ce5108-lecture 6 rational for excavation regulation nov 2010

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CE5108 Lecture 6Rational for Excavation

Requirements for Safety andRequirements for Safety and Economy

OCT 2010

By Prof Harry Tan

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Summary Report to BCASummary Report to BCA

Technical Expert Panel

Tan Siew Ann (NUS)Wong Kai Sin (NTU)

Malcolm Bolton (Cambridge UK)Malcolm Bolton (Cambridge UK)Andrew Whittle (MIT USA)

January 13th 2009

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Background• Primary constraints

– Stability of ERSS (safety)• Principal condition for Greenfield sites

– Damage control (adjacent structures) – serviceabilityP i i l diti f b it ( dj t f iliti & tiliti )• Principal condition for urban sites (adjacent facilities &utilities)

• Ground conditions– ‘Favorable’: soils covered by BS8002

– ‘Unfavorable’: specific to Singapore• Kallang formation (deep soft clays, extending to or below formation)

• Deep (clayey) fills/reclamation sites

l f l /• Principle of regulation using w/H– Stability: Mobilization of shear strength in soil

• strain controlled

– Damage: empirically guided• prior projects

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Table 1 Panel recommendations for permissible maximum wall deflection ratios

Limiting values of δw/H Facilities Located in: Ground Conditions: Zone 1

(x/H < 1) Zone 2

(1 ≤ x/H ≤ 2) Zone 3 (x/H>2)

Type A: Favourable OC stiff clays & silts Residual soils

0.5% 0.7% 0.7%

Medium-dense sands

Type B: Unfavourable Soft clays, silts or organic soils extending to or below formation (e.g., Kallang formation) Loose - fills

0.5% 1.0% (TEP) 0.7% (BCA)

1.5% (TEP) 1.0% (BCA)

Notes:

1. Shaded cells indicate parameters controlled by stability of ERSS, other cells are limited to prevent

damage to adjacent facilities

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damage to adjacent facilities

2. (TEP) - represent limits proposed by the Technical Expert Panel to meet stability requirements. These

can be considered as long term regulatory goals.

3. (BCA) – represent limits proposed by BCA and agreed by the Panel as practical limits that would be

appropriate for revision of current regulations.

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Influence Zone

Zone 1 Zone 2 Zone 3

Ground Condition Buildings located Buildings located Building located

Proposed Movement Control Limits

Ground Condition Buildings located within distance H

Buildings located within H to 2H

Building located outside of 2H

Kallang etc. (TEP) 0.5%H ^ 1.00%H 1.5%H*

Kallang etc. (BCA) 0.5%H ^ 0.75%H ^ 1.0%H *

Others 0.5%H ^ 0.75%H* 0.75%H *

Note:

* ‐ stability of ERSS

^ ‐ protection damage control5

Mobilization factors corresponding with /H values

Influence Zone: X < H H < X < 2H X > 2HInfluence Zone:

Soil Type:

X < H H < X < 2H X > 2H

Kallang formation or similar soft clays found from original ground surface to f ti l l

2.0 on cu 1.5 on cu 1.2 on cu

formation level

Others1.5 on cu

1.2 on tanϕ’1.2 on cu

1.1 on tanϕ’1.2 on cu

1.1 on tanϕ’

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Justification of Mobilization Factors M

BS8002 allows two possible interpretations for Kallang soils:

• Use of M = 2.0 on undrained strength cucorresponds to /H = 0.5% – Control of damage to adjacent structures

• Use of M = 1.2 on cu for total stress design & [M = 1.1 on tan ′ for effective stress design]:[ g ]– Stability requirements

– Justified by control procedures of inspection, monitoring and check calculation.

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Control Strategy– Worst Case (WC) defines work suspension level (WSL)

– Several alternatives for Alert Level (AL) and Check Level (CL)

Level Option 1 Option 2 Option 3

WSL WC WC WC

AL Best Est. 70% WC 70% WC

CL 70% Best Est. 50% WC Continuous*

* Performance based monitoring needed for JGP 8

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Control Strategyi. Stop Level SL, at which excavation work with be stopped due to

ground movements exceeding the designer’s worst case predictions

w,WC and pending a reassessment of the state of the ground and the

structure; ;

ii. Alert Level AL, at which a significant proportion of the maximum

anticipated wall movements will have occurred. Updated predictions

of future performance should then be made to consider the possible

need to re-engineer the remaining works.

iii. Check Level CL, at which early recognition of the behavior of the

ground and the structure can lead to a confirmation or recalibration of

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g

the design assumptions.

