plugs of this type are constructed to seal off unexpected inrushes of water either temporarily or permanently. No means of access to the sealed-off areas is provided in such plugs and they are usually designed to withstand full hydrostatic pressure from surface level.

Temporary or consolidation plugs

Plugs which allow inflow water to be controlled or stopped while simultaneously providing the resistance for high pressure grouting and consolidation operations are known as temporary or consolidation plugs. They are normally removed after the water pressure zones are sealed. Full hydrostatic pressure from surface level may again be the dominant design parameter for these plugs.

Factors to be considered in the design of plugs

When designing underground plugs the following factors need to be considered: (1) the purpose for which the plug is to be constructed; (2) the type of excavation in which the plug is to be installed (shaft or tunnel); (3) where the plug is to be sited in relation to the prevailing rock and working conditions; (4) plug shape; (5) head of water to be withstood by the plug; (6) the condition of, and the stress in, the rock surrounding the plug; (7) the strength of, and stresses in, the material of the plug; and (8) the method of plug construction.

Purpose for which the plug is to be constructed

Each of the four categories of plug described above has a different specific function and the form of a particular plug will be dependent upon the prevailing situation.

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Type of excavation in which the plug is to be installed (tunnel or shaft)

Undisturbed ground stress conditions alter locally in the areas surrounding an excavation. The adjusted stresses differ depending upon whether the excavation is for a vertical shaft or a horizontal tunnel. A more uniformly distributed stress occurs around a shaft excavation whereas, for a tunnel, the vertical ground pressure may be different from the horizontal causing stress variations around the perimeter. In highly stressed ground a fracture zone may surround the excavation and its extent will also depend upon whether it encompasses a shaft or a tunnel. Therefore the installation of a plug in a shaft will require different design considerations than for construction in a tunnel.

Where the plug is to be sited in relation to the prevailing rock and working conditions

One of the most important factors in deciding where to place a plug is the condition of the surrounding rock. Preferably the ground should be free from geological disturbances which could provide leakage paths for water. However, there could be limitations for the choice of site and the presence of faults or dykes in the immediate vicinity may have to be accommo- dated.

It is not advisable to site plugs in or near the fracture zones of highly stressed ground resulting from mining excavations although it is probably impossible always to avoid such situations. Control plugs may have to be located near mined out areas to restrict outflow of water and the distance ofboundary plugs from the workings depends upon the width of the boundary pillars in which they are installed. Boundary plugs need careful inspection at all times as boundary pillars which are too thin will be pervious to water and the danger of plug failure could be present under high hydrostatic pressures.

Plugs should be sited in ground which is not likely to be affected by subsequent ground movements resulting from mining operations. Damage to both the plug and the surrounding strata would annul the grout sealing integrity and introduce fresh leakage paths. The prevailing working conditions could also influence the choice of plug site. When there is an inrush of water, depending on the amount of water flowing into the workings, preference will be shown for a site which can be temporarily dammed upstream providing relatively dry construction conditions for the plug.

Ventilation would be another criteria to be considered particularly for an underground environment where high temperatures prevail. However, in an emergency, an adequately ventilated site may not necessarily be forthcoming.

Plug shape

Three basic forms of solid plug can be considered (Fig. 1). The first consists of a thin reinforced concrete wall (Fig. la) or unreinforced arch (Fig. lb) keyed into the excavation all around the perimeter in contact with the ground. Design of the slab involves calculation of bending moments, shear forces and axial forces, sufficient strength being incorporated in the structure to resist the applied pressure. The amount of keyed-in area is related to the bearing resistance of the surrounding ground. A solid plug of the second type possesses a longer

Design of underground plugs

Possible water {eokage paths through f - - ~ ~ strata

o i Water

, ~ ~ pressure I

PossLbte water LeaKage paths through strata

(a)

u ~ ~

Io:.' : :1 . . . . . . .

(e)

Possible water keakage paths through strata

I,ol /, :,/ .......

Posslb{e water Leakage paths through strata

(b)

w , # o o a e

' D Water

I' ' " ";!1

(d)

193

Steel bulkhead door

pressure

Stee~ {oo.d transfer cylinder

(e) (f)

Steel butkheod door \

/ A \ \ "x- / / A \ ~" / / ~-x \X / ~-,\

Steel Load tronsfer cyhnder

Fig. 1. Plug shapes. (a) Reinforced concrete slab in rectangular opening (adequate strength but insufficient leakage resistance). (b) Unreinforced concrete arch in rectangular opening (adequate strength but insufficient leakage resistance). (c) Unreinforced concrete tapered plug in rectangular opening (adequate strength and leakage resistance but uneconomical). (d) Unreinforced concrete parallel plug in rectangular opening (economical with adequate strength and leakage resistance). (e) Unreinforced concrete cylindrical parallel plug, with man access, in circular opening. (f) Unreinforced concrete cylindrical parallel plug, with roadway access, in circular opening.

length, no reinforcement and incorporates a taper to provide the ground bearing area (Fig. lc). Parallel plugs are the third type (Fig. ld) and resistance to the applied end hydrostatic pressure is achieved through mechanical interlock with the rough excavation face of the surrounding rock.

Garrett and Campbell Pitt (1958, 1961) consider plug length to be governed more by leakage resistance around the sides and through the surrounding rock than by structural

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strength. The longer length required for leakage sealing also ensures low shearing or bearing stresses at the concrete to rock interface. Thin barriers, although economic on materials, have very short, unsealable leakage paths at their extremities so are not suitable for underground plugs. Tapered plugs, when compared with parallel plugs, require more rock excavation which introduces further rock destressing, extra construction time and added cost. Increased quantities of concrete are involved and tapered plugs are subjected to larger pressures resulting from the greater projected end area at the maximum cross-section dimensions. Such factors are a disadvantage when plugs are required to be installed under emergency conditions. Although the leakage resistance paths are adequate with tapered plugs the other factors are prohibitive.

Garrett and Campbell Pitt (1958, 1961) have reported the results of tests in South Africa on an experimental plug at West Driefontein, on plugs constructed at West Driefontein and on a Virginia/Merriespruit boundary plug which show conclusively that parallel but rough- sided excavations will retain a plug without any sign of failure under very heavy load conditions. On this basis, all further discussion on plugs in this paper is focussed predominantly on parallel plugs. The section 'plug length based on bearing strength of concrete or rock at the interface' (p. 197), does however contain tapered plug design theory. Solid plugs installed in shafts or tunnels will have a cross-section of the excavation shape in which they are constructed. Shaft plugs will generally be circular in cross-section whereas for drifts, roadways or tunnels the shape may be square, rectangular, D-shaped, circular or otherwise. For precautionary plugs with access ways through them, either purely for man entry (Fig. le) or roadway access for materials transportation (Fig. if), a different concept is required.

To resist high strata-grouting pressures, which are applied in the transverse direction for leakage-sealing purposes, only the circular shape provides adequate strength. Precautionary plugs with access through therefore need to be in the form of concrete cylinders with sufficient length for leakage resistance, adequate mechanical interlock automatically being provided. In plugs incorporating access roadways, the dimensions required for clearance govern the inner diameter while strength to resist radial grout pressure determines the wall thickness. These two criteria apply for part of the length in a plug which is only provided with man access, as a structural concrete wall can be incorporated integrally with the concrete cylinder at the upstream face (Fig. le). In this case the strength of the wall is adequate, the concrete cylinder acting as a sufficiently long sleeve to provide leakage resistance and mechanical interlock with the surrounding ground.

In addition to the concrete cylinder, two other steel components are necessary for the successful operation of a precautionary plug with an access way. One is the bulkhead door for sealing off the plug in an emergency and the other is a load transfer cylinder (Figs. le, f). The steel load transfer cylinder allows the bulkhead door pressure to be carried by the concrete cylindrical plug through bearing on the ring flanges. Enough flanges are provided to reduce the bearing stresses to permissible limits.

