INTERNATIONAL SOCIETY FOR SOIL MECHANICS AND GEOTECHNICAL ENGINEERING

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This is an open-access database that archives thousands of papers published under the Auspices of the ISSMGE and maintained by the Innovation and Development Committee of ISSMGE. T hem e lecture: B ored tunnelling in the urban enviro nm ent

E xposé sur le thèm e: Tunnels forés dans un environn em ent urbain

R. J. M air - Geotechnical Consulting Group, London, UK

R. N.Taylor - City University, London, UK

A BSTRA CT: The Report reviews the state-of-the-art and recent developments in geotechnical aspects o f bored tunnelling in the urban environment. Various advances in tunnel construction techniques are highlighted, notably the increasing use o f slurry and earth pressure balance shields. Methods o f calculating the stability o f tunnels are reviewed. In the context o f tunnelling in the urban environment, particular emphasis is given in the Report to ground movements associated with tunnel construction, their modelling and prediction, and their effects on buildings. Reference is made to case histories o f tunnels constructed in a wide variety o f ground conditions. Ground loading acting on tunnel linings is discussed. The Report focuses on three areas in which there have been significant developments in ground treatment in recent years: face reinforcement, slurry and earth pressure balance shield technology, and compensation grouting.

RESUM E: Ce rapport examine l’état de l’art et les dévelopments récents concernant les aspects géotechniques des tunnels creusés en milieu urbain. Un certain nombre de progrès dans les techniques de construction sont illustrés, notamment l’augmentation de l’utilisation des boucliers à pression de boue et des boucliers à contrepression de terre. Les différentes méthodes de calcul de stabilité des tunnels sont examinées. Dans le contexte du creusement de tunnel en milieu urbain, une attention toute particulière est accordée dans ce rapport aux mouvements du sol associés à la construction du tunnel, leur modélisation, leur prédiction et leurs effets sur les bâtiments. Il est fait référence à plusieurs projets de tunnels construits dans une grande variété de type de terrain. Le chargement sur le soutènement du tunnel par le sol est traité. Le rapport se concentre sur trois domaines du traitement des sols dans lesquels des développements significatifs ont eu lieu au cours des dernières années: le renforcement du front, la technique des boucliers à pression de boue et des boucliers à contrepression de terre, et les injections de compensation.

1. INTRODUCTION • Modelling and prediction o f ground movements • Effects o f ground movements on buildings The title o f Plenary Session 4 o f this Conference, Underground • Ground loading on tunnel linings Works in the Urban Environment, principally involves bored • Developments in ground treatment tunnel construction and deep excavations. This Report focuses only All o f these topics are o f particular importance in the context on bored tunnel construction, primarily because Professor o f tunnelling in the urban environment. Nussbaumer in his Report for Plenary Session 3 has addressed the subject o f Retaining Structures and Excavated Slopes, and the 2 ADVANCES IN TUNNEL CONSTRUCTION TECHNIQUES associated Discussion Sessions have been concerned with the subject o f deep excavations. In recent years there have been considerable advances in It should be made clear that the term “bored” tunnel techniques of bored tunnel construction in soft ground. It is construction means any kind o f mined tunnel, as distinct from cut- convenient to highlight some o f these in terms o f open face and and-cover tunnels. closed face tunnelling. Open face tunnelling covers all cases where At the 7lh ICSM FE in Mexico in 1969, Professor Ralp h Peck there is easy access to the tunnel face, in contrast to closed face presented his seminal state-of-the-art paper on Deep Excavations tunnelling. and Tunnelling in Soft Ground. Since then there have been significant developments in the theory and practice o f geotechnical (a) Open Face Tunnelling engineering applied to underground construction, and there have been various major review papers devoted to the subject (Peck, There is increasing use of sprayed concrete in soft ground 1969; Cording and Hansmire, 1975; Clough and Schmid t, 1981; tunnelling to form linings, particularly for tunnels of shorter Ward and Pender, 1981; O ’Reilly and New, 1982; Schlosser et al, lengths and of non-circular cross-section. These are usually 1985; Attewell et al, 1986; Konda, 1987; Rankin, 1988; Uriel and temporary, but may be the sole means o f support for significant Sagaseta, 1989; Clough and Leca, 1989; Cording, 1991; Fujita, periods (sometimes more than a year) before the permanent lining 1989, 1994). There have also been various International Tunnelling is installed. Recent developments have involved composite sprayed Association Conferences (held annually) and National Tunnelling concrete linings, in which sprayed concrete is used for both the Conferences. temporary and permanent linings (e.g. Wittke, 1995; Grose and In particular there have been two recent Symposia organized by Eddie, 1996; Negro et al, 1996). Non-circular cross-sections and ISSM FE Technical Committee 28: the first in New Delhi in 1994 divided faces are often adopted when using sprayed concrete, as and the second in London in 1996 (Fujita and Kusakabe, 1995; shown in Figure 1, and this allows considerable flexibility in terms Mair and Taylor, 1996). of modifying the construction sequence in response to The topics addressed in this Report are: observations. The use of sprayed concrete to form linings is • Advances in tunnel construction techniques sometimes referred to as the New Austrian Tunnelling Method • Principal design and construction requirements (NATM). • Stability Ground treatment is more easily undertaken from within tunnels • Ground movements with open faces. Advances have been made in reinforcement o f the

2353 Cutter driving Erector driving motor \ motor Tai l seal

Sl ur r y suppl y

Sl ur r y ret urn

Figure 1. Example o f divided tunnel face using sprayed concrete linings (Institution o f Civil Engineers, 1996) / Agitator bnieia \ Erect or Cutter face Segments soil ahead o f the face to improve stability and to control ground Figure 3. Principle o f the slurry shield machine (Fujita, 1989) movements (Figure 2a). Improvements in jet grouting techniques are being made to form “umbrella arches” (Figure 2b) as a pre­ lining in difficult ground conditions. universally for all types o f unstable ground; the principle is shown An extension o f the concept of the umbrella-arch is the pre-‘ in Figure 4. By controlling the entry o f soil and w ater through the vault, developed in France and Italy, which is sometimes referred cutter face by means o f earth pressure balance doors, and by to as the mechanical pre-cutting method (Cazenave and Le Goer, conditioning the spoil so that it can easily be removed through a 1996). This involves the cutting o f overlapping slots around the screw conveyor, it is possible to control the pressure of the tunnel periphery in advance o f the excavation, and filling them by excavated soil in the chamber to balance the earth and water means o f sprayed concrete (Figure 2c). The technique is often used pressures in the ground. in conjunction with face reinforcement and other forms o f ground Recent developments have centred around the injection of treatment. special slurries, foams and other materials in EPB machines to improve the properties o f the excavated soil and facilitate the (b) Closed Face Tunnelling proper control o f the pressure in the chamber. This is essential for the control o f face stability and for minimising ground movements. Considerable advances have been made over the last decade in the In the case o f slurry shields, more is now understood about the use o f sophisticated closed face tunnelling machines which operate factors affecting the efficiency o f the slurry support o f the tunnel on the principle o f a pressurized face. These machines are used in face. These are discussed further in Section 4. unstable ground conditions where the face requires support at all times; this principally applies to permeable ground below the water table (i.e. mainly sands or mixtures o f sands, silts and clays) or soft Cutter dri vi ng Tai l seal clays. The slurry shield machine, illustrated in Figure 3, is most ■mot er Scr ew conveyor commonly used in water bearing granular soils. The face is dri vi ng mot or supported by a pressurized bentonite or polymer based slurry, which is circulated so that it and the excavated soil are removed to a separation plant. Earth pressure balance (EPB) machines are being used more

Bel t conveyor Gat e j ack F Bulkhead \ Erector Segments (a) Face reinforcement Cut t er face Cutter frame Shield jacks

Figure 4. Principle o f the earth pressure balance m achine (Fujita, 1989)

3. PRINCIPAL DESIGN AND CONSTRUCTION REQUIREMENTS (b) Jet grouting “umbrella arches”

As presented by Peck (1969), and developed by Ward and Pender (1981), the three most important requirements for the successful design and construction o f a tunnel can be summarised as follow s:

(i) S tability

The choice of excavation and construction technique must be (c) Pre-vault (pre-cutting) suited to the ground conditions so that it is feasible to build the Figure 2. Ground treatment and pre-lining techniques tunnel safely. O f prime importance is the stability o f the opening (after Schlosser and Guilloux, 1995) prior to installation o f the lining.

2354 (ii) Ground movements and their effects , tunnel lining O f particular relevance to the urban environment, construction o f the tunnel should not cause unacceptable damage to surrounding ... :i i i i i ♦ 4 or overlying structures and services. Prior to construction the ground movements should be predicted and their effects on the *-QT structures and services assessed. , r t i i i r r — (iii) Performance o f linings

The tunnel lining, whether it be temporary or permanent, must be capable of withstanding all the influences to which it may be (a) subjected during its design life. This requires predictions o f the soil loading acting on the lining and o f the deformations o f the lining, the latter being o f particular significance in the case o f external influences such as adjacent tunnel construction. These three principal requirements, which are closely related, form the basis of this Report. The first two are of particular importance in the urban environment and therefore more emphasis is placed on these. The subject of this Report is restricted to tunnelling in soft ground, but the meaning o f “soft” ground can differ according to the perspectives o f the tunnelling engineer and the geotechnical engineer. It is appropriate to define “ soft” ground (b) as that which requires some form o f support in the tunnel to prevent instability o f the ground, either in the short term or long Figure 5. (a) Tunnel heading in soft ground; (b) tw o-dimensional term. In this context even materials such as stiff to hard clays, and idealization o f tunnel heading weak rocks, are classified as “soft” ground.

4.2 Undrained S tability 4. STA BILITY

Based on the concept introduced by Broms and Bennermark (1967) 4.1 Introduction the stability ratio, N can be defined as;

In any tunnelling project adequate stability during construction is yz - aT N (1) clearly o f prime importance, and this is particularly the case in urban environments where the consequences o f a majo r tunnel collapse can be catastrophic. In recent years there have been a where number o f major tunnel collapses associated with the use of Y = unit weight o f the soil sprayed concrete linings (Anderson, 1996; HSE, 1996), although depth to the tunnel axis collapses also occur with other forms o f tunnelling. (=C + D/ 2, see Figure 5) Stability o f a tunnel heading (with circular cross-section) can be ° s = surface surcharge pressure (if any) tunnel support pressure (if any) considered in terms o f the idealized geometry shown in Figure °T = undrained shear strength at tunnel axis level 5(a). The heading may be supported by a fluid pressure, Op, such as s„ = compressed air or pressurized slurry (in the case o f a slurry shield). It may be excavated in free air in open face mode, in which case o T On the basis o f laboratory extrusion tests and field observations, Broms and Bennermark (1967) concluded that the critical stability = 0. The dimension P represents the distance from the face to the point where stiff support is provided; in the absence o f a tunnelling ratio at collapse, Nc , is about 6. Similar conclusions were reached shield, this is the distance from the face to the lining. In most cases by Peck (1969). when a tunnelling shield is in use in ground o f low stability, the Davis et al (1980) derived plasticity solutions employing ground is in contact with the shield and therefore P can be taken to kinematic upper bounds and statically admissible lower bounds for the two-dimensional idealization o f a tunnel heading shown in be zero. An exception is when a shield is being used in ground of Figure 5(b), assuming constant undrained shear strength profiles. higher stability, such as stiff clays; in this case there is often an Reasonably close agreement was obtained between the upper and oversized cutting edge at the front o f the shield to ensure a gap lower bounds. Their results are summarized on Figure 6, which between the ground and the length o f the shield and thereby shows the derived values o f critical stability ratio, Nc , in terms o f facilitate easy steerage. Determination o f P then requires some the dimensionless ratios C/ D and yD/ su. Slight improvements to judgement. the upper bound for the case o f weightless soil (yD/ su=0) have The issue o f whether undrained or drained conditions are more been made by Antao et al (1995). A comprehensive set o f solutions applicable to the tunnel stability problem depends principally on for the more general two-dimensional case where the undrained the permeability o f the soil, the excavation advance rate, and the shear strength increases with depth was derived by Sloan and size o f the tunnel. Based on parametric studies o f seepage flow into Assadi (1993), using finite element formulations o f the upper and tunnel excavations, Anagnostou and Kovari (1996) concluded that lower bound plasticity theorems. for most tunnels drained conditions are to be expected when the O f most relevance to practical tunnel stability problems is the soil permeability is higher than 10'7 to 10'6 m/ s and the excavation three-dimensional heading, shown in Figure 5(a). Based on advance rate is 0.1-lm/ hr or less. Hence, in a predominantly sandy centrifuge model tests, the design curves in Figure 7 showing the soil, drained stability should be considered. In low permeability critical stability ratio, Nc , in terms o f the dimensionless ratios P/ D clayey soils undrained stability is o f more importance during tunnel and C/ D were proposed (Mair, 1979; Kimura and Mair, 1981). The excavation, but in the case o f a standstill drained conditions could special case of P/ D=0 is o f particular relevance, and Figure 8 become more relevant. Factors influencing whether undrained or shows part o f the design line taken from Figure 7, together with drained conditions are more applicable are also discussed by Negro data from laboratory and centrifuge tests, and also from back- and Eisenstein (1991).

2355 ------Low er bound

------U pper bound Centrifuge model te sts i 1g model test

0 0-2 0-4 0-6 0-8 1 0 1-2 1-4 1-6 1-8 C/D 0 1 2 c/ D 3 4 Figure 8. Critical stability ratio values for lined tunnel headings Figure 6. Upper and lower bound critical stability ratios for plane (P/ D=0) with thin clay cover (Mair, 1993) strain circular tunnel (Davis et al, 1980)

prevent collapse was only slightly lower than for the two analysis o f tunnel heading failures in the field (M air, 1993). Also dimensional case (P/ D = °°). Their experimental results, together shown on Figure 8 is an envelope of lower bound plasticity with their plasticity solutions for (J)’ = 50° (w hich was evaluated by solutions derived by Davis et al (1980); in contrast to the two- the authors for the low stress levels around the tunnel), are shown dimensional case, upper bound solutions (not shown) give Nc on Figure 10. The experimental data and the plasticity solutions values which are much higher than those given by the lower bound indicate the support pressure to be independent o f the ratio C/ D. solutions. Similar high upper bound values were obtained by Leca Similar findings are reported by Chambón and Corté (1994), and Dormieux (1992). who performed centrifuge tests on lined tunnel headings in dry The stability numbers summarized in Figures 7 and 8 can also sand; their results are shown in Figure 11. As found by Atkinson be used to estimate the risk o f “blow -out”, which can occur if the and Potts, the required support pressure was almost independent o f tunnel face pressure is too high in soft clays (Mair, 1987; De Moor C/ D. and Taylor, 1991). Upper and lower bound plasticity solutions for the stability o f the three-dimensional tunnel heading (P/ D = 0) in dry soils in 4.3 Drained S tability terms o f c’ and 4>’ were derived by Leca (1989) and Leca and Dormieux (1990). Their results are also presented in Figure 9 for Atkinson and Potts (1977) derived kinematic upper bound and the case of c’ = 0, (J)’ = 35° for comparison with the two- statically admissible lower bound plasticity solutions for the two- dimensional results of Atkinson and Potts (1977). Leca and dimensional idealization shown in Figure 5(a) for dry cohesionless Dormieux found that their upper bound solution was in good soils. Their results are presented in Figure 9, for the case o f 4>’ = agreement with earlier centrifuge model tests reported by Chambón 35°, where the tunnel support pressure oT required to prevent and Corté (1989), the support pressure required to prevent collapse collapse is expressed as the dimensionless ratio oT/ yD plotted being independent o f the ratio C/ D. It is o f interest to note that, in against C/ D. It should be noted that the upper bound plasticity contrast to the presently available undrained (<)) = 0) stability solution (being inherently unsafe) gives a lower value o f oT/ yD solutions, the lower bound solution for the tunnel heading shown than the lower bound solution which is inherently safe. Centrifuge on Figure 9 gives significantly higher support pressures than the model tests in dry sands reported by Atkinson and Potts were upper bound, and these increase with C/ D, differing from consistent with these plasticity solutions, and tests on model tunnel headings (P/ D = 0) showed that the support pressure required to Lined tunnel heading

