ISSN 0453-4972
BRI Research Paper No. 147

SUMMARY OF RESEARCH ON
CONCRETE-FILLED STRUCTURAL STEEL TUBE COLUMN SYSTEM CARRIED OUT UNDER
THE US-JAPAN COOPERATIVE RESEARCH PROGRAM ON
COMPOSITE AND HYBRID STRUCTURES

byIsao Nishiyama, Shosuke Morino, Kenji Sakino, Hiroyuki Nakahara, Toshiaki Fujimoto, Akiyoshi Mukai,
Eiichi Inai, Makoto Kai, Hiroyoshi Tokinoya, Toshiyuki Fukumoto, Koji Mori, Kenzo Yoshioka,
Osamu Mori, Kenji Yonezawa, Mizuaki Uchikoshi and Yukio Hayashi

Building Research InstituteJAPAN

FORWARD

From the fiscal year of 1993, a five-year research project on concrete-filled structural steel tube (CFT) column
system was carried out as part of the fifth phase of the US-Japan Cooperative Earthquake Research Program on
Composite and Hybrid Structures. At the beginning of the project, a number of topics related to research issues
were raised and prioritized by cooperation work of the researchers and practitioners involved in the project.

Inthose days, a structural system having concrete-filled structural steel tube used for columns had just come to be
used in real buildings in Japan.

Thus, the CFT column system was selectively studied on the Japanese side,

since expansion of the application scope beyond previous work would be very useful for practitioners.

Namely,the studies were performed in case of larger width (diameter)-to-thickness ratio of steel tubes and higher
concrete / structural steel strengths. This paper presents a summary of the research completed on the Japaneseside. I believe that the results are of great use and importance for practitioners.

Other research projects on 1) New materials, elements and systems, 2) Reinforced concrete (RC) column / steel
beam system, and 3) RC / SRC (steel reinforced concrete) wall system were also carried out in parallel to theresearch project on CFT column system under the same phase of the US-Japan Cooperative Research Program.
Summaries on those research projects will be presented in different opportunity.

Lists of the individual

technical papers are included in Appendix for the advanced use by the researchers and practitioners.

January 2002
Hiroyuki YAMANOUCHI
Chief Executive
Building Research InstituteTABLES OF CONTENTS

ABSTRACT ……….………………………………………………………………………………………

i

LIST OF TABLES...………………………………………………………………………………………

ii

LIST OF FIGURES ………………………………………………………………………………………

iv

CHAPTER
1. INTRODUCTION AND OVERVIEW OF RESEARCH……...……………………………

1

2. BEHAVIOR OF CENTRALLY LOADED SHORT COLUMNS………..…………………

8

3. BEHAVIOR OF ECCENTRICALLY LOADED SHORT COLUMNS.……………………

16

4. BEHAVIOR OF BEAM-COLUMNS.………………………………………………………

25

5. BEHAVIOR OF BEAM-TO-COLUMN SUBASSEMBLADGES…………………………

32

6. PROPOSED DESIGN FORMULAS FOR COLUMNS AND BEAM-COLUMNS..………

43

7. TRIAL DESIGN AND MERITS OF CFT COLUMN SYSTEM...…………………………

51

8. SUMMARIES AND FUTURE RESEARCH NEEDS…...…………………………………

57

ACKNOWLEDGEMENTS………………………………………………………………………………

62

REFERENCES……………………………………………………………………………………………

63

TABLES …………………………………………………………………………………………………..

66

FIGURES………………………………………………………………………………………………….

92

APPENDIX….…………………………………………………………………………………………….

168

SUMMARY OF RESEARCH ON
CONCRETE-FILLED STRUCTURAL STEEL TUBE COLUMN SYSTEM CARRIED OUT UNDER
THE US-JAPAN COOPERATIVE RESEARCH PROGRAM ON
COMPOSITE AND HYBRID STRUCTURES

Editors: Isao NISHIYAMA *1 and Shosuke MORINO *2

ABSTRACTThe US-Japan Cooperative Earthquake Research Program began in 1979 following the recommendations as
outlined in the final report of the US-Japan Planning Group for the program [“Recommendations for a US-Japan
Cooperative Research Program utilizing Large-Scale Testing Facilities”, Report No. UCB/EERC 79-26,
September 1979]. First four phases of the program have been implemented on 1) Reinforce Concrete, 2)
Structural Steel, 3) Masonry, and 4) Precast / Prestressed Concrete. Phase 5 of the US-Japan Cooperative
Research Program on Composite and Hybrid Structures was carried out from the fiscal year of 1993 as afive-year research program. In this phase of research program, the following four research topics were executed:1) New materials, elements and systems, 2) Concrete-filled structural steel tube column system, 3) Reinforced
concrete column and structural steel beam system, and 4) Reinforced concrete and steel reinforced concrete wall
system.

In this paper, the technical research results conducted in the Japanese side related to the research topic of theconcrete-filled structural steel system is overviewed and summarized. It includes experimental results (columns,
beam-columns and beam-to-column assemblies), proposed constitutive model for steel and concrete consideringsynergistic action, analytical results, and design implication study. Draft of the manuscripts was shared among
the researchers directly involved in the research project, and it was summarized in this research paper by theeditors.

As for the research results on the other three topics, summary paper has not yet been completed. The available
technical papers, which cover a part of the research topics, are listed in the appendix for the reference to thereaders.

________________________________________________________________________________*1 Advanced Research Engineer, Department of Construction Engineering, Building Research Institute*2 Professor, Dept. of Architecture, Faculty of Engineering, Mie University-i-LIST OF TABLES

Table 1.1

High Priority Research Topics Recommended in the Joint Planning Workshop [1.5]66Table 1.2

Medium Priority Research Topics Recommended in the Joint Planning Workshop

67

[1.5]
Table 1.3

Test Program of Centrally and Eccentrically Loaded Stub Columns

68

Table 1.4

Test Program of CFT Beam-Columns

69

Table 1.5

Test Program of Beam-to-Column Connections

70

Table 2.1

Results of Axial Compressive Test of Steel Tubes Used for CFT Specimens

71

Table 2.2(a) Experimental Variables and Test Results (Circular CFT)

72(b) Experimental Variables and Test Results (Square CFT)73

Table 2.3

Specific Values for the Stress vs. Strain Modes for Concrete

74

Table 2.4

Specific Values for the Stress vs. Strain Models for Steel Tube in Square CFT

75

Columns
Table 3.1(a) Summary of Eccentrically Loaded Stub Column Specimens (Circular Tubes)

76(b) Summary of Eccentrically Loaded Stub Column Specimens (Square Tubes)

77(a) Material Properties of Structural Steel Tubes (Circular Tubes)

78(b) Material Properties of Structural Steel Tubes (Square Tubes)

78(a) Material Properties of Concrete (Circular Tubes)

78(b) Material Properties of Concrete (Square Tubes)

78(a) Experimental and Analytical Results (Circular Tubes)

79(b) Experimental and Analytical Results (Square Tubes)

80

Table 4.1

Summary of Beam-Column Specimens

81

Table 4.2

Material Properties of Structural Steel Tubes

82

Table 4.3

Material Properties of Concrete

82

Table 4.4

Summary of Test Results

83

Table 5.1

Summary of Beam-to-Column Connection Specimens

84

Table 5.2

Material Properties of Panel Zone Steel and Filled Concrete

84

Table 5.3(a) Test Results of Rectangular Tube Specimens

85(b) Test Results of Circular Tube Specimens

85

Table 5.4

Virtual Specimens for Numerical Simulation

86

Table 5.5

Results of Numerical Simulation

86

Table 6.1

Characteristic Parameters of Concrete Stress Block for Circular CFT Section

87

Table 6.2

Characteristic Parameters of Concrete Stress Block for Square CFT Section

87

Table 6.3

Relationships Between Ductility Grades and Limit Rotation Angles of Beam-Column s

87

Table 6.4

Summary of Database on Stiffness Degrading Ratio ay

88

Table 7.1

Investigation Items and Material Properties of Each Theme Structure

89

Table 7.2

Design Loads

89

Table 7.3

Gravity and Seismic Loads of 40-Story Frames

89

Table 7.4

List of Members for 40-Story Frames

89

Table 3.2

Table 3.3

Table 3.4

- ii -

Table 7.5

Input Seismic Ground Motions for 24-Story Frames

90

Table 7.6

Hysteresis Models for 24-Story Frames

90

Table 7.7

Weight and Stiffness of 40-Story Frames

90

Table 7.8

Story Drift of 40-Story Frames

90

Table 7.9

First Natural Period

91

Table 7.10

Absorbed Energy

91

Table 7.11

Results of Response Analyses

91

Table 7.12

Cost Estimation of CFT and Steel Frames

91- iii -

LIST OF FIGURES

Figure 1.1

Classification of Composite and Hybrid Structures

92

Figure 1.2

Typical CFT Column System and Concrete Filling

93

Figure 1.3

Overall Research Operation System of the US-Japan Cooperative Research Program94

on Composite and Hybrid Structures
Figure 1.4

Element Tests of CFT Column System

95

Figure 1.5

Key Plan of CFT Theme Structure

96

Figure 2.1

Test Setup Used for Centrally Loaded Stub Columns

97

Figure 2.2

Sizes of Specimens

98

Figure 2.3

Scale Effect on Compressive Strengths of Circular Plain Concrete Columns

98

Figure 2.4

Stress States of Steel Tube and Concrete

99

Figure 2.5

Experimental Axial Compressive Strengths of Circular CFT Columns

99

Figure 2.6

Comparisons Between Experimental and Calculated Strengths (Circular CFT

100

Columns)
Figure 2.7

Compressive Strength Factors of Hollow Square Steel Tubes

100

Figure 2.8

Comparisons Between Experimental and Calculated Strengths (Square CFT

101

Column s)
Figure 2.9

Stress vs. Strain Models for Concrete in CFT Column

101

Figure 2.10

Stress vs. Strain Models for Steel Tube in Circular CFT Column

102

Figure 2.11

Stress vs. Strain Models for Steel Tube in Square CFT Column

102

Figure 2.12

Comparisons Between Experimental Results and Proposed Models

103

Figure 2.13

Comparisons Between Experimental Results and Proposed Models

104

Figure 3.1

Dimensions of Eccentrically Loaded Stub Columns

105

Figure 3.2

Loading Condition of Eccentrically Loaded Stub Columns

105

Figure 3.3(a) Moment vs. Curvature Relations (Circular Specimens)

106(b) Moment vs. Curvature Relations (Square Specimens)

107

Figure 3.4

Stress vs. Strain Relations of Filled Concrete

108

Figure 3.5

Stress vs. Strain Relations of Steel Tube

109

Figure 3.6(a) Comparisons Between Experimental and Analytical Results (Circular Tubes)

110(b) Comparisons Between Experimental and Analytical Results (Square Tubes)

111

Figure 3.7

Assumed Stress Blocks

112

Figure 3.8

Values of

M u / M cal.1

112

Figure 3.9

Values of

M u / M cal.2

112

Figure 3.10

Values of

M u / M cal.3

112

Figure 4.1

Details of Beam-Column Specimens

113

Figure 4.2

Test Setup and Loading Condition for Beam-Column Specimens

114

Figure 4.3

Loading History for Beam-Column Specimens

114- iv -

Figure 4.4

Rules for Axial Force Loading in Beam-Column Tests

115

Figure 4.5

Test Results of Circular Interior Beam-Column Specimens

116

Figure 4.6

Test Results of Circular Exterior Beam-Column Specimens

117

Figure 4.7

Test Results of Square Interior Beam-Column Specimens

118

Figure 4.8

Test Results of Square Interior Beam-Column Specimens Subjected to Biaxial

119

Bending
Figure 4.9

Test Results of Square Exterior Beam-Column Specimens

119

Figure 4.10

Effect of Material Strength on Ductility

120

Figure 4.11

Effects of Material Strength on Enhancement in Flexural Strength

120

Figure 4.12

Effects of Loading Direction on Flexural Strength

120

Figure 4.13

Model of Beam-Column, and Assumed Curvature and Axial Strain Distributions

121

Figure 4.14

Analytical Fiber Elements at the Critical Sections of Square and Circular CFT

122

Figure 4.15

Typical Stress vs. Strain Relationships of Concrete and Steel Tube

122

Figure 4.16

Analytical Results of Circular Interior Beam-Column Specimens

123

Figure 4.17

Analytical Results of Circular Exterior Beam-Column Specimens

124

Figure 4.18

Analytical Results of Square Interior Beam-Column Specimens

125

Figure 4.19

Analytical Results of Square Interior Beam-Column Specimens Subjected to Biaxial

126

Bending
Figure 4.20

Analytical Results of Square Exterior Beam-Column Specimens

126

Figure 5.1

Shapes of Specimens

127

Figure 5.2

Details of Panel Zone

128

Figure 5.3

Test Setup for Specimens R1 and R2

129

Figure 5.4

Test Setup for Specimen R6

129

Figure 5.5

Loading History for Beam-to-Column Specimens

130

Figure 5.6

Rules for Axial Force Loading for Beam-to-Column Specimens

130

Figure 5.7

Instruments for Measurements of Story Drift and Displacement of Panel Zone

131

Figure 5.8

Crack Patterns of Filled Concrete in Panel Zone

132

Figure 5.9(a) Story Shear vs. Story Drift Angle Relations of Rectangular Specimens

133(b) Story Shear vs. Story Drift Angle Relations of Circular Specimens

134(a) Hysteresis Curves of Panel Zone, Column and Beam of Rectangular Specimens

135(b) Hysteresis Curves of Panel Zone, Column and Beam of Circular Specimens

137(a) Contribution of Deformation of Panel Zone, Beam and Colu mn to Overall Story

138

Figure 5.10

Figure 5.11

Drift in Rectangular Specimens(b) Contribution of Deformation of Panel Zone, Beam and Column to Overall Story

139

Drift in Circular Specimens
Figure 5.12

Story Shear vs. Story Drift Angle Relation of Three Dimensional Specimen R6

140

Figure 5.13

Hysteresis Curves of Panel Zone, Column and Beam of Three Dimensional

140

Specimen R6
Figure 5.14

Contribution of Deformation of Panel Zone, Beam and Column to Overall Story
Drift in Rectangular Specimens

-v-

141

Figure 5.15

Normalized Shear Force of Panel Zone vs. Normalized Story Drift Angle Relations

141

Figure 5.16

Yield and Ultimate Shear Strengths of Panel Zone of Two Dimensional Specimens

142

Figure 5.17

Yield and Ultimate Shear Strengths of Panel Zone of Three Dimensional Specimen

142

Figure 5.18

Assumed Tri-Linear Skeleton Curves for Steel and Filled Concrete in Panel Zone

143

Figure 5.19

Arch Mechanisms in Panel Zone

143

Figure 5.20

Comparisons of Calculated and Experimental Skeleton Curves

144

Figure 5.21

Analytical Meshes for Finite Element Method

145

Figure 5.22

Stress vs. Strain Relationships of Concrete

145

Figure 5.23

Story Shear Force vs. Story Drift Angle Relationships of Beam-to-Column

146

Connection Specimens
Figure 5.24p

Q c / Ac / f c '−γ Relationships

147

Figure 5.25p

Q s − γ Relationships

147

Figure 5.26

Contour and Flow of Principal Compressive Concrete Stress in Panel Zones

Figure 5.27p

Q c / Ac / f c '− H / D c Relationships

148

Figure 5.28'p

Q c / Ac / f c '−H / D Relationships

148

Figure 6.1

Biaxial Stress State of Steel Tube under Yielding

149

Figure 6.2

Comparisons Between Calculated Ultimate Axial Loads and Experimental Ones of

150

148

Circular Columns
Figure 6.3

Local Buckling Modes of Square Steel Tube

151

Figure 6.4

Comparisons Between Calculated Ult imate Axial Loads and Experimental Ones of152

Square Columns
Figure 6.5

Stress vs. Strain Curves for Plain and Confined Concrete

153

Figure 6.6

Stress Blocks for Concrete and Steel in Circular CFT Section (Prescribed in the153

AIJ-CFT Recommendations [6.1])
Figure 6.7

Stress Blocks for Concrete in Circular CFT Section (Proposed by Sun and Sakino

153[6.9])
Figure 6.8

Comparisons Between Calculated Ultimate Flexural Strengths and Experimental

154

Ones of Circular Columns
Figure 6.9

Stress Blocks for Concrete and Steel in Square CFT Section (Prescribed in the AIJ

155–CFT Recommendations [6.1])
Figure 6.10

Proposed Stress Blocks for Concrete and Steel in Square CFT Section with High155

Strength Concrete and / or Thin Steel Tube Wall
Figure 6.11

Comparisons Between Calculated Ultimate Flexural Strengths and Experimental

156

Ones of Square Columns
Figure 6.12

Beam-Column under Combined Forces

157

Figure 6.13

Definition of Limit Rotation Angle

157

Figure 6.14

Comparisons Between Calculated Ultimate Flexural Strengths and Experimental

158

Ones

- vi -

Figure 6.15

Comparisons Between Calculated Ultimate Flexural Strengths and Experimental

159

Ones Attained within Rotation Angle Limitation of 0.01 radian
Figure 6.16

Comparisons Between Predicted Limit Rotation and Experimental One (Circular

160

Columns)
Figure 6.17

Comparisons Between Predicted Limit Rotation and Experimental One (Square

160

Columns)
Figure 6.18

Tri-linear Skeleton Model

161

Figure 6.19

Definition of Experimental ay

161

Figure 6.20

Effect of N/No on ay (Circular CFT)

161

Figure 6.21

Effect of N/No on ay (Square CFT)

161

Figure 6.22

Comparisons Between Tri-linear Skeleton Model and Experimental Results

161

Figure 6.23

Comparisons Between Normal Tri-linear Hysteretic Model and Experimental Results

162

Figure 7.1

Elevations of Theme Structures

163

Figure 7.2

Floor Plan of Theme Structures

163

Figure 7.3

Story Displacements Caused by Beam and Column Deformations

164

Figure 7.4

Vibrational Mode Shapes

164

Figure 7.5

Maximum Responses of Shear Coefficient (El Centro NS: 25cm/sec)

165

Figure 7.6

Q − δ Relations of 40-Story Frames

165

Figure 7.7

Maximum Responses of Shear Coefficients, Overturning Moments and Story Drifts

166

Figure 7.8

Maximum Responses of Ductility Factors of Beams and Columns (HACHINOHE:

166

50cm/sec)
Figure 7.9

Formation of Plastic Hinges in Beams and Columns

167

Figure 7.10

Comparison of Steel Amount Between CFT and Structural Steel Flames167- vii -

CHAPTER 1: INTRODUCTION AND OVERVIEW OF RESEARCH

1.1 Introduction
It is widely recognized that innovative uses of two or more different materials in a structure generally lead to amore efficient system for resisting seismic force. Use of such structural system has increased in the US and inJapan during the past ten years. Despite the amount of research and development work, especially by Japaneseconstruction companies, not enough was known regarding their seismic behavior or performance. Design
procedures and codes for their usage in typical design offices were non-existent. Therefore, a five-year research
program on such structural systems was initiated as the fifth phase of the US-Japan Cooperative Research
Program, which originally started in 1979 following the recommendations outlined in the final report of the
US-Japan Planning Group for the program [1.1]. Because of the diverse and broad scope of the subject area, it
was recommended that the research program in the fifth phase should be organized into the following fourgroups: 1) New materials, elements and systems, 2) Concrete-filled structural steel tube (CFT) column system,
3) Reinforced concrete (RC) column / steel beam system, and 4) RC / SRC (steel reinforced concrete) wall
system.

