ORIGINAL RESEARCH PAPER

Shaking table tests on seismic retrofitting of rammed-earth structures

Yaan Wang1,2 • Ming Wang1,2 • Kai Liu1,2 • Wen Pan3 •

Xiaodong Yang3

Received: 11 April 2016 /Accepted: 28 August 2016� Springer Science+Business Media Dordrecht 2016

Abstract Rammed-earth dwellings are widely used in rural areas worldwide; with poor

mechanical properties and weak seismic resistance, these structures are vulnerable to

earthquakes. This study presents a method for reinforcing existing rammed-earth buildings

by strengthening the walls with externally bonded fibers. Shaking table tests were con-

ducted to study the seismic performance of two rammed-earth model structures—with and

without reinforcement. We observed the crack patterns, failure modes, changes in dynamic

properties, in-plane deformation, wall response, and roof–wall interaction in the models.

The results confirm that the seismic resistance of rammed-earth structures reinforced using

the proposed method was substantially improved.

Keywords Rammed-earth � Seismic retrofitting � Shake table test � Dynamic response

1 Introduction

Rammed-earth structures are commonly found in some of the world’s most earthquake-

prone and less developed, rural areas. Rammed-earth structures are vulnerable to earth-

quake forces because of their poor mechanical properties and the lack of effective seismic-

resistant design. In the Mw 5.7 Yiliang Earthquake that occurred on September 7, 2012, in

Yunnan Province, China, over 30,600 houses collapsed and 80 people died. In the Mw 6.5

& Ming Wangwangming@bnu.edu.cn

1 State Key Laboratory of Earth Surface Processes and Resource Ecology, Beijing NormalUniversity, Beijing 100875, China

2 Academy of Disaster Reduction and Emergency Management, Ministry of Civil Affairs andMinistry of Education, Beijing 100875, China

3 School of Civil Engineering, Kunming University of Science and Technology, Kunming 650500,China

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Bull Earthquake EngDOI 10.1007/s10518-016-9996-2

Ludian Earthquake that occurred in the same province on August 3, 2014, totally 80,900

houses were severely damaged or collapsed, and 617 people died. The impact of the

disaster was huge compared with other earthquakes of similar magnitudes and seismic

intensity. This is mainly because 90 % of the residential houses in these rural areas were

rammed-earth structures, which lack seismic resistance (Liu et al. 2015). Similar tragedies

occurred in other countries where people, especially in rural areas, still inhabit earth

buildings. The Mw 6.0 earthquake on March 8, 2010, at Elazig, Turkey, caused widespread

destruction in rural villages near the epicenter, causing 57 deaths. Other earthquakes that

caused the collapse of a large number of rammed-earth structures include the Mw 8.0

earthquake on August 15, 2007, in Concepcion, Peru, and the Mw 6.3 earthquake on 26

December 2003 in Bam, Iran (Miccoli et al. 2014). As rammed-earth structures are gen-

erally used by less developed populations, the existing rammed-earth structures will

continue to be used for a long time in the future. Thus, retrofitting of existing rammed-earth

structures is an important step to improve the disaster resilience of rural communities by

increasing their ability to withstand seismic events.

Much research has been undertaken to investigate the seismic performance of newly

built earth buildings. Some methods for strengthening rammed-earth structures focus on

improving the mechanical properties of the rammed-earth; for example, mixing the earth

with plant fiber, changing the proportion of the sand in the mix, etc. (Habibulla et al. 2012;

Chen 2009; Liu et al. 2010). Other methods improve the seismic mechanical behavior of

earth walls by adding reinforcements using bamboo, wicker, and fiber materials, which

have a relatively high tensile strength, thereby increasing the soil strength (Liu et al. 2015;

Hu et al. 2010). In many countries, national and local building codes have been adopted to

ensure safe seismic-resistant design and construction of rammed-earth buildings. For

example, the New Mexico Earthen Building Code, U.S.A., specifies the mechanical

properties of the soil, the proper ratio of height to thickness of walls, and anti-permeability

measures (Walker 1991). The Engineering Design of Earth Buildings codes of New

Zealand specifies the standards of mechanical soil testing, structural seismic and functional

design, etc. (NZS 4297 1998; NZS 4298 1998; NZS 4299 1998). China, Germany, Aus-

tralia, and India have also published code documents for the construction of earth buildings

(JGJ161 2008; DIN 18945 2012; DIN 18946 2012; Norton J 2001; IS13827 1993).