____ upper bound Leca (1989),

------low er bound J Leca and Dorm ieux (1990)

••x. lim it equilibrium Anagnostou and

K ovari (1996)

o ti P lane strain tunnel

upper bound Atkinson and ^ ^

ctt lower bound / Potts (1977) yd

1.5

1 2 3 C/D C/D

Figure 7. Dependence o f critical stability ratio on tunnel heading Figure 9. Drained stability solutions for plane strain tunnels geometry (Mair, 1979; Kimura and Mair, 1981) (P/ D=°°) and lined tunnel headings (P/ D=0); 4>’=35°

2356 0.2 œntrifuge models filter cake CTT/ nyD laboratory models slurry shield

water pressure, u

upper bound

Stability depends on: C/ D • excess pressure Ap = p - u Figure 10. Centrifuge model tests on lined tunnel headings in dry • yield strength of slurry xf sand (Atkinson and Potts, 1977) • grain size d10

membrane model experimental observations. Anagnostou and Kovari (1996) present (Ap = 40 kPa) a limit equilibrium solution for a lined tunnel heading in terms o f c ’ and <}>’. Their solution for c ’=0, 4>’=35° is shown in Figure 9 and shows the support pressure being independent o f C/ D, as obtained by Leca and Dormieux (1990) for their upper bound plasticity solution. The principal conclusion arising from the stability solutions illustrated in Figure 9, and from the centrifuge model test data in Figures 10 and 11, is that the effective support pressure required to prevent collapse o f a tunnel in dry cohesionless soil is very small, irrespective o f whether it is a tw o-dimensional circular tunnel or a three-dimensional heading. It is also independent o f tunnel depth. All o f the foregoing applies to dry cohesionless soils, which are not often encountered in practice in tunnelling. In many cases even grain size d10 (mm) soils above the water table contain sufficient moisture to behave as sand gravel if they exhibited significant values o f c’ ; then even smaller support pressures are required, as shown by the stability solutions o f Leca A: Ap = 20 kPa, 4% bentonite (xf = 15 Pa) and Dormieux (1990) for c’, (J)’ soils. In practice, tunnelling is frequently undertaken below the water B: Ap = 40 kPa, 4% bentonite (xf = 15 Pa) table, in which case positive measures are necessary to prevent C: Ap = 20 kPa, 7% bentonite (tf = 80 Pa) water inflow and ensure adequate drained stability o f gravels, sands Figure 12. Safety factor against face instability for a slurry shield and silts. Compressed air can be used for this purpose, although it (after Anagnostou and Kovari, 1996) has become less attractive in recent years as more is understood about its potential adverse effects on tunnel workers (especially at high pressures). Its decline in use is also associated with the An example of the approach by Anagnostou and Kovari is increasing adoption o f full face slurry or EPB machines. illustrated in Figure 12, in which the safety facto r against face Face stability in slurry and EPB tunnelling is considered by instability for a 10m diameter tunnel is shown to be a function o f Anagnostou and Kovari (1996). Based on a limit equilibrium the excess slurry pressure (Ap), the concentration o f bentonite and approach, they developed computational models which provide a the associated yield strength o f the slurry (xf), and the characteristic useful framework for quantifying the mechanics o f tunnel face grain size o f the soil (d10). The following practical points emerge: failure. For a slurry machine, the stabilising force capable o f being 1. An increase in safety can be achieved by raising the excess exerted by the slurry depends on the extent o f its infiltration into slurry pressure, but only in the finer grained soils. A bove a d10 size the ground. Under optimum conditions, the slurry seals the face by o f approximately 2mm, increasing the slurry pressure results in forming a filter cake and acts like a membrane; the support force deeper infiltration and fluid loss, and is o f little benefit. then results from the excess slurry pressure over and above the 2. Increasing the bentonite content o f the slurry for the coarser ground water pressure. When the soils are o f high permeability, or soils results in more effective support. when the shear resistance o f the slurry is low, the slurry will In EPB machines, the muck chamber is filled with excavated penetrate the ground and the force it is capable o f exerting depends soil under pressure. Face instability is potentially a problem only on the depth o f infiltration. Stability solutions for face stability o f when the piezometric head in the muck chamber is lower than the slurry machines, based on limit equilibrium methods, have also piezometric head in the ground, so that seepage forces act towards been derived by Jancsecz and Steiner (1994). the tunnel face. Anagnostou and Kovari provide normalised charts for assessing the tunnel face stability under these conditions. As noted by Cording (1991), tunnelling machines are most

centrifuge acceleration ng efficient when operated in a single ground condition. For this case, (jj/ rryD o n = 50 it is possible to quantify the tunnel face stability using the soil ■ n = 100 a n = 130 mechanics principles and the methods o f analyses reviewed in this Report. For mixed face conditions, however, the situation is very different. A particularly adverse combination is when potentially unstable soils, such as sands and silts below the w ater table, are encountered together with harder materials such as very stiff clays or rock (boulders) which are difficult to excavate. In these circumstances, the stability o f the tunnel face may be difficult to quantify, although it is often possible to establish upper and lower Figure 11. Centrifuge model tests on lined tunnel headings in dry limits by making simplifying assumptions about the geometry of sand (Chambón and Corté, 1994) the different strata and applying the methods o f analyses referred

2357 5.2 Components o f Ground Movement

The primary components of ground movement associated with shield tunnelling are depicted in Figure 14. These are as follows: 1. Deformation o f the ground towards the face resulting from stress relief. 2. Passage o f the shield: the presence o f an over-cutting edge (bead) combined with any tendency o f the machine to plough or yaw will lead to radial ground movements. 3. Tail void: the existence o f a gap between the tailskin o f the shield and the lining means that there will be a tendency for further b radial ground movements into this gap. (a) cla ys (b) sands 4. Deflection o f the lining as ground loading develops. Figure 13. Observed failure mechanisms based on centrifuge 5. Consolidation: as pore water pressures in the ground change to model tests (Mair, 1979; Chambón and Corté, 1994) their long-term equilibrium values the associated changes in effective stress lead to additional ground movements. to above. The difficulty o f applying soil mechanics theories to Component 1 is o f major importance in many cases, particularly practical geotechnical engineering involving mixed ground in open-faced tunnelling in clays. In London Clay, for example, conditions is a common problem not unique to tunnelling, and significant ground movements are observed ahead o f the face due considerable judgement is often required. to stress relief (Ward, 1969; Mair and Taylor, 1993). When pressurized face tunnelling machines are used, however, either in 4.4 Failure Mechanisms slurry shield or EPB mode, this component can be negligible if the face pressure is carefully controlled. Component 2 can be The geometries o f failure mechanisms for tunnels in clays and appreciable, particularly if the over-cutting edge (bead) is of sands or gravels are markedly different. Figure 13 illustrates the significant thickness, and if there are steering problems in type o f observed failure mechanism based on centrifuge model maintaining the alignment o f the shield. Component 3 can be tests on tunnel headings in clays (Mair, 1979) and in sands minimised by immediate grouting to fill the tail void (or, in the (Chambón and Corté, 1994). In the case o f clays, the mechanism case o f expanded linings, by expanding the lining against the soil propagates upwards and outwards from the tunnel invert becoming at the earliest opportunity). Component 4 is generally small in significantly wider than the tunnel diameter. In contrast, failure in comparison to the other components once the lining ring is sands involves a narrow “chimney” , propagating almost vertically completed. Component 5 can be o f importance, particularly when from the tunnel up to the ground surface. This same narrow tunnelling in soft clays, as discussed in Section 5.9. funnelling behaviour has also been observed in laboratory lg In cases where there is no tunnelling shield, for example when model studies on tunnels in sands by Cording et al (1976) and Potts sprayed concrete linings are used, components 1, 4 and 5 are still (1976). Experience o f cases o f tunnel failures in the field are applicable. generally consistent with the mechanisms depicted in Figure 13. 5.3 Surface Settlement 5. GROUND MOVEMENTS For the case o f a single tunnel in “green field” conditions, the 5.1 Introduction development o f the surface settlement trough above and ahead o f the advancing heading is as shown in Figure 15. Following work The construction o f bored tunnels in soft ground inevitably causes by Martos (1958) on field observations o f settlements above mine ground movements. In the urban environment these may be of openings, Schmidt (1969), Peck (1969) and subsequently many particular significance, because o f their influence on buildings, other authors have shown that the transverse settlement trough other tunnels and services. The prediction o f ground movements and the assessment o f the potential effects on the infrastructure is therefore an essential aspect of the planning, design and construction o f a tunnelling project in the urban environment. The important subject o f ground movements associated with bored tunnel construction in soft ground has been addressed by many authors. Notable review papers have been produced by Peck (1969), Cording and Hansmire (1975), Clough and Schmidt (1981), Ward and Pender (1981), O’Reilly and New (1982), Attewell et al (1986), Rankin (1988), Uriel and Sagaseta (1989), Cording (1991), New and O’Reilly (1991), and Fujita (1989, 1994). In 1996, ISSM FE Technical Committee 28 organised an International Symposium on Geotechnical Aspects o f Underground Construction in Soft Ground in London; 35 papers submitted to the Symposium concerning settlement effects of bored tunnels were reviewed by Mair (1996).

Figure 14. Primary components o f ground movement associated with shield tunnelling (after Cording, 1991) et al, 1986)

2358 horizontal and Schmidt, 1981; Fujita, 1981; O’Reilly and New, 1982; -2i -i 0 j 2i distance, y Rankin, 1988). In a survey o f UK tunnelling data O’ Reilly and New (1982) showed that i is an approximately linear function of the depth of tunnel, z0 , and is broadly independent o f tunnel

0.6Smax, point of construction method and o f tunnel diameter (except for very inflexion shallow tunnels where the cover to diameter ratio is less than one). They proposed the simple approximate relationship: Smax settlement Figure 16. Gaussian curve used to describe the transverse i = Kzn (5) settlement trough

where K is a trough width parameter, and they recommended that immediately follow ing tunnel construction is well-d escribed by a for practical purposes K could be taken as 0.5 for tunnels in clays Gaussian distribution curve (shown in Figure 16) as: and 0.25 for tunnels in sands and gravels; the database for tunnels in sands was confined to shallow tunnels with depths to axis level 5 v = e xp (-y2/2 i2) (2) in the range 6-10m. The validity of equation (5) was generally confirmed by Rankin (1988) for a wide variety o f tunnels and for most soil types from around the world. Most o f the data presented where Sv = settlement by Rankin (and subsequently by Lake et al, 1992), together with Smaj( = maximum settlement on the tunnel centre-line more recent data, are shown in Figure 18 for tunnels in clays and y = horizontal distance from the tunnel centre-line in Figure 19 for tunnels in sands and gravels. The details o f the i = horizontal distance from the tunnel centre-line to tunnels, the ground conditions and the excavation methods are the point o f inflexion o f the settlement trough given in Tables 1 and 2 respectively. The reference numbering on the figures and the tables are consistent with that used by Lake et The volume o f the surface settlement trough (per metre length o f al (1992), but data have only been included on Figures 18 and 19 tunnel), V s , can be evaluated by integrating equation (2) to give and the tables where the ground profile is either predominantly clays or predominantly sands and gravels. Case histories have been V = JTk i Sm S' m (3) excluded where, for example, a tunnel in clay is overlain by significant thicknesses o f granular deposits, or vice-versa.

The volume loss, V ,, (sometimes referred to as ground loss) is Offset to point of inflection, i (m) the amount of ground lost in the region close to the tunnel, 5 10 15 20 primarily due to one or more o f the components 1-4 in Figure 14. When tunnelling under drained conditions, for examp le in dense sands, Vs is less than V, because of dilation (Cording and E. Hansmire, 1975). When tunnelling in clays, ground movements usually occur under undrained (constant volume) conditions, in (J) which case Vs = V, . Whatever the soil type, it is convenient to express the volume loss in terms o f the volume of the surface 0) settlement trough, V s , expressed as a percentage fraction, V ,, of c the excavated area o f the tunnel, i.e. for a circular tunnel

TlP 2 V. = V, (4) Q. 4 a

Peck (1969) suggested a relationship between the parameter i, tunnel depth and tunnel diameter, depending on the ground conditions, as shown in Figure 17. Many authors have investigated similar relationships (e.g. Cording and Hansmire, 1975; Clough

i/ D no. Reference 0.5 1.5 2.5 • 1 Hanya (1977) O 2 Attewell and Farmer (1974) ■ 3 Attewell (1978) □ 4 Glossop et al. (1979) ♦ 5 Toombs (1980) O 6 West et a I (1981) A 7 Attewell et al (1978) A 8 Muir Wood and Gibb (1971) + 9 Glossop and O'Reilly (1982) ► 11 Eden and Bozozuk (1968) t> 12 Henry (1974) * 14 Moretto (1969) © 17 Lake et al (1992) ▼ 19 Hanya (1977) v 20 Peck (1969) * 21 Peck (1969) 0 22 O’Reilly and New (1982) H 23 O'Reilly and New (1982) * 24 O'Reilly and New (1982) * 25 Attewell (1978) * 26 Barrati and Tyler (1976) ® 27 McCaul (1978) o 28 New and Bowers (1994) + 30 Kuwamura (1997) • 31 Shirlaw et al (1988) Figure 17. Relation between settlement trough width parameter and depth o f tunnel for different ground conditions Figure 18. Variation in surface settlement trough w idth parameter (Peck, 1969) with tunnel depth for tunnels in clays

