assessment of the level of building damage induced by ... · arising from tunnel construction, is...

International Journal of Applied Engineering Research ISSN 0973-4562 Volume 11, Number 16 (2016) pp 8882-8891

© Research India Publications. http://www.ripublication.com

8882

Assessment of the Level of Building Damage Induced by Tunnel

Construction in Soft Soil

Nouaman Tafraouti1, Rhali Benamar1 and Nouzha Lamdouar1

1ERSIM Laboratory, Mohammadia Engineering School, Mohammed V University, Rabat, Morocco.

Abstract

This paper investigate the building response to ground

deformations which happen during tunnel construction in

terms of damage level. Firstly, ground movements resulting

from tunnelling are evaluated assuming greenfield conditions,

i.e. ignoring the presence of the overlying structure, by using

the empirical methods of Peck 1969 and O'Reilly & New

1982. Therefore, assessment of the potential building damage

is performed by using various approach’s namely, the limiting

tensile strain method, the method’s of Rankin 1988, Potts &

Addenbrooke 1997 and Franzius 2003, in order to obtain an

assessment of the building damage level. The results obtained

indicate that the limiting tensile strain method and Rankin

1988 method, leads to conservative damage assessment

compared to the method’s of Potts & Addenbrooke 1997 and

Franzius 2003.

Keywords: Tunnel, building damage level, settlement trough,

ground movements.

INTRODUCTION

Expanding urbanization, has resulted in a crucial reduction of

available urban space, leading to an increase in the use of the

underground space as alternative solution for mitigating the

shortage of surface area.

But, tunnel construction in urban environment is generally

carried out in bounded zones with a lot of surface structures

located in the influence zone of the tunnel, with various types

of foundation.

In the other hand, tunnel construction induces ground

movements that may affect and distort the neighboring

buildings in surface, leading generally to aesthetic or

serviceability damage and in some cases to structural building

damage. Hence, costly repairs, construction delay and

increase in operating cost of tunnel project can result from

such problem.

Consequently, assessment of the level of building damage

arising from tunnel construction, is one of the more important

aspects of the tunnel design. Because, based on this level,

protective measures can be taken previously in order to reduce

the risk of damage of neighboring buildings in surface.

PRESENTATION OF THE PROBLEM UNDER

CONSIDERATION

In the problem under investigation, the tunnel have a circular

section with a diameter D = 6m, a depth Z0 = 20m and its

excavated in clayey soil. The soil have the following

properties: The unit weight is ϒ=18 kN/m3, the Young’s

Modulus is E=20 Mpa, the cohesion is c=10 Kpa and the

internal friction angle is ϕ =22°.

The tunnel lining is constituted by reinforced concrete

segments with a thickness of 30 cm.

The neighboring building at surface, is a one level concrete

frame structure, the building width is B = 20m and the

length of building is 10m .The distance from the midpoint of

the building to the tunnel centerline is 8m.The analysis of

building potential damage that can be induced by tunnelling is

carried out by using various methods.

The methods used in practice for the assessment of damage to

buildings caused by settlements due to tunnelling are

generally based on two separate and distinct procedures

[1,2].In the first step, the ground movements due to tunnel

excavation are evaluated without taking into account the

presence of the building. Then, the obtained settlement trough

is applied to the building for evaluating the potential building

damage [3].

EVALUATION OF GREENFIELD GROUND

MOVEMENTS INDUCED BY TUNNELLING

In this work, we use different methods, in order to evaluate

the level of building damage that can be induced by

tunnelling.

But, firstly we need to evaluate the greenfield ground

movements, for the reason that, their magnitude is a key issue

in urban environments [4] and particularly the resulting

ground settlements which is a vital step in assessing the

potential impact of tunnelling [5].

In the first step we determinate the surface settlement. For the

assessment of the greenfield (without presence of any building

in surface) settlement trough due to tunnelling ,we use Peck’s

1969 method in which the shape of the settlement trough is

supposed to approximate closely a normal Gaussian

distribution curve, because this method is the most common

available methods for the assessment of the greenfield

settlement due to tunnelling.