Head of water to be withstood by the plug

For the majority of plugs the design head of water will be that from ground surface to the level of plug installation. This should be taken as normal for all designs unless very clearly

Design of underground plugs 195

defined lower overflow levels are shown to exist below ground surface which produce heads of water that cannot under any circumstances be exceeded.

Condition of, and the stress in, the rock surrounding the plug

The successful sealing of water flow by the introduction of a plug depends on the capacity of the surrounding rock to prevent leakage. Any discontinuities in the strata will make the job of sealing off more difficult. Fissures of geological origin or fracture planes resulting from high ground stress could endanger plug performance and installation of plugs in such areas should be avoided wherever possible, as indicated above in the section on plug siting.

The type of rock in which a plug is constructed is also a very important factor in governing how well leakage paths can be sealed or how efficient the shearing resistance or bearing capacity will be along the concrete to rock interface. The presence of weak beds of shale, clay, sandstones or conglomerates will increase leakage potential and reduce interface shearing resistance or bearing capacity.

As indicated previously, undisturbed ground stresses alter once excavation takes place and the magnitude and variation of such stresses around an opening differ for shafts and tunnels. High ground stresses, which cause rock fracture, depend upon the following factors: (1) the depth below the surface; (2) the size of the opening; (3) the proximity of other mining excavations; and (4) the proximity of geological disturbances which may introduce tectonic stresses.

The subject of stress evaluation around underground openings is a complex one and is too large a topic to be introduced into this paper. Nevertheless, it is a subject which must be fully understood if a true evaluation of concrete plug to rock interaction is to be formulated and more research into this area is required.

Strength of, and stresses in, the material of the plug

Five points warrant consideration when evaluating the stresses in, and strength of, underground plugs: (1) concrete compressive strength; (2) the early age development of strength; (3) the shear or bearing stress at the plug to rock interface; (4) the pore water pressure in the concrete; and (5) the possible end spaUing of the plug due to high stresses set up by ground pressure.

Provided the recommendations of current Codes of Practice for structural concrete (in the UK, British Standards Institution, 1969b, 1972, 1976) are followed, with Grade 25 concrete (characteristic strength 25 N mm -2) as the minimum specified requirement, then dense, impermeable and durable concrete ought to be achieved easily. On this basis, in conjunction with the length required for sealing which ensures low stresses, no problems of strength should be encountered.

Early age strength development is important from two aspects. First it is essential that plugs develop their specified strength without any detrimental effects occuring from shrinkage, thermal changes or ground pressures. Provided care is taken to overcome these factors, then the integrity of the concrete mass will be protected and leakage paths through plugs minimized. The second aspect of importance in relation to early age strength develop-

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ment is how quickly a plug needs to be sealed. It is possible to use higher strength concrete mixes than are required purely for design strength. This allows higher strengths to be achieved at earlier ages and hence the problem of sealing can be tackled more quickly.

The factor of safety against shearing or bearing failure in the rock or concrete of the plug at their interface depends upon the magnitude of the induced stresses, which in turn is related to plug length. Since the length of a plug should be determined with leakage sealing in mind, which means providing a longer length than is necessary for structural strength purposes, relatively low interface stresses are inherent in good plug design. Knowledge of pore-water pressure behaviour within a concrete plug is limited. A pressure gradient exists from the hydrostatic pressure at the face in contact with the impounded water to zero at the opposite end. How the pressure and induced stresses are dissipated throughout the system and into the surrounding ground is a matter for conjecture at the present time and this area, in conjunction with rock stress evaluation, needs further research. However, it is unlikely that spalling of the free face of a plug will occur due to pore water pressure unless nonhomogeneous irregularities occur in the concrete mass.

It is conceivable that high localized ground stresses at the ends of plugs could cause spalling at these points, reducing the effective resistance to applied pressure and leakage. Careful choice of site related to a study of the induced rock stresses and rQck strength can avoid or minimize this risk. Additional plug length would also contribute to solving this problem.

Method of plug construction

For precautionary, control and temporary or consolidation plugs, which can generally be constructed in phase with and under normal mine-operating conditions in a relatively dry environment, the method of construction has little influence on design. However, in conditions of emergency, materials access problems and water inflow quantities may require consideration of different methods of construction. Normal concrete transportation, placing and compaction can be replaced by grouted concrete in which a mixture of cement, sand and water is introduced into pre-placed aggregate. This technique is particularly suitable for the construction of plugs in areas where access is difficult or for plugs installed under water in flooded shafts. Concrete can also be placed by tremie under water if necessary. Resulting from the chosen method of construction, different concrete to rock interface allowable shear or bearing stresses may have to be used depending upon how dense and impermeable the plug mass is expected to be and how integral a contact can be achieved with the surrounding rock.

Design calculations

Formulae for calculating plug length and strength

Plug length based on shear strength of concrete or rock at the interface. Garrett and Campbell Pitt (1961) quote the following formula which can be applied to parallel-sided plugs with rectangular cross-sections if interface shearing is accepted as the governing failure


mechanism

pbh = 2(b + h)lpp~ (la)

wherep is the intensity of appliedpressure; b is the width of the plug; h is the height of the plug; l is the length of the plug; and Poe is the permissible punching shear stress of the rock or concrete at the interface.

By transposing Equation 1, the length of plug can be obtained

l - pbh 2(b + h)pp~" (lb)

For a square cross-section, Equation lb becomes

l = (pb)/(4pp~). (lc)

The length of circular plugs, of radius r, can be obtained from

pTrr 2 = 27rrlpp~ (2a)

giving

l = (pr)/(2ppe) (2b)

Plug length based on bearing strength o f concrete or rock at the interface. Although the shear strength concept of the previous section can be employed, Garrett and Campbell Pitt (1961) also considered that, alternatively, the mechanism of interaction between concrete plugs and the surrounding rock could be more in the form of direct bearing rather than shearing at the interface. Mechanical interlocking action is achieved at an excavation face through the various inclined planes of its surface. Orientation of these planes can be in any direction lying between the extremes parallel with or normal to the general direction of the excavation face. An assumption can be made that half of the inclined planes resist movement by direct bearing (Fig. 2a) while the others are subjected to tensile stresses and therefore can be neglected.

For a parallel plug, consider an element of the excavation face ABC (Fig. 2b) with a horizontal length AC = l', which contributes to the plug bearing resistance over the element of length 1'/2. The permissible bearing stress in the concrete or the rock ispb e and Fb represents the total bearing resistance over the element of plug bearing length BC, inclined at an angle of ~ to AC such that

BC cos ~ = l'/2. (3)

In the triangle of forces (Fig. 2b) P' is the element of applied horizontal force which is resisted by the horizontally resolved component of Fb. Therefore

!

P' = F b sin ~ (4) however

Fb = PbeBC. (5)

198

4 t

[engfl~ =

(a)

AUld

. / n

F j Z Compression component

(b)

Fig. 2. Evaluation of parallel plug length based on bearing strength of concrete or rock at the interface (Garrett and Campbell Pitt, 1961). (a) Plug bearing resistance. (b) Element of plug bearing resistance.

From Equation 3

BC = I'/(2 cos ~)

and combining Equations 4, 5 and 6 gives

e' = (pbj ' /2)tan ~.

Summing all the forces on the plug results in

l' l ]~e' = e =pbh = Xpbe~ tan~ = 2(b + h) ~Pb, tan

(6)

(7)

(8a)


from which

pbh (8b) l = (b +h)pb~ tan u

Since the surface planes will be inclined at angles of between 0 and 90 to the direction of thrust, Garrett and Campbell Pitt (1961) considered the assumption that the average inclination a = 45 for a parallel-sided plug was justified. Equation 8b becomes

l - pbh (8c) (b + h)Pbe "

For a square cross-section, Equation 8c reduces to

l = pb/2pbe (8d)

The length of circular parallel plugs can be obtained from

l pnr 2 = 2nr ~pbe tan e (9a)

giving

l = pr/Pbe (9b)

for a = 45 . Tapered plugs can also be considered if appropriate amendments are made to Equations 3-9 (Fig. 3).

