3. experimental tests on adobe...

3. EXPERIMENTAL TESTS ON ADOBE MATERIAL

For several years, the Pontificia Universidad Católica del Perú (PUCP) has dealt with the characterization of the adobe material and the seismic reinforcement for earthen constructions. In 1978-9 the PUCP carried out tests on adobe specimens in order to compute strength values, mainly compression strength and shear strength. Dynamic tests on the shake table were also carried out for understanding the seismic behaviour of unreinforced and reinforced adobe buildings.

This chapter summarise the results from static (compression test, diagonal compression test and shear test), pseudo-static and dynamic tests carried out at the PUCP (these two last carried out in 2004-5) analyzing the results for calibrating the adobe material parameters to be used in numerical models.

3.1 COMPRESSION TESTS ON ADOBE CUBES

Adobe brick units of different nominal dimensions were fabricated for axial compression tests, as reported by Blondet and Vargas [1978]. The two types of units had 0.2x0.4x0.08 and 0.3x0.6x0.08 m nominal dimensions. The bricks were made of clay soil mixed with straw, and dried under the sun for approximately 2 weeks (Figure 3.1).

Figure 3.1. Dry process of adobe bricks.

In order to reduce variability, all the adobe bricks were fabricated by one person. The adobe cubes with a 0.08 m side were obtained directly from the adobe bricks.

Sabino Nicola Tarque Ruíz

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The results of the axial compression tests on the adobe cubes are shown in Table 3.1. It is seen that the compression strength of the adobe cubes did not vary substantially over the time.

Table 3.1. Results of compression tests on adobe cubes.

Compression strength (MPa) Specimen ID

Age > 1 year Age= 1 month

1 1.80 1.30 2 1.40 1.60 3 1.20 1.40 4 1.50 1.70 5 1.30 1.50 6 1.50 1.50

Mean 1.44 1.50

3.2 COMPRESSION TESTS ON ADOBE PRISMS

The objective of these tests was to compute the maximum compression strength, fc, of the composite adobe and mortar, the elasticity modulus, E , and the strain at failure.

3.2.1 Specimens

A total of 120 adobe prisms were built by Blondet and Vargas [1978] and Vargas and Ottazzi [1981]. The dimension of the specimens varied according to the slenderness ratio thickness:height (1:1, 1:1.5, 1:2, 1:3, 1:4 and 1:5). The adobe bricks used had dimensions of 0.20x0.40x0.08 m, and were laid on top each other with mortar in between: 89 specimens were built with mud mortar and 31 with a combination of cement, gypsum and mud for mortar.

In this thesis only the adobe prisms built with mud mortar are reported. The irregularity of the top part of each prism was corrected adding a cement/sand mortar, thus obtaining a horizontal surface was obtained. Then, two steel plates of 0.20x0.40x0.02 m were placed at both ends of each pile and then loaded axially. During the specimen’s fabrication some 0.01 m steel bars were left inside each pile. Deformeters where placed at the end of each bar, and one or two were placed at the top of each specimen, as seen in Figure 3.2.

Numerical modelling of the seismic behaviour of adobe buildings

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Figure 3.2. Compression test on an adobe prism.

3.2.2 Testing

The axial load was applied perpendicular to the mortar joints with 2.45 kN increments up to failure of the specimen. The test was force controlled. The axial deformation was measured with the deformeters left in each adobe prism. In all cases, the observed failure was brittle, and cracks did not follow a common pattern. For example, in some specimens the cracks started at the central part and went diagonally to the upper part, while in the others the cracks were parallel to the load application (vertical). It should be known that it is hard to observe a non britlle failure under force control.

3.2.3 Results

Table 3.2 shows in detail the results of the compression tests carried out by Blondet and Vargas [1978]. The strain measure by the lateral deformeters is represented by p, and the one by the upper deformeters by s . However, the strain measures by s are not reliable due to initial relative movement of the steel plate and adobe prisms that can influence the strain readings. All the adobe prisms in Table 3.2 had a constant slenderness ratio 1:4 and were tested after one month construction.

Table 3.3 shows a summary of the results reported by Vargas and Ottazzi [1981], including the mean value obtained from Table 3.2. In these tests the variation of slenderness ratio did not considerably affect too much on the compression strength. However, it was recommended to test prisms of slenderness ratio 1:4. It was also observed that the age of the specimens can be an important factor; however, more tests should be carried out to study this matter. As a preliminary conclusion it can be established that the compression strength for prisms of slenderness 1:4 is between 0.80 and 1.20 MPa, depending on the specimen age.


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Table 3.2. Results of compression tests on adobe prisms carried out by Blondet and Vargas [1978].