Effectively, it has been found that the shape of the surface

settlement troughs developing during tunnel construction is

reasonably well represented by a Gaussian distribution [6], so

greenfield settlements are generally approximated empirically

by a Gaussian curve in a direction perpendicular to the tunnel

axis [7,8]. Furthermore, this method has been intensively used

all over the world [9].

Peck 1969 proposed that the formula for the assessment of

surface settlement can be expressed as [10]:

mailto:[email protected]



8883

S(x) = Smax exp (−x2

2i2) (1)

With:

S(x): Settlement in surface at a given horizontal distance x

Smax: Maximum surface settlement

x : horizontal distance from the tunnel centerline

i : Trough width parameter, this is the horizontal distance

from the tunnel centerline to the point of inflection on the

settlement trough

One of the important parameter for evaluating ground

movements caused by tunnelling, is the volume loss "Vl"

defined as the volume of the settlement trough per linear

meter "Vs" divided by the volume of the tunnel per linear

meter:

Vl =4 Vs

π D2 (2)

The volume of settlement trough per linear meter "Vs" is

[22]:

Vs = ∫ S(x) dx+∞

−∞ = √2π i Smax (3)

Consequently, in the method of Peck 1969, the volume loss

can be expressed as:

Vl = √32

π i Smax

D2 (4)

In the literature, there is a large number of empirical

relationships for estimating the trough width parameter “i”.

The best known of them is the relationship proposed by

O'Reilly & New 1982.

In 1982, O'Reilly and New had proposed to estimate the

trough width parameter, the following empirical relationship

[11].

i = k Z0 (5)

With Z0 is depth to tunnel axis and k is a coefficient,

depending on the nature of the soil.

O'Reilly and New reported that k = 0.5 is an appropriate value

for clay [12] and Rankin found in 1988, that

k = 0.5 was a reasonable value to fit most of the field

measurements above tunnels [13].

After Mair et al 1993 [6], it is reasonable to estimate the

trough width parameter by the formula presented above, so we

adopt the relationship proposed by O'Reilly & New 1982 to

estimate this parameter.

In our case, the depth of tunnel is 20 m and the kind of soil is

clay, so the coefficient k can be taken as 0.5

thus i = 10 m.

In terms of volume loss "Vl" ,it mainly depends on the type of

soil, the excavation method, the support and the quality of

works [14].

Several authors have estimated values that can be taken for

the volume loss Vl.

According to Attewell 1977 volume loss lies in the range 1%

and 5% and after O'Reilly and New 1982, volume loss Vl lies

in the range 0.5% and 3% [15].

McCabe et al 2012 [11] have proposed the following

empirical relationships to estimate the volume loss due to

tunnelling in Irish ground:

Vl = 0,008 (Z0

D) (6) in fine soils

Vl = 0,011 (Z0

D) (7) in coarse soils

With D=6m and Z0 = 20m ,we find Vl = 2,6% by using

McCabe et al 2012 method .

Augarde 1997 [2] state that the magnitude of volume loss, for

tunnelling in clay is usually in the range 1% to 3% and

according to Burd et al 2000 [7] a volume loss of 2% , is

typical of real tunnelling operations.

So in our case, we adopt for volume loss a value of Vl = 2%.

The maximum surface settlement, can be determined from

equation n°4, so:

Smax = √π

32 Vl D2

i (8)

Therefore, we find Smax = 22,56 mm .

The greenfield settlement trough evaluated by using Peck

1969 method is reported in the graph below.



8884

Figure 1: Surface settlement trough obtained by the Peck 1969 method

Other important aspect of ground movements induced by

tunnelling, is the horizontal surface soil displacements,

because, damage of buildings can arise from horizontal

ground displacements [16] and despite the fact that horizontal

movements often do not receive as much attention as

settlements, but they can be equally damaging to buildings

[17].

O'Reilly & New 1982 showed that the horizontal surface soil

displacement in the transverse direction can be determined by

using the following formula [12].

Sh(x) = −x

Z0

S(x) (9)

With

Sh(x): horizontal soil surface displacement

S(x) : Settlement in surface

x : horizontal distance from the tunnel centerline

Z0: Tunnel depth

The horizontal surface displacement is reported in the graph

below.