The project aimed at developing design guidelines (for a unified code development) through cooperative studies
to determine the relationship among full-scale tests, small-scale tests, component tests, and related analyticaland design implication studies. Because the CFT column system was becoming popular in Japan at the
beginning of the project, research topics, thought to be useful in expanding the practical application scope, wereselectively studied. These include research on larger width (diameter)-to-thickness ratio of structural steel tube,
higher strength concrete than normal concrete and higher strength steel than normal structural steel. In this paper,summary of the research done in Japan on CFT column system is presented.

Sharing of the manuscript of each chapter is as follows.

Chapter 1: I. Nishiyama*1, S. Morino*2
Chapter 2: K. Sakino*3, H. Nakahara*4
Chapter 3: T. Fujimoto*5, A. Mukai*1, I. Nishiyama, K. Sakino
Chapter 4: E. Inai*6, A. Mukai, M. Kai*7, H. Tokinoya*8, T. Fukumoto*9, K. Mori*10
Chapter 5: K. Yoshioka*11, T. Fujimoto, M. Kai, K. Mori, O. Mori*12, E. Inai, A. Mukai, K. Yonezawa*8,
T. Fukumoto
Chapter 6: K. Sakino, E. Inai
Chapter 7: M. Uchikoshi*13, Y. Hayashi*14, S. Morino*1: Advanced Research Engineer, Building Research Institute*2: Professor, Department of Architecture, Faculty of Engineering, Mie University*3: Professor, Faculty of Human-Environmental Studies, Kyusyu University*4: Lecturer, Department of Architecture, Faculty of Engineering, Kagoshima University*5: Research Engineer, Research Center, ANDO Corporation-1-*6: Associate Professor, Faculty of Engineering, Yamaguchi University*7: Research Engineer, Technical Research Institute, Matsumura Corporation*8: Research Engineer, Technical Research Institute, Obayashi Corporation*9: Senior Research Engineer, Technical Research Institute, Kajima Corporation*10: Research Engineer, Technical Research Institute, Asanuma Corporation*11: Professor, Department of Architectural Engineering, Dai-Ichi University*12: Research Engineer, Technical Research Institute, Toyo Corporation*13: Senior Engineer, Structural Engineering Department, KUME SEKKEI Co. Ltd.
*14: General Manager, Structural Engineering Department, Kajima Corporation

Members of the Technical Sub-Committee on CFT column system (refer to the overall research organization insection 1.5) are listed below. Affiliations are those at the time of the project.

Chairman:

S. Morino (Mie University)

Members:

K. Sakino (Kyusyu University)
K. Yoshioka (dispatched from Building Contractors Society)
T. Fukumoto (dispatched from Building Contractors Society)
A. Tomita (dispatched from Building Contractors Society)
E. Inai (dispatched from Building Contractors Society)
T. Fujimoto (dispatched from Building Contractors Society)
S. Gokan (dispatched from Japan Structural Consultants Association)
Y. Hayashi (dispatched from Japan Structural Consultants Association)
M. Uchikoshi (dispatched from Japan Structural Consultants Association)
F. Ohsugi (dispatched from Japan Structural Consultants Association)
T. Nakamura (dispatched from Japan Structural Consultants Association)
H. Sawada (dispatched from Japan Structural Consultants Association)
I. Nishiyama (Building Research Institute)
A. Mukai (Building Research Institute)

1.2 Definition of Composite and Hybrid Structures and Concrete-Filled Structural Steel Tube Column
SystemRecently in the field of building construction in Japan, it has become the goal to develop new types of structural
systems combining existing and / or new materials in better coordination to add structural and architecturaladvantages, which were difficult to be attained by traditional structural systems. These types of new structural
systems are called “Composite and Hybrid Structures (CHS)”. Among possible variations of CHS, the mostfrequently investigated and practically feasible one is the combination of RC and structural steel.

The advantages of CHS include attaining better “structural functions” and “productability and constructability”,
which are very difficult to be achieved by the existing structural systems. For example, largely flexible space-2-

from longer span beams, reduction of story and / or building height because of shallower floor systems, andimprovement of habitability by an increase in lateral stiffness are some advantages in structural functions. The
advantages in productability and constructability include the freedom of structural planning, and the shorteningof construction periods and quality improvements due to the shift to prefabrication.

From the viewpoint of structural technology, the control of failure modes of structures in seismic design and therational uses of structural members in a structure are easily achieved with CHS systems. In short, the structural
system can be made clearer in the mechanical viewpoint.

The CHS system can be classified into three categories according to their geometry as shown schematically in
Figure 1.1. These three categories correspond to 1-D, 2-D and 3-D CHS systems. The first category is thecomposite and hybrid member. Single members such as beams and columns are composed of different materials.
CFT, fiber reinforced concrete and SRC members are typical examples of this category. The second category is
the composite and hybrid frame. Even if the individual members such as the beam or column are not composite
members, the combination of different types of single members makes a composite and hybrid structure. The
RC column and steel beam system is an example in this category. The third category is a structure made ofdifferent materials and members, combined three-dimensionally. A typical example of this category is the RC
core wall with an exterior steel frame system, or a RC frame in the lower stories with a steel frame in the upper
stories. The CFT column system presented in this paper is in the category of a 1-D CHS system.

1.3 Status of CFT Column System in Japan
CFT has been used for columns carrying large axial force since it was first used in Great Britain for theconstruction of road bridges in the late 1870's. In Japan, the first design recommendations for CFT column were
established in 1967 by the Architectural Institute of Japan (AIJ). Then it was revised in 1981 [1.2]. In theserecommendations, CFT columns were considered a kind of SRC column because of their similarity in usage andmechanical characteristics. The recommendations made it possible to utilize CFT columns in practice. However,the limit of the width (diameter)-to-thickness ratio for structural steel tubes was not relaxed in comparison withthe ratio for bare structural steel tube. Also, the filled concrete strength was not increased in the estimate ofactual strength even though it was confined by a steel tube and was under tri-axial stress conditions. Moreover,
the CFT column without concrete cover could not be used practically without a special approval. The BuildingStandard Law in Japan [1.3] said that covering concrete for a CFT column was only an assumption to beconsidered to be a SRC column. Thus, the diffusion of CFT column into real constructions was quite slow. Thissituation has not changed even after the design method for CFT was included in the current AIJ-SRC standards
revised in 1987 [1.4], in which both the width (diameter)-to-thickness ratio and the strength of the confinedconcrete were improved reflecting the real behavior verified by further studies.

In the “New-Urban Housing Project” organized in 1985 by the Ministry of Construction, very broad researchand development (R&D) were conducted. In particular, quantitative estimation of the strength and ductility of
CFT beam-columns brought by the synergistic action of structural steel tube and filled concrete was studied andformulated into design formulas. The mixture of filling concrete and the casting methods were also investigated.
The fire resistance of CFT columns was also extensively examined to study the possibility of the omission of-3-

covering concrete considering thermal capacity of filled concrete. However, the research results weremonopolized only by the private companies involved in this R&D.

Based on this current status of CFT column systems, standardized design method usable for ordinary
practitioners is strongly required. This covers not only the general results by the New-Urban Housing Project,
but also much larger width (diameter)-to-thickness ratios and material strengths. Figure 1.2 shows a typical CFT
column system and concrete filling.

1.4 Prioritized Research Topics Recommended in US-Japan Joint Workshop
Prior to the initiation of the fifth phase of the research, a Joint Planning Workshop [1.5] was held to identify
research issues. A number of topics related to research issues of CFT column system were raised and discussed,
some of which were interrelated and others were ultimately divided into more specific issues. As a consequenceof the discussions, a number of prioritized (high, medium, or low) research topics were recommended asfollows.

(1) High Priority Research Topics
The topics are summarized in Table 1.1.
1) Beam-Column s
Determination of the effect of confinement vs. composite action.
Determination of methods of evaluating axial, shear, and flexural stiffness, ductility and resistance undermonotonic and cyclic loading.
Determination of bond stress and shear transfer mechanisms between the structural steel tube and theconcrete.
2) Columns
Evaluation of creep and shrinkage effects on CFT columns of high-rise buildings.
3) Connections
Determination of force and moment transfer mechanisms and design models for FR, PR and pinned
connections for connections between braces, beams, and columns.
Determination of the effect of panel zone deformation on the seismic performance of CFT columns.
4) Frames
Analytical studies combining the above connection and member behavior to evaluate the seismic
performance of moment resisting frames and braced frames with CFT columns. This work should include
hysteretic behavior, damping, and dynamic characteristics of the structural system. It should evaluate
factors such as weak column - strong beam behavior and strong column - weak beam behavior, and bedirected toward determination of Ds and Rw factors and the seismic design forces.
5) Very High Strength ConcreteThis is a general research topic related to all composite or hybrid structures, but it also has impact on CFTcolumns.
6) Literature Review
Literature review related to CFT construction with a complete list of publications, brief abstracts ofindividual papers and reports to facilitate the exchange of information between the US and Japan.

-4-(2) Medium Priority Research Topics
These topics are summarized in Table 1.2.
1) Columns
Determination of the buckling behavior of CFT columns.
2) Nontraditional Connection Design
Development of nontraditional connections for CFT construction through the research for innovation
initiative.
3) Construction PracticeDevelopment of rules and guidelines for casting of concrete in CFT structures.
Establish the differences between as-built vs. design of practical CFT structures.
Determination of the effect of construction loads on CFT structures.
4) Frames
Evaluation of the relative merits of designing for composite action only, for confinement only, and for
combined composite action and confinement.

(3) Low Priority Research Topics
1) Repair and Retrofit
Establish methods and guidelines for repairing or retrofitting CFT composite and structures.

1.5 Japanese Research Program: Items Plan and Organization
The overall research operation system is shown in Figure 1.3. Domestic cooperative research by the Building
Research Institute, the Building Contractors Society, the Japan Structural Consultants Association, the Kozai
Club, and the Building Center of Japan was organized with research collaboration from universities. The
Technical Coordinating Committee (chaired by Dr. H. Aoyama : Professor Emeritus of Univ. of Tokyo) played
the role of the decision of the domestic research and the research adjustment with the US side. In the TechnicalSub-committee on CFT column systems (chaired by Dr. S. Morino: Professor of Mie Univ.), the detailed
research plan and its execution were carried out.

As mentioned in Section 1.3, the previous research in Japan was biased to the experimental research on
beam-columns and the strength and ductility were empirically estimated. Therefore, much basic study on thesection behavior of CFT columns required further investigation for applicability. The usable material strength in
the field of building construction was increasing both in structural steel and concrete. Therefore, the research
result applicable to them was strongly desired. The confinement effect of the structural steel tube for filledconcrete was thought to be extraordinary large. Hence, the usage of thin structural steel tubes was the trend in
CFT column system. Therefore, the possibility of utilization of thin structural steel tube was an important issue.
Finally, the dissemination of research result to the practitioners should be accelerated by giving design examples
and rational usage of this system. Considering the above, the following four items were stressed in determining
the research plan of the Japanese side.

1) More fundamental experimental / analytical studies2) Larger material strength ( σ u = 400~780MPa, Fc = 20~90MPa)-5-3) More slender steel tube section (B/t = 19~74, D/t = 17~152; B = width of square steel tube, D = diameterof steel tube, t = thickness of steel tube)
4) Trial design for practitioners

1.5.1 Experimental / Analytical Investigations
Four series of tests were planned as shown in Figure 1.4: a) centrally loaded stub columns, b) eccentricallyloaded stub columns, c) beam-columns, and d) beam-to-column connections. The objectives of these testing
were to clarify the synergistic interaction between structural steel tube and filled concrete and the stress transfer
mechanism, and to derive methods to evaluate stiffness, strength and ductility of CFT column system.

Study parameters were as follows: 1) tube shapes (square and circular); 2) nominal tensile strength of steel tube
(400, 590 and 780MPa); 3) width (diameter)-to-thickness ratio B/t (D/t) of steel tube; 4) design standard
strength of concrete (20, 40, 80 and 90 MPa); 5) axial load ratio n = N/No (N = axial load, and No = squash loadof CFT section); and vi) connection details. The width (diameter)-to-thickness ratio was classified into threeranks (FA, FC and FD) from the viewpoint of energy dissipation capacity of plastic hinges forming in hollow
structural steel tubes. In determining the range of parameters, the emphasis was placed on obtaining a wide
range of test data usable to establish a design method for CFT column systems. Structural steel tubes werecold-formed, and the value of B/t or D/t was controlled by changing the size of the tube.

Actual mechanical properties of steel and concrete are different from the corresponding nominal and design
values. In this paper, actual or nominal/design values are used case by case in each section.

(1) Centrally Loaded Stub Columns
The main objective of the centrally loaded stub column tests was to clarify the confining effect of the structuralsteel tube on the concrete strength and the restraining effect of filled concrete on local buckling of the steel tube,
and to establish the constitutive laws for steel and concrete which can be used for the analysis of CFT members.
Table 1.3 summarizes the test program of centrally loaded stub columns.

The constitutive laws for concrete and steel in a CFT column were established from the test results of concrete
cylinders, CFT stub columns, and hollow structural steel tube stub columns. The following phenomena weretaken into account: 1) increase in concrete strength due to confinement, 2) scale effect on concrete strength, 3)
strain softening in concrete, 4) increase in tension strength and decrease in compression strength of steel tube
due to ring tension stress, 5) local buckling of steel tube, 6) effect of concrete restraining the progress of local
buckling deformation, and 7) strain hardening of steel.

(2) Eccentrically Loaded Stub Columns
The main objective of the eccentrically loaded stub column tests was 1) to check the accuracy of the moment vs.
curvature relation numerically calculated on the basis of the proposed constitutive laws for steel and concrete,and 2) to derive the formula to evaluate the ultimate strength of the cross section subjected to combined axial
force and bending moment. Test program of eccentrically loaded stub columns is summarized in Table 1.3.

-6-

(3) Tests of CFT Beam-Columns
Tests of CFT beam-columns were conducted with the main objectives of 1) checking the accuracy of the methodof analysis based on the proposed constitutive laws for steel and concrete, 2) deriving formulas to evaluate therotation capacity, and 3) developing a model for the restoring force characteristics used in the dynamic response
analysis of CFT column systems. In this test series, some of the beam-columns were loaded under varying axial
force so as to increase the verification accuracy of the proposed analytical methods, which were rarely tested
previously. Table 1.4 summarizes the test program of CFT beam-columns.

(4) Tests of Beam-to-Column Connections
In order to clarify the performance of beam and column subassemblages in which the connection panel fails inshear, and to develop the design formulas for the connection, several connection tests were carried out. In thetest, a constant vertical load on the column and cyclic shear force at the beam ends was applied. Table 1.5
summarizes the test program of beam-to-column connections.

1.5.2 Trial Design
Structural designs of 10, 24 and 40-story CFT moment frame buildings were carried out on the theme structure
which was 38.4m in X-direction (6 spans) and 35.2m in Y-direction (3 spans) shown in Figure 1.5. In thesestructural design, allowable stress design, ultimate strength design and dynamic analysis of designed frame were
performed. The buildings with the same floor plans were also designed as pure steel frames. The total amount ofstructural steel is an index for estimating the economy of buildings. The amount of structural steel used for CFTbuildings was compared with that used for pure steel buildings. In this comparison, CFT system showedadvantages over pure steel especially for higher buildings. The detailed discussion will be made in Chapter 7.

-7-

CHAPTER 2: BEHAVIOR OF CENTRALLY LOADED SHORT COLUMNSThis chapter presents the ultimate loads and load vs. deformation relationships of centrally loaded CFT(concrete-filled structural steel tube) short columns.