However, compared with the strengthening techniques of newly built rammed-earth

structures, methods for retrofitting existing earth structures are relatively limited. Tolles

et al. (2000) strengthened an adobe structure by adding polyethylene strips to the outside

walls to increase its integrity and tested the structure’s response to seismic forces. They

found that the structure performed better for improved structural integrity than for

increased structural strength. Huang (2007) improved the seismic performance of rammed-

earth walls by using sandwiched walls made of cement mortar with steel mesh or glass

fiber, anchored by bar reinforcements penetrating the rammed-earth walls. Bartolome et al.

(2004) investigated the use of grouting on the external part of the building as a retrofitting

technique and confirmed its effectiveness. However, practical applications of these tech-

niques may be restricted because they may not be affordable or easy to use, particularly in

the least developed rural communities.

Liu et al. (2015) proposed an effective retrofitting technique to increase the lateral load

capacity and ductility of a rammed-earth wall by using externally bonded fibers. This

technique has the advantages of lower cost, easy implementation, and low level of inter-

ference with the existing structure. In this research, shaking table tests were conducted to

further investigate the capabilities of rammed-earth dwellings to resist seismic loadings,

with and without application of the technique proposed by Liu et al. This paper aims to

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expand the knowledge of the seismic behavior of rammed-earth structures while assessing

the effectiveness of the proposed retrofit technique under a series of seismic loads.

2 Rammed-earth structures and retrofit technique

2.1 Rammed-earth structures

Rammed-earth structures are found all over the world, including China, Southeast

Asia, Latin America, Africa, the Indian subcontinent, the Middle East, and Southern

Europe (Blondet et al. 2003). Along with the different economies and cultures in those

regions, the dimensions, roof structures, and exteriors of the rammed-earth structures

vary considerably. Nevertheless, the construction techniques, namely manually

ramming large amounts of local soil with a certain scope of moisture content to form

walls, remain similar. Rammed-earth walls usually have low mechanical strength and

discrete and creeping characteristics that leave shrinkage cracks during the drying

process.

The rammed-earth dwellings in rural Yunnan Province, China, shown in Fig. 1 are

typical examples of earth structures that may suffer severe earthquake damage. They have

various load-bearing systems with or without wooden frames. For those without wooden

frames, the rammed-earth walls bear the roof loads directly and are usually highly vul-

nerable to horizontal seismic forces. Typical damage modes of existing rammed-earth

structures are damage to the walls, which include out-of-plane fall-over (Fig. 2a–c) and

vertical cracking at the corners of the building (Fig. 2d) and at loading points (Fig. 2e).

The out-of-plane fall-over of rammed-earth walls usually causes severe injuries and death

as well as property losses. This can be structural collapse (Fig. 2a), gable wall fall-over

(Fig. 2b), or longitudinal wall fall-over (Fig. 2c). Gable walls, with poor lateral bending

resistance, bear the lateral loads delivered by the purlins during an earthquake, resulting in

damage to the headpiece or the entire gable wall. The longitudinal walls bear the main

vertical load from the rafters and the upper parts of these walls often collapse during an

earthquake. The corners of the gable walls and the longitudinal walls are vulnerable

because of the concentration of forces acting on them. Without reinforcement, vertical

cracks often form around the corners of the walls. The concentrated forces at the loading

Fig. 1 Typical rural rammed-earth dwellings in Yunnan Province, China

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points where rafters directly sit on the wall often make cracks extend vertically to form

through cracks during an earthquake.