Define: The "worst case" prediction is the largest expected wall deflection determined through a sensitivity study that includes possible scenarios of the worst credible strength, stiffness, thickness of weak layers, and loadings in the analysis.Define: “Best Estimate” prediction is based on “Moderately Conservative” parameters about one standard deviation less than mean values; so that approximately 85% chance that you would not exceed this value

Control Strategy• Option 1 (which is similar to the proposal tabled by LTA); sets the Alert Level at the designer’s best estimate of maximum wall displacement dw,BE, and the Check Level at 0.7dw,BE. This approach offers the benefit of allowing designers to make their own estimates of dw,WC, and dw,BE and of permitting the latter to trigger the Alert Level. It is clear that the variability between the expected and worst‐case ground strength profile, on its own, should cause a rational designer to set dwBE considerably lower than dwWC, onits own, should cause a rational designer to set dw,BE considerably lower than dw,WC, on the grounds that mobilization factor M would be commensurately variable. However, the Panel also noted that designers might be tempted to advance the rational selection of dw,BE towards dw,WC so as to attempt to avoid triggering Alert Level checks during construction.• This scenario is mitigated in Option 2, where the Alert and Check Levels are simply defined as proportions of the designer’s worst case prediction of wall movements. In this way, the BCA could be better assured that two careful stages of assessment would precede the triggering of a stop order, and that the later of these would give ample

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opportunity to the BCA and the engineers responsible on site to re‐engineer the works.• Finally, the TEP recommends that Option 3, featuring continuous monitoring, be adopted where the design depends on brittle materials or where the construction process is more uncertain than usual. Good examples could include projects that make extensive use of soil stabilization techniques. In this case, field measurements are essential for validating the bulk performance of the improved soil mass. This was well illustrated by the use of inclinometer data to interpret compression of JGP layers in forensic investigations for the Nicoll Highway collapse.

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Jet Grout Piles• Soil improvement does not change the classification of ground type

• Two cases

– Gravity structures

• Follow BS8002

• No tension internally within improved soil mass

Shear plugs– Shear plugs

• Performance based design

• Monitor at all stages

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Appendix A

Empirical Data on ERSS Wall DeflectionDeflection

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EMPIRICAL DATABASE ON WALL DEFLECTION RATIO, dw/H

• For excavations in favourable ground conditions, the published data (Clough and O’Rourke, 1990; Yoo and Kim 1999; Wong and Poh 1996; andYoo and Kim, 1999; Wong and Poh, 1996; and Wong et al., 2001) indicate that most of the successfully completed excavations yielded wall deflection ratios below 0.5%H as can be seen in Figures A1 to A4. The exceptions are mainly related to soldier pile walls. These cases usually involved running sand or squeezing soils. The termrunning sand or squeezing soils. The term “favourable ground condition” refers to stiff over‐consolidated clay and silt, sand and stiff residual soils.

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Figure 1. Typical database of surface settlements caused by excavation (Clough & O’Rourke, 1990). The data are from subway projects in Oslo, Chicago and San Francisco.

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•Kallang sites can be broadly classified as ‘unfavourable’ ground conditions due to occurrence of soils with low shear strength and stiffness that extend down to the formation level or below it. Design parameters for these unfavorable ground conditions are not considered explicitly in BS8002.• According to conventional definitions a ‘soft clay’ has undrained shear strength, 12 < su < 25kPa, while ‘medium’ refers to the range 25 < su < 50kPa. Normally consolidated clays typically have undrained shear strength proportional to the in situ vertical effective stress, su ≈ 0.20±0.05s’v0.

Favourable Soils Experience

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Fig. A1 Observed maximum wall deflections in stiff clays, residual soils and sand(Clough & O’Rourke, 1990)

Fig. A2 Measured maximum wall deflections – Korean Experience (Yoo& Kim, 1999)

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Favourable Soils Experience – Singapore

Fi A4 M i W ll D fl i i

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Fig. A3 Maximum Wall Deflections – Singapore Experience (Poh & Wong, 1996 )

Fig. A4 Maximum Wall Deflections in Stiff Soil Condition at NEL (Wong et al., 2001)


• For excavations in unfavourable ground conditions, published data (Mana and Clough, 1981; Long, 2001; and Moormann 2004) indicate that most2001; and Moormann, 2004) indicate that most excavations yielded wall deflection ratios below 2%H as can be seen in Figures A5 to A8. The term “unfavourable ground condition” refers to a soil profile similar to that in Kallang formation with a thick deposit of soft clay. It should be noted that the factors of safety in these graphs are based onthe factors of safety in these graphs are based on Terzaghi’s method without considering the wall penetration below formation level.

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UNFavourableSoils Experience

• Maximum Wall Deflections against Basal gHeave in Soft Clays

• (Mana & Clough, 1977)

0.5%H

Fig. A5 Wall deflection ratio for excavations with fixed toe in soft clay(Mana and Clough, 1981)

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Fig. A6 Wall deflection ratio for excavations with free toe in soft clay(Mana and Clough, 1981)

• Comparison of Results with FEA by Mana

• (Mana, 1976)

UNFavourable Soils Experience

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0.5%H

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Fig. A7 Wall deflection ratio for excavations in soft clay (Moormann, 2004)

Fig.A8 Propped walls with Low FOS on basal heave (Long, 2001)

Idealized Clough and O’Rourke Chart

Fig.A9 Propped walls with Low FOS on basal heave (Long, 2001)

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• In cases where there is a low FOS against base heave, large movements (dhmax to 3.2%H) have been recorded in the literature.• The data mostly fall within the limiting values suggested by Mana and Clough (1981), and it is suggested that the relationships between movement, system stiffness, and FOS proposedby Clough et al. (1989) form a good starting point for preliminary estimates of the performance of such systems.