2.1 Experimental Investigation
2.1.1 Specimens, Parameters and Test Procedure
A total of 90 specimens was fabricated and tested in the first phase of experimental investigations on centrally
loaded hollow and CFT short columns. The objectives of these tests were to investigate confining effect of steel
tube on concrete strength and restraining effect of filled concrete on local buckling of steel tube, and to derive
methods to evaluate ultimate loads and load vs. deformation relationships. In order to confirm the observations
obtained from the first phase tests, a total of 24 specimens with square section was tested in the second phase ofexperiments.

Study parameters for the first phase tests are as follows: 1) tube shapes (circular and square), 2) nominal tube
tensile strength (400, 590, 780MPa), 3) tube diameter (width)-to-thickness ratio (rank FA, FC, FD) and 4)
design concrete strength (20, 40, 80MPa). The diameter (width)-to-thickness ratios D/t (or B/t) were classified
into three ranks (FA, FC, FD) from a viewpoint of energy dissipation capacity of plastic hinges formed inhollow steel tubes. Hollow steel tube with rank FA is supposed to possess the ductility factor of 4 under purecompression, and that of rank FC the ductility factor of 1, which means that the tube locally buckles at the yieldstress. Hollow steel tube with rank FD buckles elastically. In determining the range of parameters, the emphasis
was placed on obtaining a wide range of test data usable to establish a generally applicable design method of
CFT column systems. Steel tubes were cold-formed with three different nominal wall thickness: 3.0, 4.5 and6.0mm. The value of D/t ratio was controlled by the diameter of circular tube (122~450mm). The B/t ratio wascontrolled by the width of square tube (120~324mm). Both ends of steel tube were welded to the end plates withthe thickness of 20mm for centrally loaded columns. Ratio of the clear height of specimens to D (or B) was 3.0.
The details of a typical specimen are shown in Figure 2.1. Material properties of steel tubes were obtained fromtensile tests of coupons taken from each steel plate before tube-forming. Yield ratio of the steel plates, which isdefined as the ratio of the yield stress to the tensile strength, strongly depends on the steel grades: 0.64~0.69 forgrade 400MPa steel; 0.9 for 590MPa; and 0.95 for 780MPa. Specimens with the same concrete strength werefilled from one batch of a ready mixed concrete plant, and the tests were conducted about three months after
concrete casting. Properties of all specimens are summarized in Table 2.1, 2.2(a) and 2.2(b).

The test setup shown in Figure 2.1 was used to apply the axial load. Averaged longitudinal strains were obtained
from axial shortenings between two end plates measured by four LVDT’s, and strains in steel tubes were
measured by two-element rosette strain gauges mounted on outer steel tube surface.

2.1.2 Test Results(1) Hollow Steel Tube Columns
The compressive stress σ su at the maximum axial load of hollow steel tube columns is given in Table 2.1,

-8-

which includes the compressive yield stress σ scy defined as the stress corresponding to an offset strain of
0.2%, as well as the tensile yield stress σ sy obtained from tensile coupon tests. The compressive yield stress isavailable for all the circular columns except for specimen CC4-D-0. In many cases of square columns with B/tratio of rank FC or FD, however, compressive yield stress was not obtained from tests, because an abrupt dropin axial load due to the local buckling occurred before yielding.

(2) Concrete Filled Steel Tube Columns

N u of circular and square CFT columns are summarized in Table 2.2(a) and 2.2(b),respectively. The non-dimensional maximum axial loads in the form of N u / N o ( N exp is used instead of
N u in Tables 2.2(a) and 2.2(b)) are also given in Tables, where N o is the nominal squash load given by
The maximum axial loads

Equation 2.1.

N o = N so + N co = As ⋅ σ sy + Ac ⋅ σ cp = As ⋅σ sy + Ac ⋅ γ U ⋅ f c'
where…..(2.1)

As and Ac are cross sectional area of steel tube and filled concrete, respectively, and σ cp isthe compressive strength of filled concrete which is estimated by multiplying the compressive strength of
10φ x 20cm cylinder

f c' by γ U . γ U is a reduction factor introduced to take a scale effect into

consideration, and it will be discussed in detail later. In the case of circular CFT columns, the value of thecompressive yield stress σ scy given in Table 2.1 was used to evaluate

N o in Equation 2.1, insteadof σ sy . The value of σ scy of specimen CC4-D, for which σ scy is not available as mentioned before,
was assumed to be equal to that of specimen CC4-C fabricated by using the same steel plate as CC4-D. In
the case of square CFT columns, the tensile yield stress given in Table 2.1 is used as the yield stressbecause of the following two reasons: 1) the compressive yield stress is not available for many CFT
columns with B/t ratio of rank FC or FD, and 2) an effect of cold-forming during fabrication of steel
tubes is expected to be relatively small in square steel tubes.

As shown in Table 2.2(a), the maximum axial load

N u is greater than the nominal squash load N o in mostof the circular CFT columns. A main reason for this augmentation of axial load capacity is attributed to aconfinement effect of steel tube on the concrete strength. On the other hand, the maximum load is less than thenominal squash load in the square CFT columns with B/t ratio of rank FD as shown in Table 2.2(b). A main
reason for this reduction of axial load capacity is attributed to the local buckling of steel tube. Methods toestimate the ultimate axial load will be discussed in the following section after discussing the scale effect on theconcrete strength.

(3) Scale Effect
One of the most important parameters in the experimental program is the D/t (or B/t) ratio of the steel tube as
mentioned previously. Steel tubes were cold-formed from the steel plate with the same thickness for each gradeof steel. The values of D/t and B/t ratio were controlled by the outside diameter of circular steel tube and thewidth of square steel tube. This resulted in the great differences in diameter or width of specimens as shown in
Figure 2.2. As expected from Figure 2.2, it seemed to be necessary to take a scale effect on the compressive
strength of concrete into consideration. Based on the careful investigations, we have reached a conclusion that-9-the concrete compressive cylinder strength should be modified according to the test results obtained by Blanks
et al. [2.1], which is shown in Figure 2.3. Symbols “A” to “D” in Figure 2.3 denote the diameter (or width) andsectional shape (circular or square) of CFT stub column specimens shown by corresponding symbols of “A” to“D” in Figure 2.2. Figure 2.3 shows that the concrete compressive strength of the largest specimen shown by thesymbol “A” should be considered to be about 85% of that of the smallest specimen whose size is almost thesame as concrete cylinder with 4 inches in diameter and 8 inches in height. Although the conclusion on the scaleeffect shown in Figure 2.3 was limited within circular columns, we applied Blanks’ result to the square CFT
columns by replacing them into the equivalent circular columns.

2.2 Analytical Investigation
In the next phase of the study, analytical models to estimate the ultimate strength of CFT short columns havebeen developed. Based on these models, stress vs. strain relations for filled concrete and steel tube have been
proposed independently, so that the proposed stress vs. strain relations can be applicable to a moment vs.
curvature analysis of CFT columns.

2.2.1 Ultimate Strength of Centrally Loaded Short Columns(1) Circular Columns
In the initial stages of loading of the circular CFT columns subjected to concentric axial load, Poisson’s ratio forthe concrete is lower than that for steel. Therefore, a separation between steel tube and concrete core takes placeprovided that the adhesive bond between the steel and concrete does not work. As the load increases furthermore
and the longitudinal strain reaches to a certain critical strain, the lateral deformations of the concrete catch up
with those of the steel tube. When the load increases furthermore, the hoop stress in the steel tube becomes intension, and the concrete core is subjected to triaxial compression. This phenomenon results in the increase of
axial compressive load of concrete. The equation for axial compressive load capacity is obtained by thefollowing procedure:First, the strength formula for concrete is assumed by Equation 2.2.

σ ccB = γ U ・ f c' ＋ k σ r…..(2.2)

where, σ ccB = confined concrete strength,

γ U = strength reduction factor for concrete = 1.67 Dc−0. 112 ,Dc = diameter of concrete core (in mm),f c' = concrete cylinder strength,k = confinement coefficient = 4.1 [2.2], andσ r = confining stress (lateral pressure).
The hoop stress σ sθ and axial stress σ sz of the steel tube at ultimate load are assumed by Equation 2.3.

σ sθ = α u ⋅σ sy , σ sz = β uc ⋅σ sy…..(2.3)- 10 -

where,α u , β uc = coefficients determined based on experimental results, assumed to be independent ofmaterial properties and dimensions of columns.
Referring to Figure 2.4, the relation between the hoop stress σ sθ and the lateral pressure σ r is given by
Equation 2.4.

σr =

− 2t
⋅σ
D − 2t sθ…..(2.4)

In the course of the evaluation of confining effect on concrete strength, it is assumed that the difference between
the ultimate strength N u and the nominal squash load

N o is provided by the confining effect on concrete

strength, and this gain depends upon the tube strength N so .

N u − No = λ ⋅ N so ;where,

NuN= 1.0 + λ soNoNo…..(2.5)

N so = axial yield strength of steel tube (= As ⋅σ sy ), andλ = augmentation factor.
From Equations 2.2 and 2.3,

N u is given as follow.

N u = As ⋅σ sz + Ac ⋅ σ ccB…..(2.6)

Substituting Equations 2.2, 2.3 and 2.4 into Equation 2.6, and using Equation 2.1 leads to Equation 2.7.

()

N u − N o = As ⋅ β uc ⋅σ sy + Ac γ U ⋅ f c' + k ⋅ σ r − As ⋅σ sy − Ac ⋅ γ U ⋅ f c'
2t
= As ⋅ σ sy ( β uc − 1) − Ac ⋅ k ⋅⋅ σ sθ
D − 2tA
2t= As ⋅ σ sy  β uc − 1 − c ⋅ k ⋅⋅α u 
As
D − 2tComparing Equation 2.7 with Equation 2.5, the factor

λ=…..(2.7)

λ is given by Equation 2.8.

Nu − No( D − 2t )= β uc − 1 −⋅ k ⋅α u
N so
2 (D − t )…..(2.8)

Equation 2.8 shows that the value of λ becomes constant if the values of coefficients k , α u andβ uc areλ defines the normalized axial compressive load capacity N u / N o as a linear
function of the parameter N so / N o . The value of λ was determined by a regression analysis based on
constant. The value of- 11 -

available experimental data as described below. The relation between the coefficientsα u and β uc is

obtained from the assumption that steel stresses at the ultimate stage given by Equation 2.3 satisfy the vonMises’ yield criterion given by Equation 2.9.

σ sz2 − σ sz ⋅σ sθ + σ s2θ = σ sy2where,…..(2.9)

σ sz = axial stress of steel tube under yield condition, andσ sθ = hoop stress of steel tube under yield condition, and thus
α u2 − α u ⋅ β uc + β uc2 = 1.....(2.10)

λ is fixed, the values of α u and β uc are determined by solving Equations 2.8 and 2.10,
where k = 4.1 as described before, and D/t = 50 as a representative to avoid dependency of the values of α uand β uc on D/t ratio.
Once the value of

Figure 2.5 shows the relationships between experimental axial load capacity
load of steel tube

N u of CFT columns and yield

N so , both divided by nominal squash load N o . The open circles show the existing

experimental results obtained elsewhere in Japan. The design formula recommended in the “Recommendations
for Design and Construction of Concrete Filled Steel Tubular Structures” [2.3] was proposed based on the open
circled data, where the slope of the dotted line

λ is defined as 0.27 in Equation 2.5. The solid circles show theresults presented in this chapter (US-Japan data). The slope of the solid line based on the US-Japan data isslightly lower than that of dotted line. However, it is concluded that the design formula already recommended is
not necessary to be revised. The value of

λ equal to 0.27 gives the values to the coefficients α u and β uc

as -0.19 and 0.89 respectively, from Equations 2.8 and 2.10.

Figure 2.6 shows the comparisons between experimental results on axial load capacity of the circular CFT stubcolumns and calculated capacities by Equation 2.11 which is obtained from the above procedure.

N u = No + λ ⋅ N so = N o + 0.27 N so.....(2.11)(2) Square Columns
In the case of square columns, it is necessary to take into consideration a capacity reduction due to the local
buckling of steel tube of the column with large B/t ratio rather than the confinement effect of the steel tube.
Figure 2.7 shows the relationships between the axial load capacity factor of the steel tube
normalized width-to-thickness ratio

S and the( B / t ) σ sy / Es , where S denotes the ultimate compressive strength

divided by the yield axial strength of steel tube. The axial load capacity factor of the hollow steel tube stubcolumns and the steel tube in CFT stub columns shown in Figure 2.7 are given by Equations 2.12, 2.13 and 2.14,
where

N su is the axial compression strength of square steel tube portion.

N su = As ⋅ σ scr , σ scr = min (σ sy , S ⋅ σ sy ).....(2.12)- 12 -

2
1
 B  σ sywhere,= 0.698 + 0.128  ⋅S
 t  Es

for hollow steel tube stub columns

21 B  σ sy 4.00
= 0.698 + 0.128  ⋅×S
 t  Es 6.97

for steel tube in CFT stub columns.....(2.13).....(2.14)

Equation 2.13 was obtained by a regression analysis using the experimental results of the hollow steel tube stubcolumns, and modified into Equation 2.14 by multiplying 4.00/6.97. This modification is based on an elastic
buckling theory by considering the difference in boundary conditions (or buckling modes) between the hollow
steel tube (simply-supported plate) and steel tube in CFT columns (clamped plate) shown in Figure 2.7.

The axial load capacity of CFT short columns can be estimated by Equation 2.15.

N u = N su + N co = As ⋅ σ scr + Ac ⋅ γ U ⋅ f c'.....(2.15)

Figure 2.8 shows the comparisons between experimental results on axial load capacity of the square CFT stubcolumns and calculated capacities obtained by Equation 2.15 which gives a slightly conservative value to thecolumns with small B/t ratio. The reason for this is considered to be a strain-hardening effect of steel tubes
rather than the confinement effect.

2.2.2 Stress vs. Strain Models for Filled Concrete
To predict the load vs. deformation relationships of centrally loaded CFT columns, a stress vs. strain curvemodel of confined concrete is necessary. Sakino and Sun [2.4] have proposed a unified stress vs. strain model
for concrete confined by steel tube and / or conventional rectilinear hoop. This model is expressed by Equation
2.16.

Y=

VX + (W − 1) X 2
1 + (V − 2 ) X + WX 2.....(2.16)

X and Y are concrete stress (σ c ) and strain ( ε c ) normalized by corresponding peak values
σ ccB and ε cco , respectively.

where,The stress vs. strain curve can be determined if the strengths of unconfined (plain) and confined concrete are
given, which can be seen from Table 2.3. “Original” model in Table 2.3 [2.4] has been expanded for confinedconcrete in circular or square CFT column by Narahara et al. [2.5]. In the case of circular CFT columns, thestrength enhancement factor
assumed to be γ U f

K = 1.0 − k ⋅

K

σ

/σ

σ

defined as
ccB
cp ( cp is the strength of unconfined concrete, and'c ) is given by Equation 2.17 which is obtained from Equations 2.2 and 2.4.

2t
σ
⋅ sθ
(D − 2t ) σ cp.....(2.17)- 13 -

Once the values ofk and σ sθ are fixed as 4.1 and − 0.19σ sy as described before, Equation 2.17 canbe transformed into Equation 2.18 where D/t=50 as a representative for the simplicity.

K = 1.0 + 0.032

σ sy.....(2.18)

σ cp

In the case of square CFT columns, the value ofK factor should be 1.0 from the viewpoint of axial load

capacity of centrally loaded short columns, in other words the confinement effect on ultimate axial load cannot
be expected. It is expected, however, that the axial load deformation capacity of filled concrete after reachingultimate axial load can be improved by confinement effect of square steel tubes in CFT columns as observed inmany experimental results. A slope of falling branch of stress vs. strain curve for the confined concrete is

σ cp and effective lateral pressureindex σ re defined in Table 2.3. The empirical formula to estimate the value of W for a square steel tube
governed by the constant W in Equation 2.16, which is the function ofacting only as the transverse reinforcement (referred to as the steel jacket) has also been proposed by Sakino andSun [2.4], and is shown in Table 2.3 as “Original” model. In this paper, it is assumed that the value of W for
the square steel tube in CFT column is equal to that for the square steel jacket. Table 2.3 gives all theinformation to obtain stress vs. strain curves for concrete confined by square steel jackets and steel tubes incircular and square CFT columns. Figure 2.9 shows these stress vs. strain curves along with the curve forunconfined concrete. The unconfined concrete strength and the yield strength of circular (D/t=60) and square(B/t=60) steel tubes shown in Figure 2.9 are 20MPa and 300MPa, respectively.

2.2.3 Stress vs. Strain Models for Steel Tube
For the circular CFT columns, the stress vs. strain relationship of the steel tube is developed as elastic-perfectly
plastic relation model as shown in Figure 2.10. The maximum stress of the steel tube is 0.89σ sy as described
before. In the case of square CFT columns, the stress vs. strain models for steel tube in CFT columns areproposed as shown in Figure 2.11. The three types of multi-linear model are described in the figure, where
Type-1 is the model for steel tube with small B/t ratio of which maximum stress is expected to be larger than theyield stress due to the strain-hardening effect, while the maximum stress of the steel tube with large B/t ratio(Type-3) does not reach the yield stress due to the local buckling. In the case of the steel tube with the medium
B/t ratio (Type-2), the maximum stress of the steel tube is defined as the yield stress. The classification for themodeling is according to the value of generalized B/t ratio,

α s as shown in Figure 2.7. The specific valuesof σ sB , ε sB , ε sE , σ sT , ε sT are calculated by using the equations summarized in Table 2.4 for each type.