Figure 2: Greenfield horizontal surface displacement profile

-25

-20

-15

-10

-5

0-50 -40 -30 -20 -10 0 10 20 30 40 50

Settlement (mm)

Offset from tunnel centreline (m)

Greenfieldsettlement (Peck…

-8

-6

-4

-2

0

2

4

6

8

-50 -40 -30 -20 -10 0 10 20 30 40 50

Distance to tunnel centreline (m)

Greenfield horizontal surfacedisplacement (mm)

Horizontal surface displacement(O'Reilly & New 1982 method)



8885

Also, we need to evaluate the horizontal strain, because the

horizontal strains induced by tunneling interact with the

neighboring structure in surface and can lead to severe

building damage [18;19].

The horizontal strain in the transverse direction is obtained by

differentiating the horizontal displacement with respect to x

[12] .

εh(x) =dSh

dx= (

x2

i2− 1)

S(x)

Z0

(10)

The horizontal strain profile is reported in the graph below.

Figure 3: Profile of greenfield horizontal strain

ASSESSMENT OF POTENTIAL BUILDING DAMAGE

INDUCED BY TUNNELLING BY MEANS OF

DIFFERENT APPROACH’S

We have determined the ground movements induced by

tunnelling, so we can use these results to assess the potential

building damage by mains of various methods.

Rankin’s 1988 method

In 1988 Rankin proposed a method for preliminary

assessment of building damage. This method is based on two

parameters, the first is the building settlement and the second

parameter is the slope of building which is according to

Franzius 2003 [12] describes the change in gradient of the

straight line defined by two reference points embedded in the

structure.

Based on the method of Rankin 1988, for a building

experiencing a maximum slope 𝛳 less than 1/500 and a

maximum settlement less than 10 mm, the building damage

category is 1. Besides, for a building experiencing a maximum

slope 𝛳 comprising between 1/500 to 1/200 and a maximum

settlement between 10 mm to 50 mm, the building damage

category is 2. [20]

In our case and using the settlement trough profile determined

previously, the maximum magnitude of settlement is Smax = 22,6 mm while the maximum building slope is θ = 1/1000,

so the building damage category is 2 and the description of the

degree of damage is slight.

The limiting tensile strain method:

Burland and Wroth 1974 [21], showed that the beginning of

visible cracking is related to a critical tensile strain. They

applied a concept of limiting tensile strain to elastic beam

representing building.

In the limiting tensile strain method, the elastic beam

representing building is characterized by an equivalent axial

stiffness (EA) and equivalent bending stiffness (EI).

In this approach the building is supposed to follow the

greenfield settlement, and a deflection ratio Δ L⁄ was defined

to study the building deformation.

In the method of Burland and Wroth 1974, the maximum

bending strain εbmax and the maximum diagonal strain εdmax

are related to deflection ratio Δ L⁄ , according to the following

equations [22]:

𝛥

𝐿= (

𝐿

12𝑡+

3𝐼

2𝑡𝐿𝐻 𝐸

𝐺) 𝜀𝑏𝑚𝑎𝑥 (11)

-0.0012

-0.0010

-0.0008

-0.0006

-0.0004

-0.0002

0.0000

0.0002

0.0004

0.0006

-40 -20 0 20 40Distance to tunnelcentreline (m)

Horizontal strain

Greenfiled horizontal strain



8886

𝛥

𝐿= (1 +

𝐻𝐿2

18𝐼 𝐺

𝐸) 𝜀𝑑𝑚𝑎𝑥 (12)

With

L : Width of building in sagging / hogging E

G⁄ : Ratio between Young’s modulus and shear modulus of

the building

H : height of building

t : The furthest distance from the neutral axis to the edge

of the equivalent beam

I : Moment of inertia

In the sagging we have I = H3 12⁄ (13)

and t = H 2 ⁄ (14) [23]

In the hogging we have I = H3 3⁄ (15)

and t = H (16) [23]

For the ratio between Young’s modulus and shear modulus of

the building E G⁄ = 12,5 for framed building [24].

We have studied in the precedent part, the greenfield

deformations of ground and we have plotted the

corresponding graphics.

The deflection ratios are calculated in both hogging and

sagging zones which are partitioned by the point of inflexion

of the settlement trough.