ID Dimension

(m)

Maximum load (kN)

Compression strength

(MPa)

p

(mm/mm x10-3)

s

(mm/mm x10-3)

C-1 0.19 x 0.385 x 0.79 62 0.85 10.09 12.70 *C-2 0.20 x 0.40 x 0.80 59 0.74 -- -- C-3 0.20 x 0.39 x 0.80 62 0.8 11.53 15.26 C-4 0.195 x 0.39 x 0.785 57 0.74 9.30 9.64 C-5 0.195 x 0.39 x 0.80 65 0.86 8.67 13.90 C-6 0.20 x 0.39 x 0.79 65 0.83 10.61 13.65 C-7 0.19 x 0.39 x 0.79 62 0.84 7.68 13.25 C-8 0.20 x 0.40 x 0.80 59 0.73 10.64 9.88 C-9 0.195 x 0.39 x 0.80 66.25 0.87 8.57 10.51 C-10 0.195 x 0.39 x 0.785 65 0.85 8.20 11.02 C-11 0.195 x 0.39 x 0.78 73 0.96 10.87 11.96 C-12 0.195 x 0.39 x 0.785 70 0.92 9.36 10.17

Mean 0.83 9.60

Table 3.3. Results of compression tests on adobe prisms carried out by Vargas and Ottazzi [1981].

ID Slenderness Mean

Maximum load (kN)

Mean Compression

strength

(MPa)

Age # of

specimens

J 1:1 (2 bricks) 94.13 1.18 1 month 8 I 1:1.5 (3 bricks) 68.2 0.85 1 month 10 H 1:2 (4 bricks) 62.97 0.79 1 month 10 G 1:3 (6 bricks) 64.25 0.80 1 month 9 C 1:4 (8 bricks) 63.98 0.83 1 month 12 K 1:5 (10 bricks) 61.8 0.73 1 month 10

CP3 1:4 (8 bricks) 96.2 1.20 11 months 7 CMB 1:4 (8 bricks) 97.6 1.22 7 months 23

Figure 3.3 shows some plots of the stress-strain curves obtained from the previous tests. It is clear that these tests were force controlled and the inelastic part could not traced.


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The load was applied with a hydraulic jack operate manually. The maximum peak strains seem to be larger than the ones for masonry and concrete. More tests, displacement controlled, should be carried out to investigate the inelastic properties of adobe.

0.00

0.20

0.40

0.60

0.80

1.00

0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 0.016

Strain (mm/mm)

Com

pres

sion

str

engt

h (M

Pa)

F-1 and F-2F-3

a) Compression strength vs. strain, specimen C-1

0.00

0.20

0.40

0.60

0.80

1.00

0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 0.016

Strain (mm/mm)

Com

pres

sion

str

engt

h (M

Pa)

F-1 and F-2F-3

c) Sketch of the adobe prisms

b) Compression strength vs. strain, specimen C-3 Figure 3.3. Stress-strain curves for axial compression tests on adobe prisms (modified from Blondet

and Vargas [1978]). F-1 and F-2 refers to the lateral deformeters, while F-3 refers to the upper deformeters.

3.2.3.1 Elasticity modulus

The elasticity modulus was computed from the elastic part of the stress-strain curves, taking the 50% of the maximum compression load and its correspondent deformation. From the specimens tested by Blondet and Vargas [1978] a mean elasticity modulus E= 100 MPa was obtained from readings of lateral deformeters. The values computed from


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the upper were discarded because the readings included relative movements between the prisms and the steal headings.

Blondet and Vargas [1978] suggest using values around 170 MPa for E, which were indirectly computed from full adobe wall tests.

3.3 DIAGONAL COMPRESSION TEST ON ADOBE WALLETS

The diagonal compression tests were carried out to determine the tensile strength, ft, which produces diagonal cracking on the composite adobe and mortar joints.

3.3.1 Specimens

In a first campaign 10 square wallets of 0.6x0.6x0.2 m were built using 0.2x0.40x0.08 m adobe bricks, which imply 6 layers of 1 ½ adobe bricks (Figure 3.4). The load was applied at two opposite corners of the wallet. Instrumentation to measure the diagonal deformations were left in each adobe panel, which are used to compute the shear modulusG . For the second group of tests [Vargas and Ottazzi 1981], 7 panels were built and tested vertically. More precise equipment for load application and to read deformations was used.

Figure 3.4. Scheme of the diagonal compression test (modified from NMX-C-085-ONNCCE

[2002]).

3.3.2 Testing

In the first group of tests, the 9 wallets were built horizontally and tested in the same position, as seen in Figure 3.5a. The maximum load capacity of the hydraulic jack was 100 kN. The load was applied manually at 0.5 kN increments at the wall corners until the specimen failure. Deformations at the diagonals were properly measured; however, some errors were expected due to the movement of the steel heads placed at the wallet corners.


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For the second group of tests the panels were tested vertically as seen in Figure 3.5b. The diagonal load (at the panel corners) was applied which a velocity of 2 kN/min.

a) Masonry horizontally tested b) Masonry vertically tested

[Blondet and Vargas 1978] [Vargas and Ottazzi 1981]

Figure 3.5. Diagonal compression tests on adobe wallets.

3.3.3 Results

Table 3.4 shows the summarized results obtained from the diagonal compression tests. The maximum tensile strength, tf , is evaluated using the following Equation [Brignola et al. 2008].

tPfl tmax1

2 (3.1)

where Pmax is the maximum applied load, l is the lateral dimension of the square wallet, and t is the wall thickness.

Table 3.4. Results of diagonal compression tests on adobe wallets carried out by Vargas and Ottazzi [1981].

ID Maximum tensile

stress (MPa) Shear modulus

(MPa)

D-1 0.03 67.0 D-2 0.03 19.4

D-3A 0.033 25.0 D-3B 0.024 15.8


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Table 3.4. Continuation. Results of diagonal compression tests on adobe wallets carried out by Vargas and Ottazzi [1981].