σ sB and ε sB are the stress and strain at the local buckling, respectively, and ε sE is strain at
elastic limit. In the case of Type-2 and Type-3, ε sE is the same as ε sB , and σ sT and ε sT are the stress andIn the table,

strain at the termination point of falling branch.

2.3 Comparisons of the Experiments and Analytical Models
The proposed analytical load vs. deformation curves are compared with six experimental results for each shapeof CFT columns in Figure 2.12 and in Figure 2.13. These figures show the relations between experimental or
calculated axial load of CFT columns divided by the nominal squash load

- 14 -

N o and longitudinal strain. Thethick solid lines with experimental plots show experimental results and the thin solid lines show the analyticalcurves. The dashed line and chained line show calculated loads of the filled concrete and steel tube, respectively.
In each figure, specimens with different D/t or B/t ratio are shown together for comparisons. Good agreement isobserved between the predicted and experimental behavior.

- 15 -

CHAPTER 3: BEHAVIOR OF ECCENTRICALLY LOADED SHORT COLUMNSThis chapter presents the results of experimental investigation and analytical simulation of the behavior of CFTstub columns subjected to eccentric axial load.

3.1 Experimental Investigation
3.1.1 Specimens and Parameters
The details of a typical specimen are shown in Figure 3.1. Circular tube was cold-formed by press-bending andwelding. Square tube was fabricated by welding together two pieces of channel section, which were cold-formed
from flat plate. Both ends of steel tube were butt-welded with backing plates to the end plates of 40mm inthickness. Holes only for anchor bolts subjected to tension force were opened in the end plates, and a hole forconcrete casting was opened in the top end plate. Ratio of the clear height of specimen to diameter or width was
3.0.

A total of sixty-five specimens was tested: thirty-three circular and thirty-two square CFT specimens. The test
parameters were selected as follows: 1) nominal tensile strength of steel tube σ u (400, 590, 780MPa), 2)
design concrete strength Fc (20, 40, 80MPa), 3) diameter (width)-to-thickness ratio of steel tube D/t or B/t (rankFA, FC, FD), and 4) axial force ratio. The diameter (width)-to-thickness ratio is classified into four ranks(FA~FD) in the Japanese design practice from a viewpoint of deformation capacity of plastic hinges forming inhollow steel tubes. The ductility factor of 4 is guaranteed in the rank FA, 1 in the rank FC, and the elastic local
buckling occurs in the rank FD with strength deterioration. Steel tubes were cold-formed with three different
nominal wall thickness: 3.0, 4.5 and 6.0mm. The value of D/t (B/t) ratio was controlled by the outside diameterof circular tube (122~450mm) or the width of square tube (120~324mm). Although three different loading
schemes were used in the tests which are explained later in this chapter, the eccentricity of the axial load or theratio of the end moment to the axial load was determined in such a way that the ultimate strength of thespecimen would hit a certain point selected on the M-N interaction curve which was calculated beforehand.
Table 3.1 shows the dimensions of the tubes, material strengths and axial load ratios. More detailed informationon the material properties of steel tube and concrete is summarized in Tables 3.2 and 3.3, respectively, in which
the values are the average of 3 tensile coupon tests for steel and 3 cylinder compression tests for concrete.
Material properties of steel tubes were obtained from tensile coupon tests taken from each steel plate before
manufacture. Yield ratio of steel plates, which is defined as the ratio of the yield stress to the tensile strength,strongly depends on the steel grades: 0.64~0.69 for grade 400MPa steel, 0.90~0.92 for 590MPa, and 0.95 for
780MPa. Specimens with the same concrete strength were filled from one batch of a ready mixed concrete plant,
and the eccentric stub column tests were conducted about three months after concrete casting.

3.1.2 Loading Condition
Monotonic loading tests were carried out at three research institutes , so the loading conditions were not identical.

- 16 -

Circular columns were tested under the loading condition shown in Figure 3.2(a), and square columns were
tested under the conditions shown in Figures 3.2(b) and 3.2(c). In the case of loading conditions shown in
Figures 3.2(a) and 3.2(c), the magnitude of axial force in the column was maintained constant during the test. In
the case of the condition shown in Figure 3.2(b), however, the axial force and bending moment were increased
proportionally. Average curvature was obtained by measuring relative rotation between two sections, which wereapart by the distance of two times the outside diameter for circular colu mns and three times the width for squarecolumns. Longitudinal and lateral displacements and curvature were measured by LVDTs, and longitudinal andtransverse strains of steel tubes were measured by wire strain gages.

3.1.3 Test Results
The test results are shown in Table 3.4. Figure 3.3 shows the moment
specimens, where

M vs. curvature φ relationships of the

M and φ denote bending moment at the mid-height of the specimen and averagecurvature obtained by the rotation measurements, respectively. In the value of

M , Pδ moment evaluated at

the mid-height is included. In some of the specimens made of grade 590 or 780MPa steel, crack in the welds
between the end plate and steel tube was observed at the instance marked by “x” in Figure 3.3, and the test was
terminated. This failure occurred not only in the circular but also in the square CFT specimens. Reliable data on
the ultimate moment were not available for 9 specimens of circular columns and 8 square columns failing inpremature crack in the welds, of which maximum values of bending moment obtained in the tests are indicatedin the parenthes es in Table 3.4.

Figure 3.3(a) shows the moment vs. curvature relationships of the circular CFT columns. Each specimen shows
stable moment vs. curvature relation with large ductility, except for those fracturing at the weld, and four
specimens in the lower two figures using grade 400MPa steel in Figure 3.3(a). These four specimens showed
severe strength deterioration after the maximum strength reached, since the specimen with D/t = 101 and Fc =80MPa were subjected to large axial load but confining effect was not much expected from the steel tube of
400MPa grade steel, and D/t ratio was too large and the effect of local buckling clearly appeared in the case of
specimens with D/t = 152 and Fc = 40MPa. It must be careful to use the high strength concrete combined withthe low strength steel tube. It is generally understood that the usage of the high strength concrete causes the
reduction in the deformation capacity. While, the deformation capacity is improved by using high strength steel
tube.

Figure 3.3(b) shows the moment vs. curvature relationships of the square CFT columns. In each graph,

M − φ relations of two or three specimens with different axial force ratio of N / N o areshown, and they are quite similar, although the maximum strength is clearly affected by N / N o ratio. The
experimental

strength reduction due to local buckling is more or less observed in each specimen except for those in the top
two graphs, but the reduction is not very large. In some cases, the crashed concrete filled the gap between
locally deformed steel tube and concrete, and the load increased again, as observed in the specimen with B/t =33, Fc = 20MPa and grade 590MPa steel ( σ sy = 618MPa). Severe strength deterioration observed in thespecimen with grade 780MPa steel ( σ sy = 835MPa) was mainly caused by Pδ effect. Effects of the concrete
strength on the column behavior are not clear.

- 17 -

3.2 Analytical Investigation
3.2.1 Moment-Thrust-Curvature Relation
Moment-thrust-curvature relation of a CFT column was numerically derived by the fiber analysis based on the
stress vs. strain relations of concrete and steel.

The stress vs. strain relations of concrete are discussed in Chapter 2 and are shown schematically in Figure 3.4,
where the confining effect of circular steel tube is considered in the strength increase and the behavior after the
maximum strength for circular section, while it is only considered in the descending behavior after the
maximum strength in the case of square section, and the tensile strength of concrete is ignored. The
mathematical expressions of the stress vs. strain relation of concrete are shown below following Chapter 2 [3.1,
3.2]:

Y=

VX + (W − 1) X 2
1 + (V − 2 ) X + WX 2…..(3.1)

X = ε c / ε cco

for circular CFT…..(3.2-a)

X = ε c / ε cofor square CFT….. (3.2-b)

Y = σ c / σ ccB

for circular CFT….. (3.3-a)

Y = σ c / σ cp

for square CFT….. (3.3-b)

V = Ec ⋅ ε cco σ ccB

for circular CFT….. (3.4-a)

V = E c ⋅ ε co σ cp

for square CFT….. (3.4-b)

W = 1.5 − 17.1σ cp ×10 −3 + 2.39 σ re….. (3.5)

σ ccB = σ cp + k ⋅ σ r….. (3.6)

K = σ ccB σ cp….. (3.7)

σ cp = f c' × γ U , γ U = 1.67 Dc−0.112….. (3.8)

ε cco  1 + 4.7(K − 1)=ε co 3.35 + 20( K − 1.5)….. (3.9)(

K ≤ 1 .5
K > 1.5)

Ec = 6.90 + 3.32 σ cp ×103….. (3.10)- 18 -

ε co = 0.94(σ cp ) ×10 −3
14….. (3.11)

2t ⋅ 0.19 ⋅ σ sy
D − 2tk
σ re = σ r , k = 4.1 , k e = 23 for circular CFTke
σr =σ re =

where,1tρ h ⋅ σ sy  2bfor square CFT….. (3.12)….. (3.13-a)….. (3.13-b)

σ c : axial stress of concrete,
ε c : axial strain of concrete,
σ ccB : compressive strength of confined concrete,
ε cco : axial strain at the maximum strength of confined concrete,
σ cp : compressive strength of plain concrete = γ U ⋅ f c' ,f c' : compressive strength of concrete cylinder,ε co : axial strain at the maximum strength of plain concrete,
Ec : Young’s modulus of concrete,k : confinement coefficient,σ sy : yield strength of steel tube,
γ U : factor for scale effect,
D : diameter of steel tube,
t : thickness of steel tube,
σ r : confining stress,
ρ h : volumetric ratio of steel tube ( ρ h = 4( B − t ) / b 2 )B : Outside width of steel tube , andb : inner width of steel tube.

Fundamental discussions on the stress vs. strain relations of steel tube are also made in Chapter 2. Here, thestress vs. strain relations used for the analysis are shown schematically in Figure 3.5. The steel tube resists axial
load and bending moment, as well as it provides confinement. Therefore, the tube is subjected to both axial
stress and transverse stress. In order to take this effect of the biaxial stress state into account in the fiber analysis
of a circular CFT, it was assumed that the biaxial stresses at the initial yield satisfied the von Mises’ yield
criterion, and the compressive yield stress was reduced to 0.91σ sy and the tensile yield stress was increased
to 1.08σ sy , where σ sy denotes the yield stress obtained from the material coupon tests. Finally, b i-linear
stress vs. strain relation considering the strain-hardening effect was shown in Figure 3.5(a), where the local
buckling is not considered. The slope in the region of the strain-hardening can be chosen from the material tests.

On the other hand, the effect of local buckling is more pronounced in the case of a square tube, and the model
shown in Figure 3.5(b) was used for a square steel tube. The compression stress of the tube with small B/t ratio- 19 -

(rank FA) increases to the value larger than σ sy due to strain-hardening, and then the local buckling occurs
with strength deterioration. It is assumed that the local buckling of the tube with medium B/t ratio (rank FC)
occurs at the stress just reaching σ sy , and that of the tube with larger B/t ratio occurs at the elastic buckling
stress given by the following formula.

π 2 kc E s  t 
σ scr =
 
12 1 − ν 2  B 

(

2

)

…..(3.14)

where, Es and ν denote the Young’s modulus and Poisson’s ratio of steel, andk c is the plate buckling coefficient.
It is assumed that k c =6.97, since the shape of the cross section of a buckled square CFT was observed to be
similar to the clamped plate. In every case of ranks FA, FC and FD, the compressive stress reaches the stablepart in the range of large strain. The bi-linear relation is assumed for the tensile stress vs. strain relation withnominal yield stress equal to 1.1σ sy , considering confining effect similar to circular CFT and strain-hardening
in a simple manner.
The mathematical expressions of the stress vs. strain relations of circular and square tubes are given below [3.1,3.3]:a) Circular CFT− 0.91ε sy < ε < 1.08ε sy ( −0.91σ sy < σ < 1.08σ sy )σ = Es ⋅ εwhere,…..(3.15)

ε sy : yield strain of steel tube(= σ sy / E s ),ε : strain of steel tube,
σ sy : yield stress of steel tube,

σ : stress of steel tube, andEs : Young’s modulus.
ε < −0.91ε sy (σ < −0.91σ sy )σ =(σ − σ )(ε − ε ) ⋅ (ε + 0.91ε ) − 0.91E ⋅ ε
st

sy

su

sy

sy

…..(3.16)s

where, σ st : tensile stress of steel tube, andε su : strain at the tensile stress.

- 20 -

1.08ε sy < ε (1.08σ sy < σ )σ =(σ − σ )(ε − ε ) ⋅ (ε − 1.08ε ) + 1.08 E ⋅ ε
st

sy

su

sy

sy

…..(3.17)sb) Square CFTb-1) The tensile stress vs. strain relationε < 1.1ε sy (σ < 1.1σ sy )σ = Es ⋅ ε…..(3.18)

1.1ε sy < ε (1.1σ sy < σ )σ = 1.1⋅ σ sy…..(3.19)b-2) The compressive stress vs. strain relationlrank FA (ε sy ⋅

B≤ 1.54 )t….. Type-1 in Table 2.4 and Figure 2.7− ε sy < ε

σ = Es ⋅ ε− ε sB < ε < −ε sy (1 − S )(ε + ε sy ) − 1σ = σ sy  (ε sy − ε sB )where,…..(3.20)


ε sB = ε sy  6.06 2 − 0.801 + 1.10 
α
s
αs



S = 1 (0.698 + 0.128 ⋅ α s )…..(3.22)( )α s = ε sy B t…..(3.21)

2…..(3.23)− ε sT < ε < −ε sB- 21 -

σ=(σ − S ⋅σ ) (ε + ε ) − S ⋅ σsTsy(ε sT − ε sB )where,sB…..(3.24)

sy

ε sT = 3.59ε sy + ε sB…..(3.25)(

σ sT = σ sy 1.19 − 0.207 α s)…..(3.26)

ε < −ε sT
σ = −σ sTlrank FC( 1.54 ≤…..(3.27)

ε sy ⋅

B≤ 2.03 )t….. Type-2 in Table 2.4 and Figure 2.7− ε sy < ε

σ = Es ⋅ ε− ε sT < ε < −ε syσ =(σ − σ )(ε − ε ) (ε + ε ) − σ
sT

sy

sT

sy

sy

where,….. (3.28)

sy

ε sT = 4.59ε sy….. (3.29)

ε < −ε sT
σ = −σ sTlrank F D( 2.03 ≤ε sy ⋅

B)t….. Type-3 in Table 2.4 and Figure 2.7− Sε sy < ε

σ = Es ⋅ εwhere,

4.00 

S = 1  0.698 + 0.128 ⋅αs 
6.97 − ε sT < ε < − Sε sy- 22 -…..(3.30)

σ =(σ − S ⋅σ ) (ε + ε ) − S ⋅ σ(ε − S ⋅ ε )where,

sT

sy

sT

sy

sB

…..(3.31)

sy

ε sT = 4.59 S ⋅ ε sy…..(3.32)

ε < −ε sT
σ = −σ sT

Figures 3.6(a) and 3.6(b) show the comparisons between the moment vs. curvature relations obtained from thetests and the analyses of circular and square CFT specimens, respectively. In Figure 3.6(a) for circular CFT
specimens, it is observed that the theoretical initial elastic stiffness quite well agrees with the test results in allcases. In some cases, the analysis gives unsafe side estimate to the maximum strength obtained in the tests, andin some other cases, the strength deterioration after the maximum strength attained is more pronounced in theanalyses than in the tests. However, the analysis generally well traces the experimental behavior. It seems thattheM − φ relation of a circular CFT column is strongly affected by the stress vs. strain relation of concrete,

rather than that of steel tube.

On the contrary to those observations made for circular CFT columns, it is observed in Figure 3.6(b) for square
CFT columns that the analysis gives too conservative estimate to the maximum strength of specimens withgrade 780MPa steel ( σ sy =835MPa), and in these specimens the strength deterioration after the maximum
strength is much severer in the experiments. This may be attributed to the analysis in which the effect ofwork-hardening of the cold-formed square steel tube was not considered. In general, the flexural behavior ofsquare CFT column obtained by the analysis was strongly affected by the stress vs. strain relation of the steel
tube, which in contrary to the case of a circular CFT column.

3.2.2 Ultimate Bending Strength

M u denotes the test result, and M cal. 1 , M cal.2 ,and M cal.3 denote theoretical ultimate moments. The value of M cal. 1 was calculated from the rectangular
Table 3.4 shows the value of the ultimate moment, wherestress blocks assumed for both steel and concrete shown in Figure 3.7, where neither the confining effect on
concrete strength nor the local buckling of steel tube were considered. The value of

M cal.2 was calculated in

the same way, but the reduction of the concrete strength due to scale effect is considered as shown in Equation
3.8. The value of

M cal.3 was obtained from the peak point of the moment vs. curvature relation, which was

determined by the fiber analysis, based on the stress vs. strain relations given by Equations 3.1 and 3.15 - 3.32.

Comparison between the ratio of experimental ultimate moment

M u to theoretical full plastic moment

M cal. 1 and D/t (or B/t) ratio is shown in Figure 3.8 for all specimens except for those failed in premature
failure due to cracking. In the case of circular CFT column s shown in Figure 3.8(a), the effect of confined- 23 -

concrete and strain-hardening of steel are more pronounced in the specimens with D/t ratio smaller than 75, andM u / M cal. 1 exceeds 1.0, while it cannot reach 1.0 in the specimen with D/t ratio larger than 75 due to thescale effect on concrete. The average of M u / M cal. 1 of all circular specimens was 1.087 with the coefficientof variation equal to 0.04. On the other hand, the value of M u / M cal. 1 is less than 1.0 in most squarethus

specimens due to the scale effect and the local buckling, as shown in Figure 3.8(b). A few specimens with small
B/t ratio show the value ofM u / M cal. 1 larger than 1.0, but it is much smaller than those of circular specimens

with small D/t ratio, and thus the increase in bending strength due to the confined concrete is not expected insquare CFT column s. The average ofM u / M cal. 1 of all square specimens was 0.908 with the coefficient ofvariation equal to 0.014.