Figure 4: Definition of hogging and sagging modes of

deformation, after Loganathan 2011 [20].

Based on the settlement trough obtained by using Peck 1969

formula. The deflection in sagging and hogging are

respectively 𝛥𝑠𝑎𝑔 = 2,98 𝑚𝑚 and 𝛥ℎ𝑜𝑔 = 0,61 𝑚𝑚 . The

length of building in the sagging and hogging zone are

respectively 𝐿𝑠𝑎𝑔 = 12𝑚 and 𝐿ℎ𝑜𝑔 = 8𝑚 .

Hence, we find that the maximum bending strain is :

εbmax = 1,7151 10−4 in sagging

εbmax = 2,9716 10−5 in hogging

The maximum diagonal strain is:

εdmax = 1,3399 10−4 in sagging

εdmax = 6,9646 10−5 in hogging

The Burland and Wroth approach was improved by Boscardin

and Cording 1989, by including horizontal tensile strain in the

analysis.

The resultant bending strain εbr and the resultant diagonal

strain εdr are given by the equations below [12,22]:

εbr = εbmax + εh (17)

εdr = εh (1 − ν)

2+ √εh

2 (1 − ν

2)

2

+ εdmax2 (18)

With εh horizontal strain and ν the poisson's ratio.

The poisson's ratio is taken equal to ν = 0,2.

In the previous part of this paper, we have plotted the

horizontal strain profile, so:

εh = 1,128 10−3 in sagging

εh = 5,03 10−4 in hogging

Hence, the resultant bending strain is:

εbr = 1,2995 10−3 in sagging

εbr = 5,3271 10−4 in hogging

And the resultant diagonal strain is:

εdr = 9,2187 10−4 in sagging

εdr = 4,1411 10−4 in hogging

The larger value of εbr and εdr defines the critical tensile

strain

εcritical = max(εbr; εdr) (19)

So:

εcritical = 1,2995 10−3 in sagging

εcritical = 5,3271 10−4 in hogging

Hence, the maximal critical tensile strain is

εcritical,max = 1,2995 10−3

Now, we can determine the damage level of building by

comparing the maximal critical strain to the criteria given in

literature.

Boscardin and Cording 1989 have proposed a classification of

category of building damage by using the limiting tensile

strain as damage criterion.



8887

Table 1 Classification of building damage proposed by

Boscardin and Cording 1989, based on limiting tensile strain

[25]

Category of

damage

Normal degree of

severity

Limiting tensile

strain (%)

0 Negligible 0,000-0,050

1 Very slight 0,050-0,075

2 Slight 0,075-0,150

3 Moderate 0,150-0,300

4 to 5 Severe to very

severe >0,300

By using this classification, we find that the category of

damage of building is 2 and the normal degree of severity is

slight .

Potts and Addenbrooke 1997 approach:

Potts and Addenbrooke proposed in 1997 a design approach

for the assessment of the potential building damage induced

by tunnelling . They performed a parametric study by using

numerical analysis. The deformation criterion adopted in their

study are the deflection ratios and the horizontal strains

[22,25].

Deflection ratios are evaluated in the sagging and hogging

zones corresponding to the building’s location.

By using the settlement trough profile ,we find that ,the length

of building affected by sagging is 𝐿𝑠𝑎𝑔 = 12𝑚 and the length

of building affected by hogging is 𝐿ℎ𝑜𝑔 = 8𝑚. On the other

hand, the relative sagging deflection is 𝛥𝑠𝑎𝑔 = 2,98 𝑚𝑚 and

the relative hogging deflection is 𝛥ℎ𝑜𝑔 = 0,61 𝑚𝑚.

