report on rcc

7/29/2019 report on rcc

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Chapter 3

E X PE R IM E N T A L ST U D Y

3.1 I ntroduction

This chapter describes the design, construction and testing of two interior Corcon rib

beam-column subassemblages tested in the Francis Laboratory at The University of

Melbourne. The second test specimen is the Carbon Fibre Reinforced Polymer (CFRP)

repaired version of the damaged original specimen after the first test.

The first section of this chapter describes the design of prototype model structure and the

test specimen. Details of test specimen, the material properties used and the design

parameters used in designing of prototype structure are presented. The details of test set up

and instrumentation used in the test are presented and discussed.

The approximate dead and live loading of the prototype structure was represented by

adjusting the end reactions. The column axial loading was simulated by prestressing

column ends of the test specimen. The specimens were tested in the reaction frame, with

increasing ratios of quasi-static cyclic drift being applied.

Details of CFRP rectification work of the damaged original test specimen are presented in

section 3.7. The related CFRP design process and the additional instrumentation for the

second test are also presented.

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3.2 Design

The planning of the experimental program was done considering configuration of the

existing reaction frame at Francis Laboratory at The University of Melbourne. This was

one of the criteria in planning the test procedure and configuration of the test specimen

due to limited funding available for modification or rebuilding a new test set up. This

reaction frame has been designed to investigate seismic performance of reinforced

concrete wide band beam frame interior and exterior connections (Abdouka, 2003; Stehle,

2002). As found in chapter 2, the existing reaction frame is more or less similar to the

reaction frames used by other researchers for testing beam-column subassemblages. Since

one of the main objective was to assess current Australian design practice and to provide

design guidelines for these beam-slab-column systems constructed with the Corcon form

work, it was decided that the first test specimen should be designed in accordance with

Australian code requirements (AS-1170.4, 1993; AS-3600, 2001). However, basic design

guidelines for rib slabs are not available in Australian code (AS-3600, 2001) and therefore,

main design assumptions and procedure were gathered from British and US codes (ACI-

318, 1999; BS-8110, 1995). The Corcon reinforced concrete rib beam section design was

done using Australian standards. Details of design adopted are presented in Appendix A.

A typical four-storey, six bay ordinary moment resisting framed building constructed with

Corcon system as shown in Figure 3-1, was considered as a prototype model structure for

the study. A mainframe spacing of 6.0 m in the transverse direction was assumed. A

column size used of 500 mm square and beams of 894 mm total depth and 2400 mm wide

flange slab were adopted. These sizes are consistent with typical Corcon formwork



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dimensions. As shown in Figure 3-2, a one way spanning 170 mm thick solid slab was

used between the main flange beams. A half scale of this prototype was used in testing.

The building was designed as an Ordinary Moment Resisting Frame (OMRF), according

to current Australian codes (AS-1170.0, 2002; AS-1170.1, 2002; AS-1170.2, 2002; AS-

1170.4, 1993; AS-3600, 2001), which makes no special detailing requirements mandatory

for these type of frames.

The main frame shown in Figure 3-1 is designed to resist wind and earthquake lateral

loading, while in the transverse direction, lateral loading was assumed to resist by

symmetrically placed shear walls or a perimeter frame. As the prototype structure

considered is symmetrical and no torsional effects. Therefore inelastic analysis can be

limited to simplified two-dimensional analysis.



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Interior subassemblage

8.4 m 9.6 m9.6 m 8.4 m9.6 m 9.6 m

4.2 m

3.4 m

3.4 m

3.4 m

Transverse mainframe spacing 6.0 m

Beams: 2400 mm wide flange, 894 mm deep rib beam

Columns: 500 mm square

Slab: 170 mm one way spanning solid slab between main beams.

Figure 3-1: Prototype frame dimensions.

2.4 m

6.0 m

894 mm

170 mm

Figure 3-2:Di mensions of mainf rame beam section

In order to determine various loading factors, it was assumed that the frame was a office

building (for imposed live load consideration) situated on a rock site in Newcastle,

Australia. It should be noted that Newcastle has the highest peak ground acceleration

coefficient. Wind and earthquake loading parameters for selected cities are listed in Table

3-1. The design loading parameters adopted for the structure as situated in Newcastle, are

given in Table 3-2. The basic appropriate load combinations for the ultimate limit states

used in checking strength of the prototype frame, as specified in Australian code (AS-

1170.0, 2002), are given in Equation 3-1.