The element of bearing length BC (Fig. 3b) is now inclined at an angle a + fl to the horizontal and

BC cos ~ = 1'/2 cos fl (10)

where,/? is the angle of plug taper. From the triangle of forces (Fig. 3b)

P' = Fb sin(a + /?) (11)

however

l' Fb = Pbe BC = Pbe 2 COSe COSfl " (12)

Combining Equations 11 and 12 gives

l' sin(e + //) l' P' = Pbe 2 COSe COS/? -- Pbe ~ (tan e + tan/?). (13)

Summing all the forces on the plug results in

l' ~,P' P pbmaxhmax Pbe ~ (tan c~ + tan fl) 2(bay + l . . . . . h.v)~pbe(tan ~ + tan fl) (14a)

where: bma is the maximum plug width at the water face; hma is the maximum plug height at

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[

= Water presiure end tlrea

= Toto l fo rce (P)

- - - .~- - ~._=_--'_--'~ 2 ~"

// ~ Total effective reslstonce : t T (I -tan(~ton,~} J Length z

(a)

Y J

J L . . . . . . . . .

L ~

F~ [ompresslon component

Pbe/

~'// .I y(1 - tanc, t(ln~)

(b)

Fig. 3. Evaluation of tapered plug length based on bearing strength of concrete or rock at the interface (developed from Garrett and Campbell Pitt, 1961). (a) Plug bearing resistance. (b) Element of plug bearing resistance.

the water face; bav is the average plug width along its length; and hay is the average plug height along its length.

From Equation 14a

Pbmaxhm~x l = (bay + hav)Pbe(tan ~ + tan fl) " (14b)

For ~ = 45 , Equation 14b becomes


pbm~hmax (14c) l= (bav + ha0Pbe(1 + tan 8)

and for a square section

l= pb2max 2b.vpb~(1 + tan fl)" (14d)

The length of circular tapered plugs can be obtained from

2 rr(rma x q- rmin) -t- (rma x -- rmin) Pbe(tan at + tan 8) (15a) p~rma x =

where rm~ x is the maximum plug radius at the water face and rmin is the minimum plug radius at the face remote from water. Equation 15a gives

p rmax - (rmax- rmin) (15b) l = 2 (rmax + rmin)2pb2e(1 + tan/~)2

for ct = 45 . An alternative form of bearing calculation for tapered plugs is that for a smooth-faced

wedge driven into an opening. On this basis the whole surface area acts in bearing and the element of bearing length becomes AC (Fig. 3b), inclined at an angle/? to the horizontal where

AC cos/3 = l' (16)

and

P' = F b sin fl

however

F~ = PbeAC = pbel'/cos 8.

Combining Equations 17 and 18 gives

P' = PbJ' tan fl

Summing all the forces on the plug results in:

~,P' -- P = Pbmaxhm~x = ~'PbJ' tan fl = 2(bay + hav)lpb e tan fl

from which

1 = pbm~xhmax 2(bav + hav)Pbe tan/~

is derived. For a square section

z .......... pb ax 4bavpb e tan/~ "

(17)

(18)

(19)

(20a)

(20b)

(20c)

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Comparing Equations 20b and 20c with Equations 14c and 14d, respectively, if a = 45 is replaced by ~ = 0 in the latter two equations then compatibility is achieved except for the anomaly of reducing the length by half in the case of the wedge theory due to using the full bearing area.

The equivalent length of circular tapered plugs based on the smooth-wedge principle can be obtained from

pnrZmax = 7r(rma x + rmin)[l 2 + (rma x -- rmin)2J1/2pbe tan/J

where

(21a)

I 2 4 2] 1/2 P rrnax l "~- (rma x + rminX2-2e)Pb tan /~2 - (/'max -- rrnin) (21b)

Cylindrical plug strength. The strength of cylindrical parallel plugs (Figs. le,f) can be determined using the standard Lam6 elastic design theory for thick cylinders (Auld, 1979, 1982a)

2pr(t + ri) 2 _< tr = t(t + 2r3 ~Pc (22)

where a is the maximum tangential stress in the concrete cyclinder wall, occurring at the inside face; Pc is the permissible concrete compression stress; Pr is the externally applied radial pressure; ri is the inside radius of the cylinder; and t is the concrete cylinder wall thickness.

Bearing strength o f cylinder walls. Plugs of the type shown in Fig. le, which carry load from a circular face wall back through a cylindrical rear section and thence into the surrounding rock, must have sufficient strength at the interconnection between the wall and cylinder. The cylinder end area must be sufficiently large to reduce the bearing stress imposed by the end wall to a value within the permissible limit. Hence, the calculated concrete bearing stress,

P ~< Pb (23a) fb - n(r2o _ r~)

where Pb is the permissible concrete bearing stress; and ro is the outside radius of the cylinder, and P is the horizontal applied force on the cylindrical rear section.

i.e. P = pnr2o - [the concrete or rock permissible surface resistance over length l* of the front wall]. (23b)

Combined stress at the interconnection between the face wall and the cylindrical rear section. The cylinder stress and the bearing stress determined from the above sections act together at the interconnection to produce a combined compressive stress situation. Care should be

Design o f underground plugs 203

taken to ensure that the calculated combined compression and bearing stress,

f~c = (a 2 + f~)1/2

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Direct bearing of end load on cylinder wall. The bearing area of the cylinder wall must be capable of transferring the end load from the bulkhead door to the bearing flanges without overstressing, i.e. steel bearing stress,

where Pbs is the permissible steel bearing stress.

Resistance to axial compression. In addition to end bearing, the cylindrical wall must be capable of acting as a column between the flanges to allow flange bearing to be effective. Timoshenko and Gere (1961) give the critical load Pcr for a column with built-in ends, which is a conservative approach for a cylinder wall as it neglects additional strength due to curvature, as

where E is the modulus of elasticity for steel (2.1 1011 Pa); and I is the moment of inertia of the cylinder wall equal to t3/12 per unit length.

For the cylinder, multiplying Per by the circumference and dividing by the total applied end pressure will produce the factor of safety of

Provided the flanges are not spaced too far apart, satisfying the criterion of the section on 'Compression resistance to radial grouting pressure' will automatically produce a large factor of safety in Equation 32.

Bulkhead door design

Generally bulkhead door pressures will be relatively large and therefore the best shape to resist the load is a spherical segment. The shell thickness for such doors can be determined on the basis that the meridional and hoop forces per unit length of the shell are equal to pa/2 (F1/igge, 1967) Where a is the radius of curvature of the shell. Dividing this value by the shell thickness t gives the compressive stress in the steel, fc~, as

It should be noted that thin shell domes are prone to buckling, and stiffening for the door should be provided to avoid any possibility of instability under load.

The subject of steel bulkhead door design lies in the specialist field of pressure vessels and is outside the scope of this paper. Operation and sealing of such doors are prime parameters to be considered in design and recourse should be made to specialist design and fabrication manufacturers for the supply of such .elements.


Rock, concrete and steel permissible stresses

Permissible shear and bearing stresses for rock and concrete at the plug interface. The proposed formulae for determining the length of plugs, either on the basis of shear strength or on one of the two bearing strength philosophies are a very simplified form of a much more complex stress system. Both the rock and the concrete are in a confined state along their interface. The compression strength of concrete in the UK is quoted on the basis of 150 mm cubes tested at 28 days in an unconfined compression testing machine (British Standards Institution, 1970). It is known that concrete when tested in a confined state shows an increase in strength over the unconfined condition (Jaeger and Cook, 1979). The confining action of the surrounding rock against the concrete plug could modify the bearing force calculated by the formulae in the section entitled 'Plug length based on bearing strength of concrete or rock at the interface'. However, the true resistance probably lies somewhere between that given by the bearing capacity and the resistance provided by shear. Shearing in this context would be of a punching nature, as opposed to the traditional structural engineering form of beam shear, and even this could be modified depending upon the magnitude of the interface confining stresses. Hence the ultimate validity of the permissible stress values for shearing and bearing will depend upon the effectiveness by which the concrete of the plug is confined by the surrounding rock.