ID Maximum tensile

stress (MPa) Shear modulus

(MPa)

D-4 0.027 17.0 D-5 0.027 90.7 D-6 0.027 11.0 D-7 0.027 90.6 D-8 0.027 17.1 D-9 0.024 24.6

CDPM-1 0.032 50.0 CDPM-2 0.026 34.0 CDPM-3 0.027 40.3 CDPM-4 0.019 --- CDPM-5 0.017 51.8 CDPM-6 0.013 46.7 CDPM-7 0.025 35.8

Mean 0.026 39.8 COV 7.6% 61.4%

3.3.3.1 Shear modulus G

The shear modulus G is evaluated using Equation (3.2), considering the 50% of the applied load and the corresponding tangential strain, , of the stress-strain curve.

PGA

0.707 (3.2)

The tangential strain is given as the sum of the tensile t and compressive c strain -Equation (3.3)- evaluated from the relative displacement between two controls points in each wallet diagonal.

t c (3.3)

Blondet and Vargas [1978] suggest using a shear module G MPa70 (varying from 36 to 90 MPa), and a Poisson module 0.2 (varying from 0.15 to 0.25).


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3.4 STATIC SHEAR WALL TESTS

The objective of these tests was to evaluate the masonry shear strength under the influence of overburden compressional loads and seismic forces. The shear strength involves the action of shear bond strength and shear friction.

3.4.1 Specimens

A total of 18 walls using 0.2x0.3x0.08 and 0.3x0.6x0.08 m adobe bricks were built in a first group of tests [Blondet and Vargas 1978]. From this group, 10 walls had no reinforcement. The wall dimensions were 2.40x2.40 and 4.00x2.40 m, with wall thickness of 0.2, 0.3 and 0.4 m. Two walls had window openings and one had a perpendicular wall. The characteristics of the wall are given in Table 3.5 and the scheme is shown in Figure 3.6, where it is seen that the load is applied at 2/3 of the wall height.

Table 3.5. Characteristics of adobe walls for static shear test [Blondet and Vargas 1978].

ID Wall

thickness (m)

Wall dimension

(m)

Load applied at

Characteristic Overburden

load

11Ea 0.40 0.40x2.40x2.40 2/3 h --- No 11Eb 0.20 0.20x2.40x2.40 2/3 h --- No 12E 0.20 0.20x2.40x2.40 2/3 h --- No 13E 0.30 0.30x4.00x2.40 5/6 h --- No 15E 0.30 0.30x4.00x2.40 2/3 h --- No 21E 0.30 0.30x4.00x2.40 2/3 h Windows opening Yes 22E 0.30 0.30x4.00x2.40 9/10 h Windows opening Yes 23E 0.30 0.30x4.00x2.40 2/3 h --- Yes 24E 0.30 0.30x4.00x2.40 2/3 h --- Yes 25E 0.30 0.30x4.00x2.40 2/3 h T shape No

Another group of tests was performed by Ottazzi et al. [1989] and Vargas and Ottazzi [1981] who used a more precise equipment for load application and deformation reading. In this case 4 shear walls were built with 0.20x0.40x0.08 m adobe bricks. The final dimensions of the walls were 0.20x2.0x2.0 m. A sketch of the tested walls is shown in Figure 3.7. The tests were force controlled, with the horizontal load applied at the top wall by a hydraulic actuator. Horizontal load increments of 0.35 kN were applied for 2 walls, considering a vertical load of 72 kN, and 1 kN increment for the other 2 walls with vertical loads of 10 kN.


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Figure 3.6. Scheme of the adobe walls for static shear test carried out by Blondet and Vargas

[1978].

Figure 3.7. Scheme of the adobe walls for static shear test carried out by Vargas and Ottazzi [1981]

3.4.2 Tests

For the first group of walls, the horizontal load was applied horizontally, in one direction, with a hydraulic jack controlled with a manometer. The location of the hydraulic jack was at 2/3 and 9/10 of the wall height for 9 and 1 walls, respectively (see Table 3.5). The load was distributed on the wall through a 0.02x0.2x0.4 m steel plate and a 0.05x0.3x0.8 m wooden plate. Pre compression load was considered in some walls before application of the horizontal load (Figure 3.8).

For the second group of walls, the horizontal load was applied monotonically with a hydraulic actuator at the wall top (Figure 3.7).


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Figure 3.8. Adobe walls for static shear tests [Blondet and Vargas 1978].

3.4.3 Results

Blondet and Vargas [1978] gives a relation based on Coulomb-like criterion to compute the shear strength of adobe walls related to the compression strength . The expression is reported in Equation (3.4) to be used in MPa. They conclude that this expression can be also used for reinforced walls.

c 0.01 0.55 (3.4)

c is the shear strength under zero compression stress, is the friction coefficient and is the average normal stress on the compression area.

3.5 CYCLIC TESTS

The objective of these tests was the evaluation of the seismic in-plane behaviour of the adobe walls, through the evaluation of the Force-Displacement curve and the evolution of diagonal cracks at different levels of horizontal top displacements. Blondet et al. [2005] carried out tests on adobe walls, with and without reinforcement. In this thesis, only the seismic behaviour of the unreinforced adobe walls is reported.