2.2 Retrofitting technique

The proposed retrofitting technique uses externally bonded fiber materials to improve the

seismic resistance of existing rammed-earth buildings. In a previous study, laboratory tests

(Liu et al. 2015) were conducted to investigate, evaluate, and select appropriate fiber and

Fig. 2 Typical damage modes in existing rammed-earth structures: a structural collapse, b gable wall fall-over, c longitudinal wall fall-over, d vertical cracking near the corner, e vertical cracks originating at loadingpoints

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bonding materials that can provide sufficient strength and are compatible with rammed-

earth walls, considering both the cost and the reinforcement effect. Double-layered tar-

paulin with non-organic fast-hardening compound (NF compound1) was used to retrofit

existing rammed-earth walls, and a series of monotonic lateral loading tests were con-

ducted on the model structures with and without reinforcement to validate the effectiveness

of the proposed technique (Liu et al. 2015). The results showed that the proposed retro-

fitting technique increased the lateral load capacity of the rammed-earth walls by up to

38 % and the maximum horizontal displacement by up to 75 %. The retrofitting procedure

is as follows.

1. The wall surfaces are roughly ground, and the surface to be bonded is coated with a

layer of the NF compound.

2. The tarpaulins are cut into a predetermined size (usually strips 20 cm wide) and coated

with the NF compound.

3. The coated wet tarpaulin strips are attached to the walls with the NF compound.

4. The tarpaulin strips are covered with another layer of NF compound.

5. A second layer of tarpaulin strips is applied and then another layer of NF compound.

6. After the two layers of strips are completely dry (usually 24 h), an exterior protection

layer is applied using lime mortar or other locally available materials.

3 Experimental setup

The shaking table tests were performed at the laboratory of the Kunming University of

Science and Technology in Yunnan Province, China. The shaking table is customized by

SERVOTEST Testing Systems Ltd. It has a dimension of 4 m 9 4 m with two degrees of

freedom and has a maximum loading capacity of 30 t. The maximum acceleration is 1.0 g,

and the maximum displacement is 250 mm in the horizontal directions.

3.1 Test specimen and material properties

The model geometry adopted in the tests was restricted by the dimensions and loading

capacity of the shaking table. Two full-scale single-floor one-room models of dimensions

2600 9 2400 9 2100 mm and wall thickness of 400 mm were built by local masons using

local soil; the size of each model is similar to a small room of a local single-story rammed-

earth dwelling in Kunming. To avoid variation in the stresses and the accelerations through

the scale reduction, we used a full-scale small room model rather than a reduced-scale

large room model. Nevertheless, the one-room model was sufficient to show the dynamic

behavior and failure modes of a rammed-earth structure under seismic excitation, partic-

ularly when the tests were performed on models with and without reinforcement. Basic

tests were conducted to determine materials properties. Basic tests were conducted to

determine materials properties. By conducting compression tests on 600-mm cubes of three

rammed-earth samples, the average compressive strength of 1.357 MPa was obtained. The

average elastic moduli of 57 Mpa were derived from the compressive strength results. The

wet unit weight and moisture content were found to be 19.7–20.2 kN/m3 and 26 %,

respectively. Both models were constructed in situ on a 300-mm-thick concrete footing.

1 The NF compound is composed of sand, sodium silicate and hardener with a specific ratio. The applicationof a patent for the NF compound is currently in pending.

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Coarse stones were inserted into the surface of the concrete footing to improve the integrity

of the walls and the foundations. To simulate the design of actual dwellings, 1-m-long

wooden cantilever beams were inserted near the top edge of the longitudinal walls. A

wooden roof structure covered with tiles was laid directly on the gable walls and the

cantilever beams. Nails were used to connect the roof to the cantilever beams. The door

opening is 700 9 400 9 1400 mm, as shown in Fig. 3. A wooden lintel

(1200 9 400 9 20 mm) was placed over the door opening to support the structure above

it.

One model was unreinforced as a reference specimen, and the other was reinforced

using the proposed technique. The reinforced model was retrofitted by horizontally

bonding double layers of tarpaulin strips around the outside and inside surfaces of the

walls. The material properties of the tarpaulin and NF compound are listed in Table 1. The

strips (20 cm wide 9 60 cm long) were laid horizontally approximately 270 mm apart.