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QEGe

WGd

QWGc

QGb

Ga

u

u

u

4.00.1)

9.0)

4.02.1)

5.12.1)

35.1)

Equation 3-1

T able 3-1:Acceleration coefficient for major cities in Australia

City U ltimate wind velocity(m/sec)

Peak ground accelerationcoefficient (g)

Melbourne

SydneyAdelaide

Brisbane

Perth

Hobart

Canberra

Newcastle

Alice Spring

Darwin

50

5050

60

50

50

50

50

50

70

0.08

0.080.10

0.06

0.09

0.05

0.08

0.11

0.09

0.08



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T able 3-2:Design values adopted

Design parameter V alue

Gravity

Superimposed dead load

Live load

Ultimate wind velocity

Region

Terrain category

Topographic multiplier

Shielding multiplier

Importance multiplier (for wind)

Earthquake acceleration coefficient

Site factor

Structural response modification factor

Importance factor (for earthquakes)

9.81 m/s2

1.5 kPa

4.0 kPa

50 m/s

A

2

1.0

1.0

1.0

0.11g

1.0

4.0

1.0

Prototype frame loading was evaluated using actual member self-weights, based on the

sectional dimensions. Earthquake and wind loading was calculated based on the

parameters given in Table 3-2. A limit state design was employed, considering the entire

load combinations specified in Equation 3-1. A structural modification factor of 4.0 was

used, representing a frame of limited ductility. A fundamental period of 0.31 s was

calculated using the Australian code (AS-1170.4, 1993) specified formula (i.e. T=h/46).

As seen from Australian code method, the equivalent static earthquake load, mainly

depends on the accuracy of the fundamental period of the structure. As discussed in

chapter 2, there are different formulae presented in different codes and textbooks. In

general, use of code specified methods to calculate the earthquake force is more

conservative than the calculations based on inelastic dynamic analysis methods. Code

requirements, however, prevent the earthquake force being reduced below 80 % of that



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determined using the codes formula (AS-1170.4, 1993) for fundamental period. The

calculation of equivalent earthquake force in Australian code is similar to the method used

in UBC (1997). Details of earthquake and wind load calculations are presented in

Appendix A.

The frame was found to have higher earthquake loading than wind loading. It was found

that ultimate gravity load combination governed the design of most members except

columns, for which the ultimate earthquake load combination governed. It should be noted

that the area of reinforcement required for the negative bending moments at supports were

calculated ignoring the area of slab reinforcement. The structure was also designed the

capacity design method to prevent the formation of column sideway mechanism (i.e. beam

hinges were designed to form before the formation of column hinges).

3.3 T est Specimen

The two-dimensional prototype frame analysis revealed that the first interior beam-column

joint in the first floor as shown in Figure 3-1, is more critical in terms of magnitude of the

out of balance moment applied to the joint. Therefore, it was selected to be tested in this

investigation.

3.3.1 Scale

In terms of test subassemblage dimensions, half scale testing was chosen, as this was the

maximum possible scale that could be tested in the reaction frame available, which was

used by Stehle (2002) and Abdouka (2003) for testing of wide band beam-column joints.

There was no adjustment of reinforcement bar spacing of the prototype and the test

specimen as per the Code (AS-3600, 2001) requirements. Table 3-3 shows the



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reinforcement detailing used in the test specimen. The nominal aggregate size was scaled

down from 20 to 10 mm and the cover provided to reinforcement in beam and column was

scaled down to half. The scaling factors for the respective actions and dimensions are

presented in Table 3-4. It should be noted that a 2:1 scale factor is required for material

density since gravitational acceleration cannot be scaled. This effect was tackled by

including extra weight in the applied dead loading.

T able 3-3: R einforcement details of beam and column (T est specimen)

T est subassembly

Rib beam Top reinforcement over the supportRib beam Bottom reinforcement

Rib beam flange slab reinforcement

Rib beam shear ligatures

2Y201Y20

F 72 (Mesh)

R6-200 Crs

Column main reinforcement

Column shear links

12Y16

R6-175 Crs

Table 3-4: Consistent Scaling relationship -After (Stehle, 2002)

DimensionScale factor

(Modal : Prototype)

Stress, pressure 1:1

Length, displacement 1:2

Area, bar area 1:4

Volume 1:8

Force, shear 1:4

Moment, Torsion 1:8

Density 2:1



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3.3.2 Specimen details

An isolated half-scale Corcon interior beam-slab-column subassembly, taken from the

prototype model frame structure is shown in Figure 3-3. The test specimen represents a

half scale model of prototype. It was terminated at column mid height and beam mid span,

representing approximate locations of points of contraflexure under lateral loading of the

prototype. The height of column is 1900 mm and the beam length is 4800 mm. The width

of the rib beam is 1200 mm, which is the typical rib beam width of Corcon slab system.