The plug concrete can be considered as a homogeneous material on the assumption that good construction practice has been observed. However, the surrounding rock will be any- thing but homogeneous, being cracked and fissured before excavation takes place. De- stressing also occurs during and subsequent to excavation and therefore, when grouting and hydrostatic pressures are applied to the rock, movement inevitably will occur. The direct strains will be accompanied by movement in the direction of cracks and bedding planes and the effectiveness of the confining action will be dependent on this movement.

Irrespective of what theory is applied to define the stress conditions, the governing factor remains the stress in the rock. As indicated previously in this paper more research is needed to understand how the stresses in the surrounding rock are modified by confined plugs subjected to end pressure. Until this aspect is investigated in detail the validity of any formulae utilized in defining plug and rock stress conditions will be in question. At the present time, with the formulae available, it will be necessary to check the shear and bearing stresses for both the concrete and the rock and to base the design on the weaker material.

Concrete permissible stresses are contained in Table 1 based on the current UK Codes of Practice (British Standards Institution, 1969b, 1972). The values are all related to the concrete characteristic strength, this being the lower limit below which not more than 5% of the cube test results would fall based on a statistical analysis of samples tested. Both of the Codes of Practice are specifically for reinforced concrete and neither treats the unreinforced concrete situation realistically, particularly with regard to punching shear philosophy. How- ever, Manning (1961) quotes the safe punching shear stress to be about one-fifth of the safe compressive stress and this has been included in Table 1. The maximum allowable values for pc, Pb, Pbe, Pp and ppe are heavily outlined in Table 1 as these are the suggested values to be adopted in design. The reason for using a factor of safety equal to 4 for Pbe and Ppe is explained later.

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It is much more difficult to propose realistic permissible stresses for rock. The strengths of rocks are normally determined by testing cylindrical samples and, as a result of the non- homogeneity of the material, normally it is only the best pieces from which the samples are obtained. It must always be remembered that strengths of rocks which are determined from such testing will not be typical of the actual strength in situ and appropriate adjustments should be made to allow for this.

Assuming the grouting process for strata water sealing is carried out methodically and conscientiously, most of the rock bedding planes and fissures local to the plug interface should be filled and consolidated. This, allied with the confining action of the surrounding rock, could allow the lower strength concrete permissible stresses to be taken as being representative of the rock also. For the purposes of design, this would be an alternative if no actual data was forthcoming. The practice in South Africa is to use a permissible shear stress value of 0.59 N mm -2 (85 lb in -z) for concrete placed in the normal manner and 0.83 N mm -2 (120 lb in -z) for grouted concrete where positive contact between the concrete and the surrounding rock is assured by subsequent grouting (Lancaster, 1964). This is a general rule and is not related specifically to concrete or rock strength neither does it take into account the rock condition. The values are therefore unrealistic, particularly with regard to the increased concrete strengths currently being achieved in underground construction due to the improved workability and quality control procedures adopted in conjunction with better batching, transportation and placing techniques. Therefore it is considered that the values in Table 1 are more appropriate.

Permissible concrete stresses other than at the plug interface with the rock. These are also covered by Table 1.

Permissible steel stresses for load transfer cylinders and bulkhead doors. Typical permissible steel stress values for steel (Grade 43) are contained in Table 2. These are taken from the current UK Code of Practice for the use of structural steel in building (British Standards Institution, 1969a).

Table 2. Steel (Grade 43) permissible stresses (BS449: Part 2:1969; British Standards Institution, 1969)

Bending, Axial Type of Pms compression, Bearing, stress (N mm -2) Pcs (N mm -z) Pbs (N mm -2)

Up to and including 40 mm thick 165 155 190

Over 40 mm thick 150 140 190

208 Auld

Factors of safety

The factors of safety for structural concrete quoted by the UK Codes of Practice (British Standards Institution, 1969b, 1972) are given in Table 1. CP114:1969 introduces a value of 2.73 to relate the characteristic strength to the permissible compression stress in bendingpc.

CPl10:Part 1:1972 is more specific in its breakdown of safety factor. The actual compression strength of concrete in a structure is equivalent to 0.85 characteristic cylinder strength (Comit6 Europ6en du Beton, 1970), where for plug construction the 0.85 factor takes account of the difference between instantaneous loads on cylinders at an age of 28 days and loads applied for a longer duration on specimens of the same age. Since British Standard practice (BS 1881:Part 4:1970) uses the 28 day cube test as a means of strength control, a correction factor of 0.8 is needed to convert the cube strength to the equivalent cylinder strength. Hence, in relation to the characteristic cube strength, fcu, the actual in situ strength of concrete is represented by 0.68 characteristic cube strength, the value of 0.68 being equal to 0.85 0.8. A figure of 0.67 is employed in CPl l0:Part 1:1972.

Partial safety factors for load rf and strength ~m are used in the ultimate limit state approach to the design of concrete structures which was adopted for CP110:Part 1:1972. On the basis of it normally being of a long-term permanent nature, the value of yf for hydrostatic loading can be taken as 1.4. For Ym, which is introduced to account for possible strength differences between test specimens and the actual structure caused by such aspects as insufficient compaction and differences in curing, the specified value is 1.5. The effective factor of safety in accordance with CPll0:Part 1:1972 is therefore 1.4 z 1.5 = 2.1 when related to the actual strength of concrete in situ, or (1.4 z 1.5)/0.67 = 3.13 when compared against the characteristic strength as given by 28 day cube test results.

For the steel stresses in Table 2, the factor of safety to yield will be approximately 1.5 with probably the same again to failure. This gives a probable minimum factor of safety to failure equal to 2.25. The factors of safety for concrete and steel which have been built in to the permissible stresses, in the 2-3 region, are acceptable because the performance of the material under load is well established and quality control ensures consistency. For the mechanism of resistance at the plug to rock interface, the true behaviour is not understood fully and the rock shear and bearing permissible stresses cannot be established realistically. On this basis, and because plugs are normally installed as a safety measure, it would be prudent to adopt a higher safety factor when determining plug length using the shear or bearing resistance criteria. A minimum factor of safety of 4 is recommended in line with South African practice (Lancaster, 1964) and this has been introduced into Table 1 for the Pb~ and pp~ values.

Construction aspects

Batching, transporting and placing concrete

In any concrete construction work it is necessary to have the right batching plant, geared to the demand. This is particularly important for underground construction where pours must be completed with minimum interference from external factors. Rates of pouring are

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governed by the physical restrictions in particular areas underground and by the problem of access for materials to those areas. A surface batching plant is preferred because aggregate and cement weighing, together with the metering and addition of water and admixtures, can be controlled in the most effective manner. However, such a system depends on being able to transport the pre-mixed concrete underground and, in some circumstances, an underground batching plant may be necessary. This does not relieve the problem of having to transport the concrete mix constituents underground as separate items. For grouted concrete, again a surface grout mixing set-up would be preferred. Normally with surface batching plants large quantities of concrete or grout can be mixed and transported underground rapidly to give a constant uninterrupted supply. This is particularly advantageous for the construction of emergency plugs.

Current trends in UK mining development have favoured the employment of established ready-mixed concrete suppliers. By adopting such suppliers high quality is achieved through the utilization of their specialist expertise in the production of concrete. Quality control measures for the use of concrete underground are the same as those for surface works and are in accordance with the relevant Code of Practice (CP110:Part 1:1972). Independent approved organizations (British Ready Mixed Concrete Association, 1978) are normally employed for cube testing, or any other testing of the hardened concrete which is required.

The preferred method of transporting and placing concrete underground is pumping. With shaft access, concrete can be dropped down a vertical pipe for further transportation underground by pumps situated at the bottom of the shaft. Drift access allows pumping from the surface, down the incline and then further underground directly to the point of plug installation. Once pipelines are installed minimum interference with mining and plug construction operations is achieved and large volumes of concrete can be delivered and placed rapidly.