2359 Table 1. Details o f tunnels in clays for which data are plotted in Figure 18 (based on Lake et al, 1992)

Depth to tunnel Inflection no. Source Location Ground conditions Excavation methods diameter (m) axis, zo (m) Smax (mm) offset, i (m) 1 Hanya (1977) Japan la stiff cohesive soil slurry shield 7.5 18.3 16 9.2 1b stiff cohesive soil slurry shield 7.5 21.8 43 11.2 1c stiff cohesive soil shield, hand excavated 10.7 19.0 22 6.6 2 Attewell and Farmer (1974) Green Park. UK stiff OC clay shield, hand excavated 4.2 29.3 6 12.6 3 Attewell (1978) Hebburn. UK soft NC clay shield, hand excavated 2.0 7.5 8 3.9 4 Glossop et al. (1979) Belfast. Ireland 4a soft silty clay shield, hand excavated 2.7 4.9 17 2.7 4b soft silty clay shield, hand excavated 2.7 4.5 20 2.4 5 Toombs (1980) Avonmouth. UK soft alluvial deposit shield, hand excavated 3.4 6.0 13 4.8 6 West et al (1981) York Way. UK stiff OC clay no shield, hand excavated 4.1 14.1 3 7.5 7 Attewell et al (1978) Tyneside. UK soft silty alluvial clay shield, hand excavated 4.3 13.5 23 6.5 8 Muir Wood and Gibb (1971) Heathrow. UK stiff OC clay (London Clay) shield, hand excavated 10.9 12.9 11 6.6 9 Glossop and O’Reilly (1982) Haycroft. UK 9a section B soft silty clay shield, hand excavated 3.0 5.5 40 3.6 9b section A soft silty clay shield, hand excavated 3.0 6.5 60 3.6 9c section C soft silty clay shield, hand excavated 3.0 8.0 68 4.3 11 Eden and Bozozuk (1968) Ottawa. Canada firm OC clay (Leda Clay) extremely sensitive TBM, rotary 3.0 18.3 6 8.4 12 Henry (1974) Grangemouth. UK soft to very soft laminated silty clay shield, hand excavated 2.45 to 2 10.0 25 4.1 14 Moretto (1969) Buenos Aires. Argentina soft to firm clay shield, mechanical digger 4.7 16.4 150 7.1 17 Lake et al (1992) Gateshead. UK firm clay, laminate d hand excavated with tim ber lagging 1.5x1.5 5.3 37 3.7 19 Hanya (1977) Japan firm cohesive soil shield, hand excavated 7.3 16.6 63 6.3 20 Peck (1969) Chicago. USA medium clay (Chicago Leda Clay) hand excavated 6.1 11.9 23 4.9 21 Peck (1969) Toronto. Canada glacial till hand excavated 5.3 13.1 9 6.1 22 O'Reilly and New (1982) Newcastle. UK firm stiff silty clay, glacial till partial face machine 5.2 14.2 8 7.0 23 O'Reilly and New (1982) Sutton. UK 23a stiff fissured clay (London Clay) hand excavated 1.8 17.1 4 10.0 23b firm to stiff weathered clay (London Clay) hand exca vated 1.8 3.4 4 2.0 23c firm to stiff weathered clay (London Clay) full face micro TBM 1.5 4.9 7 3.0 24 O'Reilly and New (1982) Oxford. UK stiff fissured clay TBM, full face 2.8 11.7 2 5.0 25 Attewell (1978) Howden. UK stiff boulder clay hand excavated 3.6 14.2 11 6.9 26 Barratt and Tyler (1976) Regents Park. UK 26a southbound stiff OC clay (London Clay) shield, hand excavated 4.2 34.0 5 15.2 26b northbound stiff OC clay (London Clay) shield, hand excavated 4.2 20.0 7 10.3 27 McCaul (1978) Stockton-on-Tees. UK 27a soft to very soft silty clay w ith sand lenses shield, hand excavated 1.3 6.3 44 3.5 27b soft to very soft silty clay w ith sand lenses shield, hand excavated 1.3 5.9 56 3.7 28 New and Bowers (1994) Heathrow. UK 28a stiff OC clay (London Clay) sprayed concrete lining type I (sidewall only) 22.0 21 9.5 28b stiff OC clay (London Clay) sprayed concrete lining type I (complete) 11.3 22.0 28 9.9 28c stiff OC clay (London Clay) sprayed concrete lining type II (sidewall only) 22.0 15 10.6 28d stiff OC clay (London Clay) sprayed concrete lininq type II (complete) 11.3 22.0 27 9.9 30 Kuwamura (1997) Chicago. USA soft to firm silty cla y (Chicaqo clay) shield 3.7 10.7 18-30 5.0 31 Shirlaw et al (1988) Singapore. Malaysia southbound tunnel, line 7 very stiff to hard clay, w ith boulders sprayed concrete lining '6.0 20.0 6 10.1

Offset to point of inflection, i (m) Figure 18 confirms the conclusion o f O ’Reilly and N ew (1982) 5 10 15 20 ■ that for the majority of cases i = 0.5^ for practical purposes, irrespective o f whether the tunnel is in soft or stiff clay. There is some scatter in the data, generally within the envelope bounded by E, i = 0.4Z(, and i = 0.6z„. The expression i = 0.5z„ for tunnels in clays is reasonably consistent with the findings o f Fujita (1981), who examined data from a large number o f case histories in Japan for tunnels constructed using hand mined shields, blind shields, slurry <1> shields and EPB shields. Fujita generally confirmed the conclusion C C of O’Reilly and New (1982) that the width of the surface 3 settlement profile above tunnels in clays is independent of construction method. f It should be noted that equations (2) and (5) are normally a applied to the immediate surface settlements associated with tunnel construction. Additional post-construction settlement due to consolidation tends to cause wider settlement troughs (as discussed in Section 5.9) and this complicates the interpretation o f the settlement data. Softer clays are more susceptible to appreciable consolidation settlement, which often develops rapidly and can be difficult to separate from the immediate construction settlement; this may partly explain the observation by Peck (1969) that wider settlement troughs are observed above tunnels in soft clays than in (a) bel ow the water table stiff clays. Boden and McCaul (1974) The data for tunnels in sands and gravels in Figure 19 exhibit O’Reilly et al (1981) O'Reilly et al (1981) rather more scatter than for the tunnels in clays. Tw o o f the data Eadie (1977) points (Peck, 1969; Yoshikoshi et al, 1978), show considerably Yoshikoshi etal (1978) O'Reilly e ta l (1981) wider troughs. Nevertheless, the majority o f the data fall w ithin the 11 Peck (1969) 15 Moh etal (1996) bounds o f i = 0.25z„ and i = 0.45z„, with a mean line o f i = 0.35z,,. 16 Cording and Hansmire (1975) Cording (1991) has noted that the width of the transverse (b) above the water table or dewat ered settlement trough above tunnels in granular soils depends to some O 2 Butler and Hampton (1975) extent on the magnitude o f settlement, with larger settlements □ 5 MacPherson (1978) a 7 Yoshikoshi et al (1978) tending to cause a narrower overall width o f the trough consistent O 9 Chambosse (1972) with the “chimney” mechanism in sand shown in Figure 13. Figure v 10 Vinnel and Herman (1969) * 11 Peck (1969) 19 shows that there appears to be no significant difference between © 12 MacPherson (1978) tunnels below or above the water table, contrary to the suggestion E3 13 MacPherson (1978) O 17 Cording and Hansmire (1975) by Peck (1969). Tunnels are often constructed in layered strata com prising both Figure 19. Variation in surface settlement trough w idth parameter clay and granular soils. It has been suggested by Selby (1988) and with tunnel depth for tunnels in sands and gravels New and O ’Reilly (1991), that the equations for tunnels in clays

2360 Table 2. Details o f tunnels in sands and gravels for which data are plotted in Figure 19 (based on Lake et al, 1992)

Tunnel Depth to tunnel Inflection no. Source Location Ground conditions Excavation methods diameter (m) axis, zo (m) Smax (mm) offset, i (m) 1 Bodon and McCaul (1974) London, UK sand, medium to coarse, some gravel bentonite shield 4.1 10.1 22 5.0 2 Butler and Hampton (1975) W ashington. USA 2a gravel and silty sand shield with bucket digger 6.5 14.4 96 2.9 2b Lafayette Square gravel with interbedded silty sand, sa shield with bucket digger 6.4 11.6 113 4.5 3 O'Reilly et al (1980) Warrington. UK 3a Lumb Brook sewer sand, loose to medium with some gra vel shield, hand excavated 3.6 4.7 78 2.4 3b sand, medium to dense w ith some clay, cover of ver shield, hand excavated 3.6 9.0 19 2.5 4 O'Reilly et al (1980) W arrington, UK (mersey st.) sand, loose and silty shield, hand excavated, compressed air 2.0 8.4 28 3.2 5 MacPherson (1978) Washington. USA 5a section a sand, medium dense silty with interbedded sandy cla shield, articulated w ith digger arm 6.4 20.9 6 5.1 5b section b sand, medium dense silty w ith interbedded sandy cla shield, articulated w ith digger arm 6.4 23.0 3 5.1 6 Eadie (1977) Ayrshire. UK 6a Ayrshire joint drainage scheme sand, silty, fine to medium shield, hand excavated 2.9 5.1 14 1.4 6b loose silty sand with little gravel shield, hand ex cavated 2.9 5.7 16 1.6 7 Yoshlkoshi et al (1978) Tokyo. Japan 7a site II sand, fine silty open shield 3.0 8.5 11 7.0 7b site III sand, fine silty blind shield 3.7 22.1 32 8.2 7c site IV sand, fine silty blind shield 3.9 10.9 19 4.1 8 O'Reilly et al (1980) Warrington. UK 8a Acton Grange sewer sand, fine to medium, fairly uniform, occasional gr av bentonite shield 2.9 6.8 20 1.8 8b bentonite shield 2.9 6.8 14 2.0 8c bentonite shield 2.9 6.0 42 1.1 8d bentonite shield 2.9 6.0 19 1.8 8e bentonite shield 2.9 5.8 25 1.6 8f fully stabilized bentonite shield 2.9 6.9 2 3.8 9 Chambosse (1972) Frankfurt. Germany 9a Franfurt metro T-9 sand with some limestone and clay marl lenses shield 6.5 12.4 70 4.9 9b Frankfurt m etro, Dominikanerqa sand w ith some limes tone and clay marl lenses shield 6.5 10.5 140 3.9 10 Vinnel and Herman (1969) Brussels. Belgium sand, unif orm (top half of face), clayey sand to inver shield , hand excavation 10.0 16.0 150 5.5 11 Peck (1969) Toronto. Canada sand,dense shield, hand excavated 5.4 12.0 85 2.0 sand at crown, invert in till shield, hand excavated 5.4 10.7 12 7.6 12 MacPherson (1978) Washington. USA 12a sand, medium dense silty w ith interbedded sandy cla shield, articulated with digger arm 6.0 11.0 46 3.5 12b sand, medium dense silty w ith interbedded sandy cla shield, articulated with digger arm 6.0 11.0 56 5.9 13 MacPherson (1978) Illinois, USA 13a Rockford, III sand, medium dense with some gravels mechanical shield 3.0 10.8 25 2.0 13c mechanical shield 3.0 14.4 46 3.2 13d mechanical shield 3.0 13.8 23 4.1 13e mechanical shield 3.0 10.8 38 2.0 13« mechanical shield 3.0 10.2 15 2.8 15 Moh et al (1996) Taipei, Korea sand, silty (upper 1/3 of tunnel face),lower in silty cl EPBM 6.1 18.5 26 7.4 16 Cording and Hansmire (1975) W ashington, USA 16a A line sands and gravels mechanical shield 6.4 14.6 76 5.4 16b B line sands and gravels, medium dense, interbedded w ith m echanical shield 6.4 14.6 139 4.2 16c C line sands and gravels, medium dense, interbedded w ith m echanical shie'd 6.4 14.6 152 4.5 17 Cording and Hansmire (1975) W ashington, USA sands and qraveis shield 6.4 11.6 280 1.9

and tunnels in sands be simply combined taking acco unt o f the Trough width parameter, K = thicknesses o f the different strata, so that for a tw o-layered system (20-Z) 0.0 0.5 1.0 1.5 2.0 = K \z i K 2z2 (6) 0.0 £ where K, is the trough width factor for the soil type in layer 1 o f 0.2 thickness z,, and K2 is the trough width factor for the soil type in layer 2 o f thickness z,. Field observations o f surface settlement profiles above stratified soils where the tunnel is in sands overlain 9 ° 0.4 by clay layers (e.g. Ata, 1996; Atahan et al, 1996) indicate wider • profiles than would be obtained if the tunnels were only in sands. ^ ° There is less evidence, however, o f cohesionless layers overlying 0.6 ß tunnels in clays causing a narrowing of the surface settlement T\ s equation 8 profile, as implied by equation (6). Indeed, centrifuge model X > / M A / studies by Grant and Taylor (1996) indicate that in the case o f a • tunnel in soft clay overlain by sand, the surface settlement profile X is wider than would be the case if the tunnel were only in the soft i.yj -*------clay. This is probably a consequence o f the overlying sand layer being significantly stiffer than the soft clay, and also the reduced influence of movements in the sand resembling a “chimney” Location Soil type D(m) Zo(m) Reference • Green Park London Clay 4.2 29 Attewell and Farmer (1974) failure mechanism, which can affect the overall settlement trough. ▲ RegertsPark London Clay 4.2 20 Barrai and Tyler (1976) (ncrtrtxxfxO

5.4 Subsurface Settlements ▼ RegertsPark London Qay 4.2 34 Barrat and Tyler (1976) (soutttxxnd)

■ Wellington Quay Soft day 4.3 13.5 Q osscp (1978) In the urban environment, constraints o f existing tunnels and deep ♦ Heathrcw Express London Clay 11.3 22 New and BcMers (1994) foundations often result in new tunnels having to be constructed X S t James's Park London Qay 4.85 31 Nyren (1998) close beneath such structures. It is becoming increasingly (westbound)

O Centrifuge* Soft day 006 0.13 Mair (1979) important to predict how subsurface settlement profiles develop modei 2DP

and how they relate to surface settlement troughs. Mair et al (1993) □ Centrifuge* Soft day 006 022 Mair (1979) analysed subsurface data from various tunnel projects in stiff and Model 2DV soft clays, together with centrifuge model test data in soft clays. ‘ Models tested at 75g equivalent fJI-scale D = 4 5m. Zq = 98m(2CP). 16.5m (20V) They showed that subsurface settlement profiles can also be reasonably approximated in the form o f a Gaussian distribution in Figure 20. Variation o f trough width parameter K with depth for the same way as surface settlement profiles. A t a depth z below the subsurface settlement profiles above tunnels in clays ground surface, above a tunnel at depth z„, the trough width (after M air et al, 1993) parameter i can be expressed as

i = K(z0 - z) (?)

2361 and the value o f K increases with depth as shown in Figure 20. This shows that subsurface settlement profiles at depth are significantly wider than would be predicted by assuming a constant value o f K. Figure 20 is taken from Mair et al (1993) with some additional data points obtained from the Heathrow Express trial tunnel (New and Bow ers, 1994), and the Jubilee Line Extension project in London (Nyren, 1998). The expression for K proposed by Mair et al (1993), and shown in Figure 20, is 0.175 + 0.325(1 - z/z„) (a) Vectors directed tow ards axis (Attewell, 1978; O'Reilly and New, 1982) (Taylor, 19 95b) K ------^ (8) 1 - z/zn Figure 22. Direction o f ground displacement vectors above tunnels in clays Similar observations o f subsurface settlement profiles above tunnels in silty sands below the water table in Taipei were made by around tunnels in clays. A variation of K with depth, such as Moh et al (1996), and above a tunnel in loose sands overlain by a illustrated in Figure 20 and defined in equation (8), affects the firm to stiff clay layer (Dyer et al, 1996). Figure 21 shows the related vertical and horizontal strains. For a constant volume variation with depth o f K (defined as in equation 7) presented by condition, applicable to tunnelling in clays, it turns out that for this Dyer et al, together with the variation derived from the data variation in K the displacement vectors should be directed towards reported by Moh et al. A similar pattern o f increasing K with depth a point on the tunnel centre line 0.175 Zq/0.325 below tunnel axis is evident, as observed for tunnels in clays. level (Taylor, 1995b), as shown in Figure 22b. This gives horizontal movements o f 65% o f those that would be obtained by K = i/(z0 - z) assuming the ground movements to be directed towards the tunnel 0 0 .2 0.4 0.6 0.8 1 1.2 axis. Deane and Bassett (1995) analysed subsurface movement measurements for two sections o f the Heathrow Express trial tunnels in London Clay. They concluded that the displacement vectors were directed towards a point midway between tunnel axis level and invert level in one case, and towards a point at (or possibly even below ) the invert in the second case. For tunnels in sands, even the assumption o f ground movements being directed towards the tunnel axis may lead to significant underestimates o f horizontal ground movement at the ground surface near the edge o f the settlement trough (Cording, 1991). This was also observed by Hong and Bae (1995) for a 10m diameter NATM tunnel in predominantly sandy strata in Korea. Their data are presented in Figure 23, together with equation (10). The distribution is in reasonable agreement with equation (10) Figure 21. Variation o f trough width parameter K with depth for except in the region near the edge o f the settlement trough (2i < y subsurface settlement profiles above tunnels in sands < 3i). However this is usually of little practical significance because the magnitude o f both horizontal and vertical ground surface movements are generally very small near the edge o f the 5.5 H orizontal Movements settlement trough. Subsurface horizontal ground movements at tunnel axis level in Damage to structures and services can arise from horizontal stiff clays are generally inwards towards the tunnel, when open movements. However, there are relatively few case histories of face tunnelling methods are used. Figure 24 shows measurements tunnels where horizontal ground or structure movements are for 5 tunnels o f approximately 4m in diameter constructed in measured. Attewell (1978) and O ’Reilly and New (1982) proposed London Clay and lined with segments; 4 of the tunnels were that, for tunnels in clays, ground displacement vectors are directed constructed using a shield, and one of them by hand methods towards the tunnel axis, as shown in Figure 22a. This leads to the without a shield. For undrained and axisymmetric conditions, the simple relation: ground movement 6 at a radius r would be proportional to 1/ r;

shh = — s v (9) zo

This assumption leads to the distribution o f surface horizontal ground movement given by