The reinforcement details of rib beam and column are shown in Figure 3-4. Shear ligatures

were provided to entire length of beam at uniform spacing, except in the beam-column

joint area. Similarly, column ligatures were provided to full height except within beam

depth. Figure 3-5 illustrates the reinforcement provided in the flange slab. It should be

noted that main top reinforcement has been curtailed 1000 mm from column centre. This

curtailment point was taken as per the critical bending moment envelope corresponding to

the gravity load case (i.e. 1.25G+1.5Q).



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1900 mm

1200 mm

4800 mm

447 mm

50 mm thick steel plates.

Figure 3-3: Dimensions of test sub-assemblage.

Figure 3-4: Beam and column cross-section of test subassembly

M ain Top R/FF72 M esh

Figure 3-5: T op view of flange slab wi th r einforcement.



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3.3.3 M aterial properties

The nominal strength and the measured strength of reinforcement bars used in the test

specimen are given in Table 3-5. Two types of high strength twisted/ribbed reinforcement

types were used, and defined in terms of strength. Those types are specified as: Y bars

and T bars. Y bars have nominal yield strength of 400 MPa and T bars have a

nominal strength of 500 MPa. These reinforcements were used in rib beam and column as

main bars. Other type of reinforcement, specified as R round bars, has a nominal strength

of 250 MPa, used in beam and column, as shear ligatures.

Nominal concrete compression cylinder strengths at 28 days of 40 MPa and 32 MPa were

used for the design of column and rib beam respectively. The measured cylinder

compressive strengths at the time of testing subassemblage are presented in Table 3-6.

T able 3-5: Reinforcement properties

Bar T ype Y 16 Y 20 T 7 R 6

Nominal bar diameter (mm)

Area (mm2)

Nominal yield stress (MPa)

Actual yield stress (MPa)

Actual ultimate stress (MPa)

16

201

400

442

542

20

314

400

448

539

6.75

35

500

510

684

6

28

250

345

502



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T able 3-6: Uniaxial compr essive str ength of concr ete.

M ember T arget 28 day compr essivestr ength (M Pa)

M easured str ength at thetime of testing (M Pa)1

Rib Beam 32 40.6

Column 40 47.4

3.4 T est configuration

3.4.1 Specimen loading

According to the load combinations given in AS 1170.0 (2002), the gravity loading to be

taken at ultimate limit state, in an event of earthquake is assumed to be 100% dead load

and 40% of live load on the structure (1.0G+0.4Q). The performance of the test specimen

under lateral loading has to be investigated with the scaled portion of gravity loading on

the model, so that conditions present during testing are same as prototype structure.

Since the purpose of the test is to investigate the performance of the specimen during an

earthquake event, only the earthquake load combination was simulated on the test

specimen during testing. Bending moments, shear forces and axial loads in beam and

column are shown in Figure 3-6 to Figure 3-8 for the prototype structure and from Figure

3-9 to Figure 3-11 for half scale test specimen. These diagrams were obtained from a 2-

dimensional analysis conducted using Program Space Gass. The results of the analysis

are presented in Appendix B. Axial force diagram for beam and bending moment and

shear force diagram for the column are not shown since very small values were obtained.

1Concrete strength measured following day of testing. Testing was done one month after casting of beam

and 45 days after casting of upper column.



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A similar set of diagrams was presented from Figure 3-12 to Figure 3-19 for the static

earthquake loading applied according to AS 1170.4 (1993).

-500

-400

-300

-200

-100

0

100

200

300

-4.8 -4 -3.2 -2.4 -1.6 -0.8 0 0.8 1.6 2.4 3.2 4 4.8

Distance along span (m)

BendingMoment(kNm)

Figure 3-6: Bending moment diagram for beams - full scale gravity loading

-300

-200

-100

0

100

200

300

-4.8 -4 -3.2 -2.4 -1.6 -0.8 0 0.8 1.6 2.4 3.2 4 4.8


Shearforce(kN)

Figure 3-7: Shear force diagram for beams - full scale gravity loading



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-2.1

-1.8

-1.5

-1.2

-0.9

-0.6

-0.3

00.3

0.6

0.9

1.2

1.5

0 500 1000 1500 2000 2500

Axial force (kN)

Distancealongcolumn

height(m)

Figure 3-8: Axial force diagram for columns full scale gravity loading

-100

-80

-60

-40

-20

0

20

40

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


BendingMoment(kNm)

Ms

Figure 3-9: Bending moment diagram for beams - half scale gravity loading



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-100

-80

-60

-40

-20

0

20

40

60

80

100

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


Shearforce(kN)

Figure 3-10: Shear force diagram for beams - half scale gravity loading

-1.05

-0.75

-0.45

-0.15

0.15

0.45

0.75

0 100 200 300 400 500 600

Axial force (kN)

Distancealongcolumnheight(m)