In Figure 3.9, the ultimate moment data are plotted by changing the theoretical ultimate moment toscale effect on concrete strength is considered in the calculation of

M cal.2 . The

M cal.2 . The difference in the data

distributions in Figures 3.8(a) and 3.9(a), both for circular CFT column s, is quite clear; the value ofM u / M cal.2 is larger than or equal to 1.0 except for one or two specimens, and it may be said that the safe side
evaluation of the strength is improved by taking the scale effect into account. The average and coefficient ofvariation of

M u / M cal.2 became 1.164 and 0.035, respectively. However, Figures 3.8(b) and 3.9(b) for square

CFT column s show almost the same data distributions, although a little improvement of the strength evaluation
is observed; the average and coefficient of variation of

M u / M cal.2 became 0.949 and 0.011, respectively.

The reduction of the ultimate strength is thus mainly caused by the local buckling in the case of square CFT
column s.

Comparison between the ratio of experimental ultimate moment(or B/t) ratio is shown in Figure 3.10. In the calculation ofM u to calculated moment M cal.3 and D/t

M cal.3 , the following factors were considered: 1)

scale effect on concrete strength; 2) confinement effect on concrete strength and change in nominal compressive
and tensile yield stress of steel tube due to biaxial state of stress, both for a circular CFT columns; and 3)
reduction in nominal compressive yield stress of square steel tube due to local buckling. Much better accuracy isobserved in Figure 3.10, compared with Figures 3.8 and 3.9. The average and the coefficient of variation of

M u / M cal.3 are as follows: 0.998 and 0.008 for circular CFT columns; and 1.079 and 0.015 for square CFT
columns.

- 24 -

CHAPTER 4: BEHAVIOR OF BEAM-COLUMNSThis chapter presents the behavior of interior beam-columns subjected to constant axial load. It also presents
exterior beam-columns subjected to a variable axial load, or a tensile axial load, that may be caused by theoverturning moment during severe earthquakes, and the behavior of square beam-columns subjected to thetwo-dimensional lateral loads, whose significance has been revealed by recent earthquakes. A total of
thirty-three beam-column specimens was tested for these purposes. This chapter also presents the results of thesimulations based on the stress vs. strain relationships proposed in the preceding chapters.

4.1 Experimental Investigation
4.1.1 Specimens and Parameters
Table 4.1 gives a summary of the specimens. The test parameters were steel tube shape (circular and square),material strength of steel tube and concrete, width (diameter)-to-thickness ratio of steel tube, axial loading
condition (constant and variable) and lateral loading direction for the square specimens. Steel plates with thetensile strength of grade 400MPa, 590MPa, and 780MPa were used for steel tubes. The results of coupon tests
and stub column tests of steel tubes are listed in Table 4.2, where the compressive yield stress was obtained fromthe 0.2% offset method. Two classes of FA and FC were selected on the width (diameter)-to-thickness ratio of
the steel tube, based on the classification of hollow steel tubes by Japanese design practice [4.1]. A plate elementwith the width (diameter)-to-thickness ratio in FA class is supposed to possess the ductility factor of 4 under thepure compression, while that in FC class the ductility factor of 1, that is, the local buckling occurs when thecompression stress just reaches the yield stress. Concrete with the design strength (Fc) of either 40 or 90MPa
was used to fill the steel tubes. The actual cylinder strengths are listed in Table 4.3.

A constant compressive axial load of 40% of nominal compressive strength,

N o (= As ⋅σ scy + Ac ⋅ f c' ), wasAs and Ac are sectional area of steel tube and concrete,'respectively; σ scy is the compressive yield stress of steel tube; f c is cylinder strength of concrete. Thisapplied on interior column specimens, where

value was considered as the maximum of the long-term axial load in a practical design. A variable axial loadwith a range of 30% tension of nominal yield strength of steel tube,

N s ( = As ⋅σ scy ), to 70% compression of

N o was applied on exterior column specimens. The lateral loading in the direction with an angle of 22.5 or
45.0 degrees to the principal axis was applied to four square interior column specimens. The total number ofcircular specimens was thirteen, and the total number of square specimens was twenty.

As illustrated in Figure 4.1, the specimen had the footing and top stub at both column ends. The clear column
length was six times the column depth or diameter in all specimens. The steel tube run through the footing andtop stub and welded there. Square steel tubes were fabricated by welding two pieces of channel sections together,which were cold-formed from flat plates. Circular steel tubes were cold-formed from press bending. Concrete
was cast into all specimens on the same day by filling it from the hole on the top of stub. Concrete was also castinto the footing and top stub to obtain the sufficient stiffness. The loading tests were started after three months
from concrete casting. The cylinder strength of each specimen at the test age is listed in Table 4.4.

- 25 -

4.1.2 Test Setup and Procedure
Figure 4.2 illustrates the test setup and the loading condition of specimen. The test setup was designed to subject
specimen to an axial load and horizontal displacement reversals in a double curvature condition with the point of
inflection occurring at the middle height. The footing and top stub were fixed to the reaction floor and theloading beam using the tension rods, respectively. Two vertical hydraulic actuators applied an axial load, andheld the top stub in parallel with the footing at the same time. A horizontal hydraulic actuator applied horizontal
load by controlling the displacement. The history of rotation angle

θ shown in Figure 4.3 was applied to all

specimens. Figure 4.4 illustrates rules on the axial force loading. During horizontal loading, the axial

N o was maintained constant for the interior column specimens. For the exterior
column specimens, the axial compressive load of 70% of N o was maintained constant in a half cycle of thepositive horizontal loading, while the axial tension load of 30% of N s was maintained constant in the othercompressive load of 40% of

half cycle of the negative horizontal loading. The axial load was changed when the horizontal load became zero,
permitting the horizontal drift of specimen during changing of axial load. The applied forces were measured
using the calibrated load cells. The specimens were instrumented to obtain the horizontal and axial
deformations.
4.1.3 Test Results
Table 4.4 gives the measured column depth (diameter) and thickness of steel tube, the actual width(diameter)-to-thickness ratio, the material strengths, the applied axial load ( N ), and the actual axial load ratio( N / N o or

N / N s ) of each specimen. The experimental flexural strengths (M ue ) are also listed there,
compared with the calculated flexural strength ( M uc ) and the analytical flexural strength ( M ua ). M uc is the'full plastic moment based on the compressive yield stress σ scy of steel tube and the cylinder strength f c .
M ua was obtained from the simulations described in the following section. Moment ( M ) vs. rotation angle( θ ) relationships and axial strain (ε ) vs. rotation angle (θ ) relationships of the specimens are shown in
Figures 4.5 through 4.9. The moment M is the end moment in the loading direction, including the Pδ
moment. The rotation angle θ is the chord rotation angle, and is obtained from dividing the measured
horizontal drift by column length. The axial strain ε expresses the average axial strain over the column length.
The symbols of square, triangle and circle in Figures 4.5 through 4.9 denote the first local buckling of the steel
tube, the first crack or fracture of steel and the maximum flexural strength, in each loading direction,
respectively.

(1) Circular Interior Column Specimens
As seen in Figure 4.5, all specimens reached the maximum strength after local buckling occurred in thecompressive flange. The specimens with grade 780MPa steel ( σ sy =771~820MPa) tube showed a moment
reduction in the loading cycle of θ = ±4% due to the crack occurring in the heat affected zone of welding at
the column end. The other specimens showed a very ductile behavior, even after the local buckling occurred.
The experimental fle xural strength is 1.16~1.60 times the calculated strength.

(2) Circular Exterior Column Specimens
A fracture of the tensile flange at the column ends occurred in the loading cycle under the axial tension in all- 26 -

specimens as shown in Figure 4.6. The specimens with grade 590MPa steel tube ( σ sy =482~504MPa) showed asignificant moment reduction due to the fracture in the loading cycle of θ = ±3〜 ± 4% , after local buckling
occurred and progressed at the compressive flange in the loading cycle under the axial compression. The
specimens with grade 780MPa steel tube also showed a reduction in moment resistance due to the fracture in theloading cycle of θ = ±2% , without the progress of the local buckling. The experimental flexural strength is
0.91~1.59 times the calculated strength in the loading cycle under the axial compression, and 0.94~1.40 times in
the loading cycle under the axial tension.

(3) Square Interior Column Specimens
The ultimate moment was observed after the local buckling occurred in the compressive flange, followed bycyclic deterioration in moment resistance in all specimens as shown in Figure 4.7. A crack was not observed insteel tube. The rotation angle at the ultimate moment θ max was 1.0~1.5% in the specimens with grade 400MPa
steel tube ( σ sy =276~295MPa), 1.5~2.0% in the specimens with grade 590MPa steel tubes ( σ sy =537~540MPa)
and 2.0~3.0% in the specimens with grade 780MPa steel tube ( σ sy =824~825MPa). The θ max of the specimen
with FA class steel tube was clearly greater than that of specimen with FC class steel tube. The axial
deformation tended to accumulate in compression, especially after local buckling occurred. The amount of theaxial shortening was much larger than that of circular specimens. The experimental flexural strength is0.95~1.29 times the calculated strength.

(4) Square Interior Column Specimens Subjected to Biaxial Bending
As shown in Figure 4.8, the behavior of the specimen was very similar to that of the specimen loaded laterally inthe principle direction up to the ultimate moment stage, then slightly larger moment deterioration was observed.
A crack was observed at the corner of square steel tube near the column end, however a drastic reduction in
moment resistance due to the crack was not observed. The experimental flexural strength is 0.95~1.20 times thecalculated strength.

(5) Square Exterior Column Specimens
Local buckling of the compressive flange was observed at the early deformation stage in the loading direction of
axial compression, then a fracture of the tensile flange at the column end occurred on the loading direction of
axial tension in the loading cycle ofθ = ±2% , which is shown in Figure 4.9. The experimental flexuralstrength is 0.75~1.19 times the calculated strength in the loading cycle under the axial compression, and1.01~1.09 times in the loading cycle under the axial tension.

4.2 Discussion on the Test Results
4.2.1 Effect of Sectional Shape
Circular steel tubes have the advantage of restraining local buckling and confining filled concrete, comparedwith square steel tubes. Under the constant axial compressive load, the flexural strength and ductility of circularspecimens are superior to those of square specimens, as shown in Figures 4.10 and 4.11. The square specimens
showed a remarkable axial shortening and deterioration in moment resistance after the local buckling, while thecircular specimens showed no or a little strength reduction and smaller axial shortening even after the local- 27 -

buckling occurred. This indicates that larger tensile strain occurred at the tension flange of circular specimens in
the large deformation stage. This causes the crack and fracture of circular steel tubes, in particular grade
780MPa steel tubes, which results in without sufficient elongation ability. Furthermore, the tensile stresses
occurring in the hoop direction of circular tubes seem to promote the crack and fracture.

4.2.2 Effect of Width (Diameter)-to-Thickness Ratio
Local buckling generally occurs at the earlier deformation stage with the larger width (diameter)-to-thickness
ratio. The ductility of the square columns becomes smaller as this ratio is larger, while this ratio has a little
influence on the ductility of circular columns, within the range of test program.

4.2.3 Effect of Material Strength
Figure 4.10 gives relationships between the θ max and the material strengths of the interior column specimens.
The ductility becomes larger as the steel strength becomes higher. It generally becomes smaller as the concrete
strength becomes higher, but concrete strength has a little influence on the behavior in case of the specimens
with high strength steel tube. This tendency indicates that high strength steel tube is effective to improve brittle
behavior of high strength concrete, and that the restraining effect of the filled concrete against the local bucklingof steel tube dose not depend on concrete strength.

Figure 4.11 gives relationships between the moment enhancement ratio

M ue / M uc which are denoted by thesymbols of ○ and □, and the material strengths of the steel tube of the interior specimens. The moment
enhancement ratios of the circular specimens are greater than those of square specimens. The confinement effectin the circular columns is clearly recognized, especially in the specimens with high strength steel tube. The
symbol of ▲ in the Figure 4.11(a) denotes the ratio of

M ue / M ucc , where M ucc is the calculated flexuralstrength considering the confinement effect by steel tube obtained from the method prescribed in the AIJ-CFT
Recommendations [4.2]. In this case, the moment enhancement ratio becomes the same level as that of squarespecimens, and is nearly 1.2 over the strength. From the test results of eccentrically loaded square CFT shortcolumns, it was concluded in Fujimoto et al. [4.3] that the moment enhancement by the confinement effect fromthe steel tube could not be expected in square CFT columns. This indicates that the other factors, such as thestrain-hardening of steel tube, the extra confinement from the stiff footing and top stub under the compressive
axial load, provides the moment enhancement of about 1.2 for the circular and square specimens equally.
4.2.4 Effect of Variable Axial Load
The compressive flange under the axial compression became the tensile flange under the axial tension.
Therefore, the flange was cyclically strained with large amplitude under the variable axial load. This causes
local buckling at the early deformation stage and the fracture of steel tube. Consequently, the ductility of
beam-columns reduces under the variable axial load.

4.2.5 Effect of Biaxial Bending
The crack occurring at the corner of square tubes was observed in the square specimens loaded laterally in the- 28 -

direction with the angle of 22.5 or 45 degrees to the principal axis. This is because the maximum tensile strain
occurs at the corner, which has experienced the plastic tensile strain during the cold-forming process. Figure
4.12 gives relationships between the experimental flexu ral strengths and the loading directions. This indicatesthat square beam-columns have almost the same performance even if the loading direction is changed, as in thecase of circular columns.

4.3 Analytical Investigation
4.3.1 Analytical Model
Figure 4.13 illustrates a column model used in the computer simulations with the curvature and axial strain
distributions assumed in the model. To represent the plastic deformation, the model has hinging portions with acertain length near the column ends, where inelastic curvature and axial strain of the critical section uniformly
distribute. The behavior of the column model including the elastic shear deformation can be expressed in thefollowing incremental form.

1
 2L
 ∆θ x (i )   6 EI + GAL=∆ θ x ( j )   − L + 1
 6 EI GAL

−L
1 
+
6 EI GAL   ∆M x (i )  + L  ∆φ x , p ( i ) 
2L
1  ∆M x ( j )  p ∆φ x , p ( j ) +
6 EI GAL …..(4.1)

1
 2L
 ∆θ y ( i )   6 EI + GAL= ∆ θ y ( j )   − L + 1
 6 EI GAL

−L
1 
+
6 EI GAL   ∆M y (i )  + L  ∆φ y , p (i ) 
2L
1  ∆M y ( j )  p ∆φ y , p ( j ) +
6 EI GAL …..(4.2)∆δ a = ∆N / EA + L p (∆ε a, p (i ) + ∆ε a, p ( j ) )…..(4.3)

where, the variables with a subscript of (i) or (j) express those related to the end i or j;∆θ x and ∆θ y

are the increments of end rotation angle in the x- and y-directions, respectively; ∆δ a is the increment∆M x and ∆M y are the increments of moment at column end in the x- andy-directions, respectively; ∆N is the increment of axial force; EI , GA and EA are the flexural,shear and axial stiffness in the elastic range, respectively; L is column length; L p is the length of
hinging portion and assumed to be column depth or diameter in the simulations; ∆φ x, p = ∆φ x − ∆φ x, e ;∆φ y , p = ∆φ y − ∆φ y ,e ; ∆ε a, p = ∆ε a − ∆ε a, e ; ∆φ x and ∆φ y are the increments of curvature at
the critical section in the x- and y-directions, respectively; ∆ε a is the increment of axial strain at thecritical section; ∆φ x, e , ∆φ y,e and ∆ε a, e are the elastic components of ∆φ x , ∆φ y and ∆ε a , andequal to ∆M x / EI , ∆M y / EI and ∆N / EA , respectively; ∆φ x, p , ∆φ y, p and ∆ε a , p are theinelastic components of ∆φ x , ∆φ y and ∆ε a , respectively.
of axial deformation;

The first component of the right hand side of Equations 4.1 through 4.3 represents the increment of elastic
deformation of the column, and the second one represents the increment of inelastic deformation. Developing- 29 -{∆φ , ∆φ , ∆ε } at the critical sections related to i and j ends
can be expressed by the incremental deformation vector {∆θ (i ), ∆θ ( j ), ∆θ (i ), ∆θ ( j ), ∆δ } .
Therefore, the force vector {M (i ), M ( j ), M ( i ), M ( j ), N } can be obtained step by step from thethese equations, the incremental strain vectors

x

y

a

x

x

x

y

x

y

y

a

y

ordinary sectional analysis for the moment vs. curvature relationships, following the idealized stress vs. strain
relationships.