The strips on the gable walls were bonded along the edge of the triangle, as shown in

Fig. 4a, b. Additionally, the corners were strengthened with 20-cm-wide diagonally

crossed strips. Finally, each model surface was painted with a very thin layer of white

pulvis talci paint for better observation of the crack propagation during the test.

The models were fixed on the shaking table with eight bolts set in the concrete footing

as shown in Fig. 3b. Six sensors were installed on the models to record the displacement

and acceleration during the experimental (Fig. 5). The sensors are all mono-axial with a

measuring range of 0–50 m/s2, with sensitivity ranging from 231.4 to 268.4 pC/m/s2. The

data were recorded by the DHDAS dynamic signal collection and analysis system

(DH5922) with a sampling frequency of 500 Hz. Four high-resolution digital cameras were

used to monitor the crack pattern development during the testing.

3.2 Test loading protocol

The El-Centro wave, recorded by the Imperial Valley Irrigation District during the

Imperial Valley Earthquake on May 18, 1940, was selected as the seismic input motion for

the shake table tests. The original recorded time series of East–West component of the El-

Centro wave and its Fourier amplitude spectrum are shown in Fig. 6. The original peak

ground acceleration was 0.349 g; this value is adjusted for different earthquake intensities.

The spectrum of the signal indicates that the main frequency band is 0–5 Hz.

The two models were subjected to the same seismic excitations of progressively

increasing intensity until the model reached the failure condition or the shake table reached

Fig. 3 a Cross section of the models (top view), b the finished model

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its maximum capacity. Each seismic excitation was applied in the X and Y directions

(perpendicular to Wall 1 and perpendicular to the door opening, respectively). Low-level

random signals (white noise) with a maximum peak ground acceleration (PGA) of 0.07 g

were applied in both horizontal directions between each seismic excitation to monitor the

evolution of the modal properties of the tested models. The seismic loading sequence is

summarized in Table 2.

Table 1 Physical properties of fiber and bonding materials

Material Compressivestrength (MPa)

Tensilestrength (MPa)

Young’smodulus (MPa)

Density(g/cm3)

Tarpaulin – 49.03 128.88 –

NF compound 9.55 1.64 – 1.75

Fig. 4 Reinforcement strips in place, a front view, b side view

Fig. 5 Layout of the sensors on the models

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4 Theoretical strength of the models

Before conducting the experimental testing, the theoretical seismic strength of the models

was assessed. The assessment of the ultimate shear capacity of the unreinforced model was

based on the Seismic technical specification for building construction in towns and villages

(JGJ161 2008). The seismic capacity of the model was evaluated by comparing the design

Fig. 6 El-Centro seismic wave used in the shake tests, a acceleration time series, b fourier amplitudespectrum

Table 2 Seismic loading sequence of the shake tests

Seismic loadingsequence

Seismic wave Excitationdirection

Intensity of inputpeak acceleration (g)

Intensity of measuredpeak acceleration (g)

S1 First white noise XY 0.07 0.07

S2 El Centro X 0.1 0.1

S3 Y 0.1 0.09

S4 Second white noise XY 0.07 0.07

S5 El Centro X 0.22 0.21

S6 Y 0.22 0.23

S7 Third white noise XY 0.07 0.07

S8 El Centro X 0.4 0.39

S9 Y 0.4 0.42

S10 Fourth white noise XY 0.07 0.07

S11 El Centro X 0.51 0.51

S12 Y 0.51 0.53

S13 Fifth white noise XY 0.07 0.07

S14 El Centro X 0.62 0.60

S15 Y 0.62 0.64

S16 Sixth white noise XY 0.07 0.07

S17 El Centro X 0.95 0.96

S18 Y 0.95 0.94

S19 Seventh white noise XY 0.07 0.07

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value of the shear resistance of the walls and the seismic forces. The shear resistance of

each wall can be assessed as

Vb � cbEfNfv;mA ð1Þ

fN ¼ 1

1:2

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

1þ 0:45r0=fvp

ð2Þ

fv;m ¼ 0:125ffiffiffiffi

f2p

ð3Þ

where Vb is the horizontal seismic force on the wall; cbE is the adjustment coefficient (0.85

for load-bearing walls and 0.95 for non-load-bearing walls); fN is the normal stress

adjustment coefficient of seismic strength for adobe bricks; A is the cross-sectional area of

the wall; f2 represents the compression strength of the rammed-earth block; r0 is the

average compressive stress of the masonry section corresponding to a representative value

of the gravity load; and fv is the design value of the shear strength for masonry without

seismic reinforcement.