3.5.1 Specimens

Three I-shape adobe walls without reinforcement were carried out at the Pontificia Universidad Católica del Perú by Blondet et al. [2005] and Blondet et al. [2008]. In this thesis the walls are identified as Wall I-1 (Figure 3.9a), Wall I-2 and Wall I-3 (Figure 3.9b). The geometrical characteristics were the same for all of them; the only difference was in Wall I-1, which had a 0.4x0.6 m central window opening. The main longitudinal wall was 3.06 m long, 1.93 m high and 0.30 m thick. All walls had two 2.48 m transverse walls (Figure 3.10).


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The geometrical configuration of Wall I-2 and I-3 is the same in Figure 3.10 but without window opening. The adobe bricks used for the wall construction had dimensions 0.13x0.10x0.30 m and 0.13x0.10x0.22 m. The brick composition for the first wall was soil, coarse sand and straw in proportion 5/1/1, and for the mud mortar, 3/1/1. The bricks were laid alternating headers and stretchers in the courses. The adobe composition for the last two walls was not reported.

a) With windows opening (1 specimen) b) Without windows opening (2 specimens)

[Blondet et al. 2005] [Blondet et al. 2008]

Figure 3.9. Adobe walls subjected to the cyclic test.

a) Front view b) Plan view

Figure 3.10. Scheme of the adobe walls for cyclic test [Blondet et al. 2005].

Each specimen was built over a reinforced concrete foundation beam (Figure 3.11a). At the top a reinforced concrete crown beam (Figure 3.11b) was built to provide the gravity loading corresponding to the roof of a typical Peruvian dwelling consisting of wooden beams, cane, straw, mud and corrugated zinc sheets. The total weight of each wall was


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approximately 135 kN, which considers the weight of the concrete beams. The top concrete beam had a 16 kN weight, the foundation beam had a 31.4 kN weight, and the adobe wall had 87.6 kN weight for the I-1 wall and 89.0 kN for the other two walls. The lintel in Wall I-1 was made of wood.

a) Foundation b) Top beam

Figure 3.11. Scheme of the reinforced concrete ring beams [Blondet et al. 2005].

The load was applied horizontally at the top concrete beam through a servo hydraulic actuator with a maximum capacity of 500 kN, placed on a rigid steel frame. In order to avoid a concentrated load, steel and wooden plates were placed at the contact of the actuator with the crown beam (Figure 3.12a). Besides, two steel rods were placed at the wall top to improve the horizontal load transmission all over the wall, to simulate a distributed horizontal load. A total of 17 Linear Variable Displacement Transducer (LVDT) was placed in the wall. One side of the wall was painted white to help the cracks visualization.

a) Contact of the actuator with the wall b) Set-up of instrumentation on the adobe wall

Figure 3.12. Detail of load application and instrumentation on the adobe walls [Blondet et al. 2005].


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3.5.2 Tests

The tests were displacement controlled. The load application velocity was incremented in each phase as reported in Table 3.6. The cyclic load was applied in 7 phases, with 2 cycles for each phase, as seen in Figure 3.13. The maximum displacements, incremented in each phase, were 0.1, 0.5, 1, 2, 5, 10 and 20 mm.

Figure 3.13. Load application history for the cyclic tests.

The crack evolution was quite similar for all the walls as described in Table 3.6. The first phase was useful for calibration of the instrumentation. The diagonal fissures (x-shape) started to appear during the third and fourth phases, with loss of strength. In the fifth phase large horizontal fissures appeared at the transversal walls and vertical fissures at the intersection of longitudinal and transversal walls, with increment of diagonal cracking in the main wall.

Table 3.6. Description of damage on the walls subjected to the cyclic tests.

Phase Wall ID (mm) Velocity

(mm/ min)

Maximum load (kN)

Crack thickness (mm)

I-1 --- --- I-2 --- --- 1 I-3

0.1 0.1 --- ---

I-1 26 --- I-2 36 --- 2 I-3

0.5 0.5 42 ---

I-1 32 0.60 I-2 49 0.10 3 I-3

1 1.0 52 0.25

Cycle 1 Cycle 2

1st Phase

Push

Pull


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Table 3.6. Continuation. Description of damage on the walls subjected to the cyclic tests.

Phase Wall ID (mm)Velocity

(mm/ min)

Maximum load (kN)

Crack thickness (mm)

I-1 38 0.60 I-2 48 0.50 4 I-3

2 2.0 53 0.50

I-1 38 4.0 I-2 40 2.5 5 I-3

5 5.0 41 2.5

I-1 34 10.0 I-2 43 6.0 6 I-3

10 10.0 39 6.0

I-1 32 50.0 I-2 44 15.0 7 I-3

20 20.0 39 15.0

During the sixth phase (corresponding to 10 mm as top displacement) a notable loss of load strength in the walls was observed with an increment of crack thickness and tensile cracking in the adobe bricks. The diagonal cracking in both directions continue growing in thickness. Horizontal cracks appeared at the base of the transversal walls, allowing a sliding behaviour of the walls. At this stage, some rigid blocks were identified. The last phase showed a complete loss of load strength with formation of cracks that cut the adobe bricks in the principal and transversal walls. The last phase involved sliding of the walls with greater crack width.