In the simulation for specimens, the double curvature condition with the inflection point occurring at the middle
height was assumed, and thusθ x (i ) = θ x ( j ) and θ y (i ) = θ y ( j ) . The concrete and steel elements as

illustrated in Figure 4.14 were used. In order to fit the elastic stiffness to the test result, elastic rotational andaxial springs were added at both column ends, which represent the local deformations of the fixed ends observed
during testing. Furthermore, to take the moment enhancement caused by the confinement effect from the footing
and top stub under the axial compressive load into consideration, the critical section is assumed to move inward
from the column end corresponding to the development of hinging under the axial compression, and finally belocated at the position with the distance of

L p / 2 fro m the column end at the ultimate stage. Considering theaxial loading condition of the specimens, this modification of the end moment was conducted as follows:

M end = M cr (1 + β ⋅ H )…..(4.4)

H = (1 − K t / Ke )…..(4.5)

6. 0

β = L p / (L − L p )for the deformation range under the constant axial compressive load…..(4.6a)β =0

for the deformation range under the constant axial tensile load, or while the axial
load is changed…..(4.6b)

M end is the moment at the column end; M cr is the moment at the critical section, which isdetermined by the stresses of the elements; H is a irreversible index of the development of hinging
with a range of 0 to 1, but does not progress while β is zero; Ke and K t are the elastic andinstantaneous flexural stiffness at the critical section, respectively; and β expressed in Equation 4.6a iswhere,

a geometric constant.

M end is equal to M cr in the elastic range. However, it becomes L / (L − L p ) times M cr after the hinge
is fully developed under the axial compression. The value of L / (L − L p ) is 1.2 in this simulation because all
specimens have L of six times column depth and L p is assumed to be column depth. Note that thismodification of the end moment was conducted in the x- and y-directions, independently.

The stress vs. strain relationships of concrete and steel in circular and square CFT columns have beenformulated based on the test results of centrally loaded short columns in Chapter 2. However, these models werebased on the average axial strain between the measurement length of three times the column depth or diameter.

- 30 -

The slope of the descending branch of the stress vs. strain relationships is generally sensitive to the measurement
lengths because the failure region is limited within a certain length. In the simulation, the failure length was
assumed to be the column depth or diameter, and a modification was conducted with multiplying the slope of thedescending branch of the formulated stress vs. strain relationships of both concrete and steel by a coefficient of
1/3. The hysteretic rules proposed by Fujii et al. [4.4] and Meng et al. [4.5] were used for concrete and steel,respectively. Typical hysteretic stress vs. strain relationships are shown in Figure 4.15. Note that the strength of
52MPa was used in modeling 40MPa concrete whose actual cylinder strength was 36~42MPa. The reason forthis is that the test results of centrally loaded short columns with the same materials as the beam-column
specimens indicated that the concrete had a much higher strength than 40MPa. Characteristics after yielding
including the strain-hardening based on the coupon tests were used in modeling steel tubes. Further details
concerning the analytical model and the used assumptions were discussed in Inai et al. [4.6].

4.3.2 Analytical Results and Discussion
Figures 4.16 to 4.20 give the analyticalM − θ relationships and ε − θ relationships of specimens. The

analytical flexural strengths ( M ua ) are listed in Table 4.4, compared with the experimental flexural strengths( M ue ).
(1) Circular Specimens
The analyticalM −θrelationships show good agreement with the test results of the interior column

specimens while there is a little difference in theε −θrelationships, which is seen from the comparison of

Figures 4.16 and 4.17. For the exterior column specimens with grade 780MPa steel tube, the analytical results
are also in good agreement with the test results. However, the analytical results can not represent the behavior
after the local buckling observed in the exterior column specimens with grade 590MPa steel tube, since theinfluence of the local buckling is not considered in the used stress vs. strain curve and hysteretic rules of steel
tubes. The ratio of

M ue / M ua of the specimens is 0.90~1.19, except for the exterior specimens with grade590MPa steel tube. The presented analytical model can represent the hysteretic behavior of circular CFT
beam-columns, except for the behavior after the local buckling under the high axial compression.
(2) Square Specimens
Figures 4.18 through 4.20 show analyticalM − θ and ε − θ relations of square specimens. The analyticalM − θ relationships show good agreement with the test results of all specimens, including the behavior after
the local buckling, and represent the effects of test parameters on the behavior. A little difference in the ε − θrelationships is observed in the interior column specimens. However, the tendency of axial shortening after the
local buckling is well represented. The ratio of

M ue / M ua of all specimens is 0.89~1.09. The presented

analytical model can represent the hysteretic behavior of square CFT beam-columns very well.

- 31 -

CHAPTER 5: BEHAVIOR OF B EAM-TO-COLUMN SUBASSEMBLADGES

Shear panels are considered having very high reservation of ultimate strength comparing with other criticalsections such as beams and column s as the concrete panel is highly confined by steel tube. Therefore, small
concern was paid to such an element to the present and there have been not so many experimental researches on
the beam-to-column subassemblies. Recently the strength of steel and concrete become very high, but theexperimental studies on such connections are very few. This chapter presents the results of the experimental
investigation on CFT column - steel beam subassemblies using high strength materials , which were made as
weak shear panel so as to investigate the failure pattern, shear force carrying capacity, shear force transfer
mechanism in the panel and the effect of confinement.

5.1 Experimental Investigation
5.1.1 Specimens and Test Program
Figure 5.1 shows the illustration of subassemblies and Figure 5.2 shows the details of panel zone. Table 5.1
shows the dimensions and materials of eleven specimens. Ten specimens were plane subassemblies to whichreversed lateral load from one direction and column axial force were applied. One specimen was a stereo
subassembly to which reversed lateral forces from two directions and constant column axial force were applied.
There were two types of connections in the plane specimens. Eight were of interior and two were of exterior
subassemblies. Additional one had a shape of exterior but subjected to constant axial force. The column sections
were rectangular or circular. The specimens were about one third scaled models of actual CFT system, and they
were all planned to fail at shear panel zone. To make shear panel weak, the thickness of steel tube in the panel
was reduced to be about one third of that of column tube except for specimen C3. The thickness of steel tube in
the panel of C3 was around one half of that of column steel tube. The variables of the test programs were theloading direction of lateral force, the types of subassemblies such as interior or exterior, the shapes of steel tube
column and the combinations of the strength of concrete and steel. Diaphragms in the panel zone were madecontinuously to beam flange to avoid fracture in the welds at the junction of the panel and the beam flange. The
specimen R4 is only a subassembly of which diaphragm was welded around the steel tube of the column(so-called exterior diaphragm) as shown in Figure 5.2.

High strength concrete around 90MPa and high strength steel around 590 or 780MPa were used for specimens.
Rectangular specimen R1 and circular specimen C1 were standard specimens in this test program, of whichconcrete strength was around 90MPa and steel strength was around 590MPa. For specimens R2 and C2, thesteel strength was the same as R1 and C1, but the concrete strength was reduced to around 40MPa. For
specimens R3 and C3, the concrete strength was the same as R1 and C1 and the steel strength was around780MPa. Specimen R4 had the exterior diaphragm which differed from R1, but the strength of steel andconcrete was the same as R1. Specimen R6 was a stereo subassembly which had basically the same details withspecimen R1. Specimens R5, R5' and C4 were the exterior subassemblies with the same material strength of R1
and C1.

- 32 -

Execution of the tests was shared at several research institutes. So, the dimensions of the specimens were
determined according to the capacity of testing setup of each research institute. For example, the length ofbeams of circular specimen C3 was 3000 mm which differ from C1, C2. Therefore, the experimental results of
the yield and ultimate strength, the story drifts and so on of the specimens were evaluated and compared using
the normalized values. Properties of the concrete and steel are shown in Table 5.2

5.1.2 Test Setup and Procedure
Figure 5.3 and 5.4 show the test setup of specimen R1 and that of R6.

Figure 5.5 shows reversed loading procedure of interior specimens. Each cycle was controlled by the story drift
angle. Constant compressive axial force equivalent to 0.2 p No was applied to the interior specimens and was
kept through the test. Varying axial force from compression to tension was applied to the exterior specimens.
Figure 5.6 shows the applying rule of varying axial force. Compressive axial force equivalent to 0.7 p No wasapplied before the beginning of the positive cycle of loading. Then, just before negative loading after positive
one, the compressive axial force was replaced by the tensile axial force equivalent to 0.3 p N so . Here, theis the nominal squash load in the panel zone and

p Nso

p No

is the axial yielded strength of steel tube in the panelzone.

An example of instruments to measure the story drift and the deflection of the shear panel is illustrated in Figure5.7. Strain measurements were carried out so as to know the yielding process of the shear panel, beams andcolumns,

5.1.3 Test Results
5.1.3.1 Two dimensional Specimens(1) Progress of Damage
The test results are shown in Tables 5.3(a) and 5.3(b). All specimens failed in the shear panel, as they werescheduled. At first, local yield ing in the shear panel occurred and it spreads to overall panel. Then, local yielding
occurred in the beam flange and the steel tube column , which did not spread widely as the moment carrying
capacity of the beam and the column were planned being much stronger than the shear capacity of the panel asshown in Tables 5.3(a) and 5.3(b) .

At the final stage of experiment, s mall exp ansion of shear panel zone was observed, but buckling or cracking ofsteel tube in the panel was not observed. Figure 5.8 shows the crack pattern of the filled concrete in the panelzone after test, where the steel tube was cut and taken off.

(2) Story Shear Force vs. Story Drift Angle Relationships
The story shear force vs. story drift angle relationships of the specimens are shown in Figures 5.9(a) and 5.9(b).
In these figures,

p Qu

and

pQy

are the ultimate and yield panel shear strengths converted to story shear force.

All specimens showed ductile behavior with large energy absorption. After maximum strength, rapid loss ofshear force carrying capacity was not observed and gradual strength reduction occurred. For exterior- 33 -

subassemblies R5 and C4, the plastic zone in the panel expanded little by little in consequence of the increased
loading cycles. The formation of a plastic hinge in the panel made the specimens unstable, as the specimens
were unsymmetrical. Therefore in the fina1 stage of experiment, the bending deformation of the specimens
occurred at the panel zone when axial compressive force of 0.7 p No was applied to the column prior to thebeginning of the positive loading cycle. This is the reason why the flow of story drift angle occurred on the zeroline of story shear force as shown in the relationships of specimen R5 and C4 in Figures 5.9(a) and 5.9(b).

(3) Behavior of Panels, Beams and Columns
Figures 5.10(a) and 5.10(b) show the shear force vs. shear deflection angle relationships of panels . The shear
deflection was measured as shown in Figure 5.7. Every shear panel showed very ductile performance. Figures5.10(a) and 5.10(b) also show the relationships of story shear force and deflection angle of beams or column s.
These figures indicated that the beams and column s were still almost elastic even though local yield ing had
occurred.

Contributions of each elements, panel, beam and column, in the overall displacement of each subassembly are
showed in Figures 5.11(a) and 5.11(b). Percentage of contribution of shear panel in overall displacement became
greater in consequence of the increase of overall displacement. In the case of R5 and C4, different contributionsof each element were observed in negative loading cycle subjected to tensile axial force. Filled concrete in shear
panel decreases the contribution in proportion to the increase of overall displacement. Therefore the stress vs.
strain relationships of specimen R5, R5' and C4 showed a little greater inelastic hysteresis comparing with otherinterior subassemblies subjected to constant compressive axial force as shown in Figures 5.10(a) and 5.10(b).

5.1.3.2 Three Dimensional Specimens(1) Progress of Damage and Behavior of Subassembly
Test results of stereo specimen R6 are also shown in Table 5.3(a). The values about Specimen R6 in the table areexpressed as of 45-degree direction. Figure 5.12 shows shear force vs. story drift angle relationship. The valuesof this figure are also converted to the values of 45-degree direction. Elastic rigidity was almost the same as thecalculation. First yield of panel occurred at R=1.5% in the story drift angle. Up to R=3.0%, overall shear panel
had yielded. Story drift angle at the maximum strength was R=3.0%, in both positive and negative loading cycle.
At this time no sign of failure of the joint panel was observed except for yielding. A fterward, the gradualexpansion of steel in the panel began. At last half cycle, cracks occurred at the corner of the panel and thewelded boundary with diaphragm. The shear force carrying capacity of the joint panel was gradually decreased
after the maximum strength. However, the specimen indicated large ductility.

(2) Behavior of Panel, Beam and Column
Figure 5.13 shows the relations of shear force of panel and shear deflection angle of shear panel in x- andy-directions respectively. The figures show very ductile behavior of the shear panel. Figure 5.l3 also shows therelationships of the story shear and the deflection angle of the beam and column . These figures indicated almost
elastic behavior of the beam and column and the elastic rigidities were almost the same with the calculation.
Shear panel, beam and column showed approximately the same hysteresis in both x- and y-d irections.

- 34 -

Figure 5.14 shows the contribution of the each elements, panel, beam and column, in the overall story drift angle.
The percentage of story drift angle by panel was very small at first, though, finally it exc eeded 70% of overall
drift angle.

5.1.3.3 Discussions on Test Results
Figure 5.15 shows the relationships of normalized shear force of panel and normalized story drift angle so as tocompare the effects of different test variables.

(1) Effect of Material Strength
Figure 5.15(a) shows the comparison of behavior of rectangular subassemblies Rl, R2 and R3 and circularsubassemblies C1, C2 and C3.

In these figures, the envelope curves of R1 vs. R2 and C1 vs. C2 show the comparison of the behavior of thespecimens made of the same steel strength and different concrete strength. Much difference did not observed
between them. However, R1 and C1, of which concrete strength were higher, showed slightly steeper
degradation of strength after the ma ximum strength.

The envelope curves of R1 vs. R3 and C1 vs. C3 show the behavior of the specimens with the same concrete
strength and different steel strength. Both specimens R3 and C3, of which steel strength were higher, showedrather steeper strength degradation and small ductility.

Circular subassemblies showed higher shear force carrying capacity than rectangular subassemblies.

(2) Effect of Axial Load
Figure 5.15(b) shows the envelope curves of normalized shear strength of the shear panel and story drift angle
relationships of specimens R1 and R5 to know the effect of axial force. In the positive lateral loading side,
specimen R5 subjected to higher axial load of 0.7 p No showed smaller ductility than R1 subjected to constantaxial load of 0.2 p No . In the negative lateral loading side, the rigidity of R5 did not decrease so much eventhough the specimen subjected to tensile axial load of 0.3 p N so . Both specimens showed ductile performance.
(3) Effect of Diaphragm
Figure 5.15(c) shows the envelope curves of specimens R1 and R4, which have different types of diaphragms.
Rigidity and strength of specimen R4 with exterior (outer) diaphragm was smaller than those of R1 with interior(through) diaphragm. Exterior diaphragm of R4 was partially yielded near the corner of the steel tube column,
however it remained being elastic at the center portion of the diaphragm during the test.

(4) Effect of Loading Directions
Figure 5.15(d) shows the envelope curves of specimens R1 and R6. R6 was the three-dimensional specimen
loaded simultaneously in x- and y-directions to give lateral force in 45-degree direction. The shear force
carrying capacity of the panel of R6 in 45-degree direction was a little greater than that of specimen R1.

- 35 -

(5) Yield / Ultimate Shear Strength of Shear Panel
Figure 5.16 shows the yield strength and ultimate strength of shear panel compared with the design strengths
calculated by AIJ-SRC Standards [5.1] as the short-term allowable strength and the ultimate strength. Only
specimen R4 with exterior diaphragm showed a little smaller yield strength than AIJ-SRC Standard. The
ultimate strength was given as 1.2 times of the short-term allowable strength in the AIJ-SRC Standard.

Figure 5.17 shows the yield and ultimate shear strengths of specimen R6 compared with R1. In this figure, theyield and ultimate shear strengths were normalized by the short-term allowable shear strength by AIJ-SRC
Standards. It can be seen that the shear panel subjected to the 45-degree directional loading showed a little
higher ultimate strength that R1 and it showed much higher reserved strength in the yield strength.

5.2 Analytical Investigation
5.2.1 Model for the Restoring Force Characteristics of Shear Panels
The discussion of this section is limited to rectangular specimens. The relationships of shear force and shear
deflection angle of steel tube and concrete are assumed to be presented by tri-linear curve as shown in Figure
5.18 [5.2].

(1) Shear Strength vs. Shear Deflection Angle of Steel Panel
The shear behavior of the steel panel is given by tri-linear curve, of which the yielding point, the softening point
and the maximum point are given by the following equations.

1) yielding point8
ps Q y = w A
9ps γ y = k

ps
w

w

σ 2y − s σ o2…..(5.1)

3

Qy…..(5.2)

A⋅ s G2) softening point
ps Q r = w A

ps

γr =

w

σ r2 − s σ o2…..(5.3)

3( Q − Q )+ γ
ps

r

w

ps

A⋅ s Gy'ps…..(5.4)y3) maximum point- 36 -

ps Q m = w A

ps

γm =

w

…..(5.5)

3( Q − Q )+ γ
ps

m

w

where,σ B2 − s σ o2w

psA⋅ s Gr''

ps…..(5.6)r

A = 2⋅w B⋅ w t ,B : width of shear panel,s G : elastic shear rigidity of steel,'''s G , s G : the second and the third rigidity of the relationship of shear stress and shear deflection
w

angle determined by Prandtle-Ruess’ stress vs. strain relationship,t : thickness of web plate of steel tube,s σ o : axial stress of steel tube,
w σ y : yield point of web plate of steel tube,
w σ r = 0.8 × ( w σ B − w σ y )+ w σ y : intermediate stiffness degradation point stress of web plate ofw

steel tube,w

σ B : tensile stress of web plate of steel tube, andk : shear stress coefficient of steel box section.
(2) Shear Strength vs. Shear Deflection Angle of Filled Concrete
It was assumed that concrete was elastic up to shear cracking and concrete continued to keep the same strength
after the maximum strength. The points of shear cracking and the maximum strength was expressed by thefollowing equations.