Table 3 lists the values of the parameters in Eqs. (1)–(3), and the horizontal seismic

shear strength of the wall Vb under different seismic excitations is shown in Table 4.

Tables 3 and 4 indicate that the unreinforced rammed-earth wall would be damaged if

subjected to a seismic wave of intensity 0.4 g. The ultimate seismic-bearing capacity of the

reinforced model was assessed according to the Code for design strengthening masonry

structures (GB50702 2011). Similarly, the seismic capacity of the model was assessed by

evaluating the ultimate shear-resistance strength of the reinforced walls:

V �Vm þ VF ð4Þ

V � 1:4aVVm ð5Þ

VF ¼ af ffX

n

i¼1

Afi cos ai ð6Þ

where V is the design value of the in-plane horizontal seismic force on the wall; Vm is the

shear-resistance strength of the wall without reinforcement; VF is the increase in the shear-

resistance strength after fiber reinforcement is applied; aV is the adjustment coefficient for

the compression stress of adobe bricks; af is the participation coefficient of the fibers; ff is

the tensile strength of the reinforced fiber (ff should be multiplied by the working coef-

ficient 0.28); Afi and n are the cross-sectional area of the working fibers and the number of

working fibers across the calculated cross-sectional area; and ai is the angel of the bonded

fibers with respect to the horizontal direction.

The values of the parameters in Eqs. (4)–(6) are listed in Table 5, and the calculated

design values of the in-plane horizontal seismic force on the wall (V) under different

seismic excitations are shown in Table 6.

Table 3 Values of the parameters in Eqs. (1)–(3)

Wall cbEfNfv;mA[KN]

cbE fN fv,m (N/mm2) r0 (N/mm2) fv (N/mm2) f2 (N/mm2)

Wall 1 or Wall 2 85.50 0.85 1.01 0.14 0.06 0.0564 1.32

Wall A 75.66 0.95 1.01 0.14 0.06 0.0564 1.32

Wall B 103.53 0.95 1.01 0.14 0.06 0.0564 1.32

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Tables 5 and 6 show that the reinforced model can resist an earthquake of intensity

0.4 g, and much closer to 0.51 g. The shear strength of Walls 1 and 2 increases by 37.1 %

and that of Walls A and B by 40.5 and 35.0 %, respectively. It should be noted that the

design code (GB50702 2011) used for the reinforced model calculation is not fully

applicable for the proposed retrofitting technique and materials. Therefore, its reinforcing

effect needs to be further tested experimentally.

5 Experimental results and discussion

The seismic performance of the models with and without reinforcement was evaluated and

compared through a series of observations and analysis. They included the crack pattern

development, the structural damage and failure modes, the changes in dynamic charac-

teristics, the in-plane deformation, the wall response, and the roof–wall interaction.

5.1 Crack patterns and failure modes

5.1.1 The unreinforced model

The unreinforced model was tested to evaluate its seismic behavior and failure modes.

When the input excitation was 0.1 g in the X direction, no cracking was observed. The

damage to the model was first observed at a PGA of 0.1 g in the Y direction when a vertical

Table 4 Horizontal seismicforce on the wall (V

b) under dif-

ferent seismic excitations

Intensity ofseismicexcitations (g)

Wall 1 orWall 2 (KN)

Wall A (KN) Wall B(KN)

0.1 24.36 20.73 27.99

0.15 38.13 32.45 43.81

0.2 47.67 40.57 54.76

0.3 72.03 61.30 82.76

0.4 95.33 81.13 109.53

Table 5 Values of the parameters in Eqs. (4)–(6)