The deflection ratios are defined by [25]:

𝐷𝑅𝑠𝑎𝑔𝑔

=𝛥𝑠𝑎𝑔

𝐿𝑠𝑎𝑔 (20)

and 𝐷𝑅ℎ𝑜𝑔𝑔

=𝛥ℎ𝑜𝑔

𝐿ℎ𝑜𝑔 (21)

So 𝐷𝑅𝑠𝑎𝑔𝑔

= 2,4833. 10−4 and 𝐷𝑅ℎ𝑜𝑔𝑔

= 7,625. 10−5

The horizontal strain profile was determined previously in the

precedent part. The maximum horizontal compressive strain

𝜀ℎ𝑐𝑔

and tensile strain 𝜀ℎ𝑡𝑔

in the section in which the building is

located are respectively:

𝜀ℎ𝑐𝑔

= −1,128. 10−3 and 𝜀ℎ𝑡𝑔

= 5,03. 10−4

Moreover, Potts and Addenbrooke 1997 used relative stiffness

parameters, which account for the stiffness of both the

building and the ground.

Relative bending stiffness [26]:

ρ∗ =EI

Es(B 2)⁄ 4 (22)

Relative axial stiffness [26] :

α∗ =EA

Es(B 2⁄ ) (23)

In the above expressions, B is the width of the building, E, A

and I are the equivalent Young’s modulus, cross-sectional

area and second moment of area of the structure. Es is a

representative soil stiffness and is defined according to Law

2012 [26] as the secant young's modulus that would be

obtained at 0.01% axial strain in a triaxial compression test

performed on a sample retrieved at the half tunnel depth.

In the study of Potts and Addenbrooke 1997 ,the building was

modeled as an elastic beam with bending and axial stiffness

properties.In the literature, the axial and bending stiffness for

the equivalent beam, are generally calculated by using the

parallel axis theorem.

The parallel axis theorem gives the axial and bending stiffness

for the equivalent beam by the following equations [26].

(EcA)structure = (m + 1)(EcA)slab (24)

(EcI)structure = Ec ∑ (Islab + AslabH2)

m+1

1

(25)

With Ec is the Young's modulus for the concrete, A is the area

of the cross-section and I the moment of inertia, H is the

distance between the neutral axis of the structure which is

assumed to be at the mid-height of the building to the

individual slab's neutral axis [26] .

In our case we have a one level framed structure, so the

stiffness properties are:

(EcA)structure = (EcA)slab (26)

(EcI)structure = Ec Islab (27)

The Young's modulus for the concrete is Ec = 30 Gpa and

the slab have a thickness tslab of 15 cm.

Second moment of area I and the area A for slab are defined

as:

Aslab = tslab L (28)

and Islab =tslab

3 L

12 (29)

Plane strain conditions give L = 1m

So Aslab = 0,15 m2 m⁄ & Islab = 2,8125. 10−4 m4 m⁄

Hence

(EcA)structure = 4,50. 106 KN m⁄

(EcI)structure = 8437,50 KNm2 m⁄

In the other hand, the soil stiffness parameter is Es = 20 Mpa.

Then, the stiffness parameters defined by Potts and

Addenbrooke 1997 are:

Relative bending stiffness ρ∗ = 4,21875. 10−5 m−1

Relative axial stiffness α∗ = 22,5



8888

In addition, we have e B⁄ = 0,4, with “e” is the eccentricity

,which is the offset between building and tunnel centerlines.

So, we can use the design curves established by Potts and

Addenbrooke 1997 to assess the modifications factors.

Figure 5: Design curves established by Potts and

Addenbrooke 1997, for modification factors of deflection

ratio (Ref [27])

Figure 6: Design curves established by Potts and

Addenbrooke 1997, for modification factors of maximal

horizontal strain (Ref [27])

By using the design curves of Potts and Addenbrooke

1997,we have the following modification factors.

Modification factors of deflection ratio:

( MDRsag = 1,06 ; MDRhog = 1,24)

Modification factors for the maximum horizontal strain:

(Mεhc = 0 ; Mεht = 0)

The modification factors introduced by Potts and

Addenbrooke 1997 characterize the maximum deflection

ratios and horizontal strains in the structure in relation to

greenfield situation.

MDRsag =DRsag

DRsagg (30) MDRhog =

DRhog

DRhog

g (31)

Mεhc =εhc

εhc

g (32) Mεht =εht

εht

g (33)

The deformation criteria of the building can be calculated by

multiplying the greenfield deformation criteria with the

corresponding modification factors:

DRsag = MDRsag . DRsagg

(34)

DRhog = MDRhog . DRhogg

(35)

εhc = Mεhc . εhcg

(36) εht = Mεht . εhtg

(37)

So , we find : DRsag = 2,632. 10−4, DRhog = 9,455. 10−5 and εhc = εht = 0

By using the diagram established by Burland 1995, we can

evaluate potential damage caused by tunnel construction in

existing building.