1) shear cracking point
Shear cracking strength and corresponding displacement angle are calculated by the equation of the principal
stress of concrete.
pc

Qc = c A

cpc γ c = 1.2pccwhere,cσ t ( cσ t + c σ 0 )…..(5.7)

Qc…..(5.8)A⋅ c Gσ t = 0.5 f c' ,A : sectional area of concrete,c G : elastic shear rigidity coefficient of concrete,c σ o : axial stress of concrete, andf c' : compressive strength of concrete.
c2) maximum strength point- 37 -

Maximum strength of concrete is calculated by using the arch mechanism in the panel zone surrounded by steel
tube and diaphragm as shown in Figure 5.19.

cB

f M pc⋅ t ⋅ f '
4
Q
=
⋅
tan
θ
+
sin
θ
pc m
c
'
 2
c
ct ⋅ fc



pc

γm =

…..(5.9)Qm
c α m ⋅ c A⋅c G / 1.2
pc…..(5.10)

2


 H
H 

 −
where, θ = tan  1 + 
,


cB
cB


H
'
+ 0.086 ,c α m = 0.00158 f c + 0.0411
c B−1

B : width of concrete in panel,
H : height of concrete in panel,θ : declined angle of concrete arch mechanism in panel,2f M pc : plastic moment of flange plate of steel tube = c B⋅ f t
fσy,f t : thickness of flange plate in panel,f σ y ; yield point of flange plate of steel tube, andc t : thickness of concrete.
c()

The estimated skeleton curves for specimens Rl, R2, R3, R4 and R5 are plotted in Figure 5.20 with experimental
ones. Specimen R5 is the exterior subassembly with varying column axial force in the positive and negative
loading cycles. Thus, the skeleton curves are shown separately in the figure.

Calculated skeleton curves show almost the same characteristics with experiments, especially in the yieldstrength of shear panels. However, the calculations are apt to evaluate the smaller maximum shear force carrying
capacities than exp eriments.

5.2.2 Three Dimensional FEM Analysis
5.2.2.1 Analyzed Specimens
It is clear that the consideration of the confinement of filled concrete by steel tube is inevitable so as tounderstand the inelastic behavior of columns, beam-columns and beam-to-column connections of CFT system.
A three dimensional finite ele ment analysis is an appropriate tool to precisely simulate this effect. Therefore,
nine specimens, R1, R2, R3, R5 and R5' from rectangular specimens and C1, C2, C3 and C4 from circularspecimens, were analyzed by the three dimensional FEM method and compared with test results .

- 38 -

5.2.2.2 Analytical Procedure1) Idealization of Specimens
The halves of the specimens were idealized into three dimensional finite elements using symmetrical conditions
as shown in Figure 5.21. The computer code “FINAL” [5.3, 5.4] developed by Obayashi Corporation was used
for this study.

2) Constitutive Relations
Isoparametric solid elements with eight nodes were used for concrete, and isoparametric shell elements withfour nodes were used for steel plates. The constitutive equations of the concrete elements were based on an
orthotropic hypoelastic model based on the equivalent uniaxial strain concept. Willam-Warnke's five-parametermodel [5.5] was applied to determine the failure criteria of concrete surface under an triaxial stress condition,
and actual values of five-parameters were given by Ohnuma and Aoyagi's coefficients [5.6]. Stress vs. strain
relationship of steel was expressed by an elasto-plastic relation, which is based on von Mises' yield criterion.

3) Stress vs. Strain Relationship of Concrete
It’s well recognized that concrete confined by a steel tube increases the ultimate strength and ductility. Empiricalequations to evaluate the confinement effect were proposed on the basis of the test results of short CFT columns.
However, cracks might be produced in the CFT panel when shear stress exceeded a critical stress level. It wasquestionable whether the empirical equations are applicable to the analysis in such case. Therefore in thisanalysis , two kinds of stress vs. strain relationships for concrete are prepared so as to take into account theductile characteristics of confined concrete as shown in Figure 5.22.

Case 1: whole stress vs. strain, including both ascending and degrading zone, was expressed by a modified
Ahmad model [5.7]Case 2: relation of the ascending zone up to the maximum stress was expressed by the Ahmad model andthe degrading zone after the maximum stress was exp ressed by Sakino's model [5.8]4) Bond Transfer
In the panel zone, a perfect bond was assumed between the concrete and steel. It was because the concrete wassurrounded and highly confined by a steel tube and diaphragms . For the column, a film element was inserted
between the concrete and steel tube to permit the bond slip on the contact surfaces.

5.2.2.3 Analytical Results
Figure 5.23 shows the story shear force vs. story drift angle relationships of specimens obtained by experimentand analysis. The analysis of case 2 gave good simulations with experiments of all specimens. The analysis bycase 1 gave a little smaller shear force carrying capacity than the experimental and showed rapid loss of shear
strength in the case of specimens R1 and C1.
Figure 5.24 shows the relationships of shear force ( p Qc ) transferred by concrete and shear deflection angle ( γ )- 39 -

in the panel obtained from the analysis.
elements. In this figure,

p Qc

p Qc

was calcu lated from the stresses of integral points of the concrete

is normalized by Ac (sectional area of concrete) andf c' (the maximum

strength of concrete). Figure 5.25 shows the relationship of shear force ( p Qs ) transferred by the steel plate andthe shear deflection angle (γ ) in the panels. It is observed that the maximum shear stress of the rectangularpanel is around

0.3 f c' for the concrete strength of 100MPa, and is around 0.35 f c' for 50MPa concrete

strength. The maximum shear stresses of the circular panel is around0.35 f c' for the concrete strength of'100MPa, and is around 0.4 f c for 50MPa concrete strength. The maximum shear stresses of both rectangularand circular panels are not in proportion to the concrete strengths of the panel. In Figure 5.25, bi-linear
relationships ofp Qs − γ

calculated by the following equations are also given. The values ofs

Q p obtained

from the FEM analysis, which included the frame -work contribution carried by the flanges of steel tube anddiaphragms in the panel, are almost the same with the bi-linear line or a little greater.

EsAs
2(1 +ν s )Gs =σ sy − σ o2…..(5.11)2p

Q sy =

where,

3⋅ As….. (5.12)

Gs : elastic shear rigidity,

Es : Young's modulus of steel,p Q sy : shear yield strength of panel,

As : effective sectional area to transfer shear stress, andfor rectangular tube As = the area of web
for circular tube
As = half area of whole sectional area
σ sy : yield point of steel, σ o : axial stress of steel, and ν s : Poisson's ratio.
Figure 5.26 shows the contour of the minimum principal stress of concrete and the flow of the principal stresses
in the rectangular and circular panels. The principal stress flows seems to be wider compared with RC joint
panels . It is because that the steel tube of the CFT column in the joint panel contributes to the confining of theconcrete.

5.3 Examples of Numerical Simulations
So as to grasp the effect of column depth to beam depth ratio and declivity coefficient D’ in the softening zoneof concrete on the shear force transfer by concrete panel, imaginary column -to-beam assemblies shown in Table
5.4 were analyzed. The analytical results on imaginary assemblies are summarized in Table 5.5 together withthose on tested assemblies. In Figure 5.27, normalized concrete shear force ( p Q c / Ac / f c' ) is plotted against thecolumn depth to beam depth ratio (H/Dc), which indicates that smaller H/Dc gives larger
5.28 shows the effect of declivity coefficient D’ on the'p Q c / Ac / f c .

- 40 -'
p Q c / Ac / f c . FigureNotes1) Equations on CFT Shear Panel of AIJ-SRC Standards [5.1](a) Short-Term Allowable Shear Strength of Shear Panelp Qy =(2⋅ j fs ⋅ j β ⋅cV +s fs ⋅sV )/sB dwhere,

jf s : short-term shear strength of concrete (kgf/cm2 ) ,= 1.5 × min( f / 30,5 + f / 100)
'
c

'
c

f s : short-term shear strength of steel (kgf/cm2 ) = F y / 3 ,for circular column;
j β : = min (2⋅s D / sB d ,4.0 )= min (2.5⋅ s D / sB d ,4.0 )for rectangular column,s D : diameter of steel tube (outer diameter for circular tube and outer width for square tube),
cV : effective volume of concrete in the joint panel zone = c A⋅ sB d
sV : effective shear volume of steel in the joint panel zone = s A⋅ sB d / 2ssB d: distance of central axis between upper and lower flanges of steel beam.

(b) Ultimate Shear Strength of Shear Panelp Qu = 1. 2× p Q y2) Modified Ahmad Model and Sakino's Model for Concrete [5.7], [5.8]

{

(

) }

{

σ = A ⋅ X + D' − 1 X 2 ⋅ σ p / 1 + ( A − 2 ) ⋅ X + D ' ⋅ X 2}

A = E / Eo , E p = σ p / ε p(a) Softening Zone in Ahmad model

X = {1 + (ε − ε p ) / ε p } , n = 0.9 + 3.4( f c' / 1000 )

2

n

{(

) }

2

D ' = 1 + 1800 ⋅ σ p / f c' − 1 / f c'(b) Softening Zone in Sakino's modela) rectangular tubular column

X = ε / ε p , σ p = f c + 23σ re'

D ' = 1.5 − 1.68 ×10 −3 f c' + 0.75 σ re()

σ re = 0.5 ρh ⋅ σ hs d '' / C ⋅ {1 − s / (2 Dc )}b) circular tubular column{}

X = ε / ε p , σ p = f c' + 0.625 ⋅ 2 ⋅ t / ( Dc − 2t ) ⋅ σ sh

D = 1.5 − 1.68 ×10
'

−3

3

{

}

f + 0.75 (K − 1) ⋅ f / 23
'
c

'
c

0 .5

K = σ p / f c'
where, σ p , ε p : the maximum stress and the corresponding strain,

σ , ε : stress and strain,
Eo : elastic rigidity,
f c' : uniaxial compressive strength,
d '' : thickness of tube,
σ sh : yield point of steel,: volume ratio of steel tube,- 41 -

ρ h : volume ratio of steel tube,
C : interior width of steel tube, ands : pitch of lateral reinforcements (in the case of CFT s=0).

- 42 -

CHAPTER 6: PROPOSED DESIGN FORMULAS FOR COLUMNS AND BEAM-COLUMNS

6.1 General
The Architectural Institute of Japan (AIJ) recently edited "Recommendations for Design and Construction ofConcrete Filled Steel Tubular Structures" [6.1] (referred to as AIJ-CFT Recommendations). The AIJ-CFT
Recommendations have following features from the viewpoint of design formulas for load carrying capacities of
CFT columns and beam-columns. The first one is that they have new design formulas for slender columns(column curves). The second one is that the design formulas for circular CFT columns take into consideration ofconfinement effect which has been taken into consideration in Chinese Code [6.2], British Code [6.3] and Euro
Code [6.4], but has been ignored for a long time in the "Standards for Structural Calculation of Steel Reinforced
Concrete Structures" [6.5] (referred to as AIJ-SRC Standards) as well as ACI Codes [6.6]. The contents of
AIJ-CFT Recommendations are expected to be included near future in the AIJ-SRC Standards, which deal withthe structures composed of concrete-encased steel members including CFT members.

The AIJ-CFT Recommendations, however, have following limitations of application concerning material
strength and thickness of steel tube wall.

Material Strength:
Compressive strength of concrete

Fc ≤ 60MPa

Tensile strength of steel

σ u ≤ 590MPa

Diameter (or width)-to-thickness ratio D/t (or B/t) of steel tubes:
Circular section

Square section

D 2.35 ×10 4
≤
× 1.5
t
F
B 7.3 ×10 2
≤
× 1.5
t
F

…..(6.1)…..(6.2)

where , F is a standard value to determine allowable stress of steel (in MPa), and is taken as smaller value of
nominal yield stress σ y , or 70% of nominal tensile strength σ u .

The main objective of the five-year research project on CFT column system carried out as a part of the fifth
phase of the US-Japan Cooperative Earthquake Research Program was widening these limitations by anexperimental work as mentioned in Chapter 1. The tests of centrally and eccentrically loaded stub columns andbeam-columns subjected to combined axial load and bending moment were planned and conducted. The design
formulas for ultimate axial load and ultimate bending moment have been proposed based on those data, which
covered the following ranges of experimental parameters:- 43 -

Material strength:
Compressive strength of concrete

20 MPa ≤ Fc ≤ 90MPa

Tensile strength of steel

400 MPa ≤ σ u ≤ 780MPa

Diameter (or width)-to-thickness ratio D/t (or B/t) of steel tubes:

D / t ≤ 152 , B / t ≤ 74
D / t ≤ 75 , B / t ≤ 50
D / t ≤ 52 , B / t ≤ 41

for grade 400MPa steel
for grade 590MPa steel
for grade 780MPa steel

The design formulas for ultimate bending moment were proved to be applicable to estimate the ultimate load
carrying capacity of beam-columns obtained by experiments subjected to combined axial load, bending momentand shear. This chapter summarizes the design formulas for load carrying capacities of CFT columns based on
experimental investigations described in the previous chapters . The emphasis is placed on the capacities of the
CFT columns with thinner steel tube walls and / or high strength material which does not satisfy the limitationsof AIJ-CFT Recommendations. Note that in this chapter the soil mechanics sign convention for stress and strain
is followed. Namely, compressive stresses and strains are positive.

6.2 Ultimate Strength Formulas for Centrally Loaded Short Columns
6.2.1 Circular Columns

An ultimate axial load of centrally loaded circular CFT short columns is given by Equation 6.3.

N u = N o + λ ⋅ N sowhere,…..(6.3)

N o = nominal squash load = N so + N co = As ⋅ σ sy + Ac ⋅ γ U ⋅ f c'…..(6.4)

λ = augmentation factor to take confinement effect on concrete strength into consideration = 0.27,
f c' = cylinder strength of concrete,
As = cross sectional area of steel tube,
Ac = cross sectional area of filled concrete,
σ sy = yield point stress of steel tube,
γ U = 1.67 Dc− 0. 112 (for small scale test specimens)…..(6.5)= 0.85 (for actual design), and

Dc : diameter of concrete core, in mm.
The value of

λ in Equation 6.3 was obtained as 0.27 by the regression analysis of the test data of centrally

loaded CFT stub columns. Based on the analytical procedure in which yield criteria for the steel and concrete

λ =0.27 gives thebiaxial stress state of the steel tube shown in Figure 6.1, i.e. a magnitude of axial stress σ sz is 0.89σ sy , andthat of hoop tensile stress is − 0.19σ sy . A reduction factor γ U for concrete strength, which has been
are assumed to be ones proposed by von Mises and Richart [6.7], respectively, the value of- 44 -

introduced as the coefficient to take into consideration a scale effect [6.8] in the original empirical formula, isassumed to be a constant of 0.85 for the actual design. Figures 6.2(a) and 6.2(b) show comparisons between
ultimate axial loads calculated by Equation 6.3 and experimental results obtained in the US-Japan Research
Program and those in database which were collected from the proceedings of Annual Meeting of Architectural
Institute of Japan published in 1977 through 1989. As observed in Figure 6.2, the design formula given by
Equation 6.3 has reasonable accuracy.
6.2.2 Square Columns
An ultimate axial load of centrally loaded square CFT short columns is given by Equation 6.6.

N u = N su + N co = As ⋅ σ scr + Ac ⋅ γ U ⋅ f c'…..(6.6)

σ scr = min (σ sy , S ⋅ σ sy )…..(6.7)1
4. 00
 B σ= 0.698 + 0.128  sy ×S t  Es 6.97
2…..(6.8)

The confinement effect of square steel tube on compressive strength of concrete is ignored in Equation 6.6, but
the restraining effect of filled concrete on local buckling of steel tube wall is taken into consideration throughS given by Equation 6.8. The factor S is a reduction factor due to the local buckling of the steel
tube. The formula for S was first empirically derived from the test data of centra1ly loaded stub column ofthe factor

hollow square tube, and then modified to the reduction factor applicable to the steel tube in a CFT column by
multiplying 4.00/6.97, considering the difference in the mode of local buckling shown in Figure 6.3. The
numbers 4.00 and 6.97 are coefficients for plate buckling corresponding to the modes shown in Figures 6.3(a)and 6.3(b), respectively. Figures 6.4(a) and 6.4(b) show comparisons between ultimate axial loads calculated by
Equation 6.6 and experimental results obtained in the US-Japan Research Program and those in database whichwere collected from the proceedings of Annual Meetings of AIJ of 1977 through 1989. As observed in Figure
6.4, the formula given by Equation 6.6 results in slightly conservative estimation for ultimate strength

6.3 Ultimate Strength Formula for Eccentrically Loaded Short Columns
The ultimate strength for eccentrically loaded short columns is presented by an interaction curve between axial
load and ultimate moment. The AIJ-CFT Recommendations prescribe that the ultimate moment of columns
under axial load be estimated as a full plastic moment. In other words, both of the concrete and steel tube areassumed to be a perfectly plastic material. On the other hand, according to stress vs. strain models for filledconcrete of circular and square CFT columns proposed based on test results of centrally loaded CFT columns,
the specific stress vs. strain curves for filled concrete of circular CFT column with D/t ratio of 100 and square
CFT column with B/t ratio of 75 are rather brittle as shown in Figure 6.5, where the concrete compressive
strength and yield stress of steel tube are assumed to be 80MPa and 300MPa, respectively. Then, it is necessary
to introduce an ultimate (usable) strain and shape factors for concrete stress block for the evaluation of thestrength of CFT columns composed of high-strength concrete and thin-walled steel tube, which the AIJ-CFT- 45 -

Recommendations cannot be applied to.