Walls Vm ? VF (KN) 1:4aVVm

(KN)Vm (KN) VF (KN) af ff (MPa) Afi (mm2) cos ai n

Wall 1 orWall 2

117.30 119.70 85.50 31.79 0.57 14 1992 1 2

Wall A 106.34 105.91 75.65 30.68 0.55 14 1992 1 2

Wall B 139.78 144.94 103.53 36.25 0.65 14 1992 1 2

Table 6 Design values of in-plane horizontal seismic force onthe wall (V) under differentseismic excitations

Intensity of seismicexcitations (g)

Wall 1 or Wall 2(KN)

Wall A(KN)

Wall B(KN)

0.4 95.33 81.13 109.53

0.51 127.11 108.18 146.04

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Fig. 7 Development of cracks with increasing PGA in the unreinforced model during the shaking table test.a The crack pattern of Wall A, b crack pattern of Wall B, c crack pattern of Wall 1, d crack pattern of Wall 2

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crack appeared above the wooden lintel in Wall A; this may be attributed to the propa-

gation of an earlier shrinkage crack. The crack size was about 1–2 mm wide. At the same

time, a long vertical crack appeared under the right cantilever beam in Wall B. When the

intensity of the PGA increased to 0.22 g in both the X and Y directions, the existing cracks

in Wall A and Wall B expanded in both width and length, while small new cracks appeared

near the bottom-left corner of Wall 1. The largest crack size increased to about 3 mm.

When the applied PGA reached 0.4 g in the X direction, the model was severely damaged

as new cracks on the walls formed quickly and the existing cracks extended further, along

with the roof tiles falling down. The maximum width of the long vertical crack on Wall B

and the crack around the wooden lintel reached 1 and 2 cm, respectively. Horizontal cracks

appeared parallel to the wooden lintel and diagonal cracks formed, extending from the

right end of the lintel to the bottom of the right cantilever beam. Meanwhile, three long

vertical cracks appeared in Wall 2; thus, Wall 2 lost some of its connection with Walls A

and B. The crack sizes in Wall 2, Wall A, and Wall B significantly increased to 5–10 mm

wide. At this stage, the unreinforced model was close to failure. When the applied PGA

reached 0.51 g, an extremely large tremor was observed through the model and most of the

roof tiles fell to the ground. A rectangular section of the wall above the wooden lintel was

visibly separated from Wall A. A long vertical crack appeared through the upper corner

that joined Wall 2 and Wall B and the two walls separated near the top corner. The cracks

penetrated the width of the wall and soil spilled from the cracks. Around the top corner that

joins Wall A and Wall 2, large pieces of the wall fell onto the shaking table. Displacement

was observed in the beams of the roof that were connected to the walls. At this stage, the

largest crack located on the corner of Wall 2 and Wall B completely penetrated through the

corner and reached 27 mm wide. The evolution of the wall crack pattern with increasing

PGA during the testing is shown in Fig. 7.

5.1.2 The retrofitted model

No cracks were formed when the applied PGA was between 0.1 and 0.62 g in both the

X and Y directions. A few tiles slid from the roof at a PGA of 0.22 g in the Y direction and

began to fall down at a PGA of 0.62 g in the X direction. Compared with the unreinforced

model, the integrity of the retrofitted model was significantly improved. At the end of the

seismic sequence (PGA = 0.95 g), the retrofitted model remained intact without visible

cracks or debonding of the reinforcement strips, and only a few roof tiles had fallen, as

shown in Fig. 8.

The experimental results of the unreinforced model agree with the theoretical calcu-

lation that predicted that the model would experience severe damage at a seismic intensity

of 0.4 g. However, the theoretical result underestimated the seismic capacity of the rein-

forced model; the theoretical estimate was 0.4 g, while in the experiment, severe damage

occurred at 0.95 g. This may be due to the contribution of a large amount of applied

bonding materials that were not accounted for in the theoretical analysis. Moreover, the

design code used for calculating the reinforced model focuses on the in-plane shear

strength of the walls, while the improvement in the overall structural integrity and stiffness

is not specifically considered. However, the experiment showed that the proposed retrofit

cFig. 8 Comparison of the two models showing areas where damage occurred in the unreinforced model(left unreinforced model; right reinforced model)

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not only increases the wall shear strength through the externally bonded fiber but also

provides significantly higher stress constraints that prevent crack development, and better

structural integrity that helps resist seismic loading of the whole system.