Concrete mix designs

Similar to any concrete construction, care must be taken to provide the right balance of ingredients, first of all to suit the particular mode of transportation and placing being used, and secondly to make sure that the optimum design is achieved. In addition to strength, the most important factor of the mix design is to obtain the correct workability. For underground work, with its restricted placing environment, it is essential that high workability mixes are used. In the author's opinion (Auld, 1982b,c), successful construction with concrete underground is entirely dependent upon the inclusion of plasticizing admixtures for transporting and placing.

Two mix designs which have been used successfully by Cementation Mining Limited are contained in Table 3. Cement replacement materials were incorporated to minimize the thermal effects which are discussed in more detail below.

Thermal and shrinkage effects

If it can be achieved, it is preferable to pour a concrete plug in one operation to avoid construction joints which are potential leakage paths through the plug itself. Although

210

Table 3. Cement replacement mixes previously used by Cementation Mining Ltd for underground plugs.

Auld

Emergency plug in roadway: Grade 30 (OPC replacement with PFA); 30 N mm -2

Site Supplier Total cementitious content Sand Sand % of total aggregate Coarse aggregate Water Water/cement ratio Slump without plasticizer Plasticizer Slump with plasticizer

British Gypsum Ltd, Sherburn Mine Topmix Ltd. 400 kg m -3 (250 kg m -30PC, 150 kg m -3 PFA) 770 kg m -3 (Elvaston Zone 2) 42% 1050 kg m -3 (Elvaston Gravel) 180 litres m -3 0.45 50 mm Flocrete N (Cementation Chemicals Ltd) 160 mm

Temporary consolidation plug in shaft: Grade 55 (OPC replacement with Cemsave ground granulated blast furnace slag); 55 N mm -2

Site Supplier Total cementitious content Sand Sand % of total aggregate Coarse aggregate Water Water/cement ratio Slump without plasticizer Plasticizer Slump with plasticizer

National Coal Board, North Selby No. 1 shaft Topmix Ltd 500 kg m -3 (30% OPC, 70% Cemsave) 595 kg m -3 (Blaxton Zone 3) 34% 1150 kg m -3 (Blaxton Gravel) 180 litres m -3 0.36 60 mm Flocrete N (Cementation Chemicals Ltd) 160 mm

normal structural concrete mixes (CP110:Part 1:1972) could be used for plug construction, the large volumes required for mass filling can be subjected to detrimental thermal effects during setting. This is dependent upon the amount of cement included in the mix. Internal build-up of heat within the mass due to the cement hydration process could induce high thermal stresses. The strength integrity of the structure would be impaired and, on cooling, thermal cracking may result. By cooling the aggregates and mixing water the ultimate temperature attained by normal mixes can be reduced but it is preferable where possible, to use cement replacement materials to minimize heat of hydration gain. Table 3 contains two such mixes.

An additional factor which assists concrete thermal control is the embedment of service, water control and grouting pipes in the plug mass. Heat will be dissipated through these pipes, particularly in the case of temporary or consolidation and emergency plugs if water is flowing through them. Shrinkage will not normally be a problem with the designed concrete mixes currently being employed in underground construction. The use of plasticizers


enables the water/cement ratio of the basic mix to be kept to a minimum therefore ensuring very little water loss during the curing stage which prevents excessive shrinkage. Three other factors also contribute to shrinkage reduction. These are the underground environment, in which no rapid drying out conditions normally prevail, the limited facial exposure to drying elements in the environment and the relatively thick concrete sections used.

Construction points of detail

Excavation. Care should be taken during excavation to minimize damage to the surrounding strata. Machine cutting and hand trimming is preferred to drilling and blasting.

Plug installation. Two factors assist in reducing leakage paths at the concrete to rock interface. Before commencing to pour concrete for a plug, the floor should be thoroughly cleaned to remove any debris or construction dust. At the roof of the plug, to ensure a tight seal, concrete must be discharged as high up as possible and a crown feed pipe, which can be withdrawn as topping up takes place from one end of the plug to the other, should preferably be installed. Air bleed pipes, which subsequently can be used for contact zone grouting, are also beneficial at roof level.

Grout seals. Where mass concrete is cast directly against rock, it is necessary to grout up the contact zone to prevent leakage through any shrinkage gaps. It is very difficult to obtain full tig]~t contact with the surrounding rock over a large surface area with grouting. There- fore, it is preferable to also provide one or more narrow chases, surrounding the plug cross-section completely, in which grout can be injected and pressurized to provide a tight ring seal.

Temporary water control. For temporary or consolidation plugs and emergency plugs, contro]l of shaft water is essential to allow good construction. In roadways, flood water will need to be temporarily dammed upstream and the water led off through valved pipes cast into the plug. Debris grills will need to be installed at the upstream ends of pipes. Temporary consolidation plugs in shafts also need to include vertical steel rising mains through which shaft inflow water can be withdrawn during construction, to prevent pressurizing of the underside of the plug during hardening.

Services. Pipes need to be installed in precautionary plugs to carry services. These pipes should be fitted with sealing glands at each end for plug watertightness when the bulkhead doors need to be closed.

Plug sealing and resistance to leakage

Grouting procedure

Design calculations can be carried out as shown previously to determine plug dimensions. However, an integral part of the successful installation of an underground plug is the means

212 Auld

by which leakage past the plug is minimized or eliminated. Grouting is the process by which this is achieved.

The science or 'art' of grouting depends very much upon the knowledge and experience of mining development contractors, and cannot be discussed in detail here. As a process, grouting consists of the pressurized injection of cement or chemical grouts into the strata to fill voids, fissures, bedding planes and any other anomolies in the rock surrounding a plug. Its purpose is to seal off all water paths and grouting of the plug itself may be needed depending upon whether construction joints are incorporated and also upon the standard of workmanship.

Injection of grout at the contact surfaces between the plug and the rock is also necessary to fill shrinkage gaps, porous zones due to placing difficulties and cracks in the rock adjacent to the plug due to destressing. Pressures of up to twice (Garrett and Campbell Pitt, 1958) and two and a half times (Garrett and Campbell Pitt, 1961; Lancaster, 1964) the pressure which the plug has to resist have been recommended for this grouting. These pressures are used in the deep gold mines of South Africa where generally strong rocks and relatively high water pressures are encountered. Even with localized fracture zones around such excavations, opening up of the cracks under high pressures to accommodate the entry of grout is not detrimental. However, in softer rocks at shallower depths as in the UK coal measures, such pressures would be damaging and are not to be recommended. Precautionary plugs of the cylindrical type should only be stressed to a maximum of one and a quarter times the hydrostatic pressure, related to surface level, this being the value by which the normal structural concrete permissible stresses can be exceeded for short term loading (CP114:1969). Hence the post-stressing of the plug and rock, which is advocated for the South African conditions (Garrett and Campbell Pitt, 1958) will generally not be as effective in UK practice for enhancing the confining action. The radial Poisson's ratio effect, resulting from the end pressure, will also be less effective with regard to increasing the interlocking resistance.

Leakage associated with plugs can occur at the following places: (1) through the plug concrete; (2) along the concrete to rock interface; (3) through the rock surrounding the plug; and (4) along the interface between the plug concrete and the steel load transfer cylinder if access through the plug is provided.

Leakage through the plug concrete

Three possible reasons exist for leakage through the plug concrete: (1) porosity of the concrete; (2) construction joints; and (3) presence of cracks.

The presence of highly porous concrete is very unlikely due to the dense, impermeable and durable mixes currently used in underground construction. High workability, achieved with the use of plasticizers, ensures full compaction and with good quality control and careful placing techniques no excess porosity problems should occur. Grouting will help to seal off the more porous zones if they do occur.

Wherever possible construction joints should be avoided but, with the current mix designs, if they are necessary, very little joint preparation is required. Good bonding should be easily achieved between consecutive pours.


Cracks in the concrete can be caused by excessive water pressure behind the plug, thermal effects during setting and maturing or excessive stresses and strains transmitted to the concrete by the surrounding rock. Provided the design is processed correctly in relation to the applied water pressure; adequate measures are introduced in the mix design to minimize the thermal effects and the chosen plug site is competent, then a homogeneous structure can be constructed free from defects.