Figure 3-11: Axial force diagram for columns - half scale gravity loading



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-250

-200

-150

-100

-50

0

50

100

150

200

250

-4.8 -4 -3.2 -2.4 -1.6 -0.8 0 0.8 1.6 2.4 3.2 4 4.8


Bendingmoment

(kNm)

Figure 3-12: Bending moment diagram for beams - full scale earthquake loading

-2.1

-1.8

-1.5

-1.2

-0.9-0.6

-0.3

0

0.3

0.6

0.9

1.2

1.5

-300 -250 -200 -150 -100 -50 0 50 100 150 200 250 300

Bending moment (kNm)


Figure 3-13: Bending moment diagram for columns -full scale earthquake loading



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-75

-50

-25

0

25

50

75

-4.8 -4 -3.2 -2.4 -1.6 -0.8 0 0.8 1.6 2.4 3.2 4 4.8


Shearforce(kN)

Figure 3-14: Shear force diagram for beams full scale earthquake loading

-2.1

-1.8

-1.5

-1.2

-0.9

-0.6

-0.3

0

0.3

0.6

0.9

1.2

1.5

50 60 70 80 90 100 110 120 130 140 150

Shear force (kN)


Figure 3-15: Shear force diagram for columns -full scale earthquake loading



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-100

-80

-60

-40

-20

0

20

40

60

80

100

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


Bendingmoment

(kNm)

Figure 3-16: Bending moment diagram for beams - half scale earthquake loading

-1.05

-0.75

-0.45

-0.15

0.15

0.45

0.75

-100 -75 -50 -25 0 25 50 75 100

Bending moment (kNm)

Distancealon

gcolumnheight(m)

Figure 3-17: Bending moment diagram for columns - half scale earthquake loading



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-50

-40

-30

-20

-10

0

10

20

30

40

50

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


Shearforce(kN)

Figure 3-18: Shear force diagram for beams - half scale earthquake loading

-1.1

-0.8

-0.5

-0.2

0.1

0.4

0.7

0 10 20 30 40 50 60 70 80 90 100

Shear force (kN)


Figure 3-19: Shear force diagram for columns - half scale earthquake loading

The lateral loading setup is shown in Figure 3-20. This setup demonstrates the boundary

conditions and the loading arrangement to simulate the prototype structure. It can be seen

from bending moment and shear force diagrams for lateral earthquake loading (Figures 3-

12 to 3-19) that mid span bending moment is zero and the shear force is constant along the

span. These bending and shear force diagrams were achieved by using pin connections at



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ends of beams and column and providing lateral loading at bottom end of the column as

shown in Figure 3-20. In this setup, axial load developed in columns due to earthquake

loading is neglected, as it is small compared with gravity loading.

+ Loading- Loading

Figure 3-20: Setup for lateral loading

The dead and live load effects were approximately modelled by adjusting the beam-end

reactions as shown in Figure 3-21. In this arrangement two roller supports were allowed to

freely deflect under beam self-weight and a further down ward reaction was applied

through the each roller pin connection. The resulting bending moment and shear force

diagrams are shown in Figure 3-22 and Figure 3-23 respectively.

The axial loading on the column was simulated by using external prestressing. The

prestressing loading framework moved with the lateral displacement of the

subassemblage. Therefore, the applied axial load was able to be kept concentric regardless

of the lateral displacement of the specimen. The required axial load on the bottom column

was calculated as 490 kN. The adopted axial load diagram is shown in Figure 3-24. This



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force was achievable with four 12.7 mm diameter-prestressing strands. The prestress force

was transferred to the column via transfer beams at the top and bottom level of the

column.

Free vertical deflection allowed

under self weight of the beam

before connected to pin roller.

RR R=15.3 kN, reaction force is

applied through pin roller

connection link.

Applied lateral load

Applied axial load

Figure 3-21: A dopted setup for lateral and gravity loading

-100

-80

-60

-40

-20

0

20

40

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


BendingMoment(kNm)

Figure 3-22: Bending moment diagram for beams - Adopted half scale gravity loading



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-50

-40

-30

-20

-10

0

10

20

30

40

50

-2.4 -2 -1.6 -1.2 -0.8 -0.4 0 0.4 0.8 1.2 1.6 2 2.4


Shearforce(kN)

Figure 3-23: Shear force moment diagram for beams - Adopted half scale gravity loading

-1.05

-0.75

-0.45

-0.15

0.15

0.45

0.75

0 100 200 300 400 500 600

Axial force (kN)

Distancealon

gcolumnheight(m)

Figure 3-24: Axial force diagram for columns - Adopted half scale gravity loading

3.4.2 T est setup

The reaction frame has been designed to provide a pin connection to the top column end,

two roller supports which allow vertical deflection under the specimen self weight, a pin

connection at the bottom of the column to which the actuator was attached and a



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prestressing system to apply the axial load to the column. This setup was originally

designed by Stehle (2002) and Abdouka (2003) for wide band beam testing. It was

modified to suit the specimen tested in this project.