5.2 Dynamic properties

White noise test results and fast Fourier transform method were used to estimate the

dynamic properties of the models. The fundamental frequency of each model in

response to rising PGA is shown in Fig. 9. As expected, a significant decrease in the

measured fundamental frequency was observed as the PGA increased; this was caused

by the development of cracks that led to a reduction in the wall stiffness. The fun-

damental frequency of the unreinforced model was 8.79 Hz and remained constant

during the first two white noise tests. The frequency dropped to 7.32 Hz after tests S5

and S6 (PGA = 0.22 g), corresponding to visible cracks that appeared in the walls

(Fig. 8). The fundamental frequency dropped to 4.88 Hz after tests S8 and S9

(PGA = 0.4 g) as full-length cracks developed through the walls. After test S10

(PGA = 0.51 g in the X direction), the model was severely damaged, and the white

noise test could not be performed. The damping ratios of the tested models were

calculated by using the acceleration records at the attenuating free vibration stage in

each test. The damping ratios were estimated to be 9.81 % for the unreinforced model

at 0.1 g. The damping ratios of the unreinforced model showed an increasing trend

with the increase in PGA, which was consistent with the observed cracking develop-

ment. The estimated damping ratios were increased to 10.95, 12.92, and 13.56 % at

0.22, 0.4, and 0.51 g, respectively.

For the reinforced model, the fundamental frequency remained constant at 13.18 Hz

during the first three white noise tests. A slight decrease in the fundamental frequency,

from 13.18 to 12.69 Hz, was observed after tests S8 and S9 (PGA = 0.4 g). The decrease

in the fundamental frequency may be attributed to microcracking formed inside the walls.

The frequency then remained constant until the end of the test (PGA = 0.95 g), indicating

that the external constraint from the reinforcement strips prevented the cracks from

developing further. The damping ratios were estimated to be 10.31 % for the reinforced

model at 0.1 g. The damping ratios of the reinforced model were in the range of

9.71–10.36 % with increase in PGA (9.71 % at 0.22 g, 10.23 % at 0.4 g, 9.88 % at 0.51 g,

10.36 % at 0.62 g and 10.11 % at 0.95 g). The results show that the proposed technique

13.18

13.18

13.18 12.69 12.69 12.69 12.69

8.79

8.79

7.32

4.88

0 2 4 6 8

10 12 14

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Freq

uenc

y (H

z)

PGA (g)

Reinfored

Unreinforced

Fig. 9 Fundamental frequencies of the two models with increasing PGA

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can effectively maintain the integrity of the rammed-earth structure and prevent crack

development and propagation. Moreover, the fundamental frequency of the intact rein-

forced model was approximately 1.5 times higher than that of the unreinforced one.

Adjusting for the added weight of the strips and adhesives, the stiffness of the model

increased by about 2.34 times, revealing that the retrofitting technique can greatly improve

the overall stiffness of the structure.

5.3 Comparison of in-plane deformation

The peak displacements at different points on the wall were extracted as shown in Fig. 10.

The maximum displacement of the unreinforced model increased almost linearly before the

seismic excitation reached 0.4 g; furthermore, the maximum displacement at the eave

height and roof height in the X direction (Fig. 10a) coincided with that at the foot of the

wall. The displacement results combined with the evolution of the fundamental frequency

show that although the stiffness of the unreinforced model decreased from 8.79 to 7.32 Hz

at 0.22 g, the model still behaved as a rigid structure. However, when the seismic intensity

reached 0.51 g in the X direction, the maximum displacements at the roof and at the eaves

Fig. 10 Absolute maximum displacements of the reinforced and unreinforced models recorded at differentheights on the walls, a X direction; b Y direction

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increased significantly from 28.6 to 63 mm and from 27.9 to 56.9 mm, respectively,

indicating that both the integrity and stiffness of the model were greatly reduced. A similar

trend was found for the maximum displacements in the Y direction (Fig. 10b). The tests

also showed that because of the inherent brittle property and low ductility of the rammed-