Figure 7: Diagram of Burland 1995, showing the relationship

of damage category to deflection ratio and horizontal tensile

strain for hogging (Ref [26])

Therefore, we find that the building damage category is 0,

and that means that the normal degree of severity is

negligible.

Franzius 2003 approach:

Franzius 2003 [12] have introduced some modifications to the

original approach of Potts and Addenbrooke 1997.

He show that the influence of building width B in the original

definition of relative bending stiffness 𝜌∗ was overestimated

and the tunnel depth 𝑍0 was not sufficiently weighted [27]

So, Franzius 2003 proposes refining the relative bending

stiffness to:

𝜌𝑚𝑜𝑑∗ =

𝐸𝐼

𝐸𝑠𝑍0𝐵2𝐿 (38)

with L is the length of building in the longitudinal direction.

Besides, Franzius 2003 modified the expression of relative

axial stiffness to take into consideration the length of building

in the longitudinal direction “L”.



8889

𝛼𝑚𝑜𝑑∗ =

𝐸𝐴

𝐸𝑠 𝐵 𝐿 (39)

Furthermore, Franzius 2003 proposed new design curves for

evaluating modification factors.

In our case, we find that the relative modified bending and

axial stiffness are:

𝜌𝑚𝑜𝑑∗ = 5,2734. 10−6 and 𝛼𝑚𝑜𝑑

∗ = 1,125

We have 𝑒𝐵⁄ = 0,4 so by using new design curves of

Franzius 2003 we can assess the modification factors.

Figure 8: Design curves for 𝑴𝑫𝑹 adopting the modified

relative bending stiffness 𝝆𝒎𝒐𝒅∗ (Ref [12])

Figure 9: Design curves for 𝑴𝜺𝒉 adopting the modified

relative axial stiffness 𝜶𝒎𝒐𝒅∗ (Ref [12])

So we find that the modifications factors of deflection ratios

are :

(𝑀𝐷𝑅𝑠𝑎𝑔 = 1,34 ; 𝑀𝐷𝑅ℎ𝑜𝑔 = 1,40)

Also, the modifications factors of maximum horizontal strain

are :

( 𝑀𝜀ℎ𝑐 = 0,021 ; 𝑀𝜀ℎ𝑡 = 0,036)

In the same manner that Potts and Addenbrooke 1997 method,

the deformation criteria of the building can be calculated by

multiplying the greenfield deformation criteria with the

corresponding modification factors:

DRsag = MDRsag . DRsagg

(40)

DRhog = MDRhog . DRhogg

(41)

εhc = Mεhc . εhcg

(42) εht = Mεht . εhtg

(43)

So, we find:

DRsag = 3,3276. 10−4 , DRhog = 10,675. 10−5

εhc = −2,3688. 10−5 , εht = 1,8108. 10−5

By using the diagram established by Burland 1995, that was

previously presented, we find that the building damage

category is 0, implying that the normal degree of severity is

negligible.

Comparison of the level obtained of building damage:

The level of damage assessed by Franzius 2003 approach

agrees with that obtained by the Potts and Addenbrooke 1997

approach.

In the same way, the level of damage determined based on the

Rankin 1988 method is in compliances with that assessed by

the limiting tensile strain method.

However, the level of damage evaluated by both the Rankin

1988 method and the limiting tensile strain method is more

conservative and this can be explained by the fact that this two

approach’s ignore the contribution of building stiffness and

suppose that the building is fully flexible and follow entirely

the greenfield settlement trough while in reality the building

stiffness is a key factor that influence significantly the

building behavior. On the contrary Potts & Addenbrooke 1997

and Franzius 2003 approach’s, takes into consideration the

building stiffness by means of the relative bending and axial



8890

stiffness parameters which expresses the relative stiffness

between the building and the ground.

In both methods used, we find that the normal degree of

severity range between negligible to slight, hence the structure

can be considered to be at low risk.