Leakage along the concrete to rock interface

Interface leakage could result from: (1) shrinkage gaps at the interface; (2) shear cracks at the interface due to plug movement under high water pressure; (3) cracks caused by excessive ground stress; (4) poor contact with the surrounding rock, caused by debris and construction dust not removed from the floor prior to casting and also as a result of air and water pockets trapped at the underside of the roof.

As discussed in the section on thermal and shrinkage effects, shrinkage in underground concrete should be minimal. The grouting process also enables gaps caused by shrinkage to be sealed up.

The possibility of plug movement will only occur if the plug length is too short and, as mentioned previously, the capacity to seal leaks is the prime factor in determining plug length. A longer length is needed to provide leakage resistance than is required for structural purposes. This will ensure that the interface shear stresses are sufficiently small to avoid any plug movement under high hydrostatic pressure.

The radial adhesion between the plug concrete and the rock at the interface will be small and highly stressed rock conditions could damage the intimate contact. Judicious choosing of the plug site could avoid such failure.

Good workmanship will prevent problems such as construction debris not being removed prior to casting the plug concrete. Provision of the correct concreting facilities should assist in attaining close contact with the roof.

Leakage through the rock surrounding the plug

Leakage through the strata can occur as a result of: (1) geological fissures or other discontinuities in the rock; and (2) cracks in the rock formed by ground stress or by strain from mining operations.

If possible, plugs should be sited away from faults in the rock. However, if they are unavoidable, the grouting process will help to seal and stabilize conditions.

The problem of rock failure under high stress is experienced when other mining excavations encroach too closely or at great depth where overburden pressures become excessive. Every effort should be made to avoid overstressed areas.

Determination of plug length required for sealing

The problem of leakage associated with underground plugs has been discussed previously on the basis of where it occurs and how it can be minimized or stopped by grouting. To

214 A uld

determine the plug length which conforms to the permissible punching shear and bearing stress values at the concrete to rock interface is relatively straightforward. Quantifying exactly the length which is required for a leakage free plug is much more difficult. Published data concerning the subject is scarce and what is available is related to specific ground conditions which cannot be applied on a general basis.

Garrett and Campbell Pitt (1958, 1961) published the results of tests on an experimental plug, 1.220 m (4 ft) square by 3.350 m (7 ft 8 in) long and situated in sound quartzite, at West Driefontein. The static water pressure was approximately 20.7 N mm -2 (3000 lb in-2). An extensive system of tapping points and holes in the rock were incorporated for studying leakage at the steel load transfer cylinder interface with the concrete, at the concrete to rock interface and through the strata. Leakage quantities were observed before grouting and after various stages of pressure grouting were completed. From the test results, Garrett and Campbell Pitt (1958) proposed certain concepts which are given below.

1. The resistance of a plug to passage of water either along its contact with rock or through the adjacent fractured rock depends on two factors, the length of the plug and the resistance of the rock to the passage of water.

2. The latter, being a condition of the rock which varies greatly with different types and mining conditions, can be regarded as the practical consideration for determining plug length.

3. The two factors can be interrelated using the pressure gradient through the rock as the linking medium.

Results from the West Driefontein test plug form the basis for the graphs contained in Fig. 4 which are reproduced from Garrett and Campbell Pitt (1958). These results only refer to the rock and pressure conditions described. The graphs are: (A) the minimum length of plug that would be required if the contact between plug and rock was ungrouted [p/l = 0.23 N mm -2 m -I (20.8 lb in -2 ft-1)]; (B) the minimum length when the contact is grouted but before the rock is grouted [p/l = 3.64 N mm -2 m -1 (161 lb in -2 ft-1)]. (C) The minimum length when normal grouting of the rock was 41.4 N mm -2 (6000 lb in -2) [p/l = 9.14 N mm -2 m -1 (404 lb in -2 ft-1)]. This is normal to South African practice, being twice the hydrostatic pressure. It is not normal to the UK. (D) This graph is similar to C but with the addition of chemicals to seal rock fissures. C is applicable in South Africa to a normally grouted plug but has no safety margin.

Garrett and Campbell Pitt (1958) suggested from this that plug length should be such that a leakage factor of safety should not be less than four and may be as much as ten. The choice depends on an assessment of many factors which include fracture of rock during excavation and subsequent destressing, porosity of the rock and its acceptance of grout. In Fig. 4, the graphs show plug lengths when factors of safety of 4, 6, 8 and 10 are applied. These depend on the plug to rock contact and the rock being grouted to at least the same pressure as that which the plug is designed to resist. Plug lengths for various square section sizes based on an Pbe value of 4.14 N m -2 (600 lb in -2) are also included on the basis of Equation 8d (shown by dotted lines). The value of 600 lb in -2 was used by Garrett and Campbell Pitt (1958).

As far as the author is aware this is the only published information which attempts to


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quantify directly leakage resistance in relation to plug length apart from records of past plugs which have been successful. It has been emphasized throughout this section that the data put forward relates only to the particular test conditions. This leaves the plug designer very much to his own initiative and experience for determining the plug length which will provide adequate sealing. Further research is therefore necessary into this area. However, if a plug i~ constructed with sufficient length to resist movement but cannot prevent leakage through the surrounding rock, the length can always be increased. In an emergency this may be important, because the plug could be constructed to prevent flooding and subsequently lengthened to reduce leakage.

Case studies

British Gypsum Ltd, Sherburn Mine, England, 1980 (Emergency plug)

In 1980 a pressure pad was constructed by British Gypsum Ltd (see Section A-A of Fig. 6) in an ',attempt to seal off water inflow into the area of the pump sump (1 East 5 South in

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Fig. 5. British Gypsum Ltd, Sherburn Mine, England. Underground layout showing position of emergency plug. Information reproduced by permission of British Gypsum Ltd.

Fig. 6. Emergency plug at British Gypsum Ltd, Sherburn Mine, England, 1980. General arrangement p Jan (a) and sections A-A(b), B-B(c) and C-C(d) showing proposed concreting sequence. Design of plug and grouting scheme for sealing leakage by Cementation Mining Ltd who also provided construction assistance to the mine and carried out the sealing. Information published by permission of British Gypsum Ltd.

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218 A uld

Fig. 5). The general dip of the strata was from left to right in Fig. 5 (West to East) and it was considered that water was following the interface between the gypsum, in which the roads were driven, and the marl bed below and making its way down the strike. Excavation of the gypsum appeared to have encroached on the marl bed and allowed water to come up through the floor. The main access to the mine was via the 1 in 4 adit which was close to the inflow position. Upon failure of the pressure pad during grouting operations Cementation Mining Ltd were asked on 21 October 1980, to design a new scheme for sealing off the water. The water inflow at that time was estimated to be 182 litres s -1 (2400 gal min -1) (Fig. 7). Various structural schemes for pressure pads and combinations of pads and plugs were considered and discarded in favour of the complete plug solution shown in Fig. 6 for simplicity, speed of construction and permanency.

Urgency was the main criteria as, within six days of Cementation Mining Ltd being called in (27 October), the water inflow had risen to 379 litres s -1 (5000 gal min -1) and it was rapidly becoming obvious that there was a danger of losing the mine.

The plug scheme adopted is detailed in Fig. 6. It was deemed prudent not to disturb the remaining sections of the original pressure pad. Another gravel bed was laid over the top in which six more water control pipes were placed in addition to the two pipes (one 200 mm diameter and one 100 mm diameter) previously installed below the original pressure pad. The additional pipes were four 200 mm diameter and two 300 mm diameter and carried the water to a new sump position adjacent to the proposed plug site. Additional rising mains were installed in the shaft to cope with the increasing inflow.

The first concrete was poured on 28 October and Fig. 6 shows the concreting stages. Because of the large mass of concrete involved, construction joints were necessary, and a low heat of hydration mix, incorporating a cement replacement material, was used (Table 3). Concrete was pumped from the surface down the 1 in 4 adit, through a 100 mm pipe, directly into position in the plug. The four week time period for placing the concrete resulted from various equipment, labour and general construction problems but once concreting had commenced the water inflow was controlled at a peak level of 606 litres s -1 (8000 gal min -1) (Fig. 7).