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Lessons from Empirical Data (Long, 2001)•A database of some 300 case histories of wall and ground movements due to deep excavations worldwide is presented. Although recognizing the weakness in the approach, a large database is used to examine general trends and patterns. • For stiff soil sites, movements are generally less than those suggested in the well known relationships proposed by Clough and his coworkers. (dH << 0.5%H)• However, for walls that retain a significant thickness of soft material but have a high factor of safety against basal heave, movements are similar to those calculated using the Clough charts. • In these cases, when soft ground is actually present at dredge level, the Clough charts will underpredict movement and need to be used with care. • For the above cases there is no discernible difference in the performance of propped or anchored systems but there is some evidence to suggest top‐down systems perform better.• In cases where there is a low factor of safety against excavation base heave, large movements can occur but the Clough charts will give reasonable preliminary estimates

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movements can occur, but the Clough charts will give reasonable preliminary estimates of the likely movement in such cases. • Cantilever walls have shown displacements that are often independent of the system stiffness. There is evidence to suggest that, in the case of cantilever walls and for all walls in stiff soils worldwide, design practice is conservative. • Finally, the inclusion of a cantilever stage at the beginning of a construction sequence seems to be the main cause of unusually large movements.

Appendix B

Relationship between safety and wall deflection – MSD method to relate wall

deflection to soil shear strain and monbilization factorsmonbilization factors

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RELATIONSHIP BETWEEN WALL DEFLECTION RATIO AND MOBILIZATION FACTOR (Bolton MSD method)

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Fig. B1 Idealised mechanisms of ground movement due to excavationA stiff diaphragm wall driven down to a hard layer as in Figure B1.a will first engage in cantilever rotation with l = L. A “floating” in situ wall will engage a succession of l values, starting with l > L due to additional soil shearing below z = L in the early stages, then with progressively smaller values of l as props are placed as shown in Figure B1.b.

Principle of MSD Method

• For greenfield sites, the regulation of wall deflections should be guided by sound principles of soil mechanics that relate the kinematic mechanisms of wall and ground deformations to the mobilization of shear strength within the soil mass.• The first lesson to draw is that the width of the zone significantly influenced by undrained excavation should correspond roughly with the height, L, of the wall itself, rather than depending on the depth H of excavationrather than depending on the depth, H, of excavation. • The second conclusion is that the average soil shear strains in the zone adjacent to the wall are likely to be roughly:

γaverage 2 dw,max/laverage (B1)

where dw,max is the largest lateral wall movement.• It should be emphasized that this is an approximation, since the location of maximum wall movement varies stage by stage. Nevertheless, it will be useful here to recognize

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g y g , gthe proper dimensionless groups involved in lateral wall movements and ground deformations.

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max 1.5H

min 0.5H

• If one considers the geometry of Figure B2 as representative of ERSS in unfavourableground conditions (Class B), then excavation would begin with lmax 1.5H, and would proceed until l i 0.5H. Accordingly, l H in Eqn B1. Then B1 becomes:

min

cement/soil plug

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proceed until lmin 0.5H. Accordingly, laverage H in Eqn B1. Then B1 becomes:

γaverage /2 dw,max/H (B1)

• Shearing within the retained soil mass is characterized by principal stress rotation and is best approximated (at the element level) by data from direct simple shear tests.


• Figure B4 summarizes the mobilized shear strength, τ/τf (=1/M) from undrained DSS tests on a variety of K0‐normally consolidated clays compiled by Whittle. • The data include results from a

0.8

1.0

1.2

reng

th,

/f =

1/M

1.0%4.0% =

u

• The data include results from a variety of medium to high plasticity marine clays comparable to those found in the Kallang formation. • For τ/τf ≤ 0.8 (i.e., M ≥ 1.2) these clays are well described by a parabolic relation proposed by Bolton:

0.2

0.4

0.6

obili

zatio

n of

Und

rain

ed S

hear

Str

wL (%)I

p (%)ClayLine

4523BBC5833GoM(RP)8059Empire8040GoM(U)

160120GoG11063SFBM

221τγ

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g = γu /M2 (B2)

where γu is the strain mobilized at peak strength.• For the high plasticity clays, γu ≈ 4.0%.