The top and side view of the test setup are shown in Figure 3-25 and Figure 3-26

respectively. The bracings have been provided at top level and on two side faces of the test

rig to provide the lateral stiffness. The pin connection at the top of the column was

provided by using a mild steel pin that was inserted into a hole in the test rig. Once pinned,

the specimen was able to hung freely at the centre of the test rig. At the base of the

column, another mild steel pin was used to connect the actuator arm. The pins at the beam-

ends have been created using vertical links, which hang down from the test rig.

T op-level br acings

Side bracings

Column exter nalpr estressing strands

C ompound cannel at columntop and bottom to tr ansfer axial for ce.

Figure 3-25: T op view of the built t est assembly



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Pin connections

Actuator

A ctuator hydraulic

controller

comput er based

data logging system

Fi gure 3-26: Side view of the built test assembly

All pin supports has been designed with 50 mm diameter mild steel pins through 52 mm

diameter holes. Hence, a pin slip at the connection was expected and the force-

displacement hysteresis results need to be adjusted. Other researchers using the same test

rig (Siah, 2001; Stehle, 2002) had encountered similar pin slip problems. The correction

procedure to take account of this slip is described later in chapter 4.

The two vertical link pin supports, one at each beam end were also supported by the

reaction frame. These vertical links allow free horizontal movement of the beam, but not

vertical movement. Beam end reactions were applied to simulate gravity loading on the

specimen so that the joint bending moment at the beginning of the test is consistent with

the prototype structure, as mentioned earlier in chapter 3.4.1. The reaction at the beam-

ends was provided by adjusting the length of the threaded bolt by screwing or unscrewing

it. A photo of the vertical link is shown in Figure 3-27. The amount of reaction applied

was measured by four strain gauges, which were installed around the circumference of the



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threaded rod that connects the top and bottom parts of the vertical link. The calibration of

these strain gauges for each vertical link was done prior to the testing of the specimen and

shown in Figure 3-28 and Figure 3-29.

A djusting Thr eaded r od

50mm pin connections

Figure 3-27: Photo of beam-end vertical link



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y = 0.2859x + 0.9694

R2

= 0.9999

-80

-60

-40

-20

0

20

40

60

80

-300 -200 -100 0 100 200 300

Average Microstrain

Force(kN)

Figure 3-28: Calibration of North vertical link

y = 0.2874x - 0.8555

R2 = 1

-80

-60

-40

-20

0

20

40

60

80

-300 -200 -100 0 100 200 300

Average Microstrain

Force(kN)

Figure 3-29: C alibr ation of South ver tical link



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3.4.3 Constr uction of test specimen

The construction details of test specimen were presented in Figure 3-3 to Figure 3-5. The

construction sequence of test specimen followed here is slightly different to the

conventional construction sequence. The followed construction sequence was mainly

determined by the specimen installation procedure to test rig. The steps of construction

were as follows:

Cut and bend column and beam reinforcement as required.

Fabrication of column cage with tie wires.

End bearings plates were welded to top and bottom end of column cage.

Fixing of strain gauges and associated wiring.

Construct formwork around column cage from top bearing plate to beam top level.

Horizontally casting top column and curing.

Remove column formwork and lift column into position in test rig and bolt to transfer

beams.

Place the reinforcement mesh for slab with temporary supports.

Erect Corcon beam and slab sheet metal formwork and formwork for the lower part of

the column.

Place pre-sized and strain gauged beam bars in position with shear links. Figure 3-30

shows the top view of beam ready for next stage concreting.

Concreting lower column and beam.

Curing.

Remove formwork.



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Upper column after

the completion of 1st

stage concreting

Figure 3-30: Specimen ready for concreting

3.5 I nstr umentation

Three types of instrumentations, i.e. Strain gauges, displacement transducers and load

cells, were used to monitor the behaviour of the specimen during the test. All the data

from the instruments were collected through a computer based data logging system.

3.5.1 Strain gauges

The strain gauges were attached to beam and column reinforcement. The strain gauges

used were Kyowa Type KFG-5-120-C1-11. The locations of strain gauges on beam and

column reinforcement are shown in Figure 3-31 and Figure 3-32 respectively.



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BTG1 BTG2

BTG3 BTG4

Beam Top view

BBG1 BBG1

Beam Side view

2400 mm 2400 mm

Fi gure 3-31: L ocation of str ain gauges on beam r einforcement

South

North

SWC1

SEC2NWC3

NEC4

N ote : Only column CornerR/F is shown for clarity.