Minimal true design was required for the Sherburn Mine plug. The depth below ground level was 48 m which resul(ed in a hydrostatic pressure of 0.47 N mm -2 (68 lb in 2). This is not excessive and the length of the plug was extremely long. However, length in this case was governed by practical considerations to suit the particular situation. The pressure gradient from one end to the other was only 0.47/35.300 = 0.013 N mm -2 m -1 (0.59 lb in -1 ft -1) and the 35.3 m (116 ft) length (see Plan in Fig. 6) was eventually extended out to the adjacent access roadway.

Fig. 7 indicates how effective the plug was in stopping water. On completion of the various concrete stages, the control pipe valves were closed and the inflow almost completely stopped. Final sealing by grouting commenced after the valves were shut off and involved a combination of grout pipe positions. Some were previously cast into the plug to reach places which would have been inaccessible by drilling from the two plug faces. These, in addition to injection at the inflow point through the water control pipes and to the contact zones through other holes drilled from both faces, enabled the water to be completely sealed off on a permanent basis. Only cement grout injection was necessary.


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22 2'7 2 7 1'2 1'7 22 2'7 1 ; 1'1 1'6 21 26 Z ; ;1 1'6 SEPT OET NOV DEC 1980

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Fig. 7. British Gypsum Ltd, Sherburn Mine, England. Water inflow quantities related to concrete poured in emergency plug, showing effectiveness of water stopping (water inflow quantities in imperial units as recorded). Reproduced by permission of the Mine Manager, Mr R. Hird, and British Gypsum Ltd.

Proposed precautionary plug, 1981, for overseas contract

Fig. 8 contains recent proposals (1981) for three precautionary plugs to be installed near the bottom of a shaft for access protection in the case of an inrush. Construction of the plugs would be in a thin limestone bed, 15 m thick, situated above and below weak, water- bearing strata zones at a depth of 542.5 m (1780 ft) [5.43 N mm -2 (787 lb in -z) hydrostatic pressure].

Design of the plug, load transfer cylinder and bulkhead door was carried out in accordance with the design calculation section, Grade 35 concrete being specified. Concrete to rock interface calculated punching shear stress was 0.63 N mm -2 (91 lb in -z) and the pressure gradient 5.43/12 = 0.45 N mm -2 m -1 (19.9 lb in -2 ft-1).

The proposed grouting scheme would depend on the actual ground conditions at the level of the plug when the shaft is sunk. However, care must be taken above and below the plug not to encroach too close to the water-bearing zones.

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222 Auld

National Coal Board North Selby Mine, England, 1982 (temporary consolidation plug)

Cementation Mining Ltd are currently (December 1982) sinking two shafts at North Selby for the National Coal Board Selby project. Both shafts have reached the stage of sinking through the Ackworth Rock, which is a Coal Measures sandstone and an aquifer, No. 1 shaft sump (Fig. 9) standing at -540 .2m (1772fl) below surface level [hydrostatic pressure 5.4 N mm -2 (783 lb in-Z)]. The previous sump level in No. 1 shaft stood 13.8 m (45 fl) above the present sump level and strata cover grouting was carried out from that level.

During the period of strata cover grouting, problems of grout standpipe installation were experienced due to the poor rock conditions and deterioration and heave of the sump took place. The length of cover grouting was also long (over 40 m) whereas the preferred maximum length is approximately 30 m. To enable the wall of the cover grouting cone to be less prone to leakage at the lower level of treatment and to guarantee satisfactory grout standpipe installation, it was decided to sink to the -540.2 level and install a concrete plug. This would be closer to the zone of strata requiring the major grout treatment and, by casting the grout standpipes into the plug, a pressure pad for the next cover of strata grouting could be provided.

Due to the potential water inflows for sinking below the plug, it was necessary to install a pump lodge (Fig. 9) for stage pumping to surface. No choice of position was available for_ the pump lodge other than immediately below the last cast section of the shaft wall.

At the time of placing the plug, shaft water inflow to the sump was approximately 11 litres s -1 (150 gal min-1). Fig. 10 shows the framework for supporting the grout pipes and the water control rising mains during casting of the plug. The concrete mix design for the plug is given in Table 3. Minimal heat of hydration existed in the concrete mass due to using the cement replacement material (Cemsave) and additional heat removal occurred through the rising mains and grout pipes. The Grade 55 concrete was the same as the shaft lining concrete. However, designing the plug on the basis of the seven-day cube test results 32- 55 = 36.7 N mm -2 (5317 lb in -z) allowed pressurizing of the plug for water stopping, at the earliest opportunity. The benefit of the 28 day strength was taken for the wall bearing resistance. The recommendations given in the design calculation section for cylindrical plugs were followed for the plug design.

Neglecting the bearing resistance of the tapered plug, the calculated punching shear stress for the concrete to rock interface was 0.89 N mm -2 (129 lb in -2) and the pressure gradient 5.4/17.3 = 0.31 N mm -2 m -1 (13.7 lb in -2 ft-1).

Grouting up of the plug started from the bottom through 50 mm grout pipes installed in the rising mains. These pipes were grouted in, leaving the bottom free for injection into the gravel bed, and also secured by high-pressure flanges bolted together at the top of the rising mains. The bottom injection was phased to follow backwall injection of the shaft wall above the plug, and controlled using the standpipes as 'tell-tales' before closing off for final pressurizing.

The shaft water make was reduced to approximately 0.45 litres s -1 (6 gal rain -1) before final tightening up, this amount being predominantly from behind the shaft lining above the pump lodge. The pump lodge was restricted to a position close to the plug and remained a


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Fig. 9. National Coal Board, North Selby Mine, England, Section through shaft, showing temporary consolidation plug (a) and plan at pump lodge level (b). Design and construction by Cementation Mining Ltd (1982). Information published by permission of the National Coal Board.

224 Auld

potential source for strata water to penetrate the shaft if it could not be sealed off by grouting.

To enable the plug to be subsequently broken out without damaging the shaft wall, the bottom surface of the wall was painted with a bond breaking agent Setcrete 11 (Cementation Chemicals Ltd), the hanging rod ends were sleeved and two water bars incorporated, the inner one protected and the outer one sacrificial for plug sealing.

Conclusions and recommendations

The first objective of this paper was to review underground plug design for the purpose of bringing the subject to prominence and more up to date. As an additional objective, design rationalization was attempted on the basis of current practice.

The author considers the first objective to have been achieved. However, much more work needs to be carried out to quantify, in greater detail, strata leakage resistance in relation to plug length before the design procedure can be regarded as being completely rationalized.

The philosophies of design included in the paper are based predominantly on the excellent work of Garrett and Campbell Pitt which was reported in 1958 and 1961. In addition to the normally accepted punching shear stress concept of design for plug interaction with the surrounding rock, they proposed a bearing stress concept which was related to the surface roughness and also carried out tests on both experimental and service plugs to quantify plug length in relation to leakage resistance. This is the only published work known to the author which relates to the latter factor. However, the work carried out by Garrett and Campbell Pitt is specifically applicable to the gold mines of South Africa where hard rocks of the quartzite type are encountered at deep levels and high water pressures are experienced (Fig. 4). The quoted data is not directly applicable to any other rock conditions, particularly those of the softer sandstone, limestone, marls and coal measures experienced in the UK (Fig. 4), where the aquifers are closer to the surface and the hydrostatic p~essures are much less. A study of the Garrett and Campbell Pitt work was essential in the paper to form the basis for applying their principles to other rock conditions, in line with modem construction Codes of Practice, as it appears that very little forward progress has been made in the subject of plug design during the last two decades.

Considering the two parallel plug length design theories, one based on punching shear stress and the other on bearing stress, which have been proposed for resistance to horizontal thrust at the concrete to rock interface, it would appear that they are incompatible.

Fig. 10. National Coal Board, North Selby Mine, England. General arrangement elevation of temporary consolidation plug (a) and sections A-A(b), B-B(c) and C-C(d) showing supporting framework for cast in grout stand pipes and water control rising mains. Design and construction by Cementation Mining Ltd. Information published by permission of the National Coal Board.