0.00 2 4 6 8 10

Mo

Shear Strain, (%)

(/u) = M

2

Fig. B4 Validation of strength mobilization for MSDBBC – Boston Blue Clay, GoM – Gulf of Mexico, GoG Gulf of Guinea, SFBM – San Francisco Bay Mud (unpublished data from MIT files provided by Whittle) Singapore marine clay has Ip = 50%, wL =55%

fu Mτγ

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• The results in Fig. B4 also show that the shear strength of some low plasticity, sensitive clays (e.g., Boston Blue Clay) is mobilized at smaller strains (γu ≈ 1.0%).• If strain to peak, gu, is considered as a characterstic material property, the eqns. B1 and B2 can be combined as:

d /H γ /2 0 5 g /M2 (B3)dw /H γavge/2 0.5 gu/M2 (B3)

• It is important to recognize that this expression is rather simplistic, since it assumes a typical geometry for the wall, brushes over the influence of wall flexibility, ignores the incremental nature of the deformation, and assumes the shape of soil stress‐strain curves. • If more accurate assessments are to be made, then either numerical methods or the fully incremental Mobilizable Strength Design method could be applied.• For regulation purposes, Eqn. B3 recognizes that wall movements must be

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For regulation purposes, Eqn. B3 recognizes that wall movements must be normalized by some appropriate dimension, that soil strains must be proportional to those normalized wall movements because of the requirement for soil‐structure compatibility, that such strains must be proportional to the reference strainγu, and that mobilization factor has an exaggerated effect on mobilized strain because of the shapes of stress‐strain curves.

For greenfield sites the Panel recommendations are based on the

assumption that a mobilization factor, M = 1.2 on undrained shear strength is

sufficient to ensure safety of ERSS for Class A and B soils. Table 2 shows

estimated values of w/H ≈ 0.7% (Class A) and 1.4% (Class B) for u = 2.0 and 4.0%,

respectively. These values are then recorded as Panel recommendations (TEP)

for Zone 3 limiting wall deflection ratios in Table 1. It is important to emphasize

Principle of MSD Method – How wall deflection is related to average soil shear strains and mobilization factors

for Zone 3 limiting wall deflection ratios in Table 1. It is important to emphasize

that these recommendations are contingent on rigorous implementation of

procedures for inspection, monitoring and checking of ERSS performance as

outlined in Section 4.

Table 2. Interpreted mobilization factors from simple model (App. B)

Limiting, w/H Mobilization Factor, M

Shear strain to

failure, u(%)1.0 1.2 1.5 2.0 3.0

dw /H γavge/2 0.5 gu/M2

, u( )

1.0 0.5% 0.3% 0.2% 0.1% --

2.0 1.0% 0.7% 0.5% 0.3% 0.1%

4.0 2.0% 1.4% 0.9% 0.5% 0.2%

Notes:

1. Reference conditions for stiff clays (UK practice) – used in BS8002, u = 2.0%

2. Best estimate of properties for Kallang clays, u ≈ 4.0%

3. Worst case scenario for stiff, brittle clays, u = 1.0% 30

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Appendix C

Application of JGP as Gravity walls and Shear plugs

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Question 1aUse of the mobilisation factor of 1.5 on undrained shear strength of soft soils, and the expected design strain in terms of excavation depth. How would this be affected if the excavation base is plugged with jet‐grouted piles (JGP), taking into consideration the COI’s comments on use of JGP and novel features?

1. For rigid wall rotation, BS8002 stipulates that by using the factored strength in the design, the resulting wall movement is likely to be less than 0.5% of the wall height (Hw) and the corresponding soil strain will be less than 0.5%. This may be applicable to rigid cantilever wall or rigid single‐propped wall.

2. However, it is difficult to relate soil strain with excavation depth H.0 5%H

H

Hw

0.5%Hw

HHw

0.5%Hw32

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Question 1a (con’t)Use of the mobilisation factor of 1.5 on undrained shear strength of soft soils, and the expected design strain in terms of excavation depth. How would this be affected if the excavation base is plugged with jet‐grouted piles (JGP), taking into consideration the COI’s comments on use of JGP and novel features?

3. If the JGP layer is treated as a hard soil, comments given in (1) remain valid.

H

0.5%Hw

H

Hw

HHw

0.5%Hw

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4. As JGP is considered a “novel” feature, the following actions should be taken:

a. Carry out trial tests.b. Implement stringent quality control during installation.c. Monitor compressive strain in JGP during excavation.d. Prepare contingency plan.

0 5%H

H

Hw

0.5%Hw

HHw

0.5%Hw34

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5. For mulit‐propped deep excavation, the soil strain varies from location to location. It is difficult to relate M=1.5 with soil strain for this class of problem.

6. Soil strain may not be the best index to assess the system

Further comments on this issue

stability.

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• Here is an example where soil strain is not relevant. WHY?• Same with Cofferdam in Water. WHY?

Slurry

Rock

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Question 1b

How high should the factor of safety be for reliance on JGP?

1. If M=1.5 is applied to JGP, the corresponding factor of safety is 1.5. This is only relevant to rigid rotation of cantilever and single propped wallspropped walls.

The real FS depends largely on the design strength of JGP.

0.5%Hw

Cantilever Wall Mode of failure toppling

Single‐Propped Wall Mode of failure toe kick‐out

H

Hw

HHw

0.5%Hw 38

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Question 1b (con’t)

How high should the factor of safety be for reliance on JGP?