3.5.3 Results

All the walls had a similar in-plane behaviour: x-shaped diagonal cracks, horizontal cracks in the transversal walls, and loss of strength after 2 mm top displacement. The two walls without windows had larger initial stiffness than the wall with opening, as seen in the hysteretic curves in Figure 3.14. In terms of lateral force, Wall I-2 and Wall I-3 resisted around 25% more than Wall I-1. The last two walls were used for another research program to investigate the effect of crack grouting. The injected cracks in the Wall I-2 and I-3 are clearly identified in the right photos in Figure 3.14.

It seems that at the beginning of loading there is an adjustment of the hydraulic actuator and instrumentation, giving an apparent initial stiffness greater than the real one. To


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avoid this problem, the initial stiffness should be computed from the hysteretic curves when the maximum displacement at the top is around 1mm. Beyond this displacement some fissures started to appear in all the walls.

-60

-40

-20

0

20

40

60

-25 -20 -15 -10 -5 0 5 10 15 20 25

Displacement (mm)

Forc

e (k

N)

a) Load-displacement response and crack patterns of the wall I-1

-60

-40

-20

0

20

40

60

-25 -20 -15 -10 -5 0 5 10 15 20 25

Displacement (mm)

Forc

e (k

N)

b) Load-displacement response and crack patterns of the wall I-2

-60

-40

-20

0

20

40

60

-25 -20 -15 -10 -5 0 5 10 15 20 25

Displacement (mm)

Forc

e (k

N)

c) Load-displacement response and crack patterns of the wall I-3

Figure 3.14. Hysteretic curves and crack pattern of the adobe walls subjected to the cyclic tests [Blondet et al. 2005; 2008]. All the right figures refer to the end of the cyclic tests.


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The hysteretic curves yield useful results up to a 10 mm top displacement. Beyond this value sliding behaviour of the wall due to the relative movement between rigid blocks along the cracks was observed. Figure 3.14 shows the hysteretic curves for each wall and a picture representing the crack pattern at the end of the cyclic tests. For Wall I-1, the x-shape cracks started at the window corners and grew diagonally to the top and base of the wall. The wider cracks were formed when the wall was pushed. Unlike the previous wall, Wall I-2 and Wall I-3 were repaired and tested again; however, the responses of the repaired walls are not reported here. Figure 3.14b and Figure 3.14c show the repaired crack patterns.

Figure 3.15 shows the envelopes of the experimental hysteretic curves, in both positive and negative directions. Since the wall was first pushed, the positive branch of each hysteretic curve is stronger than the negative branch, which represents the behaviour of the wall when it is pulled. To compute the initial part of the envelope curve, the data obtained before 1mm top displacement was neglected since it can contain noise due to the equipment calibration.

0

10

20

30

40

50

60

0 1 2 3 4 5 6 7 8 9 10Displacement (mm)

Forc

e(k

N)

Wall I-1, push

Wall I-2, push

Wall I-3, push

0

10

20

30

40

50

60

0 2 4 6 8 10Displacement (mm)

Wall I-1, pullWall I-2, pullWall I-3, pull

a) Positive branch of the hysteretic curves b) Negative branch of the hysteretic curves

Figure 3.15. Envelope of the hysteretic curves (positive and negative branch) from the cyclic tests.

3.5.4 Evaluation of the elasticity modulus

Some useful equations for the evaluation of the wall stiffness, considering deformations due to shear and bending, are reported as follows:

Stiffness when the wall is double fixed:

v

Kh hE I G A

31

12

(3.5)


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Stiffness for cantilever walls (one end is fixed and the other is free):

v

Kh hE I G A

31

3

(3.6)

Total stiffness obtained from contribution of walls connected in series:

n

iT iK K1

1 1 (3.7)

Total stiffness obtained from contribution of walls connected in parallel:

n

T ii

K K1

(3.8)

In the previous equations h is the wall height, E is the elasticity modulus, I is the area moment of inertia, G is the shear modulus taken as E0.4 and vA is the effective shear area. This last value has a shear deformation factor of 1.2.

3.5.4.1 Wall I-1

The experimental initial stiffness K can be computed from the pushover curves showed before. For wall I-1 the initial stiffness is estimated as:

FK kN mm1 28 /

The influence of the central window should be taken into account for computing analytically the stiffness as follows:

1) The complete wall stiffness is due to the contribution of 4 parts, as seen in Figure 3.16. The lower part can be assumed as double fixed, while the other three parts (where the 2 blocks next to the windows are in parallel) can be assumed as cantilever walls. The total initial stiffness is given by the contribution of each part.

Part A. Considering Equation (3.5) with h mm876

AEK E347

0.0000174 0.00286


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It can be seen that the deformation due to bending is almost negligible. Therefore, considering a cantilever beam or a double fixed element will not influence the final evaluation of the initial stiffness. For Part A, considered as a cantilever beam, AK is 341E.

Figure 3.16. Scheme for evaluating the lateral stiffness of the adobe wall I-1.

Part B and Part C. Considering Equation (3.6) with h mm600

B CEK K E219.1

0.00005564 0.004511

Again, it is seen that the influence of the bending deformation is small comparing it with the shear deformation. For Part B and C, considered as a double fixed beam,

BK is 221E.