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226 Auld

Comparing Equations lc and 8d, giving I = pb/4ppe and l = pb/2pb e respectively, using the value forpb e = 3.75Pr ~ from Table 1 indicates that the length based on permissible punching shear stress, as given by Equation lc, will always be the longer by a factor of 1.875. Based on the concept of length being a priority for resistance to leakage, the bearing stress concept can be neglected in the design of parallel plugs. It should, however, be pointed out that although the permissible shear stress concept is recommended for determining length, in order to assist sealing by increasing the leakage resistance, the acutal strength will be greater because of the bearing action.

As already mentioned in the design calculations section, the two tapered plug design theories based on different bearing stress concepts are also not compatible. In this case, the longer length is given by the Garrett and Campbell Pitt rough surface-bearing resistance philosophy, as opposed to the smooth-faced wedge principle, and the former is therefore the recommended approach based on the longer length required for leakage resistance.

With regard to the permissible stresses quoted in Table 1, the values of pc, Pb and pp are realistic for the current types of concrete now being used underground. The factor of safety equal to 4 used in connection with the Pb~ and pp~ values at the concrete to rock interface is also probably realistic. However, care should always be taken to study rock strength and condition to confirm the values. It is interesting to note that the permissible shear stress values for the interface, which are quoted by South African practice (see p. 207), are less than the values recommended in Table 1. Although the South African values are not related directly to concrete or rock strengths, nor the rock condition, they result in longer plug lengths which err on the safe side for leakage resistance. On this basis, it can be seen that the stronger, and better quality, concretes now being employed in underground construction will give shorter plug design lengths for strength but could have inherent leakage problems if sufficient length is not provided.

At the present time it is-not possible to define the exact length which is needed for sealing in relation to any particular ground conditions. The pressure gradient concept of Garrett and Campbell Pitt (1958, 1961) would appear to be a practical means for quantifying the resistance of rock to the passage of water through specific lengths but insufficient data is available as yet for general application of the principle. The allowable pressure gradient of 9.14 N mm -2 m -1 (404 lb in -2 ft-1), which the South Africans would accept for normally grouted rock, should not be adopted in the UK as it is based on plug to rock interface grouting pressures of 2 to 2 times hydrostatic. Such high pressures would not be adopted in the UK, values of 1.25 to 1.5 being more representative.

Comparing the pressure gradients from the case studies with the Garrett and Campbell Pitt data in Fig. 4, the Sherburn Mine emergency plug value of 0.013 N mm -2 m -1 (0.59 lb in -2 ft -1) is much less than that given by graph A [0.23 N mm -2 m -1 (20.8 lb in -2 ft-a)] showing it possessed a satisfactory leakage resistance without grouting.

For the proposed precautionary plug the pressure gradient of 0.45 N mm -2 m -1 (19.9 lb in -2 ft -1) was much less than that given by graph B [3.64 N mm -2 m -1 (161 lb in -2 ft-~)]. This indicates that although leakage would occur before grouting of the contact zone it would not leak after grouting the interface. The North Selby temporary consolidation plug was also in this category, possessing a greatly reduced pressure gradient [0.31 N mm -z m -1 (13.7 Ib in -2 ft-x)] than given by graph B.


It would appear that the Garrett and Campbell Pitt pressure gradient of 3.64 N mm-2 m-1 (161 lb in -2 ft -~) could be used as an upper limit in the UK for plugs with the contact zones and strata grouted. However, much lower pressure gradients will result in the ability to seal off leakage more easily.

Each plug scheme will generally be an individual design tailored to the particular situation. The above recommendations for pressure gradients should be used with caution and the rock leakage resistance in situ, which is associated with each design, must be investigated as thoroughly as possible prior to preparing any scheme. Successful plug design therefore, will rely heavily on the mining contractor's experience and knowledge.

Current concrete mix designs, using plasticizers for high workability, are much more easy to place and provide much tighter contact with the surrounding rock. Improved sealing will be achieved and leakage resistance much greater. Increased pressure gradients should be capable of being withstood by shorter lengths of plug and therefore, in future, the quantifying of such data by experiment and in situ monitoring is essential to progress and improve underground plug design.

Understanding of plug mechanisms of resistance to horizontal thrust, when confined by the surrounding rock, can be enhanced by further studies into rock stresses resulting from excavations. Modification of these stresses during interface pressure grouting and the accompanying plug stressing needs to be investigated. Finally study of the effects of end pressures on such a combined stress system would lead to knowledge of how stresses are dissipated throughout the whole and possibly a clearer picture of the interface ultimate behaviour under load would emerge. Future research and experiment are therefore imperative to advance the state of the art of plug design.

Acknowledgements

The author wishes to thank Mr J.C. Black, Managing Director of Cementation Mining Ltd, for permission to publish the paper. Illustrations and details are included from the Selby New Mine Project and the author is indebted to Mr C.T. Massey, Deputy Director (Mining) - Selby Project, of the National Coal Board for his permission to use this information. Details of the Sherburn Mine incident are published by courtesy of British Gypsum Ltd, and thanks are due to Mr W.S. Gibson, Chief Mining Engineer, British Gypsum Ltd, and Mr R. Hird, the Mine Manager at the time of the incident, for their permission to include such data. Further thanks are extended to Mrs M. Mordue, who typed the manuscript.

References

Auld, F.A. (1979) Design of concrete shaft linings, in Proceedings of the Institution of Civil Engineers, Part 2, Vol. 67, Sept., pp. 817-32.

Auld, F.A. (1982a) Ultimate strength of concrete shaft linings and its influence on design. Proceedings of the Symposium on Strata Mechanics, Newcastle upon Tyne, Elsevier Scientific Publishing Company, Amsterdam, pp. 134-40.

228 Auld

Auld, F.A. (1982b)Concrete in underground works, Concrete Society Technical Report No. 105, The Concrete Society, London.

Auld, F.A. (1982c) Concrete in underground works in The Concrete Society North West Region Symposium, Concrete in the Energy Industry.

British Ready Mixed Concrete Association (1978) Register of Commercial Test Houses, 4th edn, British Ready Mixed Concrete Association Ltd, London.

British Standards Institution (1969a) BS 449:Part 2:1969. Specification for th e use of structural steel in building, Part 2, Metric units, British Standards Institution, London.

British Standards Institution (1969b) CP 114:i969, The Structural use of reinforced concrete in buildings, British Standards Institution, London.

British Standards Institution (1970) BS 1881:Part 4:1970, Methods of testing concrete, Part 4, Methods of testing concrete for strength, British Standards Institution, London.

British Standards Institution (1972) CP ll0:Part 1:1972, The structural use of concrete, Part 1, Design, materials and workmanship, British Standards Institution, London.

British Standards Institution. (1976) BS 5337:1976, Code o f practice for the structural use of concrete for retaining aqueous liquids, British Standards Institution, London.

Comite Europeen du Beton - Federation Internationale de la Precontrainte (1970) International recommendations for the design and construction of concrete structures. Principles and Recommendations, FIP Sixth Congress, Prague.

Fliigge, W. (1967) Stresses in Shells, p. 24, Springer-Verlag, Berlin. Garrett, W.S. and Campbell Pitt, L.T. (1958) Tests on an experimental underground bulkhead for

high pressures, Journal of the South African Institution of Mining and Metallurgy 59, 123-43. Garrett, W. S. and Campbell Pitt, L. T. (1961) Design and construction of underground bulkheads and

water barriers, Paper presented to the Seventh Commonwealth Mining and Metallurgical Congress, Johannesburg.

Jaeger, J. C. and Cook, N.G.W. (1979) Fundamentals ofRockMechanics, 3rd edn, Chapman and Hall, London. pp. 86-8.

Lancaster, F.H. (1964) Report on Research into Underground Plugs, Transvaal and Orange Free State Chamber of Mines Research Report No. 27/64.

Manning, G.P. (1961) Reinforced Concrete Design, 2nd edn. Longma

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