Further comments on this issue

2. JGP provides two major functions in deep excavation:p j pa. As compression member to reduce wall deflection.b. As part of the anchorage system to hold down the soil to

minimise base heave and hence reduce wall deflection.

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Further comments on safety factor involving JGP

3. JGP as permanent compression member

a. The actual safety factor is very high because of

Mode of failure: crushing

the low design strength adopted in practice.

b. The failure is progressive which provides ample of telltale signs of impending danger.

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7

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Specification:

q 0 9 MPa

0

1

2

3

4

5

6

<0.9 0.9-1 1-2 2-3 3-4 4-5 5-6 6-7 7-8 8-9 9-10

Unconfined Compressive Strength of JGP (MPa)

No.

of

Sam

ples

qu = 0.9 MPa

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JGP strength (14 days)

Project JGP Method

Design qu(kPa)

Mean Measured qu

(kPa)

Source

Singapore River Double 500 1225 Chia & Tan (1993)Singapore River Double tube

500 1225 Chia & Tan (1993)

Geylang River Single tube

500 1843 Liang et al. (1993)

Clarke Quay MRT Station

- 600 2520 Shirlaw et al. (2000)

Tunnel at Race Course Rd

- 600 2024 Shirlaw et al. (2000)

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Tunnel at Race Course Rd

- 600 1290 Wen (2005)

C824 – Nicoll Highway Double tube

900 5826 Wong (2006)

C824 – Nicoll Highway Triple tube

900 3584 Wong (2006)

Question b (con’t)

Further comments on safety factor involving JGP

3. JGP as sacrificial compression member

a. The strength selection of the sacrificial layer is

Mode of failure: Strut over‐stressed

critical.

b. If the design strength is too low, it can lead to under‐estimation of the strut forces.

6

7

8

Specification:

q 0 9 MPa

0

1

2

3

4

5

6

<0.9 0.9-1 1-2 2-3 3-4 4-5 5-6 6-7 7-8 8-9 9-10

Unconfined Compressive Strength of JGP (MPa)

No.

of

Sam

ples

qu = 0.9 MPa

42

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Further comments on safety factor with JGP

3. JGP as part of anchoring system

a. The actual safety factor is depends on the

Mode of failure: Shearing, bending

interface shear strength.

b. If the piles are far apart, JGP can fail in bending.

c. A floating wall system is more risky as it lack a second line of defense to prevent basal heaveheave.

d. Full penetrating wall is more robust because the system is safe against basal heave.

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Question 1c

How will the JGP change the movement limit value to beallowed in design? What is the recommended design strain limitfor the JGP? How can this be taken into account in numericalmethod?

1. No change 0.5%H for favourable ground otherwise use 1.0%H.

2. Deformation (SLS) and safety (ULS)should be treated separately whenever possible.

3. For some problems such as walls with rigid rotation (M=1.5) or design of pile capacity (FS=2.5), it may be possible to satisfy both g p p y ( ) y p ySLS and ULS in one simple analysis.

4. For complicated soil‐structure interaction problem such as multi‐propped deep excavations, it is more appropriate to conduct separate SLS and ULS analyses.

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Question 1c (con’t)

How will the JGP change the movement limit value to be allowed in design? What is the recommended d i t i li it f th JGP?design strain limit for the JGP? How can this be taken into account in numerical method?

5. Typical failure strain from JGP samples is between 0.5% to 1%.

Core sample

6. Failure strain for JGP mass likely to be > 0.5%.

7. Recommended design strain limit is 0.5%. JGP mass

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How will the JGP change the movement limit value to be allowed in design? What is the recommended d i t i li it f th JGP?design strain limit for the JGP? How can this be taken into account in numerical method?

8. Eu/cu = 200 failure strain = 0.5%300 0.33%400 0.25%00 0 20%500 0.20%

9. Use Eu /cu ≥ 200 in FEA. The design Eu and cu must be validated against test results.

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Further comments on FEA involving JGP

10. Brittle behaviour of JGP can lead to progressive failure.

11. Must check the extend of yielding of JGP at each stage of excavation.

12. If yielding is excessive, it may be necessary to model the strain softening behaviour.

1 – 3

Real behaviour FE simulation using Mohr‐Coulomb Model

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stress‐strain curves of clay‐cement mix under different confining pressures

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Deflection profiles at the south wall at different stages of excavation

105

Back‐Analysis of C824 at Nicoll Highway

50

55

60

65

70

75

80

85

90

95

100

Red

uce

d L

eve

l (m

)

Level 3

Level 4

Level 5

Level 6

Level 7

Level 8

Level 9

Level 10

50

60

70

80

90

100

Red

uce

d L

evel

(m

)

50

60

70

80

90

100

Red

uced

Leve

l (m

)

50

60

70

80

90

100

Red

uce

d L

evel

(m

)

Level 1

Level 2

Level 3

Level 4

Level 5

Level 6

Level 7

Level 8

Level 9

Level 10

49

50

0 50 100 150 200 250 300 350 400

Wall Deflection (mm)