Part D. Considering Equation (3.6) with h mm454

DEK E6 669.7

9.87 10 0.00148

The total initial stiffness is computed combining Equations (3.7) and (3.8) as follows:

T A B C DK K K K K1 1 1 1

Evaluating the previous equation, the total stiffness is TK E150 . So E MPa187


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3.5.4.2 Wall I-2 and I-3

1) In this case there is not a window opening. The wall can be assumed as fixed at both ends with a distributed horizontal load at the top. The experimental initial stiffness computed from the pushover curves is around TK kN mm34 /

The area moment of inertia is:

I mm3 3

12 42480 3060 1090 24602 3.217 1012 12

The effective shear area is given by the wall web:

v wA A mm25 5 2460 600 300 7650006 6

Solving for Equation (3.5):

K E E/ 0.0001862 0.0063 154 . So, E MPa220 .

2) Rotation at the top part of the wall is accepted (cantilever beam); so, the initial stiffness is evaluating with Equation (3.6), resulting in E MPa240 .

3) Paulay and Priestley [1992] suggest to compute an effective width on wide-flanged walls to evaluate the seismic capacity of wall subjected to shear forces, as seen in Figure 3.17, this is based on the assumption that vertical forces due to shear stresses introduced by the web of the wall into the tension flange spread out at a slope of 1:2.

According to Figure 3.17, the effective width of the tension flange is expressed as:

eff w wb h b b (3.9)

where wh is the height wall, wb is the in-plane wall width and b is the length of the flange wall.

The effective width in compression is given by:

eff w wb h b b0.3 (3.10)

With the two expressions written above, the effective widths for the flanges are 2230 mm and 879 mm for the tension and compression zone, respectively.

The coordinates of the gravity centre along the application of the force is:


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y mm2230 300 150 2460 300 1530 879 300 2910 1189.222230 300 2460 300 879 300

The area moment of inertia is aI mm12 41.97 10

The effective shear area is given by the wall web:

v wA A mm25 5 2460 600 300 7650006 6

Solving for Equation (3.6): K E E/ 0.0003 0.006307 151 . So, E MPa224 .

Figure 3.17. Estimation of effective flange widths in structural walls (modified from Paulay and

Priestley [1992]).

It is seen that the analytical computed elasticity values for the elasticity modulus are around 220 MPa. Blondet and Vargas [1978] have shown high variability in the elasticity modulus, varying from 100 MPa to 250 MPa, but they suggest using values around 170 MPa. The elasticity modulus is more sensitive to shear deformation than flexural deformation. As previously discussed, it seems that there is not major difference in assuming a double fixed wall or a cantilever wall. From the analysis presented here it can be concluded that a E MPa200 220 can be used for the numerical analysis.


50

3.6 DYNAMIC TESTS

The objective of these tests was the complete evaluation of the seismic capacity of adobe structures, considering simultaneously the in-plane and out-of-plane actions in the adobe walls, and the evaluation of failure patterns. The seismic demand is represented by a displacement signal applied at the base of the structure and related to an acceleration record of a Peruvian earthquake occurred in 1970. Blondet et al. [2006] and Blondet et al. [2005] carried out a group of dynamic tests on typical adobe modules (see Figure 3.18) at the PUCP; one of them did not consider seismic reinforcement and it will be evaluated in this thesis.

MURO WMURO E

MURO N

MURO S

3.20

2.710.250.25

Ventana

Ventana

1.12

0.96

1.12

Puerta HILADAIMPAR

a) Plan view b) Front wall

c) Rear wall d) Right and left wall

Figure 3.18. Scheme of the adobe module subjected to a dynamic test [Blondet et al. 2005].

3.6.1 Specimens

The module was built over a reinforced concrete foundation and had a total weight (module + foundation) of around 135 kN. The weight of the concrete beam was 30 kN.


51

The adobe bricks and the mud mortar used for the construction of the module had soil/coarse sand/straw proportions of 5/1/1 and 3/1/1; respectively. These proportions were the same as those used for construction of Wall I-1 in the cyclic test. The nominal adobe brick dimensions were 0.25x0.25x0.07 m. The idea of the module was to represent a part of a typical vernacular Peruvian adobe building within the limitations of the 4.0x4.0 m shake table sides, which allows a maximum specimen weight of 160 kN. All the adobe walls were built considering normal stretcher bond.

The module had three openings: 1 door at the frontal wall, and 1 window in each lateral wall. These lateral walls had tapered height, started at 1.98 m at the frontal wall until 2.25 m at the rear wall (Figure 3.18, Figure 3.19). The square plan of the specimen had sides of 3.21 m. The lintels for all openings were made with cane rods and mud. The seismic signal was applied perpendicular to the front and rear walls.

Figure 3.19. Views of the adobe module subjected to the dynamic test.

The roof was built as follows: 4 lateral wooden 0.05x0.10 m beams were placed on the direction of the lateral walls, 1 beam over each lateral wall and 2 beams connecting the 2 perpendicualr walls. The lateral wooden beams were attached to the walls with mud and steel nails; this connection does not allow modelling the roof as a rigid diaphragm. Over


52

the lateral beams, 0.05x0.05 m wooden joints were placed perpendicularly to support the clay tiles, as seen in Figure 3.19. A square space was left at the centre middle for facilitating transportation of the module from the open construction field to the shake table.