Fi gure 3-32: L ocation of str ain gauges on beam r einforcement



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3.5.2 D isplacement tr ansducers

Displacement transducers were used to calculate the curvatures of the central portion of

the beam at the beam-column joint. The locations of these transducers are shown in Figure

3-33.

T1 T2

T3 T4

Fi gure 3-33: L ocations of displacement tr ansducers

3.5.3 L oad cells

Load cells were used to determine the load in each column-prestressing strand. Calibration

of the load cells was done before the setup, so that the correct load could be obtained. A

total of four load cells were used for strands. The inbuilt load cell in the actuator was used

to measure the lateral load applied at different displacements.



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3.6 T esting sequence

As described in section 3.4.2, the reactions to simulate the gravity loading had to be

applied before the testing commenced. The sequence of the loading of the specimen was

as follows:

Apply column axial load of 400 kN using four prestressing stands.

Release the vertical link connection at the end of beam and allow beams to freely

deflect under its self-weight.

Reconnect the beam-ends to vertical links, such that no tension or compression

developed in the vertical links.

Apply a downward force of 15.3kN by turning the threaded bar of the vertical links.

This was done by setting the vertical link strain gauge reading to an equivalent micro-

strain value as per the calibration graphs shown in Figure 3-28 and Figure 3-29.

Finally connect the actuator arm to the bottom end of the column.

At this stage the specimen was ready for the application of cyclic loading. The specimen

cyclic loading sequence was based on specimen drift rather than the ductility index

because of the difficulty in predefining a yield displacement in beam-column

subassemblage. The lateral loading sequence used for the test is shown in Figure 3-34. It

consisted of repeated cyclic loading of increasing drift ratio up to the maximum stroke of

the actuator, which corresponds to a nominal drift of 4%. The test results and observations

are presented in chapter 4.



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-5

-4

-3

-2

-1

01

2

3

4

5

0 1a 1b 2a 2b 3a 3b 4a 4b 5a 5b 6a 6b 7a 7b 8a 8b 9a 9b 10a

Drift Ratio (%)

Cycle No.

Figure 3-34: Lateral Cyclic loading sequence

3.7 2nd test specimen

3.7.1 General

The second test specimen was a retrofitted version of the damaged first specimen. It was

very clear from the test results and observations that test specimen was experiencing

reinforcement detailing problems. These observations are described in chapter 4. Having

identified the detailing problems in the first test, it was decided to repair the damaged

specimen using a Carbon Fibre Reinforced Polymer (CFRP) system, as a cost-effective

alternative. The repaired version of the specimen would then be re-tested to ascertain its

post-repair performance to loading taken to high drift limits and to observe any

improvements that may have resulted from the proposed repair process. Figure 3-35 shows

the adopted CFRP system.



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Mortar build-upBolted steel plates

2-layers of CFRP

from A-DAn additional

layers of CFRP

from B-C

2-layers of CFRP

from E-F on each

face

Figure 3-35: Details of CFR P system used for rectification



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3.7.2 Use of externally bonded FRP for structural repair work

For this repair work, externally bonded carbon fiber in fabric form (CFRP) was selected

after performing a parametric analytical study of the performance of two different types of

FRP systems. The relevant geometrical and mechanical properties of the material chosen,

provided by the supplier, are given in Table 3-7.

T able 3-7: G eometri cal and mechanical proper ties of fi bre

Fibre Type Carbon fibre Glass fibre

Reference CF130 EG-90/10A

W idth/T hickness 300 mm/0.176 mm 670 mm/0.154 mm

E-modulus 240,000 MPa 73,000 MPa

Ultimate tensile strain 1.55 % 4.5 %

T ensile str ength 3800 MPa 3400 MPa

Design tensile force 211 kN/m @ 0.6% strain/m width 264 kN/m @ ult. strain/m width

The moment curvature relation for the (rib beam section at column face) strengthening

system shown in Figure 3-35 was obtained from sectional analysis program Response-

2000 (Bentz and P.Collins, 2000). The positive and negative moment-curvature

relationship for Carbon, Glass fibre and the reinforcement used in the first test specimen

are shown in Figures 3-36 and 3-37 respectively.