50

-100 100 300 500

Wall deflection (mm)

50

-100 0 100 200 300 400


50

-100 0 100 200 300 400


Measured

(326 mm)

(A)

no softening

(263 mm)

(B)

50% reduction

(318 mm)

(C)

80% reduction

(380 mm)

100

105

100 100 100

Deflection profiles at the north wall at different stages of excavation

Back‐Analysis of C824 at Nicoll Highway

50

55

60

65

70

75

80

85

90

95

0 20 40 60 80 100 120 140 160 180 200 220 240

Wall Deflection (mm)

Red

uce

d L

eve

l (m

)

Level 1Level 2Level 3Level 4Level 5Level 6Level 7Level 8Level 9Level 10

50

60

70

80

90

-40 0 40 80 120 160 200 240 280


Red

uce

d L

evel

(m

)

50

60

70

80

90

-40 0 40 80 120 160 200 240 280


Red

uce

d L

evel

(m

)

50

60

70

80

90

-40 0 40 80 120 160 200 240 280


Red

uce

d L

evel

(m

)

Level 1

Level 2

Level 3

Level 4

Level 5

Level 6

Level 7

Level 8

Level 9

Level 10

50

( ) ( ) a de ect o ( )

Measured

(181 mm)

(A)

no softening

(191 mm)

(B)

50% reduction

(220 mm)

(C)

80% reduction

(225 mm)

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Outline of Lecture on Nicoll Highway

• Summary of CoI Report

• Significant Events leading to Collapse

• Incorrect use of PLAXIS, Method A vs B

• Incorrect waler/strut connection design

• Other mitigating factors

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• Conclusions and Lessons Learnt

Overview of Collapse Site M3/M2

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533:46 pm

Summary of CoI Report on Causes of Collapse

• Under design of strut/waler connection• Under-design of strut/waler connection• Under design of Dwall using Method A in Plaxis

Significant contributory factors:• Incorrect Back analysis• Problems with instrumentation and monitoring

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ob e s s u e a o a d o o g

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• Significant events leading to Collapse

• Replacement of stiffener plates with C‐channels

• Removal of upper JGP enbloc over 8 b h ll th l lbays without installing 10th level Struts

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Source: NLC Expert Report, Figure 10.6.1

Source: Exhibit E93

Undrained Strength Differences in Method A and Method B for C824

RL Method A Method B Diff %Diff

MC (Lower)

MC (Upper)

Method A

56

RL Method A Method B Diff %Diff(m) (kPa) (kPa) (kPa)

MC(U) Top 97.1 16 20 -4 -20MC(U) Bot 85.6 37 24 13 56

MC(L) Top 82.1 50 29 21 71MC(L) Bot 69.0 75 50 25 49

Method B = GIM

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29

Original Field Vane Shear T GIM Li & P i A B

Strength (Cu) of Lower Marine Clay

Tests, GIM Line & Points A, B & C produced by Method A in Plaxis

Using Method A with φ’=22o

gave a design line of

57

gave a design line of Cu/p’=0.3

Strength of soft clay grossly over‐estimated

GIM

Cu/p’=0.3

Method A

Method B

Strength (Cu) of Lower Marine Clay

Figure 1 of first report of experts for LTA (Nkt = 14)

CPTu data suggests that lower part of LMC is

58

lower part of LMC is Under‐Consolidated due to more recent reclamation works

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30

Impact of Method A or B on Type K Wall

59

GOLDEN MILE TOWERGOLDEN MILE TOWER

THE CONCOURSETHE CONCOURSE

Inclinometer,I‐100, Soil

NICOLL HIGHWAY STATIONNICOLL HIGHWAY STATION CROSSOVER BOXCROSSOVER BOX

Inclinometer,I‐63, DW

60

F2 G2 Type I Type K

HG1

Type J

Inclinometers

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31

77

7880828486889092949698100102104I‐

63

I‐100

7880828486889092949698

100102104

L1L2L3L4L5L6L7Current

Type K Area (XOB)

505254565860626466687072747678

-100-50050100150200250

L1L2L3L4L5L6L7Current

54565860626466687072747678

-100 0 100 200 300 400 500

I ‐ 63 I ‐ 100

61

10050050100150200250

03-Mar-03 3768303 S 04 38233

0255075

100125150175200225250

9-M

ay-

03

8-J

un

-03

8-Ju

l-03

7-A

ug

-03

6-S

ep-0

3

6-O

ct-

03

5-N

ov-

03

5-D

ec-0

3

4-Ja

n-0

4

3-F

eb-0

4

4-M

ar-

04

3-A

pr-

04

3-M

ay-

04

2-J

un

-04

2-Ju

l-04

Lat

era

l D

efl

ecti

on

(m

m)

I100, Type K Area

Trigger @ 147mm

Design @ 210mm

050

100150200250300350400450

23-M

ar-0

3

22-A

pr-

03

22-M

ay-0

3

21-J

un

-03

21-J

ul-

03

20-A

ug

-03

19-S

ep-0

3

19-O

ct-

03

18-N

ov-

03

18-D

ec-

03

17-J

an-0

4

16-F

eb-0

4

17-M

ar-0

4

16-A

pr-

04

16-M

ay-0

4

15-J

un

-04

Lat

eral

Defl

ecti

on

(m

m)

I63, Type K Area

Trigger @ 306mm

Design @ 438mm

Type K Wall Modeling

Fill

Upper Marine Clay

Lower Marine Clay

F1

OA

F2JGP

62

OA

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32

-0.5

Displacement [m]

Method ...