= 1.65 y exp (10) 2 i2

The theoretical maximum horizontal movement, ShlIlax, occurs at .the point o f inflexion o f the settlement trough and is equal to 0.61KSmax. This is consistent with field observations by Cording and Hansmire (1975) and Attewell (1978). Assuming equation (9) to be valid, and if it is also assumed that the trough width parameter K is constant with depth, the vertical lateral distance from tunnel centreline, y/i and horizontal strains determined by differentiating expressions for vertical and horizontal movement are equal and opposite. This is Figure 23. Distribution o f horizontal ground surface movements a necessary condition for undrained (constant volume) movements above a tunnel (after Hong and Bae, 1995)

2362 settlement trough. They found that the surface settlement directly above the tunnel face generally corresponds to about 0.5Smax for tunnels constructed in stiff clays without face support. However, for tunnels constructed in soft clays with face support provided by compressed air, the surface settlement directly above the tunnel face was considerably less than 0.5Smax. Pressurized face tunnelling tends to restrict ground settlements developing ahead o f the tunnel face (indeed significant heave can be observed in soft clays). Field observations o f settlements above EPB or slurry shield tunnelling machines indicate that the majority o f the construction settlement is associated with the tail void (component 3 in Figure 14) and that the surface settlement directly above the tunnel face is generally Figure 24. Horizontal movements at axis level of tunnels in much less than 0.5Smax. London Clay (Mair and Taylor, 1993) Settlement observations on the centre-line above a 6.05m diameter EPB shield in loose silty sands and soft clay in Taipei reported by Moh et al (1996) are shown in Figure 26. Settlements Figure 24 shows the movements plotted normalised by tunnel are shown for the ground surface and at three depths. Most o f the radius (Mair and Taylor, 1993). Zero movement is implied at construction settlement is associated with the tail void; very little distances in excess o f 1.5 tunnel diameters beyond the tunnel settlement has occurred at the time when the tunnel face is directly boundary at tunnel axis level. beneath the instrumentation. Similar observations for EPB and In the case o f EPB tunnelling in soft clays, subsurface horizontal slurry shield tunnelling in predominantly sands or silts are reported movements at tunnel level may be either inwards or outw ards, by Nomoto et al (1995) in their survey on Japanese shield depending on the bulkhead pressure. Clough et al (1983) measured tunnelling. Figure 27 shows surface settlement observations above horizontal movements at 4 instrument lines for a 3.7m diameter a 9.48m diameter slurry shield in Cairo at a depth o f about 16m in EPB shield in San Francisco Bay Mud. Where the bulkhead medium to dense sands overlain by a clay layer (Ata, 1996). The pressures were high, initial outward movements exceeded the settlement above the tunnel face was found to be in the range o f subsequent inward movements into the tail void, whereas the 0.25-0.3Smlu. reverse was the case elsewhere. Both inward and outward It can be concluded that the longitudinal settlement trough movements around EPB shields in soft clay are also reported by having the form o f the cumulative probability curve illustrated in Fujita (1994). Figure 25 is generally reasonable, but it has only been validated for tunnels in clays. The surface settlement equal to 0.5Sm„ above the 5.6 Longitudinal Settlement Trough tunnel face is strictly only applicable to open-faced tunnelling techniques in stiff clays. Where there is significant face support, as In the urban environment there may be cases where a structure in EPB or slurry shield machines, the major source o f ground close to or directly above the tunnel centre-line might experience movement is further back from the face and this leads effectively more damage from the progressive longitudinal settlement trough to a translation o f the cumulative curve, as shown in Figure 25. generated ahead o f the tunnel face, as shown in Figure 25, than from the final transverse settlement profile after the tunnel face has 5.7 Volume Loss passed beneath the structure. Also, construction works may comprise various intersecting tunnels o f short lengths in different The magnitude o f volume loss, V h (defined in equations (3) and (4) directions and o f different diameters (for example an underground depends principally on the type o f ground and on the tunnelling station complex); in such cases the development o f settlement method. Many authors have reviewed volume loss values from ahead o f any particular tunnel face is o f importance, because the tunnelling projects (e.g. Peck, 1969; Cording and Hansmire, 1975; effects o f a number o f finite length tunnels can then be estimated Clough and Schmidt, 1981; O’Reilly and New, 1982; A ttew ell et and summated. al, 1986; Uriel and Sagaseta, 1989; Mair, 1996). It follow s from the assumption that the transverse settlement For tunnels in clays, Clough and Schmidt (1981) proposed a profile has a Gaussian shape that the longitudinal profile should relationship between stability ratio (or overload factor) N and V, have the form o f a cumulative probability curve, assuming all based on the closed form solution for the unloading o f a circular ground deformation takes place at constant volume (New and cavity in a linear elastic-perfectly plastic continuum under O’Reilly, 1991), which is applicable to tunnelling in clays. By axisymmetric conditions. Attewell et al (1986) and Uriel and examining a number o f case-histories of tunnel construction in Sagaseta (1989) presented field data o f volume losses related to clays, Attewell and Woodman (1982) showed the cumulative stability ratio, based on Clough and Schmidt’s proposal; the results probability curve to be reasonably valid for the longitudinal

Days after the Passing o f the Head translated longitudinal settlement profile - original ground level with face support

longitudinal settlement profile (cumulative advancing tunnel - probability form) - without face support

' tunnel face

Figure 25. Longitudinal surface settlement trough Figure 26. Settlements above EPB shield in silty sands in Taipei (Moh et al, 1996)

2363 10 TBM □ FROM TRANSVERSAL SECTIONS tl X FROM POINTS ON THE AXIS

•: 8 _ 1 1 ____ o I5IT.T31 SIT.T2 | SIT. Tl '*X. r t I r-SURFA Ci «-AA < 6 UJ * 'TERTIARY/ Uj LIJJ- ?T o> »A z UJ '/ / K. * //// UJ

Distance from TBM (m) 0 0 0,5 1,0 1,5 2,0 Figure 27. Settlements above EPB shield in sands in Cairo (Ata, 1996) Ht /D Figure 28. Influence on volume loss o f cover o f com petent soil in mixed ground conditions (M elis et al, 1997) show a very wide scatter. The wide scatter is probably associated with many construction details and differing standards of workmanship. Another important factor is the value o f undrained 2. Construction with sprayed concrete linings (NA TM ) is effective shear strength assumed in the calculation o f stability ratio, N. In a in controlling ground movements. Recent construction in London number of cases, the influence of sample disturbance and the Clay, for example, has resulted in volume losses varying from method o f laboratory testing (usually triaxial compression tests) 0.5%-1.5%, which compares favourably with well-controlled may have led to erroneous values o f su being adopted. Extension shield tunnelling in which there is little or no face support. stress paths on vertically orientated specimens are generally more 3. For closed face tunnelling, using EPB or sluny shields, a high relevant to the unloading o f the ground around and above a tunnel degree o f settlement control can be achieved, particularly in sands than compression stress paths; in soft clays, triaxial extension tests where volume losses are often as low as 0.5%. Even in soft clays, typically give significantly lower undrained shear strengths than volume losses (excluding consolidation settlements) of only 1%- triaxial compression tests. 2% have been reported. Based on centrifuge model test data and finite element analyses, Volume losses may be higher in mixed face conditions for EPB Mair et al (1981) and Mair (1989) proposed that V, should be more or slurry shields, particularly where sands or gravels overlie stiff properly related to the Load Factor, defined as N/ Nc (Nc being the clays, or where the cover o f competent soil above the tunnel crown critical stability ratio), rather than to N alone. Despite the different is low. This is illustrated by Melis et al in their paper to this stress histories and C/ D ratios, which meant that the volume losses Conference, as shown in Figure 28. A 7.4m diameter EPB shield were markedly different at the same stability ratio, there was a was used in very stiff to hard sandy clays in Madrid with varying reasonably well-defined relationship between volume loss and load cover (Ht) to overlying sands or fills. The EPB was used in open factor. O’Reilly (1988) used the approach to provide good face mode when the ratio HT/D exceeded 0.6, and they state that predictions o f volume loss at 6 different tunnelling sites in London the observed volume losses were in the range 0.03-1%. Larger Clay. volume losses were observed for lower HT/D ratios, and When closed face tunnelling methods are employed, using EPB significantly larger values (generally 2-4% ) were obtained for or slurry shields, good control o f the face pressure can result in the mixed face conditions. stability ratio being close to zero, in which case the component of ground movement resulting from stress relief at the face would be 5.8 M ultiple Tunnels very small, leading to smaller volume losses (less than 1%). In such cases in soft clays the principal cause of volume loss is When two or more tunnels are constructed it is commonly assumed usually the tail void (Broms and Shirlaw, 1989). that the ground movements that would have occurred for each Recent experiences with EPB and slurry shield machines in tunnel acting independently can be superimposed. Fo r two tunnels sands and gravels have generally shown small volume losses. constructed side by side at the same level this would lead to a Leblais and Bochon (1991) report volume losses in the range 0.2- symmetrical combined surface settlement profile. Cording and 0.9% for 9.25m diameter tunnels driven through dense fine Hansmire (1975) presented evidence o f asymmetry above twin Fontainebleau sands at depths ranging from 22m to 52m; values tunnels o f 6.4m diameter with a clear separation o f 4.6m driven o f 0.8-1.3% were observed when the tunnels were very shallow through medium dense silty sands and gravels for the Washington with the tunnel crown being only 4.1-7.2m below the ground Metro. A degree o f asymmetry o f the surface settlement profile was surface. Volume losses reported by Ata (1996) for a 9.48m also reported by Standing et al (1996) in the case o f twin tunnels o f diameter slurry shield in Cairo at a depth o f about 16m in medium 4.8m diameter constructed in London Clay with a clear separation to dense sands below the water table were in the range 0.2-1%, o f as much as 16m. Asymmetric surface settlement profiles above with a mean o f about 0.5%. twin tunnels are also reported by Lo et al (1987). In a review of 35 papers submitted to the recent TC 28 Interaction o f tunnels in close proximity was observed by Symposium in London on the subject o f settlement associated with Shirlaw et al (1988) for tunnels constructed using sprayed concrete bored tunnels (Mair, 1996), the following main conclusions were linings in a very stiff to hard clay in Singapore, as shown in Figure drawn: 29. The clear separation o f the 6m diameter tunnels was only 1,7m. 1. For open face tunnelling, volume losses in stiff clays such as The volume loss observed for the first (south bound) tunnel was in London Clay are generally between 1% and 2%. the range 0.5-1%, and the surface settlement trough width

2364 Time (days)

Figure 30. Long term settlement observations above tunnel in soft clay (after O ’Reilly et al, 1991)

settlement was observed for a considerable period w ith final equilibrium being achieved only after about 10 years. The corresponding transverse surface settlement profiles for one o f the arrays are shown in Figure 31, which shows a significant widening Distance from centre point (m) Distance from centre point (m) o f the profile with time. Similar widening has been reported by a number o f authors (e.g. Glossop, 1978; Howland, 1980; Shirlaw, 1995). The major factors influencing the development of post­ construction settlements above tunnels are as follows: 1. The magnitude and distribution o f excess pore pressure, Au, generated by construction o f the tunnels. 2. The compressibility and permeability o f the soil. 3. The pore pressure boundary conditions, particularly the • Southbound tunnel only permeability o f the tunnel lining relative to the permeability o f the x Northbound tunnel only soil. Figure 29. Settlement troughs for closely spaced tunnels in 4. The initial pore pressure distribution in the ground prior to Singapore (Shirlaw et al, 1988) tunnel construction. The magnitude and distribution o f Au depends to a large degree on whether the ground is unloaded during construction, as is often parameter i was found to be close to 0.5^. However, in the case o f the case for open face tunnelling, or whether it is subject to an the second (north bound) tunnel the volume loss observed was increase in loading, such as in slurry shield or EPB tunnelling generally much larger, in the range 2-4%, and the settlement when the face pressure exceeds the in-situ stresses or when troughs were unexpectedly wide for the shallower tunnel (i/ z„ excessive grouting pressures are used. Clough et al (1983) values o f about 1). observed evidence o f high excess pore pressures being generated Significant interaction effects are evident. It is clear that, when in soft clay near the spring line o f the tunnel as the EPB machine tunnels are very closely spaced, the ground in the region where the approached. The complexity of the dependency of the pore second tunnel is to be constructed will already have been subjected pressure response on details o f the construction process is clearly to appreciable shear strains associated with construction o f the first illustrated in Figure 32 for a 4.2m diameter tunnel constructed at tunnel, resulting in reduced stiffness, and hence a higher volume a depth o f 5.6m in soft clay in Shanghai with an EPB shield (Yi et loss is likely for the second tunnel. In the case o f 5.6m diameter al, 1993). The observed variation in excess pore pressures at tunnel tunnels in weathered schists, sands and gravels in Caracas, volume axis level with distance from the tunnel is shown for 9 different losses for the second tunnel were generally in the range 80-125% positions o f the shield face in relation to the instrumented section. larger than for the first tunnel, for a clear spacing o f between 0.5 The excess pore pressures are seen to fluctuate in response to the and 1 tunnel diameter (Perez Saiz et al, 1981). position of the shield. An increase of face pressure (between positions 5 and 6) and grouting o f the tail void (between positions 5.9 Post-construction Settlements 7 and 8) each resulted in significant increases in positive excess pore pressures. The detailed ground movement observations show Post-construction settlements can be significant, particularly in the that the subsequent consolidation settlements associated with case of tunnels in soft, compressible clays. These arise from dissipation o f these positive excess pore pressures resulted in an changes o f pore pressures (and hence effective stresses) following additional settlement trough o f similar width to the immediate construction of the tunnel, and take the form of increasing settlements (sometimes referred to as consolidation settlements) Distance to £ in metres but generally with very little increase in the horizontal component o f the ground movements. A comprehensive review o f field data 20 18 16 14 12 10 8 6 4 2 0 o f consolidation settlements above tunnels in soft clays by Shirlaw (1995) concluded that typically the increase in settlement over the long term is o f the order o f 30-90% o f the total settlement, and that in many cases a widened settlement trough develops. This is illustrated by settlement observations over a period o f 11 years reported by O’Reilly et al (1991) for a 3m diameter tunnel constructed in normally consolidated silty clay in Grimsby. The centre-line settlement at one o f the locations is shown in Figure 30. The tunnel was constructed with compressed air to provide the necessary face support for stability. On removal o f the compressed air approximately 100 days after tunnel construction, further tunnel in soft clay (O ’Reilly et al, 1991)

2365 Legend.

Driving of Up Track Tunnel 15 ( not true peak values ) Driving of Down Track Tunnel

Damaged before peak was reached 10

Probable Distribution

iÜ

— T—8-a- 10 15 20 25

Distance ( m ) Figure 34. Excess pore pressures observed at axis level of EPB tunnel in soft silty clay in Taipei (Hwang et al, 1995)

troughs, i.e. the incremental settlements developing after App roaching M oving away completion o f the tunnel. The data are shown in normalized form. Instrum ented section Three o f the case histories (the Furongjuang Sewer Line B, the 4- Singapore Subway Line TA 265-273 and Thunder Bay) involved (D tunnel construction methods where there was very significant face