Fig. 11 Dynamic amplification factor at different heights in all the seismic excitation conditions of the test.a At the eaves in the X direction, b in the roof in the X direction, c at the eaves in the Y direction

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earth material, rammed-earth structures always collapse suddenly during an earthquake,

leaving little time for residents to escape from the building. Unlike the unreinforced model,

the reinforced structure behaved as a rigid structure during the testing process except for a

slight relative movement between the eaves and the base of the structure at a seismic

excitation of 0.95 g. This confirmed that the proposed technique improved the structural

stiffness of the building while maintaining its integrity.

5.4 Wall response and roof–wall interaction

The dynamic amplification factor was calculated as the ratio of the acceleration time series

at each point to the base input acceleration. The amplification of the dynamic response of

the walls was observed. The amplification factor at the roof was higher than at the eaves of

the same wall. The dynamic amplification factor of the reinforced model was consistently

higher than that of the unreinforced model (Fig. 11), indicating that the retrofitting greatly

increased the structural stiffness of the building.

The roof–wall interaction was measured by the relative maximum displacements, cal-

culated as the difference between the absolute maximum displacement of the roof and the

foot of the structure. The results are shown in Fig. 12. For the unreinforced model, the

rammed-earth wall and the roof moved almost in phase during the first test runs. As the

simulated earthquake intensity increased, the relative displacements between the roof and

walls increased slightly, indicating that the integrity of the model was gradually weakened

as cracks formed and developed. A sudden increase in the relative movement, from 3.5 to

29 mm, was observed at the seismic excitation of 0.51 g when the model was close to

collapsing; roof tiles fell on the ground; and movement of the roof beams was observed. In

the unreinforced model, no significant relative displacement occurred between the wall and

the roof. Even at 0.95 g, the maximum relative displacement was 4.762 mm, confirming

that the proposed technique can effectively maintain the integrity of the rammed-earth

structure under strong earthquake forces.

Fig. 12 Difference between maximum displacement of the wall and the roof for the two models

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6 Conclusion

Rammed-earth structures are still widely used in less developed rural communities

worldwide. Shrinkage cracks often form in rammed-earth walls during their construction

and subsequent usage. When subject to seismic loads, existing and new cracks easily

develop, significantly reducing the structural integrity and stiffness of the structure; this

may lead to loss of stability of the walls even at low seismic intensities.

In this study, the seismic performance of rammed-earth structures was investigated, and

a retrofitting technique was presented. The technique, aimed at reinforcing rammed-earth

buildings against seismic forces, was evaluated by conducting shaking table tests of

rammed-earth dwelling models with and without reinforcement. The results show that the

proposed retrofitting technique significantly increases the seismic performance of the

rammed-earth structure by improving the load-bearing capability (from 0.4 g to over

0.95 g), the structural stiffness (improved by a factor of 2.33), and the structural integrity

of the building. The use of this technique can effectively prevent the occurrence and

propagation of cracks and make the structure less brittle. This is meaningful because even

in a strong earthquake when the building will not withstand the seismic force, the increase

in ductility can extend the time between the earthquake and the collapse of the building,

which gives the occupants crucial time to escape to a safe location.

Based on our experimental and analytical results, we conclude that the proposed ret-

rofitting technique provides a simple and practical measure to strengthen existing rammed-

earth dwellings against seismic forces with minimal disturbance to the existing structure.

The proposed retrofitting technique can be adopted in rural areas of less developed

countries where rammed-earth dwellings that are not designed to withstand seismic forces

are widely used. This can be a practical way to increase the disaster resilience and sig-

nificantly reduce casualties during earthquakes. Future research should address issues such

as the durability of the retrofit materials and the applicability of the technique to other

types of earth structures using various soil and construction techniques, before applying the

technique to rural communities worldwide.

Acknowledgments This study was supported by the Fundamental Research Funds for the CentralUniversities, the Ministry of Education, China, and the RETROFIT research project sponsored by theInternational Center for Collaborative Research on Disaster Risk Reduction (ICCR-DRR).

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