Method b

Compare Method A and B

-0.4

-0.3

-0.2

Method A

Method A*

Method A, LMC

Method B or Method A*, LMC phi’=17 deg ie Cu/P=0.25

I63 measured

0 50 100 150 200 250

-0.1

0

Time [day]

phi’=24 deg ie Cu/P=0.32

For same strength profiles, Method A* and Method B gave same results63

• Method A gave better fit to field data up to 5th level struts (about 15m depth)

D th 6th l l• Deeper than 6th level struts, wall displacement increased significantly and is closer to Method B predictions

• Results showed that a simple linear elastic model

t di t th h l

64

cannot predict the whole range of excavation response, from start to finish

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Compare Method A, B and D (HS)

-0.5

Displacement [m]

Method ...

Method B

Method A-0.4

-0.3

-0.2

Method A

Method A*

Method D(HS)

Method A, LMC phi’=24 deg ie

Method A*, LMC phi’=17 deg ie Cu/P=0.25

Method D (HS Cc=0.6, Cs=0.06), LMC phi’=24 deg

I63 measured

65

0 50 100 150 200 250

-0.1

0

Time [day]

phi 4 deg ieCu/P=0.32

Method D (HS) can gave similar results to Method B

-0.5

-0.4

Displacement [m]

Method ...

Method B

Method AMethod D (SS EI/2 Cc=0 6 Cs=0 06)

Compare Method A, B and D (SS)

-0.3

-0.2

-0.1

Method A*

Method D (SS)

Method D (SS ...

Method A, LMC phi’=24 deg ie

Method D (SS Cc=0.6, Cs=0.06), LMC phi’=24 deg

Cc=0.6, Cs=0.06), LMC phi’=24 deg

I63 measured

0 50 100 150 200 2500

Time [day]

Cu/P=0.32

Method D (SS) gave results between Method A and Method B

66

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Use of Method A and Method B

2nd JSE ‐ Cl. 7.2.7 In C824, the Contractor had adopted Method A in the Plaxis analysis for the design. All experts agreed that Method A, without checking the undrained shear strength it implied, was not appropriate.

It would have been more appropriate to use Method B in the analysis. All experts (save for experts for MCS) agreed that for this particular case, the use of Method A, instead of Method B, would result in an underestimation of the predicted bending moments and deflections by about 50%, or a factor of 2 for

67

moments and deflections by about 50%, or a factor of 2 for the original design in M3.

Impact of Method A or Bon Type M3 Wall

68

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35

M3 Symmetrical Mesh

Upper JGP removed

Lower JGP

69

90

95

100

105

110

Method AMethod B

Stage 19 – after removal of Upper JGP when excavated

65

70

75

80

85

RL

(m

)

to RL 72.5m

Deflection is under‐predicted by about 50%

70

50

55

60

-50 0 50 100 150 200 250

Wall Deflections (kNm/m)

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36

95

100

105

110

Method AMethod B BM is under‐predicted

by 50%

65

70

75

80

85

90

RL

(m

)

D/Wall would be under‐designed

71

50

55

60

-2000 -1000 0 1000 2000 3000

Bending Moments (kNm/m)

90

95

100

105

110

Method AMethod B

Total Pressures on Active Side under‐predicted by about

65

70

75

80

85

90

RL

(m

)

p y10%

72

50

55

60

65

-500 0 500 1000 1500

Earth Pressures (kN/m2/m)

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37

-2500-2000-1500-1000-5000500

1

2

Loads (kN/m)

Method A43%

21%3

4

5

6

7

Str

ut

Lev

els

Method B

31%

21%

10%

40%

5%

73

8

9

-4%

10.5%

Struts load at level 9 under‐predicted by 10%

Summary on Method A/B

• Under‐estimate undrained strength by factor of 2 (2nd JSE 7.2.7)

• Under‐estimate wall deflection and BM by factor of 2 (2nd JSE 7.2.7)

• Under‐estimate Strut Load at level 9 by about less than10% with no other changes in calculation or design (2nd JSE 7 2 11)

74

calculation or design (2 JSE 7.2.11)

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Implications of Method A/B

• Under‐design of D/Wall (less thick wall with smaller toe penetration)

• Smaller Toe Penetration into OA than required for adequate q qsafety

• Reduced redundancy causing plastic hinge to occur in D/wall panel at I104 at about excavation level after excavation to 10th

strut level, and hence reduced capability to re‐distribute strut loads below hinge level

• Large wall displacements causing JGP to be loaded beyond peak capacity invoking brittle behavior