-U -12 -10 -8 -6-4-2 0 2 U 6 8 10 (m) support, either by an EPB machine or a full-face TBM . This caused

D istance betw een shield face and instrum ented secti on appreciable positive excess pore pressures to be generated, and in Figure 32. Excess pore pressures observed during passage o f EPB all three cases dissipation o f these excess pore pressures resulted shield in soft clay in Shanghai (Yi et al, 1993) in post-construction settlement troughs very similar to the classical Gaussian curve associated with short-term settlement (assuming i=0.5zo). trough caused by tunnel construction, rather than a wider trough as The other 4 case histories on Figure 35 have much w ider post­ was observed at Grimsby (see Figure 31). Similar behaviour was construction settlement troughs. These are associated with the reported by Shirlaw and Copsey (1987) and Shirlaw (1995) for tunnel lining acting as drain and the development o f steady state EPB tunnels in soft clay in Singapore, where the face pressures seepage towards the tunnel. The result is a widespread reduction o f were 1.2-1.5 times the overburden pressure. pore pressures around and above the tunnel leading to Similar observations o f positive excess pore pressures around consolidation settlements developing over a wide area. These wider tunnels where the ground was subjected to an increase in loading consolidation settlement troughs often result in no additional are reported by Hwang et al (1996), who contrast the pore pressure damage to buildings, because the differential settlements associated responses around EPB tunnels in silty sands and silty clays in with consolidation are very small. Taipei. They concluded that, provided the face pressures are close Theoretical analyses o f pore pressures induced around tunnels to the in-situ ground stresses, the predominant source o f excess in clays where there is unloading (e.g. Clough and Schmidt, 1981; pore pressures induced in the ground is the grouting o f the tail Schmidt, 1989; Samarasekera and Eisenstein, 1992; M air and void. Whilst in sands these excess pore pressures dissipate rapidly, Taylor, 1993) demonstrate the dependence o f the magnitude and in clays they may take months to dissipate fully. A t another site in distribution o f the excess pore pressures on the degree o f unloading Taipei, shown in Figure 33, Hwang et al (1995) recorded positive and the strength and stress history o f the clay. In open face excess pore pressures around 6.05m diameter tunnels in soft silty tunnelling in clays when there is significant unloading of the clay (su = 30-40 kPa), as shown in Figure 34. The measured excess ground, the excess pore pressures induced during tunnel pore pressures at tunnel axis level were limited to a zone extending construction are generally always negative; this is in response to about 6m, or one tunnel diameter, from the edge o f the tunnels; a the unloading and, in the case o f overconsolidated soils, to a similar pattern was observed by Yi et al (1993), as indicated in Figure 32. It is evident from the above examples that two distinct patterns distance from centre line V i depth to tunnel spring line of behaviour can be distinguished for long term consolidation settlement troughs for tunnels in soft clay, as shown on Figure 35 (Shirlaw, 1995). This shows the post-construction settlement

JUR0NCJIAN6 SEWER LIKE 8 . EPS . 4.13« DIA.

SINGAPORE SUBWAY LINE T A !6 5 - !7 3 . EPB . 5.93 m DIA.

THUNOER SAT ARRAY I . TBM. 2.47» DI*.

GRIMSBY ARRAY A . OPEN SHIELD . 3 .0 0 » DIA.

GRIMSBY ARRAY B l. OPEN SHIELD a COMPRESSED AIR . 3.00 m DIA.

SINGAPORE SUBWAY UNE TA94-IO O . OPEN SHIELD 8 COMPRESSED AIR . 5.93» 0IA. TWIN TUNNELS .

.JAPAN SEWER IHANYA) B2 , TBM ft COMPRESSED AIR . 7.05 m DIA. TWIN TUNNELS .

.GAUSSIAN CURVE WITH I . 0.5 i DEPTH TO SPRIN6 LINE .

Figure 35. Normalized post-construction surface settlement Figure 33. Piezometer array around EPB tunnel in soft silty clay in troughs due to consolidation o f soft clays (Shirlaw , Taipei (Hwang et al, 1995) 1995)

2366 tendency for dilation to occur. An example of pore pressure reduction close to a 8.7m diameter tunnel lined with sprayed concrete at a depth o f 21m in London Clay is shown in Figure 36 (New and Bowers, 1994). For normally consolidated clays, significant zones o f positive excess pore pressure can be generated even for a tunnel where unloading occurs, as shown by Schmidt (1989); negative excess pore pressures are induced near the tunnel, but positive excess pore pressures due to shearing may result a short distance away. In overconsolidated clays, how ever, the excess pore pressures generated are generally always negative when the ground is unloaded during tunnel construction. The permeability o f the lining relative to the soil determines -45 ■ whether or not the tunnel acts as a drain. Ward and Pender (1981) -50 ------1------1------1. i i____ i____ i____ i____ i____ concluded that in most cases segmentally-lined tunnels in London -15 -10 -5 0 5 1 0 15 20 25 30 35 Clay acted as drains, despite the linings having been grouted; the Horizontal offset from centreline (m ) permeability o f the London Clay at the depth of the tunnels is • Long term data — Fitted Gaussian curve typically in the range 10"10 -10 " m/ sec. In contrast, O ’Reilly et al +• Short term data Fitted Gaussian curve (1991) found that long term piezometric observations showed no evidence o f reduced pore pressures close to the tunnel in Grimsby Figure 37. Immediate and post-construction surface settlements for which the long term consolidation settlements are shown in above tunnel in London Clay (Bow ers et al, 1996) Figures 30 and 31. In a back-analysis using finite elements, the closest match to these observed consolidation settlements was deflection ratio. Bow ers et al show that there is also very little obtained by assuming the permeability o f the combined primary change in horizontal strain in the 3 year period. segmental lining and secondary in-situ concrete lining to be 5 x In summary, the following conclusions can be drawn about 1 O'1'm/ sec (Mair et al, 1991). The permeability o f the clay at post-construction settlements: Grimsby deduced from in-situ constant head tests was about 10'10 1. When tunnelling in soft clays, particularly with EPB shields, m/ sec. Negro (1994) draws attention to the influence o f tunnel significant positive excess pore pressures may be generated. These lining imperfections on water infiltration into tunnels. It is clear can be induced through shearing even when unloading of the that in many cases tunnel linings in clay soils may act as drains, ground occurs in terms of reduced face support, but the most either fully or partially. In such cases, the development o f steady marked positive excess pore pressures are observed when over- state seepage towards the tunnel may result in a widespread pressurization o f the face takes place or when tail void grouting reduction o f pore pressures around and above the tunnel, with pressures are high. associated consolidation settlements developing over a wide area 2. These positive excess pressures are generated lo cally in the (Howland, 1980). ground immediately surrounding the tunnel, generally within about An example o f a wider consolidation settlement trough above one tunnel diameter. If the tunnel lining is o f low permeability a tunnel in stiff clay is given by Bow ers et al (1996). They present relative to the clay, consolidation settlements are only associated post-construction settlement data obtained over a 3 year period for with dissipation o f these local excess pore pressures, resulting in an 8.7m diameter tunnel at a depth o f 21m in London Clay, the an additional settlement trough o f similar width to the immediate tunnel lining being sprayed concrete. Figure 37 shows surface trough caused by construction o f the tunnel. settlement measurements obtained immediately after construction 3. In stiffer clays only negative excess pore pressures generally (short term data) and 3 years later just before installation o f the result from tunnel construction. If the tunnel lining is of low secondary concrete lining (long term data). There is a relatively permeability relative to the clay, swelling rather than consolidation uniform settlement increase across the whole settlement profile in would be expected, resulting in no discernible post-construction the 3 year period, similar to the observations above the tunnel in surface settlements. soft clay at Grimsby shown in Figure 31. In particular it should be 4. If the tunnel lining is permeable relative to the permeability o f noted that there are only very small increases in distortion or the clay, the tunnel acts as a drain and the resulting consolidation settlements lead to a significantly wider surface settlement trough than the short term trough associated with construction o f the tunnel. Very small increases in distortion, deflection ratio and horizontal strain are observed.

6. MODELLING AND PREDICTION OF GROUND MOVEMENTS

6.1 Em pirical Method

The method most commonly used to predict ground movements associated with tunnelling is based on the Gaussian distribution proposed by Schmidt (1969) and Peck (1969), as defined in equation (2) o f this Report. Assumptions are made about the width o f the settlement trough and the volume loss, both determined by empirical means. The width o f the settlement trough is determined Time from the parameter K, which is a function o f the tunnel depth, as Movement (IB2) Stress (SB2) Porewater pressure (SB2P) shown in Figures 18-21. The magnitude o f volume loss depends on a number of factors, as discussed in Section 5.7, and its Figure 36. Pore pressure and total horizontal stress changes selection requires considerable experience and judgement. This together with horizontal ground movement during approach is often referred to as the empirical method. At its tunnel construction in London Clay (New and Bow ers, simplest, it is used in two dimensions to predict the magnitude and 1994) shape o f the transverse surface settlement profile. The method can

2367 be easily extended to predict surface ground movements in three heading is fully three-dimensional (see Figure 15). Clough and dimensions by incorporating the longitudinal settlement trough in Leca identified a number of reasons hindering the successful the form of a cumulative probability curve, as proposed by development o f FE analysis for analysis o f this com plex problem: Attewell and Woodman (1982) and discussed in Sectio n 5.6. 1. The cost of a full 3D analysis, which properly simulates Details o f the relevant equations are given by New and Bow ers construction aspects and the 3D geometry, with a realistic non­ (1994) and shown to be capable o f realistic predictions o f surface linear constitutive soil model, is substantial and difficult to justify ground movements. for many tunnelling applications. (This remains the case at present Subsurface ground settlements can also be predicted by despite the rapid increase in computer power and reduction in empirical methods if account is taken o f the observed variation o f costs.) trough width parameter K with depth (see Figures 20 and 21). The 2. Many o f the parameters influencing the results are difficult to horizontal components o f the surface and subsurface ground define, for example tunnel lining properties, tail void size and soil movements can be calculated from the simple assumption that all model parameters (for the particular constitutive model assumed). ground movement vectors are directed towards the tunnel axis, or 3. Multiple analyses would often be required for any given project, as a refinement, towards a point below the axis for tunnels in clays, in view o f the usual changes o f geology and alignment geometry as indicated in Figure 22b. An alternative assumption, in which the along the length o f the tunnel. ground movements are directed towards a ribbon-shaped zone of 4. No constitutive soil model has been shown to be successful at ground at tunnel invert level, has been proposed by New and simulating all aspects o f soil behaviour important to tunnelling. Bow ers (1994) and shown to correlate well with field observations. All o f these reasons still prevail, but considerable progress in FE The limitations o f the empirical method are that it is generally modelling has been made in recent years. Two-dimensional (2D) only applicable to: FE modelling remains at present more common. Many 2D analyses • Single tunnels, or multiple tunnels for which there is no are based on the ground reaction curve concept (Peck, 1969), significant interaction. sometimes referred to as the Convergence-Confinement method • Short term construction ground movements in the case of (Panet and Guenot, 1982). As shown in Figure 38, 3D effects are tunnels in clays (post construction consolidation settlements approximated by reducing a proportion o f the stresses imposed by are less amenable to prediction by empirical methods). the soil to be excavated acting on the tunnel boundary, and then • Greenfield sites (the presence o f structures of significant installing the tunnel lining. This can be simplified by considering stiffness are difficult to take into account). the radial stresses applied to the tunnel boundary, a,. Following the A further limitation of the empirical method is that good concept o f Panet and Guenot (1982), this can be expressed as judgement is required in the selection o f an appropriate value o f volume loss. This requires consideration o f such aspects as the °r = (1 - *) (n) tunnelling technique, geometry o f the tunnel in relation to the where o 0 is the initial total ground stress prior to tunnelling, and A. various ground strata, groundwater conditions and soil properties. is an unloading parameter (0 < X < 1). The stress removed from the Nevertheless, the empirical method has considerable practical soil prior to installation o f the lining is Xo0, and correspondingly value, particularly in instances where there are previous case the stress applied to the installed lining is (1 - X)a0. As the stress is histories o f tunnelling in similar ground conditions using similar construction techniques.

6.2 Closed Form Solutions

Predictive methods based on the closed form solution for unloading o f a circular cavity in a linear elastic-perfectly plastic continuum under axisymmetric conditions have been described by ______c Clough and Schmidt (1981) and Lo et al (1984). Mair and Taylor —► (1993) describe a simple approach for predicting ground deformations ahead o f an advancing tunnel face in terms o f the closed form solution for unloading o f a spherical cavity in a linear elastic-perfectly plastic continuum; the approach w as found to : <*r ! provide a reasonable approximation for prediction o f axial ground \jS 'V movements ahead o f an advancing tunnel face in London Clay. A limitation o f these approaches is the assumption o f axisymmetry, which is often not applicable - particularly for shallow tunnels. Sagaseta (1987) presented a two-dimensional theoretical analysis of ground deformations towards tunnels based on solutions for incompressible fluid flow (the strain path method proposed by Baligh, 1985). The analysis requires an assumption o f a value for the volume loss. The predicted lateral spreading and width o f the surface settlement trough are considerably greater than observed in practice (Schmidt, 1988); a possible way o f accounting for this is proposed by Sagaseta (1988), but it involves assuming volumetric dilatant strains for the soil which is clearly in conflict with the behaviour o f clay soils under undrained conditions.