A mud plaster with 3/1/1 proportions was applied internally and externally to the Right Wall (lateral wall). The thickness of this wall was around 0.28 m, while in the others it was 0.25 m. Ten accelerometers and 8 LVDTs were left in the model distributed in all the walls, and 1 accelerometer and 1 LVDT were left at the shake table (Figure 3.20). The module was tested after more or less two weeks from the construction end.

Figure 3.20. Position of accelerometers and LVDTs on the adobe module.

3.6.2 Shake table description

The shake table is made with a pre-stress concrete slab that weights 160 kN. The slab is supported by 8 metallic vertical plates that are pinned at both ends to allow the horizontal movement of the table. In plan, the shake table is 4x4 m and the maximum supported weight is around 160 kN.

The slab is pushed back and forth by a hydraulic actuator, which it is fixed at one side of the table and at one side of a reaction wall that weights 6180 kN. The actuator has a maximum force capacity of around 313 kN and is displacement controlled. The total displacement of the actuator is 300 mm ( 150 mm).

The seismic signal used for the dynamic tests was based on the horizontal acceleration record from the May 31st, 1970, Peruvian earthquake, component N08W recorded in Lima (seismic station of the Geophysics Peruvian Institute, IGP, Figure 3.21). This


53

earthquake had a Mw 7.9 magnitude, a maximum intensity XI in MMI and generated an avalanche in Huaraz (northern Peruvian city). The total number of fatalities was around 66000. Normally Peruvian earthquakes are caused by the interaction between the Nazca and South American tectonic plates (subduction process).

-0.15

-0.1

-0.05

0

0.05

0.1

0.15

0 5 10 15 20 25 30 35 40 45

Time (s)

Acc

eler

atio

n (g

)

Figure 3.21. Horizontal acceleration record from May 1970 Peruvian earthquake, component N08W,

registered in Lima.

The recorded acceleration signal was filtered to remove frequencies lower than 0.15 Hz. Then, the signal was integrated twice to compute the displacement record, and filtered again to remove frequencies larger than 15 Hz. A time segment of 27 s was taken and smoothed at the beginning and end of the signal. The displacement signal can be scaled to have different maximum absolute displacements and to simulate different earthquake intensities.

3.6.3 Test

The adobe module was subjected to three levels of displacement signal, which was scaled to have maximum displacements of 30, 80 and 120 mm at the base (Figure 3.22). The input displacement signals correspond to PGA of 0.3, 0.8 and 1.2g, respectively. These earthquake levels (phases) intend to represent the effects of a frequent, moderate and severe earthquake on the adobe buildings. The displacement signal was applied parallel to the two walls with window (identified as Right and Left Wall). Before each phase, a rectangular pulse was applied to measure the free vibration motion of the module.During the first phase (30 mm maximum displacement at the base) few slightly diagonal and vertical fissures, appeared at the right and left walls (which are parallel to the signal). The lateral wooden beams of the roof started to loose the mechanical connection with the walls demonstrating that the nails and mud were not sufficient to maintain the roof-wall integrity. However, no relative displacement was seen between roof and walls. In phase 2, complete vertical cracks appeared at the wall corners, allowing the separation of the walls. The diagonal cracks incremented in width and new cracks appeared in all the walls.


54

a) Adobe module after the Phase 1

b) Snapshots of the adobe module during the Phase 2

c) Adobe module after Phase 3

Figure 3.22. Views of the adobe module during and after the dynamic test.

Cracks due to vertical and horizontal bending were also observed in the perpendicular walls (identified as Front and Rear Wall) with initiation of rocking of some adobe


55

masonry blocks. A complete separation of the roof from the walls was observed; however, the roof was maintained in position just by its own weight. At the end of the second phase the LVDTs were removed in order to prevent any damage to them in the next phase. During phase 3, the perpendicular walls collapsed and the parallel walls became instable. The roof was supported by the parallel walls.

3.6.4 Results.

The period of vibrations were measured in each wall during a free vibration test before imposing any load. Before the Phase 1, the computed periods were 0.167 s, 0.151 s, 0.121 s, and 0.167 s measured along the direction of the movement with accelerometers A1, A2, A4, and A5, respectively (Figure 3.20). The observed damage pattern revealed the typical failure modes of adobe walls subjected to in-plane and out-of-plane actions. During the first phase, with a maximum displacement at the base of 30 mm (Figure 3.23), no loss of wall rigidity was observed. During the second phase (with maximum displacement at the base of 80 mm) the adobe walls were separated by vertical cracks and diagonal in-plane cracks on the Left and Right walls. The first vertical cracks appeared at the Right Wall. The Front and Rear walls cracked due to horizontal and vertical bending. The Right Wall, which had stucco, was stiffer than the other walls. The difference in stiffness from the two parallel walls allows the structure to undergo some torsion in the second and third phases.

-40-30-20-10

010203040

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-23.0

30.6

Figure 3.23. Displacement input at the base, Phase 1.