CONCLUSION In this paper, the building response to tunnel construction is

studied in term of potential damage level. The analysis of

building damage that can be arising from ground movements

generated by tunnelling is performed by using various

approach’s.

The comparison of the results obtained show that the level of

building damage obtained from the limiting tensile strain

method agree with that obtained by using the method’s of

Rankin 1988.

In such way the level of building damage obtained from the

Potts and Addenbrooke 1997 method agree with that obtained

by using the method of Franzius 2003.

Nevertheless, the level of damage evaluated by the first two

method’s is incompatible with that obtained by the others. For

the reason that, the limiting tensile strain method and Rankin

1988 method disregard the influence of the building stiffness

on the ground movements arising from tunnel construction

and this explain the conservative level of damage obtained by

these method’s compared to the method’s of Potts &

Addenbrooke 1997 and Franzius 2003 which take into

account the building stiffness in the assessment of building

damage.

REFERENCES

[1] Burd, H.J., Houlsby,G.T., Chow, L., Augarde, C.E.,

Liu, G.,1994. Analysis of settlement damage to

masonry structures. Numerical methods in geotechnical

engineering, Smith (ed.), 203-208.

[2] Augarde, C.E., 1997.Numerical modeling of tunnelling

processes for assessment of damage to buildings. PhD

Thesis. University of Oxford.

[3] Liu, G., Houlsby, G.T., Augarde, C.E., 2000. 2

dimensional analysis of settlement damage to masonry

buildings caused by tunnelling.The structural Engineer

79 (1).19-25.

[4] Mair, R.J., 1998. Recent experiences of tunnelling and

deep excavations in London. Proceedings of Fourth

International Conference on Case Histories in

Geotechnical Engineering, St. Louis, Missouri,

pp.1193-1206.

[5] Murphy, J., Gaynor, S., Laefer, D.F., 2010. Estimated

tunnel-induced settlement and damage to Findlater’s

church with respect to freefield and constructed side

considerations. In: Advances in analysis, modeling &

design, Proceedings of the GeoFlorida 2010

Conference.

[6] Mair, R.J., Taylor, R.N., Bracegirdle, A., 1993.

Subsurface settlements profiles above tunnels in clays.

Géotechnique 43 (2), 315-320.

[7] Burd, H.J., Houlsby, G.T., Augarde C.E., Liu, G.,

2000. Modelling tunnelling-induced

[8] settlement of masonry buildings. Proceedings of the

Institution of Civil Engineers, Geotechnical

Engineering. Bloodworth, A.G., Houlsby, G.T., Burd,

H.J., Augarde, C.E., 2001. Three-dimensional

Modelling of the Interaction Between Buildings and

Tunnelling Operations. CIRIA Conference on

Settlement Damage to Buildings, London.

[9] Hertog, R.W.D., 2005. Boring Esfahan’s North-South

Line. Master thesis. Delft University Of Technology.

[10] Xie, X., Yang, Y., Ji, M., 2016. Analysis of ground

surface settlement induced by the construction of a

large-diameter shield-driven tunnel in Shanghai, China.

Tunnelling and Underground Space Technology

51,120–132.

[11] McCabe, B.A., Orr, T.L.L., Reilly, C.C., Curran, B.G.,

2012. Settlement trough parameters for tunnels in Irish

glacial tills.Tunnelling and Underground Space

Technology 27,1–12.

[12] Franzius, J.N., 2003. Behaviour of buildings due to

tunnel induced subsidence. PhD Thesis. University of

London.

[13] Gui, M.W., Chen, S.L., 2013. Estimation of transverse

ground surface settlement induced by DOT shield

tunnelling. Tunnelling and Underground Space

Technology 33,119–130.

[14] Serratrice, J.F., Magnan, J.P., 2002.Analyse et

estimation des tassements de surface pendant le

creusement du tunnel nord de la traversée souterraine

deToulon. Bulletin des Laboratoires des Ponts et

Chaussées 237, 5-36.

[15] Mroueh, H., Shahrour, I., 2008. A simplified 3D model

for tunnel construction using tunnel boring machines.

Tunnelling and Underground Space Technology 23,38–

45.