6.3 Finite Element Analysis

Finite element (FE) analysis offers considerable po ssibilities o f modelling many aspects of bored tunnel construction. A comprehensive review o f FE analysis o f bored tunnel construction Figure 38. Application to 2D FE analyses of principle of was undertaken by Clough and Leca (1989). It is evident that the convergence-confinement method (Panet and Guenot, development o f ground movements around an advancing tunnel 1982)

2368 removed from the tunnel boundary, radial displacements occur Horizontal distance x Tunnel diameter D which are equivalent to volume loss. The resulting volume loss V, is related to the reduction in stress, Xa0. The remaining excavation o f soil continues after installation o f the lining, but in cases where the completed lining ring is installed further changes in soil deformation or lining stresses are generally very small. Two approaches are commonly used. The first is to install the tunnel lining after a prescribed amount o f unloading, i.e. to select an appropriate value o f X. This requires considerable judgement and experience; axisymmetric analyses can usefully assist the selection o f X (e.g. Panet and Guenot, 1982). The resulting volume loss will depend on the value o f A. selected. The second approach is to prescribe the volume loss, V„ and install the tunnel lining after a suitable degree o f unloading to achieve the prescribed volume loss. In this case the resulting lining stress will depend on the prescribed volume loss. Whichever approach is used, it is important that the calculated lining stress (1 - A)o0 and volume loss Horizontal distance x_ V, should be reasonably compatible with field measurements for Tunnel diameter D similar tunnels in comparable ground conditions. They should lie within the realistic range o f values expected for the tunnel in the ground conditions analysed; if they do not, the analysis should be viewed with caution. An alternative method o f analysis, known as the Progressive Softening approach (Swoboda, 1979), involves reducing the stiffness o f the soil within the tunnel area prior to excavation and placement o f the tunnel lining. The amount by which the stiffness is reduced requires considerable experience and judgement. Rowe et al (1983) proposed the concept o f the ‘‘gap” parameter to define the ground displacements that should be prescribed in a 2D FE analysis prior to installing the lining. This is illustrated in Figure 39. The “ gap” represents the physical clearance between the outer skin o f the shield and the lining (Gp) plus an allowance for out-of-plane (3D) ground movements (u3D), together with an allowance for workmanship (co). Lee et al (1992) used 3D elasto- Figure 40. FE predictions by Lee and Rowe (1989) of surface plastic finite element analysis to develop a means o f quantifying settlement troughs observed in centrifuge models of the gap parameter for use in a 2D analysis; these w ere shown by tunnels in soft clay (Mair, 1979) Row e and Lee (1992a) to lead to reasonable predictions o f field observations using 2D analyses. When closed face techniques such as EPB or slurry shields are used to provide full support to the face, Gunn (1993) found that isotropic non-linear elastic perfectly the gap parameter is simply the physical clearance between the plastic “small-strain” stiffness soil models (Simpson, 1979; outer skin o f the shield and the lining (i.e. u3D=0) and thus the 2D Jardine et al, 1986) improved the 2D FE predictions for tunnels in idealization becomes closer to reality. However, where stress relief heavily overconsolidated London Clay tunnels compared with occurs at the face, as in open faced tunnelling, 2D idealizations are linear elastic-perfectly plastic models, but even these predicted inevitably more approximate. wider settlement troughs than observed in practice. In contrast, In 2D FE analysis, considerable attention has been paid to the Simpson et al (1996) reported 2D FE predictions for a tunnel in difficulty in predicting the Gaussian surface settlement distribution London Clay showing the shape o f the settlement trough to be little defined in equation (2). Isotropic linear elastic-perfectly plastic soil influenced by non-linearity but substantially influenced by shear models lead to predictions of much wider surface settlement modulus anisotropy (as suggested by Lee and Rowe, 1989). They troughs than the observed Gaussian distribution (Mair et al, 1981; reported bender element tests on undisturbed London Clay samples Lee and Rowe, 1989; Gunn, 1993). Lee and Rowe (1989) showed and referred to in-situ shear wave tests in London Clay, both of that the use o f anisotropic elastic properties significantly improves which demonstrate significantly anisotropic shear moduli; the in- the prediction. Figure 40 shows their 2D FE predictions of situ shear wave data are shown in Figure 41. centrifuge model tests on tunnels in lightly overconsolidated soft Figure 42 shows a series o f 2D FE predictions by Addenbrooke clay (Mair, 1979). Their isotropic linear elastic-p lastic model et al (1997) o f surface settlement profiles above a 4.75m diameter (Gvh/Ev = 1/ 3) predicts a significantly wider settlement trough than tunnel at a depth o f 34m in London Clay compared with the field anisotropic models with smaller Gvh/Ev ratios. measurements reported by Standing et al (1996). The same volume loss (3.3%), as measured, was prescribed for each o f the predictions. All three predictions used a non-linear elastic, perfectly plastic model, the non-linear elastic (“small strain”) GAP — Gp + U3D + w GAP response being described by Jardine et al (1986). Two of these assumed different degrees o f anisotropy. The non-linear models all predict deeper and narrower troughs, and the predictions improve with increasing degrees o f assumed anisotropy. Addenbrooke et al note, however, that the ratio Gvh/F.' = 0.2 assumed for the higher degree of anisotropy does not appear to be consistent with presently available laboratory or field test data for London Clay, such as presented in Figure 41. The apparent relative importance Figure 39. Definition o f gap parameter (Row e et al, 1983) o f anisotropy may be linked to the constitutive soil model assumed in the analysis.

2369 SHEAR W AVE VELOCITY (m/s)

50 100 150 200 250 300 centre line 25 m o i I

10mm

non-llnear isotropic. K. = 1.5

Brown London Clay non-flnear Isotropic, reduced K,

20 m m (leid data (Standing et al, 1996)

Figure 43. Influence o f K*, on FE predictions o f surface settlement Grey London Clay trough (Addenbrooke, 1996, 1997)

FE analysis o f ground movements during NATM construction in London Clay was reported by Dasari et al (1996), using a strain Figure 41. In situ shear wave tests in London Clay (from Simpson dependent Modified Cam Clay soil model, but the predicted et al, 1996) transverse surface settlement trough was significantly wider than the field observations, as in the 2D FE analyses referred to earlier. As noted by Clough and Leca (1989), NATM poses fewer The effect o f K0 on the predicted width o f surface settlement problems for 3D finite element modelling than does shield trough is surprisingly often neglected (Gens, 1995). The wider tunnelling which is generally a more complex process to model. settlement trough predicted by 2D FE analyses for tunnels in Akagi and Komiya (1996) report a 3D FE analysis of an EPB heavily overconsolidated clays may be linked to K„ generally being shield tunnel constructed in soft clay, using a critical state soil assumed to be significantly greater than 1. Reducing K, locally model. The shield machine and its advance were modelled in what either side o f the tunnel to account for unloading associated with clearly was an extremely complex analysis. Only a few field 3D effects leads to narrower, more realistic settlement troughs measurements were presented for comparison with the predictions, being predicted by 2D FE analyses (Addenbrooke, 1996). Figure but the analysis appeared capable o f reproducing many of the 43 illustrates the influence of K0 on the predicted surface features observed in practice. De Borst et al (1996) describe 2D settlement profile for the same tunnel for which results o f analyses and 3D FE analyses o f a pipe jack tunnel in soft clay. 3D FE are shown on Figure 42. Kq was reduced to 0.5 in a zone o f depth analyses are also reported by Row e and Lee (1992b) and Gioda et D and width 1.5D either side o f the tunnel o f diameter D, and this al (1994). had a marked influence on the shape of the settlement trough 3D FE analyses are reported in several papers to this (Addenbrooke, 1997). Conference, but there are still very few examples o f published 3D Stallebrass et al (1994) carried out a parametric study using a FE analyses of tunnel construction in the field where detailed three-surface kinematic hardening soil model; they concluded that comparisons are made between the predictions and the the recent stress history of the soil, described by anisotropic measurements. 3D FE analysis o f tunnel construction with elasto- unloading or re-loading, has an important effect on the predicted plastic soil models remains a major undertaking despite the recent surface settlement trough. However, the value o f K„ was similar in rapid increases in computer power and the decreases in cost. all their analyses. Nevertheless it is likely to become increasingly common in the In summary, it is clear that sophisticated soil models are future. required to achieve realistic predictions o f the transverse surface A general difficulty in comparing FE analyses with field data is settlement trough using 2D FE analyses, particularly for tunnels in that discrepancies could be due to one or more o f the following: (a) heavily overconsolidated clays. deficiencies in the soil model (b) the soil parameters adopted (c) Clough and Leca (1989) suggest that the use o f 2D analyses to idealizations in the modelling, especially with respect to the represent 3D effects is itself one o f the reasons for the shape o f the boundary conditions, and (d) possible uncertainties in the field settlement trough not being well predicted. Non-linear 3D FE measurements. As discussed by Clough and Leca (1989), the analyses have been undertaken of NATM construction (e.g. tunnel construction process is extremely complex, particularly if Swoboda et al, 1989; Katzenbach and Breth, 1981), but detailed shield tunnelling is involved, and it therefore represents almost the comparisons with surface settlement data were not presented. 3D ultimate challenge to geotechnical analysts.

offset from tunnel axis 6.4 Physical M odelling

25 m 50m Physical models have been used to study ground movements associated with tunnel construction, either to provide detailed data against which numerical prediction methods can be calibrated or in cases where numerical analysis is impractical. Self-weight o f the soil is a major factor influencing tunnel stability and associated 8 10 mm ground deformations, and hence centrifuge model tests are non-Bnear Isotropic appropriate, particularly for shallow tunnels. A review of the _ _ non-flnear anisotropic, m ' = 0.444 application o f centrifuge model testing to tunnels is given by . . . non-llnear anisotropic, m ' ■ 0.2 Taylor (1995a). Tests on tunnels in soft clays to investigate K, ■ 1 & m ' ■ G */E,'

Held data (Standing et al, 1996) stability and ground movement patterns are summarised by Mair et al (1984). Centrifuge model studies investigating the ground Figure 42. Influence o f anisotropy on FE predictions o f surface deformation mechanisms o f tunnels in soft clay overlain by sand settlement trough (Addenbrooke et al, 1997)

2370 7.2 Classification o f Damage and Assessment C riteria Dry sand C/D = 2

The subject of settlement damage to masonry buildings was <2>=1 Ocm addressed by Burland and Wroth (1974) and Burland et al (1977), who introduced a damage classification system, together with the concept o f limiting tensile strain. In an important development, Boscardin and Cording (1989) analysed case histories of excavation induced subsidence and showed that the damage categories put forward by Burland et al are related to the magnitude Tunnel axis o f estimated tensile strain in the building. Ranges o f strain were ■v\ 'V au lt extremity identified for different damage categories, as show n in Table 3. Figure 45 was derived for buildings with length (L) to height (H) Envelope of the ratios o f 1 in terms o f angular distortion (P) and horizontal strain; shear deformation was assumed to be dominant, and bending Face Failure mechanism strains less critical. T777////////// ///////7777 Tunnel invert Burland (1995) and Mair et al (1996) categorised building damage in terms of deflection ratios A/ L for sectio ns o f the Figure 44. Failure mechanism observed in centrifuge model test o f building in hogging or sagging mode, as shown in Figure 46. A tunnel with a pre-vault (Skiker et al, 1994) damage category chart for buildings in a hogging mode with L/ H = 1 is shown in Figure 47; this takes into account strains due to both shear and bending modes and is broadly equivalent to the layers are described by Grant and Taylor (1996). The use o f digital chart produced by Boscardin and Cording, since generally p image analysis, as described by Grant and Taylor (1996), is an approximately equals 2 - 3 times (A/ L). The evaluation o f p from important development for detailed and accurate measurements o f settlement measurements is not always straightforward because the ground movements in centrifuge model tests. Centrifuge model tests on tunnels in sands to investigate tilt o f the building needs to be identified; in practice the evaluation stability were reported by Atkinson et al (1977), Atkinson and o f A/ L is easier. Figure 48 shows the more general results given by Potts (1977) and by Chambón and Corté (1989, 1994), as described Burland (1995) (in terms o f the limiting tensile strain, elim) for a in Section 4.3. These have been extended recently by Skiker et al range o f L/ H ratios. Based on this approach, a methodology for (1994) to study the face stability and deformation mechanisms of assessment o f risk o f building damage due to bored tunnelling is tunnels constructed using the mechanical pre-cutting method described by Mair et al (1996). This approach has been used successfully for a number o f major tunnelling projects. illustrated in Figure 2(c); the failure mechanism observed in the tests is shown in Figure 44. Failure was initiated at the tunnel face and it then propagated to the crown o f the pre-vault, but there was Table 3. Relationship between category o f damage and limiting little propagation or soil deformation above the pre-vault. Further tensile strain, e,im (after Boscardin and Cording, 1989) insight into this is provided by Leca et al in their paper to this Category Normal degree Limiting tensile strain (e ^ ) Conference. The use o f centrifuge modelling to investigate specific of damage o f severity construction issues has also been undertaken by Imamura et al (%) (1996) and Nomoto et al (1996), who designed a miniature EPB 0 Negligible 0 - 0.05 tunnelling machine o f 100mm diameter for use in a centrifuge. 1 Very Slight 0.05 - 0.075 Tests were undertaken at 25g in loose dry sand, modelling a 2.5m 2 Slight 0.075 - 0.15 diameter shield machine at prototype scale. The influences o f face 3 Moderate* 0.15 -0.3 excavation and tail void closure on earth pressures induced on the tunnel lining were investigated, both independently and in 4 to 5 Severe to > 0.3 combination. A miniature tunnelling machine was also used by Very Severe Kim et al (1996) for tests at lg in soft clay with a surface surcharge ’Note: Boscardin and Cording (1989) describe the damage to investigate interaction effects o f closely spaced tunnels. corresponding to in the range 0.15 - 0.3% as "Moderate to Physical modelling has a valuable role in investigating complex Severe”. However, none o f the cases quoted by them exhibit severe three dimensional soil-structure interaction problems. An example damage for this range o f strains. There is therefore no evidence to o f this is described by Hergarden et al (1996), who used centrifuge suggest that tensile strains up to 0.3% will result in severe damage. model testing to study the influence o f tunnel construction on deformations o f adjacent pile foundations. Centrifuge testing was also used by Bolton et al (1996) to investigate the mechanics of compensation grouting close to a tunnel face in soft clay.

7. EFFECTS OF GROUND MOVEMENTS ON BUILDINGS horizontal strain 7.1 Introduction <%>

In the urban environment the effects o f underground construction on buildings are o f paramount importance. Despite the wealth o f case records o f ground movement measurements associated with 0 ,0.1 0.2 0.3 0.4 0.5 0.6 0.7 bored tunnel construction, there are remarkably few published angular distortion (%) records of the detailed performance of buildings affected by tunnelling. Figure 45. Building damage categories relating to horizontal strain and angular distortion (after Boscardin and Cording, 1989)

2371 (a) bending strain (b) diagonal strain minimum of (a) and (b)

Figure 48. Generalized building damage categories relating to horizontal strain and deflection ratio for different L/ H values (Burland, 1995)

the Mansion House in London (a fragile building more than 200 years old) when a tunnel was constructed beneath it (Frischmann et al, 1994). The observed profile is much wider than the predicted “greenfield” profile, with correspondingly much low er deflection ratios and distortions. As discussed by Boscardin and Cording Figure 46. Deformation o f building above a tunnel (Mair et al, (1989) and Geddes (1990), in many cases the foundations o f a 1996) building will modify the horizontal ground movements so that the horizontal strain induced in the building is considerably reduced. Buildings on continuous foundations such as rafts are likely to 7.3 Influence o f B uilding Stiffness experience negligible horizontal strain from bored tunnel construction; the same may apply to strip foundations, depending In assessing possible effects on buildings it is commonly assumed on their orientation. that the structure conforms to the “greenfield site” settlement Potts and Addenbrooke (1996, 1997) used finite element trough, i.e. its stiffness is neglected. This is often conservative, analyses incorporating a non-linear elastic plastic soil model to because in reality the inherent stiffness o f the building will tend to study parametrically the influence o f building stiffness on ground reduce both the deflection ratio and the horizontal strains. This was movements induced by tunnelling, as shown in Figure 51. About well illustrated by Breth and Chambosse (1974) who recorded 100 analyses were undertaken o f different configurations o f tunnel settlement behaviour o f a number o f buildings in response to tunnel and building dimensions. Tw o important dimensionless parameters construction in Frankfurt Clay. Figure 49 shows how the stiffness were introduced: the relative bending stiffness, p*, which of the buildings (of reinforced concrete framed construction) expresses the relative stiffness between the building and the modified their deformed shape compared with that o f the adjacent underlying ground, and the corresponding relative axial stiffness, ground. Figure 50 shows the actual measured settlement profile for a*. These are defined as

E l P * ESH 4 (12a)

EA EH (12b)

□ □ '□ □ □ □ □ □ < □ □ □ L. J □□□□□□□□□□□ o □ □□□□□□□□□□□ □ . nn c n n n . m a °aDDPj o

ai CD

Horizontal strain (%)

Figure 47. Building damage categories relating to horizontal strain : 49. Influence o f building stiffness on settlement profile and deflection ratio for L/ H=l, hogging mode (Burland, associated with tunnel in Frankfurt Clay (Breth and 1995; Mair et al, 1996) Chambosse, 1974)