The Rear Wall, which was itself broken more or less into 3 big blocks, had a rocking behaviour due to out-of-plane actions. During the third phase, with maximum displacement of 130 mm at the base, the perpendicular walls collapsed at the beginning of the input signal while the parallel walls were completely cracked. Since the roof was supported by the lateral walls, it did not collapse. The following figures summarize the results obtained from the dynamic test, phase 1 and phase 2. Displacements (D1, D2, D3, D4, Figure 3.24) and accelerations (A1, A2, A3, A4, Figure 3.25) histories were measured at the top of the walls, as seen in Figure 3.20. The greater relative displacements and total acceleration responses were obtained at the walls perpendicular to the movement (Front


56

and Rear walls) because they had a rocking behaviour. All the walls were disconnected due to the vertical cracks at the wall intersections.

-1.5

-1

-0.5

0

0.5

1

1.5

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-0.66

0.81

a) Right wall (the only wall with plaster)

-1.5

-1

-0.5

0

0.5

1

1.5

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-0.89

0.54

b) Left wall

-1.5

-1

-0.5

0

0.5

1

1.5

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-1.34

1.35

c) Front wall (with door opening)

-1.5

-1

-0.5

0

0.5

1

1.5

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-1.39

1.26

d) Rear wall

Figure 3.24. Relative displacement history during Phase 1.


57

-0.5

-0.25

0

0.25

0.5

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.33

0.27


-0.5

-0.25

0

0.25

0.5

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.33

0.29

b) Left wall

-0.5

-0.25

0

0.25

0.5

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.45

0.40


-0.5

-0.25

0

0.25

0.5

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.42

0.36

d) Rear wall

Figure 3.25. Total acceleration history during Phase 1.

The input displacement at the base (shake table) related to phase 2 had a maximum amplitude displacement of 80 mm (Figure 3.26). The displacement and acceleration histories for Phase 2 are shown in Figure 3.27 and Figure 3.28. All the walls have a peak displacement right after 10 s. Afterwards, the two walls parallel to the movement move as a rigid body, while the two perpendicular walls move back and forth with a rocking behaviour. This is due to the formation of vertical cracks that made possible the


58

separation of walls, allowing them to move independently. Besides, it can be seen from the displacement time history that after the walls separate, the stiffer wall was the Right Wall because it was the only wall with mortar and probably because of some torsion related to the shake table.

-100

-50

0

50

100

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

80.73

-61.04

Figure 3.26. Displacement input at the base, Phase 2.

-50

-25

0

25

50

75

100

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-14.97

7.99


-50

-25

0

25

50

75

100

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-19.51

71.07

b) Left wall

-50

-25

0

25

50

75

100

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-39.3

64.3


Figure 3.27. Relative displacement history during Phase 2.


59

-50

-25

0

25

50

75

100

0 5 10 15 20 25 30

Time (s)

Dis

plac

emen

t (m

m)

-47.4

86.8

d) Rear wall

Figure 3.27.- (Continuation). Relative displacement history during Phase 2.

-2-1.5

-1-0.5

00.5

11.5

2

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.76

0.68


-2-1.5

-1-0.5

00.5

11.5

2

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-0.94

0.93

b) Left wall

-2-1.5

-1-0.5

00.5

11.5

2

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-1.79

1.48


Figure 3.28. Total acceleration history during Phase 2.


60

-2-1.5

-1-0.5

00.5

11.5

2

0 5 10 15 20 25 30

Time (s)

Acc

eler

atio

n (g

)

-1.27

1.44

d) Rear wall

Figure 3.28. (Continuation). Total acceleration history during Phase 2.

The difference in displacement history in the Right and Left walls –both parallel to the displacement input- supports the assumption that the roof is not rigid enough to tie completely the walls. Besides, the torsion is due to the fact that the Right wall had a plaster, making it thicker and stiffer than the Left wall. Figure 3.29 shows the relative rotation between the displacement movement of the Right and Left wall. Again, it is seen that the greater effect appears right after 10 s.

-3.0E-02

-2.0E-02

-1.0E-02

0.0E+00

1.0E-02

0 5 10 15 20 25 30

Time (s)

Rot

atio

n (r

ad)

-2.2E-02

0.6E-02

Figure 3.29. Rotation history between Right and Left Wall during Phase 2.

3.7 SUMMARY

This chapter deals with the results from static, pseudo static and dynamic tests carried out at the Pontificia Universidad Católica del Perú on adobe masonry. The compression strength on adobe masonry is between 0.80 and 1.20 MPa and the tensile strength is around 0.03 MPa. An expression for computing the shear strength considering overburden load was reported by Blondet and Vargas [1978]. Since the adobe material is brittle and the static tests were made in the 80’s, there is neither information about the inelastic part of the compression curve nor information about the tensile constitutive law for characterization of the mud mortar or the adobe masonry (i.e. fracture energy).

The elasticity modulus can not be defined by a unique value because of variation in the adobe composotion. A lower bound can be 170 MPa. However, from the analysis of the pseudo static test results values between 200 to 220 MPa were computed. Besides, it was


61

observed that the soil characteristics have a great influence on the strength of the adobe masonry.

From the dynamic tests it was observed that in a building the adobe walls behave quasi independently during earthquakes because of the lack of diaphragm effect. Vertical cracks allowed the loss of connection between perpendicular walls. Taking advantage of the results from the pseudo static and the dynamic test, numerical models of the adobe wall and adobe module are developed in the next chapters to calibrate the adobe material parameters, particularly at the inelastic range (e.g. fracture energy in tension and compression).

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3. experimental tests on adobe...

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