75

peak capacity invoking brittle behavior

• 10% error in strut load estimate is not trivial in the light of a highly compromised design in‐terms of wall capacity and waler connection capacity (this is significant in view of structural steel design done with FOS of 1.2)

Numerical Modelling gof

Collapse State

76

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39

Dwall EI by 50% and JGP strength and stiffness reduced by 50%

103

101

73

78

83

88

93

98

Ele

vatio

n R

L (

m)

Excavate to 100.9m

Excavate to 98.1m

Excavate to 94.6m

Excavate to 91.1m

Excavate to 87.6m

Excavate to 84.5m

Excavate to 81.6m

Excavate to 78.3m

Excavate to 75.3m

Excavate to 72.3m,Mstage=0.4Remove S971

76

81

86

91

96

Ele

vatio

n R

L (m

)

Excavate to 100.9m

Excavate to 98.1m

Excavate to 94.6m

Excavate to 91.1m

Excavate to 87.6m

Excavate to 84.5m

Excavate to 81.6m

Excavate to 78.3m

77

63

68

73

0 0.05 0.1 0.15 0.2

Horizontal displacement (m)

Remove S9,Mstage=0.4Excavate to 72.3 m,Mstage=1

61

66

71

0 0.2 0.4 0.6

Horizontal displacements (m)

Excavate to 75.3m

Excavate to 72.3m,Mstage=0.4

Remove S9,Mstage=0.4

Excavate to 72.3 m,Mstage=1


78

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40


79

Deflection profile with brittle JGP

80

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To achieve failure at RL72.3m, and reasonable matching pattern of wall deflections at all stages, the combined effects of reduced lower Dwall EI and JGP by 50% and low strut capacity is needed in the model

81

• Using the MStage setting; the evolution of strut loads for the last stage of excavation

400

600

800

1000

1200

1400

1600

1800

Str

ut

forc

e (

kN/m

)

Strut 1

Strut 2

Strut 3

Strut 4

Strut 5

Strut 6

Strut 7

Strut 8

before failure can be computed

• The trend showed reduction of load in strut 9, sharp increase of load in strut 8, followed by gradual

0

200

400

0 0.05

0.1 0.15

0.2 0.25

0.3 0.35

0.4 0.45

0.5 0.55

0.6 0.65

0.7

Mstage

Strut 9

82

reduction of loads in strut 7 and 6

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Back Analysis(taken between the period between Feb to Apr 2004)

23rd FEB 04, I104

1st APR 04, I104

302mm > 252.7mm

2 d 9th 0159mm > 145mm

1st BA ‐ 5 MAR 04

2nd BA – 19th APR 04

83Source: Maunsell Exhibit E17.6

NLC Back Analysis:NLC Back Analysis:

•Do not compare well with I104 at every stage up to current stage of BA

• Cannot correctly predict future trends

84

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43

M3 is an Unforgiving Site

• Very Deep Excavation in Very Deep Soft Clay require use of Sacrificial JGP

• Curved D/walls: resulted in use of short walers and missing splays

• 66 kV crossings: gaps in JGP and Dwall, shifting of bored piles away from front of gaps

• Deep buried valley on south wall just west of I104:

85

p y jweaker OA with reduced passive resistance on south side

DWall Panel Positions at Collapse Site

• North Dwall moved more than South (South restrained by longer bored piles)

• Most panels seem to i d i iremained in one piece,

especially panel with I65 tube

• North wall panels laterally translated shearing off top of OA soils

• South wall panels over the 66 kV gap appear to have failed by

86

g p pp ytoe kick‐in, since bored piles were moved away from in front of gap area

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44

• One must never use FEM software like a black‐box, without sound engineering judgment

Conclusions

• Need to understand soil mechanics, constitutive model behavior, and some aspects of numerical analysis to use FEM program correctly

• Over‐reliance of design decisions on FEM analysis must be discouraged, and basic mechanics for

87

checking of FEM results must also be employed

• Design is not a one time act, but a continual process of ensuring that design assumptions are consistent with site reality by checking with proper instrumentation and monitoring

Conclusions

instrumentation and monitoring• When observed data and behavior is different from FEM analysis, we must determine rational reasons for difference, and back‐analyze with realistic models and input parameters to make sound predictions of subsequent works

• Steel works connection details must be done and

88

Steel works connection details must be done and reviewed by competent persons

• Use of any new design features must be properly investigated before implementation (e.g. C‐Channel brittle response, sacrificial JGP)

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45

Lessons Learnt

• The design of important structures should involve a review by an independent checker, in the form of another consulting team. In this way possible shortcomings may be identified at an early stage of the project.

89

Lessons Learnt

• Proper back analyses should attempt a systematic assessment of the causes of significantassessment of the causes of significant discrepancies between analysis and measurement.

• The modification of material parameters should reflect the uncertainty in the determination of these parameters and thus the range of variation has to be within sensible limits.

90

ce5108-lecture 6 rational for excavation regulation nov 2010

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