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0

25

50

75

100

125

150

175

200

225

0 25 50 75 100 125

Cu rvature (Rad/km )

M

oment(kNm)

+ve Moment-R/F

+ve Moment-

CF130(200 mm2)

+ve Moment-

Eglass(400 mm2)

Figure 3-36: Positive moment curvature wit h differ ent reinfor cing material s

0

25

50

75

100

125

150

175

200

225

0 25 50 75

C urvature (Rad/km)

M

oment(kNm)

-ve Moment -R/F

-Ve Moment-

CF130(450 mm2)-ve Moment-

Eglass(900 mm2)

Figure 3-37: N egative moment curvature with dif ferent r einfor cing materi als



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The use of FRP as a means of flexural strengthening will compromise the ductility of the

original reinforcement system. Significant increases in moment capacity with FRP sheets

are afforded at the sake of ductility. However, sections that experience a significant loss in

ductility must be considered. The approach taken by (MBrace, 2002) follows the

philosophy Appendix B of ACI 318, where a section with low ductility must compensate

with a higher strength reserve. The higher reserve of strength is achieved by applying a

strength reduction factor of 0.70 to brittle sections as opposed to 0.90 for ductile sections.

Both concrete crushing and FRP rupture before yielding of the steel are considered as both

brittle failure modes. Steel yielding followed by concrete crushing provides some level of

ductility depending on how far the steel is strained over the yield strain. Steel yielding

followed by FRP rupture is typically ductile because the level of strain needed to rupture

FRP is significantly higher than the strain level needed to yield the steel.

The moment-curvature behaviour of Carbon and Glass fibre was compared with

reinforcement, as shown in Figure 3-36 and Figure 3-37, the Carbon fibre behaviour was

closely matching with reinforcement than that of Glass fibre. It should also be noted that

the area of carbon fibre required to give same moment capacity was approximately half

compared to the glass fibre. In addition the high strength, high modulus and negligible

creep rupture behaviour make carbon fibres ideal for flexural strengthening applications.

Therefore, the carbon fibre was selected for this repair work, as the overall cost of repair

would be drastically reduced due to the less material and labour involvement. Generally,

major portion of total cost is the labour cost, as the preparation and application of FRP

requires highly trained skilled workers.



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3.7.3 Structural repair work

Since the composite element of an FRP repair system is required to be bonded onto the

concrete substrate, the efficiency of the system depends on the integrity of this bond at the

interface layer. A minimum tensile strength of 1.5 MPa is recommended for the substrate

for this type of CFRP design (MBrace, 2002). Any loose material in areas where the CFRP

system was to be applied was removed and patched with suitable mortars. All cracks

greater than 0.3 mm in width were repaired by epoxy injection. Figure 3-38 illustrates the

prepared specimen for epoxy injection. Wide cracks, similar to those shown in Figure

3-39, were repaired adopting pour techniques using low shrinkage structural grout.

Fi gure 3-38: Specimen before epoxy inj ection



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Figure 3-39: Fill ing of w ide cracks with low shr ink age str uctural grout.

It was observed that an uneven surface was created near the large cracks, after repair, due

to geometric deformation. Care was taken to flatten these areas using suitable mortars.

Alternatively grinding to flatten out the surface can also prevent possible CFRP peel-off

failure due to unevenness of the substrate.

Other important considerations when applying a CFRP repair system to a damaged

structure, concerns detailing. Material-specific use restrictions for CFRP necessitate

avoidance of sharp corners in structural elements to which they may be applied. Such

corners need to be rounded to decrease the chance of fibre fracture due to stress

concentrations induced by sharp edges. In this specimen, the column and beam dimensions

are different, therefore in order to provide smooth transition for CFRP bottom layer over

the column width, an additional mortar build-up was created near the beam column joint.

In addition to the treatment near the column-beam joint described above, four mild steel

plates were placed over the FRP layer and bolted to the rib beam. The steel plates, as



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shown in Figure 3-35 provided restraining (clamping) forces to prevent delamination of

the CFRP due to diverting forces created in the fibres from the mortar build-up.

Figure 3-40: M ortar build-up near the beam column joint

3.7.3.1 Surface preparat ion for FR P application

After completing all the structural repair work, the concrete surface needs to be prepared

to receive the FRP application. The concrete surface should be clean, sound and free of

surface moisture, any foreign matter such as dust, laitance, grease, curing compounds and

other bond inhibiting materials from the surface by blast cleaning or equivalent

mechanical means. The surface preparation was done using mechanical wire brush instead

of sand blasting.The general requirement is that the surface must present similar to 60-grit

sandpaper. Figure 3-41 shows the prepared surface at the bottom of the rib beam adjacent

to the beam column joint.



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Prepared concrete

surface by removing

excess surface grout.

Figure 3-41: Prepared concrete surface to receive FRP application

3.7.3.2 CFR P application to prepared sur face

The first step in the FRP application process was the priming of the concrete surface with

the penetrating primer prior to the application of any subsequent coatings applied using a

roller. The primer was applied uniformly in sufficient quantity to fully penetrate the

concrete and produce a non-porous film in the surface approximately 100-150 microns in

thickness after full penetration. It must be noted that the volume to be applied may vary

depending on the porosity and roughness of the concrete surface.