2372 I Tunnel Ç

6 E (Dc E 01

Predicted green field

*20 30 40 SO* D istance from north portico: m

Figure 50. Influence o f building stiffness on settlement profile associated with tunnel in London Clay (after Frischmann et al, 1994) where H is the half-width o f building (= B/ 2) and El and EA are Figure 52. Modification factors for deflection ratio according to the equivalent bending and axial stiffness o f the building. Es is a relative building stiffness (Potts and Addenbrooke, representative soil stiffness taken by Potts and Addenbrooke to be 1996) the undrained secant stiffness at 0.01% axial strain in a triaxial compression test on a soil sample at a depth o f z/ 2. The expression for p* is similar to that used by Fraser and Wardle (1976) and by stiffness o f any shear walls. Changes o f stiffness that might occur Potts and Bond (1994) in soil-structure interaction analyses o f rafts as a result of cracking are highlighted by Simpson and Grose and retaining walls respectively. The expression for a* is similar (1996). to that used by Boscardin and Cording (1989). The results of the study by Potts and Addenbrooke are 7.4 Piled Buildings summarised in Figure 52. Modification factors to the deflection ratios (A/ L) that would be obtained from the “greenfield site” The response of buildings on piled foundations to ground settlement profiles are shown as different curves for sagging and movements induced by tunnelling poses a major challenge to hogging deformation modes for different e/ B ratios (e being the engineers concerned with underground construction in the urban eccentricity o f the tunnel from the centre line o f the building, see environment. Very few published case histories exist. Forth and Figure 51). These vary with the relative bending stiffness, p*. In Thorley (1996) present a case history o f the effects of tunnel practice, many buildings have p* values exceeding 10'2 and for construction in Hong Kong on a piled building. Figure 53 shows these cases Figure 52 indicates low modification factors in the two 7.9m diameter tunnels constructed adjacent to piles using open range 0.1-0.2, i.e. the deflection ratio that would be predicted if the face tunnelling shields and compressed air in completely weathered building were perfectly flexible is reduced to only 10-20% o f that granite below the water table. The 31 storey building was founded value by the stiffness o f the building. Similar modification factors on 2m diameter bored piles varying in length from 41m to 64m. were produced by Potts and Addenbrooke for horizontal strain. The maximum recorded settlement of the building at the side The study by Potts and Addenbrooke is a valuable contribution closest to the tunnels due to tunnel construction effects alone to methods o f prediction o f potential damage due to tunnelling and (including removal of compressed air) was 12mm; this was to understanding how buildings may behave in response to probably largely due to redistribution o f shaft friction caused by tunnelling. The approach has been successfully used for prediction vertical ground movements towards the tunnels. and interpretation o f measurements o f building response for the The effects o f construction o f a 7.5 m diameter tunnel (in stiff Jubilee Line Extension in London; these measurements have been London Clay) between bored pile foundations for a six-storey the principal focus o f a major research programme described by building were reported by Mair (1993) and by Lee et al (1994). The Burland et al (1996). clear spacing between the extrados o f the tunnel and the nearest Experience and judgement are required in the assessment o f the pile (1.2 m diameter) was only 1 m, as shown in Figure 54. The equivalent bending and axial stiffness o f the building, El and EA. piled foundations were designed to take account o f the future In particular, the contribution to the bending stiffness o f different tunnel construction; slip coating was incorporated along the entire floors o f a multi-storey building will depend on the presence and length o f the pile shafts to 4 m above the base, so that the piles

Figure 51. Geometry of problem analysed by Potts and Figure 53. Tunnelling adjacent to piled building in Hong Kong Addenbrooke (1996) (Forth and Thorley, 1996)

2373 8. GROUND LOADING ON TUNNEL LININGS

8.1 Introduction

Peck (1969) defined the two principal requirements o f a tunnel lining as (a) being able to withstand the direct compressive forces that develop circumferentially as a result o f the ground loading, and (b) being able to withstand whatever bending may occur. He clearly illustrated the factors influencing the development of ground loading on the lining. Key points highlighted by Peck can be summarized as follow s: 1. By the time the lining is installed, soil displacements have inevitably occurred (the ground having moved radially towards the tunnel and axially towards the face). This is particularly the case for open face tunnelling. In closed face tunnelling, however, soil displacements could be very small. Woolwich and Reading Beds 2. The short-term ground loading experienced by the tunnel lining Figure 54. Tunnelling between piled foundations in London Clay is inversely proportional to the magnitude o f soil displacements that have occurred prior to its installation. (Mair, 1993) 3. The circumferential flexibility o f the lining may allow a further reduction in the loading, although this is generally small since most would behave essentially as end-bearing piles. Inclinometers were completed rings are generally relatively stiff circumferentially. installed in the ground, and in some o f piles. The deflected profiles 4. The short-term ground loading acting on the tunnel lining can o f the piles and ground were very similar, indicating that the piles be taken to correspond to when the tunnel face has moved away acted as slender members and deformed with the ground. The from the lining by approximately 2 tunnel diameters, i.e. when the maximum horizontal displacement o f the nearest piles to the tunnel three-dimensional effects o f the tunnel heading are no longer was only about 10 mm, and the measured settlement o f the evident. building due to tunnel construction was negligible. 5. Any further increase in ground loading on the lining depends on Piles with bases located above or close to tunnel axis level are the soil type. In the case o f sandy soils, tunnel excavation and potentially more adversely affected. Vermeer and Bonnier (1991) lining erection take place under drained conditions, and hence little undertook finite element analysis o f potential effects o f shield or no increase in ground loading would be likely (unless caused by tunnelling on piled foundations in Amsterdam. The piles in additional grouting). In clays, however, the ground loading will question were driven down to a sand layer, and the analyses increase as excess pore pressures dissipate. predicted that the piles would follow the settlement o f this sand layer. Centrifuge model tests by Hergarden et al (1996) on a similar 8.2 Linings in Clays configuration, modelling construction o f a 7 m diameter tunnel, showed piles at a distance o f 2 tunnel diameters from the edge of Data on lining load measurements for tunnels in clays were the tunnel to be unaffected; at distances in the range 0.25-1 times presented by Peck, the majority being in London Clay. Several o f tunnel diameter, pile settlements varied in proportion to the volume these measurements indicated a rapid build-up o f ground loading, loss imposed in the tests, and closer than 0.25 times tunnel within a few weeks to a year, to a value equivalent to the full diameter, severe settlements occurred. These effects, together with overburden pressure. It should be noted, however, that these reductions in pile resistance, are illustrated by Figure 55 derived measurements were on tunnels in close proximity to other tunnels from similar tests by Bezuijen and van der Schrier (1994). (typically within one tunnel diameter). It is clear that the mechanisms o f ground movement and soil- Measurements have been made by the UK Transport Research structure interaction are particularly complex in the case o f bored Laboratory o f loads in tunnel linings by installing vibrating-wire tunnel construction beneath piled foundations. There is a need for load cells between the segments to measure the thrust at intervals more field measurements. around the ring. One example o f these measurements, reported by Barratt et al (1994), is for a 4m diameter tunnel lined with expanded concrete segments with its axis at a depth o f 20m in London Clay at Regent’s Park. The measurements made for a period o f 20 years, as shown in Figure 56, indicate that the vertical loads (measured at axis level) shortly after construction were equivalent to about 30% o f the total overburden pressure, and then have steadily built up to about 60% o f the overburden pressure and appear to have almost stabilized. The horizontal load (measured at the crown ) is about 70% o f the vertical load, despite the fact that K0 would have been in the range 1.5-2 prior to tunnel construction, the London Clay being highly overconsolidated. A similar ratio o f horizontal to vertical load has been measured at another site by Bow ers and Redgers (1996, 1997), also using load cells installed between segments; details of these are given by Davies and Bowers (1996), as shown in Figure 57. These measurements confirm the important point highlighted by Peck that in reality the tunnel lining is not subjected to the original in-situ ground stresses as if it had been “wished into place” . The soil displacements that occur prior to installation of the lining clearly have a major Figure 55. Centrifuge modelling of influence of tunnel influence in reducing both the short-term and long-term ground construction on settlement o f adjacent piles (Bezuijen loading to much lower values than the original in-situ stresses. and van der Schrier, 1994) Particularly in the case o f highly overconsolidated clays for which

2374 full overburden

100

1 year

80 - 1 week after lining installed

e 60 ■ CU>

tim e in days 3 4 5 6 7 cover to diameter ratio (C/D) Figure 56. Measurements by load cells in tunnel lining in London Clay over 20 years (Barratt et al, 1994) data from load cells • Muir Wood (1969) K0 is usually considerably greater than 1, in general it is erroneous k Barratt etal (1994) to consider the tunnel lining being subjected to higher horizontal ♦ Tedd etal (1991) than vertical ground loading (Mair, 1994). ■ Bowers and Redgers (1997) Figure 58 shows a collection o f data from the monitoring of relatively short term lining loads for 12 different segmentally lined data from strain gauges v Thomas (1976) tunnels in London Clay. The lining loads were either measured > Ward and Thomas (1965) directly as a circumferential thrust by means o f load cells between < Smyth-Osboume (1969) the segments (shown as solid symbols) or interpreted from vibrating wire strain gauges. The Victoria Line data reported by Smyth-Osbome (1969) were obtained from photo-elastic stress Figure 58. Short-term ground loading on segmental tunnel linings gauges cast into the concrete segments. The lining loads are in London Clay (single tunnels) expressed as a percentage o f the load equivalent to the overburden pressure (yz) at the tunnel axis. These are plotted against the C/ D C/ D ratio. The data one year after lining installation shows a less ratio for each tunnel, where C is the cover above the tunnel crown clear trend; generally the lining loads at one year vary from about and D is the tunnel diameter. Except in the case o f the Regent’s 40-60% o f the overburden pressure. The measurements obtained Park measurements reported by Barratt et al (1994), most o f the by Barratt et al (1994) over a 20 year period, show n in Figure 56, data are only available for a period o f about a year after installation indicate only a relatively small increase in load in the years o f the lining; in several o f the cases measurements were taken for follow ing the first year. an even shorter period. The data in Figure 58 are for one week after Prediction o f the development o f lining loads with time in clay lining installation, and also (where available) for one year after soils is complex and requires, amongst other things, knowledge of installation. In all cases the linings were either for single tunnels, the drainage boundary conditions (i.e. whether or not the tunnel or where adjacent tunnels were at least one diameter apart so that acts as a drain) and o f the variations o f permeability o f the ground interaction effects were small. The tunnel diameter was about 4m with distance from the tunnel. Several authors report pore pressure in all cases, except for the Heathrow Cargo Tunnel which was measurements around tunnels in clays indicating the lining to be 10.9m. Some o f the tunnels were lined with either cast iron or acting as a drain (Terzaghi, 1942; Eden and Bozozuk, 1969; concrete bolted segments, which were grouted, and others were Palmer and Belshaw, 1980; De Lory et al, 1979; Ward and lined with expanded concrete or cast iron segments for which no Pender, 1981). However, there is a need to verify the development grouting was undertaken. and distribution o f pore pressures around tunnels, for which lining The data in Figure 58 show a trend o f decreasing immediate loads are also measured. If tunnels in clays do act as drains, as is short-term lining load (one week after installation) with increasing probably the case for many situations, the distribution of pore pressure corresponding to long term seepage will depend on the relative permeabilities o f the lining and surrounding clay. Factors SIDE ELEVATION VIEW OF LINING INTRADOS influencing lining permeability, pore pressures, water heads and flow through tunnel linings are discussed by Mueser Rutledge Steel plates 980 x 360 x 10 (1988), Atwa and Leca (1994) and Negro (1994). radiused to fit linin Load cell and bearing « ?0 Evidence o f the short-term lining loads being inversely related Extrados plate fitte in pocket in segment 360x 188x20 in 12mm recess to the delay in installation o f the lining is presented in Figure 59 by \ / Jack space 20 separation A ------1------Negro et al (1996). They reviewed measurements o f lining loads by between means o f flat jack tests, embedded strain gauges and load cells for I segments * “ 7 M24 x 100 a variety o f tunnels in Sao Paolo, Brazil, the majority being lined countersunk screw T ...... with sprayed concrete (NA TM). Most o f the tunnels w ere in stiff ? \ to hard fissured Tertiary clays. The delay in installation was Bearing plate 35 o.d steel tube defined by means o f the distance (P) behind the excavation face to with M24 internal 700 x 188x20 M24 x thread where the lining could be considered an effective ring, i.e. the 100 1000 bolt location at which the sprayed concrete invert was closed or where ------the segmental lining was erected. The lining loads in Figure 59 are All dimensions in mm plotted as proportions o f the equivalent load corresponding to full overburden pressure at tunnel axis level (yz); these are plotted Figure 57. Details o f load cells installed between tunnel lining against the ratio P/ D, where D is the equivalent tunnel diameter. A segments (Davies and Bow ers, 1996)

2375 W ater pressure

Figure 59. Relationship between average lining loads and delay in installation o f lining in stiff to hard clays (after Negro Figure 61. Total soil and water pressures acting on 5m diameter et al, 1996) tunnel lining at a depth o f 32m in clayey silty sands (Inokuma and Ishimura, 1995) reasonably clear trend is evident of decreasing loads with increasing delay in installation o f the linings; the point no. 10 on Figure 59 corresponds to a measurement in an expanded concrete tunnel where they are almost zero. Figure 61 shows the total soil segmental lining and may not be representative due to local jacking pressures and water pressures around a concrete segmental lining forces (Negro, 1997). for a 5m diameter tunnel at a depth o f 32m, also constructed with a slurry shield in loose clayey silty sands below the water table. A 8.3 Linings in Sands and Gravels similar pattern is evident, with very low effective stresses acting on the lining. Peck (1969) and Ward and Pender (1981) drew attention to the In contrast, Atahan et al (1996) report a case history in which scarcity o f published measurements o f lining loads for tunnels in measurements by strain gauges cast into the concrete segments sands and gravels. In their review o f available data, Ward and indicated the lining loads for a 3.4m diameter tunnel constructed Pender concluded that the ground loading and deformations of with a slurry shield in sands and gravels to correspond almost to tunnels in dense sand and gravelly soils are much smaller than in the full overburden pressure. The water table was only just above clays and silts, provided any adverse water conditions are dealt the tunnel, and hence the effective stress imposed by the soil was with effectively during excavation. They showed that the effective the major component o f the loading. However the tunnel was at a stress loading from the soil on the lining is often very low, and is shallow depth o f 7.5m with a layer o f silt close to the crown, and comparable to the tunnel support pressure required to maintain this may explain the high proportion o f the overburden pressure drained stability (see Section 4.3 o f this Report). In cases o f deeper experienced by the tunnel. tunnels below the water table, the majority o f the ground loading is from water pressure, the effective stress component being very 8.4 Methods o f Measurement low. This was also concluded by Ohta et al (1995) in their review o f measurements on tunnels in Japan. Ward and Pender (1981) concluded that the technique o f measuring Examples o f low effective stresses acting on tunnel linings in lining thrusts by installing load cells between segments, as sands are illustrated in two case histories reported by Inokuma and illustrated in Figure 57, is the most reliable method. It is only Ishimura (1995). Figure 60 shows total soil pressures (measured applicable to segmental linings in clayey soils; in other ground with pressure cells) and water pressures around a concrete conditions different techniques have to be used. The use o f earth segmental lining for a 7.1 m diameter tunnel, which is at a depth o f pressure cells, for example, to measure radial soil pressures acting 17m, constructed with a slurry shield in loose clayey silty sands on the lining is often problematic, as concluded by Ohta et al below the water table. The calculated effective soil pressures are (1995) in their comprehensive survey o f measurements made on also shown, and these are very low, particularly at the invert o f the tunnel linings in Japan. The relative advantages and disadvantages o f earth pressure cells and strain gauges attached to reinforcement (in the case o f concrete segments) are discussed by Ohta et al and by Barratt et al (1994). Particularly problematic is the use o f pressure cells to deduce loads in sprayed concrete linings (NA TM). Pressure cells filled with oil or mercury are often installed against the ground, as shown in Figure 62, to measure the earth pressure (i.e. radial stress o r) acting on the sprayed concrete lining. The same type o f pressure cells are often also installed to measure the hoop or tangential stress (o e) transmitted through the sprayed concrete, as shown in Figure 62. Key factors potentially affecting the measurements made by such pressure cells are: (a) cell action effects related to differences in stiffness between the cell and the surrounding material (b) temperature effects (c) stress non-uniformities in the sprayed concrete

corrected hydraulic (d) accuracy o f positioning o f the cell pressure) (e) cracking in the sprayed concrete. Figure 60. Total soil and water pressures acting on 7.1m diameter Factors (a) and (b) can usually be taken account o f in the design tunnel lining at a depth o f 17m in clayey silty sands and calibration o f the pressure cells, but factors (c), (d) and (e) are (Inokuma and Ishimura, 1995)