[16] Moller, S., 2006. Tunnel induced settlements and

structural forces in linings. PhD Thesis. University of

Stuttgart.

[17] Houlsby, G.T., Burd, H.J., Augarde, C.E., 1999.

Analysis of tunnel-induced settlement damage to

surface structures. Proceedings of the 12th European

Conference on Soil Mechanics and Geotechnical

Engineering, Amsterdam, Vol. 1, pp. 147-152.

[18] Standing, J., Burland, J., 2008.Impact of underground

works on existing infrastructure. Post-Mining 2008,

Nancy, France

[19] Laver, R.,2010. Long-term behaviour of twin tunnels in

London Clay. PhD Thesis. University of Cambridge.

[20] Loganathan, N., 2011, An innovative method for

assessing tunnelling-induced risks to adjacent

structures.Published by Parsons Brinckerhoff Inc.

[21] Burland, J.B., Wroth, C.P., 1974. Settlement of

buildings and associated damage. Proceedings of

Conference on the settlement of structures.London. pp.

611–654.

[22] Pickhaver, J.A., 2006. Numerical modeling of building



8891

response to tunnelling. PhD Thesis. University of

Oxford.

[23] Liu, G., 1997. Numerical modeling of damage to

masonry buildings due to tunnelling. PhD Thesis.

University of Oxford.

[24] Smit, G.,2010. The behaviour of modern flexible

framed structures undergoing differential settlement.

PhD Thesis. University of Southampton.

[25] Chatzigogos, C.S.T., 2002. Effects of existing surface

structures on settlements induced by tunnelling. Master

thesis. Massachusetts Institute of Technology.

[26] Law, R.H.C., 2012. Effect of existing building on

tunnelling-induced ground movements. Master thesis.

Massachusetts Institute of Technology

[27] Franzius, J.N., Potts, D.M., Burland, J.B.,2006. The

response of surface structures to tunnel construction.

Proceedings of the Institution of Civil Engineers.

Geotechnical Engineering 159, 3-17.

[28] Bloodworth, A.G., 2002. Three-dimensional analysis of

tunnelling effects on structures to develop design

methods. PhD Thesis. University of Oxford.

[29] Dimmock, P.S., Mair, R.J., 2008.Effect of building

stiffness on tunnelling-induced ground movement.

Tunnelling and Underground Space Technology 23,

438–450.

[30] Do, N.A., Dias, D., Oreste, P., Maigre, I.D., 2013. 3D

Modelling for mechanized tunnelling in soft ground-

influence of the constitutive model. American Journal

of Applied Sciences 10 (8), 863-875.

[31] Hor, B., Caudron, M., Al Heib, M., 2011. Experimental

analysis of the impact of ground movements on surface

structure. Published in "64. Canadian geotechnical

conference - 14. Pan-America conference on soil

mechanics and geotechnical engineering. Toronto,

Canada"

[32] Mroueh, H., Shahrour, I., 2003. A full 3-D finite

element analysis of tunnelling–adjacent structures

interaction. Computers and Geotechnics 30, 245–253.

[33] Keshuan, M., Lieyun, D., 2008. Finite element analysis

of tunnel–soil–building interaction using displacement

controlled model. Proceedings of international

conference on applied computer & applied

computational science, China, pp.306-311.

[34] Liu, J., Qi, T., Wu, Z., 2012. Analysis of ground

movement due to metro station driven with enlarging

shield tunnels under building and its parameter

sensitivity analysis. Tunnelling and Underground Space

Technology 28, 287–296.

[35] Pickhaver, J.A., Burd, H.J., Houlsby, G.T., 2010. An

equivalent beam method to model masonry buildings in

3D finite element analysis.Computers and Structures

88, 1049–1063.

[36] Wang, S., Luna, R., Yang, J., The Influence of Tunnel

Construction on the Performance of Shallow

Foundation Bridge. North American Tunnelling 2008

Conference Proceedings, San Francisco, California.

[37] Zymnis, D.M., 2009. Evaluation of analytical methods

to interpret ground deformations due to soft ground

tunnelling. Master thesis. Massachusetts Institute of

Technology.

assessment of the level of building damage induced by ... · arising from tunnel construction, is...

Documents