The next step was to apply an epoxy resin on the primed surface and lay the FRP sheet.

According to the manufactures guidelines, the resin has to be applied to the primed surface

using a medium nap roller (approx. 10 mm) to approximately 500 - 750 microns wet film

thickness (1.3-2 m2

per litre) or sufficient to achieve a wet-out of the FRP Fabric Sheet.



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This value will vary depending on the weight of the FRP Fabric Sheet as well as the

ambient conditions and wastage. The mixed batch resin has to be used before expiration of

its batch-life, as increased resin viscosity will prevent proper impregnation of the FRP

fabric materials. Figure 3-42 shows the application of epoxy resin over the primer layer.

Epoxy Resin application

Pr evious Pr imer layer

Figure 3-42: A pplication of E poxy r esin

FRP Fabric Sheets must be cut beforehand into required length using appropriate scissors.

The FRP Fabric Sheet was placed with the fibre side placed on the concrete surface and

work in the direction of the fibres and work from the centre of the length of the sheet to

the ends, to remove any entrapped air. The other subsequent layers of FRP were laid

similar to the first layer. Figure 3-43 shows the application of the first CFRP layer. It

should be noted that the application of resin has to be done before and after the laying of

each new FRP fabric layer. A hard roller was used to enhance the impregnation of the

fabric material. The backing polythene paper was then peeled away. The surface of

adhered fabric was squeezed in the longitudinal direction of the fibre using a ribbed roller

in order to impregnate resin into the fabric material and remove any air bubbles (see

Figure 3-44). Figure 3-45 depicts the completely repaired test specimen.



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Figure 3-43: L aying CFR P on the Epoxy applied surface

Figure 3-44: A r ibbed roller used to impregnate r esin into the fabr ic material



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Figure 3-45: CFRP repaired test specimen ready for test instrumentation

3.8 I nstr umentation for second test specimen

The instrumentation used for the second test was same as the first test. Additional strain

gauges were installed on the CFRP. The locations of strain gauges on beam top flange and

beam rib are shown in Figure 3-46. All the data from the instruments were collected

through a computer based data logging system similar to the first test.



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N S

N S

CG1 CG2 CG3 CG4 CG5 CG6 CG7 CG8

CG13 CG14 CG15 CG16

CG9 CG10 CG11 CG12

CG17 CG18 CG19 CG20Beam west side strain gaugenumberingBeam east side strain gauge

numbering

Figure 3-46: Location of strain gauges on CFRP

3.8.1 Photogrammetr y-based measur ement

A photogrammetry-based measurement setup was used to follow the deformations in the

repaired specimen during the test. Approximately 200 highly reflective photosensitive

targets were introduced on one side of the concrete beam as well as on the CFRP surfaces.

Figure 3-47 shows the test specimen with photosensitive target points. Using a purpose-

specific camera, three-dimensional digital measurements were determined from these

target locations from a series of multiple photographs taken at different stages within the

loading cycles. These measurements enabled both global and local deformation of the test

specimen to be followed during the testing.



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Fi gure 3-47: t est specimen wi th photosensitive tar get points



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Abdouka, K. (2003). The Seismic Performance of Reinforced Concrete Wide Band Beam

Frames: Exterior Connections. Melbourne, The University of Melbourne, Australia.

ACI-318 (1999). Building Code Requirements for Structural Concrete (ACI-318M-99)

and Commentary (ACI-318RM-99), American Concrete Institute, Farmington Hills, Mich.

AS-1170.0 (2002). Structural design actions, Part 0: General principals, Standard

Association of Australia, Sydney, Australia.

AS-1170.1 (2002). Structural design actions, Part 1: Permanent, Imposed and other

actions, Standard Association of Australia, Sydney, Australia.AS-1170.2 (2002). Structural design actions, Part 2: Wind loads, Standard Association of

Australia, Sydney, Australia.

AS-1170.4 (1993). Minimum Design Loads on Structures, Part 4: Earthquake loads,

Standard Association of Australia, Sydney, Australia.

AS-3600 (2001). Concrete Structures, Standard Association of Australia, Sydney,

Australia.

Bentz, E. C. and M. P.Collins (2000). RESPONSE-2000 Reinforced Concrete Sectional

Analysis Program Manual.BS-8110 (1995). Structural use of concrete, Part 1: Code of practice for design and

construction, British Standard Institute, London.MBrace (2002). Composite strengthening system,Engineering design guidelines, New

York., Watson Bowman Acme Corp.

Siah, W. L. (2001). Seismic Performance of Interior Post-Tensioned Concrete Wide

beams, The University of Melbourne, Australia.

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report on rcc

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