2376 tunnel face

cables to data lo g g e r or '/•& terminal board

I distance along nail

Figure 63. Tunnel face reinforcement using soil nails

diameter o f the nails, as well as the nail stiffness in relation to the ground stiffness. The type o f grout used also influences the nail stiffness. Barley and Graham (1997) present results o f a series o f pull-out tests on trial fibreglass and steel soil nails used for tunnel Figure 62. Typical layout o f pressure cells in sprayed concrete face support in London Clay. The more extensible fibreglass nails linings (Kimmance and Allen, 1996) were found to be less efficient than the steel nails, but they are clearly easier to cut during tunnel excavation. Numerical analysis o f soil nailing o f tunnel faces has been undertaken by a number o f more difficult to allow for. Experience shows that measurements authors (e.g. Jassionnesse et al, 1996; Peila et al, 1996). from pressure cells used for sprayed concrete linings often show considerable scatter, and interpretation o f absolute earth pressure 9.2 Slurry and Earth Pressure Balance Shield Technology or concrete stress values must be viewed with caution. The main use o f such cells is to record changes in pressure which, together Closed face tunnelling machines, either slurry shields or earth with strain gauge and lining deformation measurements, allow any pressure balance (EPB) shields, are increasingly used in soft changes o f behaviour o f the sprayed concrete lining to be closely ground. The choice between these two types o f shield (see Figures monitored. Frequency o f reading o f pressure cells is an important 3 and 4) and their successful operation depends critically on the factor during the early stages o f tunnel construction following characteristics o f the ground. Steiner (1996) reviewed the major installation o f the cells (Pacovsky, 1996). The rate o f development o f ground loading on sprayed concrete geotechnical factors influencing the choice o f machine by reference tunnel linings is o f particular importance in cases where only a thin to case histories, and stresses the importance o f A tterberg limits temporary shell is installed for reasons o f economy. The temporary and grading. Figure 65 illustrates the applicability o f the two types lining may be required for a significant period, possibly up to a o f system according to the soil grading when tunnelling below the year, before the permanent inner lining is installed. Further water table. Steiner also notes that EPB machines can be used in research is needed to develop reliable methods of measuring more granular soils than indicated in Figure 65 if suitable foam loading induced in sprayed concrete linings so that predictions and additives are injected into the tunnel face working chamber. In a performance can be compared. paper to this conference, Jancsecz discusses the applicability of slurry shields in relation to the ground properties. 9. DEVELOPMENTS IN GROUND TREATMENT Kanayasu et al (1995) describe the use o f slurry admixtures and foams to improve the performance of an 8.7in diameter EPB 9.1 Face Reinforcement machine in gravels mixed with boulders below the water table; the maximum ground movements were less than 10mm. Foams may When open face tunnelling is undertaken, there is increasing use o f also be injected into the working chamber to prevent clogging or soil reinforcement to improve face stability, particularly in France “balling” of clay soils. Additional conditioning of the excavated and Italy (e.g. Schlosser and Guilloux, 1995, Lunardi et al, 1992). The technique can also be effective in reducing ground movements, because the use o f long soil nails restrains the ground 1— i— i— I— i— r— 0 1 2 3 4 5 (m) ahead o f the tunnel face, thereby limiting deformations towards the l ■ r face. The zone of potentially significant ground movement i i i i i i immediately ahead o f the tunnel face is restrained by the soil nails, i i i i i i provided that they extend a sufficient length into the ground i i i beyond the zone, as illustrated in Figure 63. Lunardi et al (1992) □ * A * * He Kg Q describe the use o f 15m long fibreglass soil nails to reinforce the entire face area o f a 12m diameter tunnel constructed in stiff clays 1800 interbedded with layers o f water bearing silty sands. Horizontal jet grouting was used to create a horizontal grout column o f about 150mm in diameter, into which a 15m long fibreglass tube was pushed. In several experimental sections, strain gauges were attached to the fibreglass tubes at 2.5m intervals. Figure 64 shows 450 the measured strains (and the derived axial load in the tube) as the tunnel face advanced in lm lengths. The strain distribution is seen to change along the fibreglass tube as the tunnel face advanced. The efficiency o f soil nailing in limiting ground movements Figure 64. Strain measurements on instrumented fibreglass soil depends on a number o f factors, notably the spacing, length and nails (Lunardi et al, 1992)

2377 Controlledneu seiuemsettlem enteni Í ...... i..i.

LIM IT VARIES WITH PLASTICITY OF SOIL LOWER FOR MORE PLASTIC SOILS Settlem ent in absence of \ 1 / compensation grouting Compensation grouting SLURRY SHIELD WITH SPECIAL SLURRY 1 OOOI 0002 Figure 67. Basic principle o f compensation grouting

Figure 65. Applicability of slurry or EPB shield tunnelling according to soil grading (Steiner, 1996) to limit building settlements and distortions to specified amounts. The success of the technique relies on proper real time use of soils passing through the screw o f EPB machines (see Figure 4) monitoring data (Leca and Clough, 1994). can be achieved by injecting suitable additives through the screw The technique was used successfully in Baltimore USA to casing to enable smooth flow and prevent either plugging or too protect about 40 masonry buildings (Baker et al, 1983) and in high a slurry content. Minneapolis USA to protect a masonry arch culvert (Cording et al, The choice and applicability of foam and polymer additives, 1989). In both cases this was achieved by means o f compaction depending on the ground conditions, is a relatively new challenge grouting above tunnels constructed in dense sands. Fracture to geotechnical engineers concerned with closed face pressurised grouting with a fluid grout is usually adopted for compensation tunnelling machines. The environmental aspects o f disposal of grouting in clay soils and examples of its successful use are excavated soil containing foam and polymer additives are described by Pototschnik (1992), Wittke (1995), Harris et al (1996) becoming increasingly important. Mixed face tunnelling in layered and Osborne et al (1997). Injection o f a fluid grout with a high ground makes the choice o f suitable additives especially difficult. solids content into granular soils, thereby allowing the grout to An illustration of the complexity of modem slurry shield bleed rapidly (sometimes termed “intrusion grouting”), was machine technology is given in Figure 66 by Kuzuno et al (1996). successfully used to protect two sensitive masonry structures in They describe a remarkable method o f constructing a 17m wide London (Mair et al, 1994; Harris et al, 1994). station tunnel complex close beneath a building in clays and sands Figure 68 shows a cross section through the tubes a manchettes below the water table. A triple circular face slurry shield machine (TA M s) installed about 8m above the tunnels for grouting during was constructed for this purpose, in order that the entire cross construction o f two 9.1m diameter platform tunnels and a 11.8m section was simultaneously excavated. The maximum settlement diameter concourse tunnel beneath the foundations o f Waterloo o f the building amounted to only 7mm. Station in London (Harris et al, 1996). The tunnels were supported by temporary sprayed concrete linings (NA TM). A plan view o f the grouting shafts and layout o f the TA Ms is shown in Figure 69. The 9.3 Compensation G routing settlements o f the foundations were generally controlled to less than 1 Omm, compared with up to 60mm expected in the absence Compensation grouting is a relatively new technique which is o f compensation grouting. Harris et al (1996) showed that the total being used increasingly to control ground settlements during bored volume o f grout injected into the clay was well in excess o f the tunnel construction in soft ground. The principles o f the method volume loss associated with tunnel construction. Based on a likely and a review o f its applications are presented by M air and Hight volume loss o f around 1.5%, this implied an “efficiency” of about (1994). The basic principle is illustrated in Figure 67. Grout is 0.3 (efficiency being defined as the ratio o f volume o f ground loss injected between the tunnel and the building foundations (not to the volume o f grout injected). Comprehensive measurements shown) to compensate for ground loss and stress relief caused by and further research are required to explain this relatively low tunnel excavation. Grout injection is undertaken simultaneously efficiency. with tunnelling in response to detailed observations, the aim being

Foundation Level +98.0m TD

Terrace Gravels

London Clay TAM Level +88.5m ID

Platform E/B Concourse Platform W/B oc ■a aj K > 5.767m TD| #74. ID A J75.767m Tok (U lo D = 9 .!n y 135m 1 0 0 1 1 .8 m I E

Woolwich and Reading Beds Clay

(Unit: mm) Figure 68. Cross-section through tunnels and compensation grouting tubes (TA Ms) at Waterloo Station, London (Harris et al, 1996) Figure 66. Use o f triple circular face slurry machine (Kuzuno et al, 1996)

2378 overlying masonry building, as shown in Figure 70. Deep settlement pins were installed above and below the level o f the TAMs through which grout was injected during tunnel construction. Figure 71 shows the observed settlement o f these pins during enlargement of the tunnel from a 5.75m diameter pilot tunnel to the final size by hand-mining methods. The settlement pin below the level o f grouting settled by almost 90mm. In contrast the settlement pin immediately above the grouting layer, and the building, settled by no more than 20mm. In recent years the technique o f compensation grouting has been increasingly used to control ground and building movements above tunnels. Although a very effective technique, it should not be regarded as a panacea for all potential settlement problems, nor Figure 69. Shaft and grout tube layout for compensation grouting should it be regarded as a substitute for good quality tunnelling at W aterloo Station, London (Harris et al, 1996) practice aimed at minimizing settlement.

10. PRINCIPAL CONCLUSIONS

The principal conclusions arising from the main sections o f this Report are listed below:

10.1 Advances in Tunnel Construction Techniques

(i) Open face tunnelling, where there is easy access to the tunnel face, has resulted in increasing use o f sprayed concrete linings. Other advances include further developments o f the pre-cutting method and o f various ground treatment techniques such as soil nailing. (ii) There have been considerable advances in the technology of closed face tunnelling, which operates on the principle of a pressurized tunnel face. The use o f sluny and EPB shields for a wide variety of ground conditions is becoming increasingly common. Figure 70. Deep settlement pins for monitoring compensation

grouting (after Osborne et al, 1997) 10.2 S tability

10 (i) Kinematic upper bound and statically admissible lower bound plasticity solutions now exist for the stability o f tunnel headings, 0 both for undrained and drained conditions. -10 (ii) Limit equilibrium solutions are available for assessing stability -20 o f tunnel faces pressurized by slurry or EPB shields. I -30 (iii) Model tests have demonstrated markedly different failure mechanisms for tunnels in clays and sands or gravels. Failure

§ -50 mechanisms o f tunnel headings in clays are significantly wider than the narrow “chimney” or “funnelling” mechanisms observed in sands or gravels. -70

-80 10.3 Ground Movements -90 17/ 05/ 96 16/ 06/ 96 16/ 07/ 96 15/ 08/ 96 14/ 09/ 96 (i) Transverse short term surface settlement trough widths are date - linearly proportional to tunnel depth (i = Kz,,) Figure 71. Settlement o f deep pins and building during tunnelling - independent o f tunnel size and tunnelling method and compensation grouting (after Osborne et al, 1997) - similar for all clays (whether soft or stiff), with most data being reasonably consistent with K = 0.5, as proposed by O’Reilly and New (1982) An exclusion zone imposed on the grouting in close proximity - similar for sands and gravels, whether the tunnel is above or to the sprayed concrete linings is described by Harris et al (1996). below the water table, generally with smaller values of K However, field trials reported by Kimmance and Allen (1996) and than for clays (K ranging from 0.25 to 0.45, with a mean of Falk (in a paper to this Conference) indicated that the increase in 0.35). total stress imposed on the tunnel lining by compensation grouting (ii) Subsurface transverse short term settlement troughs can be operations was relatively minor. Finite element analysis by reasonably approximated as Gaussian curves in the same way as Kovacevic et al (1996) showed that compensation grouting using surface settlement troughs. For tunnels in clays and sands, K fracture grouting would have little effect on the lining if the grout increases with depth below the ground surface. tubes are more than one tunnel diameter above the tunnel. (iii) Longitudinal settlement troughs have the general form o f a Osborne et al (1997) describe compensation grouting close to cumulative probability curve. The surface settlement being 0.5 Smlx the crown o f a 10m diameter tunnel to control settlement of an above the tunnel face appears to be only applicable to open-faced

2379 tunnelling techniques in stiff clays. In cases o f significant face many cases the percentage o f overburden acting on the lining is support, as in pressurized face tunnelling machines, there is a less than 50%. The percentage of overburden reduces with translation o f the cumulative probability curve with the settlement increasing cover to diameter ratio. The long term soil loading equal to 0.5 Sm„ further back from the face. increases very slow ly in low permeability clays. (iv) Small values of volume loss (often as low as 0.5%) are (ii) Tunnel linings in sands and gravels generally experience very achievable using slurry or EPB shields in sands and gravels. Higher low effective stresses; most o f the total ground loading is water values are usually obtained for tunnels in clays. pressure. (v) Post-construction settlement troughs above tunnels in clays (iii) Measurement o f ground loading acting on sprayed concrete depend on the magnitude and distribution o f excess pore pressures linings remains problematic. induced by tunnelling and on the tunnel drainage boundary conditions. In many cases, these troughs are significantly wider 10.7 Developments in Ground Treatment than the short term settlement troughs and consequently only very small increases in distortion, deflection ratio and horizontal strain (i) Success in controlling stability and ground movements during are observed. Exceptions can be when appreciable positive excess tunnelling in the urban environment can often be ensured by the pore pressures are generated, for example in soft clays when over- application o f suitable ground treatment techniques. This Report pressurization of the face occurs or when tail void grouting has focused on 3 areas in which there have been significant pressures are high. developments in recent years: face reinforcement, slurry and earth pressure balance shield technology, and compensation grouting. 10.4 M odelling and Prediction o f Ground Movements (ii) Face reinforcement by soil nailing can be an effective means o f improving face stability and controlling ground deformations. (i) The empirical method o f predicting ground movements has The efficiency of soil nailing in limiting ground movements considerable practical value, particularly where there are previous requires further research on the influence o f spacing, length, case histories o f tunnelling in similar ground conditions using diameter and stiffness o f the nails. similar construction techniques. A major limitation is that good (iii) In closed face tunnelling there have been significant judgement is required in the selection o f an appropriate value o f developments in slurry and earth pressure balance shield volume loss. technology, particularly in the use o f special slurries, foams and (ii) Closed form solutions for unloading o f circular and spherical polymer additives. Assessment o f their applicability to different cavities in linear elastic- perfectly piastic continua under types o f ground is important. axisymmetric conditions can provide useful methods o f prediction, (iv) Compensation grouting has been increasingly used with albeit approximate. success for control of ground and building movements above (iii) 2D finite element analysis is still commonly used in present tunnels. This is o f particular relevance to tunnelling in the urban engineering practice. However, sophisticated soil models are environment. required to achieve realistic predictions o f the shape and width o f the transverse settlement trough. Non-linearity and K„ have ACKNOWLEDGEMENTS important influences on the predictions, and anisotropy may also be important. Good judgement is required in selection o f either an The authors are grateful to Rob Nyren for his assistance in appropriate value o f volume loss (as in the case o f the empirical preparation o f Figures 18 and 19 and the accompanying tables, and method) or a suitable amount o f unloading prior to installation o f to Annette Nielsen for her help in production o f this Report. the tunnel lining. (iv) 3D finite element analysis with sophisticated soil models REFEREN CES remains a major undertaking, particularly for shield tunnelling,

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