6.3.1 Circular Columns
The AIJ-CFT Recommendations prescribe that ultimate moments of circular CFT columns be calculated byusing stress blocks for filled concrete and steel tube shown in Figure 6.6. Nominal yield stress is assumed to be

0.89σ sy and − 1.08σ sy in compression and tension, respectively. These values were derived from theexperimental data, and based on the assumption that stress in the steel tube confining the concrete satisfy vonMises’ yield criterion with the presence of hoop tension σ sθ equal to− 0.19σ sy as shown in Figure 6.1.

The magnitude of compressive strength of concrete is given by Equation 6.9 which has been proposed by
Richart et al. [6.7].

σ ccB = γ U ⋅ f c' + k ⋅ σ r = 0.85 ⋅ f c' + 4.1

2t(0.19)σ sy
D − 2t…..(6.9)

where , γ U =0.85 taking the scale effect into consideration as described in 6.2.1, and k=4.1 recommended
by Richart et al.
The confining stress (lateral pressure) σ r is caused by the confinement provided by the steel tube, and related
to the hoop stress σ sθ . The relation between σ r andσ sθ depends on the D/t ratio, but the relation for

D/t=50 is used here as a representative, for the simplicity in the design practice. Finally, the concrete strength forcircular CFT column is given by Equation 6.10.

σ ccB = 0.85 f c' + 0.032σ sy…..(6.10)

The use of high-strength concrete and / or very thin-walled steel tube results in rather brittle stress vs. strain
curve for filled concrete as shown in Figure 6.5. Sun and Sakino [6.9] have proposed the ultimate strain andshape factors for stress block shown in Figure 6.7 and Table 6.1 for concrete confined by the circular steel tube,
instead of the block shown in Figure 6.6(a) recommended in AIJ-CFT Recommendations.

The ultimate moment theory, in which the ultimate strain is introduced as prescribed in ACI Code [6.6], needs
rather complicated computation procedure, because there exists an elastic part in the web of steel tube near theneutral axis. The full plastic moment of steel tube, however, can be used as approximate estimation of ultimate
moment under such complicated stress distribution, because the contribution of thin-walled steel web to theultimate moment of the CFT column is relatively small. In that ease, the ultimate strain is not necessary to
calculate the ultimate moment. It should be kept in mind, however, that the shape factors of concrete stress blockshown in Table 6.1 are formulated based on the assumed ultimate strain.

- 46 -

Figure 6.8 shows the comparisons between the experimental ultimate moments of specimens tested in the
US-Japan Research Program and calculated ones based on the stress blocks for steel tube and filled concrete
shown in Figures 6.6(b) and 6.7, respectively. As observed in Figure 6.8, the design formula has reasonableaccuracy. In the case of CFT columns whose material strength and D/t ratio are within the applicable limitationsof the AIJ-CFT Recommendations, the difference between the ultimate moment calculated by the proposed
method and method prescribed in the AIJ-CFT Recommendations is very small, because the confinement effectof steel tube with relatively thick steel tube brings very ductile behavior of filled concrete. Then the proposed
method to calculate the ultimate moment of circular CFT columns based on stress blocks shown in Figures
6.6(b) and 6.7 can be used in wide range as described in the section 6.1.

6.3.2 Square Columns
The AIJ-CFT Recommendations prescribe that the ultimate moment of square CFT columns can be estimated asa full plastic moment calculated by using stress block for concrete and steel shown in Figures 6.9(a) and 6.9(b),respectively. In the case of CFT columns using steel tube with large B/t ratio, however, the reduced compressive
stress block of steel should be introduced to take the effect of local buckling of steel tube into consideration.
Moreover, the use of high-strength concrete and / or thin-walled steel tube results in rather brittle stress vs. strain
curve for concrete, then the ultimate strain and shape factors for stress block of concrete should be introduced in
the similar manner as the case of circular CFT columns.

Based on the consideration described above, it is proposed that the stress block for concrete and steel tube
shown in Figures 6.10(a) and 6.10(b) be used to estimate the ultimate moment of the square CFT columns withhigh-strength concrete and / or thin-walled steel tube. For those columns, the confinement effect of steel tube on
the ductility of filled concrete is very small and negligible, then the stress block shown in Figure 6.10(a) andTable 6.2 for plain concrete proposed by Sun et al. [6.10] can be used. The magnitude of compressive stress instress block for steel tube is the local buckling strength σ scr discussed in preceding section and given by
Equations 6.7 and 6.8. The effect of the elastic part in the web of square steel tube near the neutral axis isignored as shown in Figure 6.10(b) because of the same reason described in preceding section for the circular
CFT columns. This results in a slight overestimation of the ultimate moment of steel tube, which is compensatedby an underestimation of that of concrete due to ignoring the confinement effect of steel tube on the ductility ofconcrete described before. It is noteworthy that the stress block shown in Figure 6.10 should be used only for
CFT columns with high-strength concrete and / or thin-walled steel tube which are out of applicable limitationsin AIJ-CFT Recommendations. Figure 6.11 shows the comparisons between the experimental ultimate momentsof columns subjected to axial load and bending moment and calculated ones based on the stress block in Figure
6.9 or 6.10. A s observed in Figure 6.11, the design formula has reasonable accuracy.
6.4 Design Formulas for Beam-Columns under Combined Compression, Bending and Shear
In order to establish a seismic design method for the CFT column systems, it is necessary to investigate abehavior of CFT columns subjected to combined forces and deformed in a double curvature pattern shown in
Figure 6.12. The maximum shear

Qmax and the limit rotation angle Ru , which are defined in Figure 6.13 as- 47 -

indices of load carrying and deformation capacities of columns, will be discussed in this section.
6.4.1 Ultimate Bending Strength
The ultimate horizontal load of the columns shown in Figure 6.12 are dominated by their flexural capacity
except for the extremely short columns such as the columns with an aspect ratio,h / D ≤ 3.0 or h / B ≤ 3.0 .

The design formulas for ultimate moment discussed in the preceding section, which are proposed based on the
behavior of eccentrically loaded CFT columns, might be used to estimate the ultimate moment at a criticalsection of the columns shown in Figure 6.12. The comparison between the experimental ultimate moments andtheoretical ones calculated by the method discussed in the preceding section is shown in Figures 6.14(a) and6.14(b). The experimental ultimate moments are defined as the column end moments at the maximum shear of
the envelope curve of hysteresis loops of shear force vs. rotation angle of columns subjected to cyclic shear
force under constant axial load. The column end moment includes a secondary moment due to axial load andlateral displacement, so-calledPδ moment. As observed in Figure 6.14, the experimental ultimate moments

are larger than theoretical ones. The main reason for the discrepancies between the experimental and theoreticalultimate moments is that the additional confinement provided by the stiff loading stub adjacent to the criticalsection would shift the critical section away from the end section to a section carrying smaller moment. The
other reason can be attributed to a strain-hardening effect of the steel tube. It can be said that the theoreticalprediction has a comfortable margin to the actual ultimate moment even though some part of this margin iscanceled by thePδ moment which is usually ignored in the actual design procedure.

The column end moment at the maximum shear is taken as the experimental ultimate load carrying capacity ofcolumns and is plotted in Figures 6.14(a) and 6.14(b). The alternative definitions of ultimate load carrying
capacity can be made according to a performance-based design procedure. For example, if the maximum
column end moment attained within a rotation angle limitation of 0.01 radian is taken as the experimental
ultimate moment, Figures 6.14(a) and 6.14(b) are replaced by Figures 6.15(a) and 6.15(b). Following remarks
can be made from Figures 6.15(a) and 6.15(b).

1) The ultimate moments of specimens with circular and square sections using high-strength steel (grade
780MPa) tube cannot reach the theoretical moment due to the larger yield strain of high-strength steel.
2) The ultimate moments of specimens with circular section hardly reach the theoretical moment except for the
specimens using mild steel (g rade 400MPa). The reason for this is that the confinement effect of steel tube
cannot be fully developed within the rotation angle of 0.01 in radian.

In order to design the CFT columns using high-strength steel tube according to the performance-based design
procedure, it is necessary to conduct further investigations on lateral load carrying capacity of the CFT columns
under loading condition shown in Figure 6.12 especially for columns with circular section.

6.4.2 Limit Rotation Angle
The limit rotation angle

Ru of CFT columns discussed in this section is a characteristic point on the envelope- 48 -

curve of hysteresis loops of shear force vs. rotation angle, and is defined as the rotation angle at which 95% of
the maximum shear is maintained after reaching the maximum shear as shown in Figure 6.13. The design
formulas forRu given by Equations 6.11 and 6.12 have been established for both of circular and square CFT

columns by regression analysis using the experimental results tested in the US-Japan Research Program andthose in the database described before. The comparisons between experimental limit rotation angles and thosepredicted by Equations 6.11 and 6.12 are shown in Figure 6.16 for circular columns and in Figure 6.17 forsquare columns.
For circular columns:

Ru (% ) = 8.8 − 6.7

ND− 0.04 − 0.012 f c'Not…..(6.11)

For square columns:

Ru (% ) =

100

t
β
N B
0.15＋3.79
N0…..(6.12)

f c' − 40.3
β = 1.0 −≤ 1.0566
where , the value off c' is given in MPa.

The Japanese Building Standard Law prescribes that the ultimate state seismic demand used in the capacity
design procedure can be decreased in accordance with a structural characteristic factor
theDs which is similar to

R factor in NEHRP Provisions [6.11]. The Ds factor depends on the deformation capacity and energy

Ds factor is between 0.3~0.55 for so-called steel reinforced
concrete (SRC) structures. The Japanese Building Standard Law prescribes the Ds factor of each story
absorption capacity of the structures. The value ofaccording to the structural system and member ductility. The members are classified in a design practice into
four classes from the viewpoint of ductility, i.e. FA (very ductile), FB (ductile), FC (semi ductile) and FD (semibrittle). Table 6.3 is proposed to classify the CFT columns into these four categories according to their limit
rotation angles given by Equation 6.11 or 6.12.

6.4.2 Hysteretic Model for Behavior of CFT Beam-Columns(1) Tri-linear Skeleton Model
As illustrated in Figure 6.18, the moment vs. rotation angle relationship of CFT beam-columns subjected tocyclic lateral loading and a constant axial load can be expressed by a tri-linear skeleton model. This model isdefined by the following five parameters: 1) the elastic stiffness Ke; 2) the first bending point moment My; 3)
the stiffness degrading ratio at the second bending point ay; 4) the second bending point moment Mu; and 5) theultimate rotation angle Ru. The behavior of CFT beam-columns is assumed to be elastic until the first bending
point. My is given by the short-term allowable flexural strength prescribed in the AIJ-SRC Standards [6.5]. Mu- 49 -

is given by the ultimate flexural strength of beam-columns discussed in the preceding sections in this chapter.
Ru can be estimated from Equations 6.11 or 6.12. ay can be determined by statistical analysis of theexperimental data.

(2) Stiffness degrading Ratio ay
Figure 6.19 gives the definition of the experimental stiffness degrading ratio ay, which was given by the ratio of
K2 to K1, where K2 was the secant stiffness at the point of 85% of the maximum flexural strength Mmax, andK1 was the experimental initial stiffness and assumed to be the secant stiffness at the point of 33% of Mmax.
The moment of 85% of Mmax was considered to correspond to the ultimate fle xural strength proposed in thispaper. The effects of structural parameters (D/t, B/t,s

σ y , c σ B and N/No) on ay were statisticallyinvestigated using the experimental data shown in Table 6.4. As a result, it was concluded that no or very weakcorrelations existed between the structural parameters and ay. 80% of experimental ay were distributed within
the range from 0.4 to 0.9 as shown in Figures 6.20 and 6.21. Accordingly, the following average values of theexperimental data were proposed as ay to be used in modeling.

αy = 0.65 for circular CFT beam-columns….. (6.13)

αy = 0. 7….. (6.14)

for square CFT beam-columns(3) Comparisons between Proposed Hysteretic Model and Experimental Results
Figure 6.22 shows comparisons between the proposed trilinear skeleton model and the test results of thespecimens in the US-Japan Cooperative Research Program. The proposed model underestimates the ultimate
moment. This is because the ultimate flexural strength proposed in this paper gives average value for
eccentrically loaded or pure bended CFT columns, and because moment enhancement by the extra confinementfrom the loading stubs is frequently observed in the tests of beam-column specimens with stiff loading stubs.
This moment enhancement is considered as a safety margin in practical structural design. Figure 6.23 shows
comparisons of the hysteretic behavior, where the normal trilinear hysteretic rule is applied in the model. The
model can give a good prediction for rotation angles within about 1%, which covers the deformation range
expected in the ordinary design of buildings, while it overestimates the hysteretic energy for drift angles over
1% because of the difference of the stiffness on unloading.

Recently, more precise analysis is required to show the performance of building structures, which frequently
requires the inelastic dynamic analysis of building structures. The proposed hysteretic model for CFT
beam-columns can be used with high reliability based on experimental verification.

- 50 -

CHAPTER 7: TRIAL DESIGN AND MERITS OF CFT COLUMN SYSTEM

An actual merit of the CFT column system in a real building structure has not yet been clearly proved. It is very
important to define the merit of CFT system in comparison with structural steel system, when it is applied to areal building structure. From this perspective, trial design of CFT theme structures based on the design formulas
presented in the preceding chapters have been performed in an attempt to achieve this objective.

7.1 Theme Structures
7.1.1 Geometry

Theme structures treated here are 10, 24 and 40-story unbraced building frames made of CFT or structural steel
system as shown in Figure 7.1, and they have a typical framing floor plan as shown in Figure 7.2 [7.1]. CFT or
structural steel is used for columns and H-shaped structural steel member is used for beams.

All frames were designed according to the current Japanese practice [7.2], that is, the allowable stress design
against the seismic shear force under moderate earthquake, and the check for the ultimate horizontal strength ofthe designed frame against severe earthquake. The story shear used in the allowable stress design is given by thefollowing design formulas [7.3].
nQi = Ci ⋅ ∑ W j…..(7.1)

Ci = Rt ⋅ Z ⋅ Ai ⋅ Co…..(7.2)

j = j +1

Ai = 1 +

1
2T
⋅αi ⋅
1 + 3Tαi…..(7.3)

where, W j : weight of the j-th story,

Rt :Rt = 1.0 when T < Tc

Rt = 1.0 − 0.2(T / Tc − 1) when Tc ≤ T < 2Tc
2

Rt = 1 .6Tc / T when 2Tc ≤ T ,T : first natural period of the building,,

Tc : natural period of the ground below the building with the value of 0.4, 0.6 or 0.8 according to thekind of ground,

Z : factor of seismic zone with the value of 0.7, 0.8, 0.9 or 1.0 according to the seismic map of Japan,and

α i : ratio of the weight supported by the i-th story to the whole building weight above the groundlevel.

- 51 -

Rt = 1.0 for 10-story frame, 0.6 for 24-story frame, 0.5 for 40-story frame, Z = 1.0 ,and Co = 0.2 . Here, the value of Rt for 24-story and 40-story frames were not calculated directly by the
It was assumed that

equation described above, and were selected by the experienced design value for each height of building,because the height of them exceed the capable range of the equation in the current Japanese practice.

The ultimate horizontal strength was calculated by the pushover analysis, and it was verified that the strength of

Ds and Fes times Qi given by Equation 7.1 with Co = 1.0 , where Ds is the
structural characteristic value and was taken equal to 0.25, and Fes is the building shape factor and was taken
each story exceeded

equal to 1.0. Table 7.1 shows the characteristics of each theme structure used in the trial design and kinds ofanalysis performed. In the trial design, plastic hinges mainly formed in beams, and the columns remain elastic
until mechanism state, except for a few cases such as the column bases in the 1st story. Executed were elastic
static and dynamic analyses using full stiffness matrix of an entire structure, pushover analysis to obtain

Q − δ relation of each story, and elasto-plastic dynamic analyses. Elasto-plastic dynamic analysis was
performed only for 24-story frames.
7.1.2 Load Conditions
Table 7.2 shows the intensities of gravity loads, which are normally employed in the design practice of a typical
office building in Japan. The intensity of live load is different for the design under the long-term gravity load,
and for the design under the short-term seismic load. Table 7.3 shows the intensity of gravity load calculated foreach story indicated, which is used for the seismic design. The value per unit floor area is approximately 8.8
kN/m2 . Shear force acting in each story

Qi was calculated from Equations 7.1 to 7.3.

7.2 Trial Design
7.2.1 Design Conditions
The objective of the trial design is to find merits of CFT system in a common design, by investigating thedifference between the behavior of CFT and structural steel systems. In a common seismic design of a building
structure, the concept of weak beam and strong column has been adopted to avoid energy concentration to a
specific story. Thus, the following design conditions were adopted in this study.

1) The ratio of the stress in the column caused by the design load to the allowable stress was kept as near to 0.8
as possible, and that of the beam as near to 1.0 as possible.
2) Story drift angles were kept within 1/200 under the design load in the allowable stress design.
3) The collapse mechanism at the ultimate state was the overall frame mechanism in which the plastic hinges
formed only in beams, and all columns remained elastic except for the specific part such as the bottom ends ofcolumns at the lowest story.

Table 7.4 shows the list of members for 40-story frames proportioned by the design conditions.

- 52 -

7.2.2 Analysis of Designed Frame(1) Elastic Stress Analysis
The following treatments and assumptions were made for the model of elastic stress analysis, which was toobtain design stresses. These were basically applied to other analyses described later in this paper.

1) Bending, shear and axial deformations were considered for columns.
2) Bending and shear deformations were considered for beams.
3) The floor of each story was assumed as a rigid horizontal diaphragm.
4) Stiffness of CFT columns was calculated as a simple sum of stiffness of steel and concrete.
5) Multiplying factor to take the effect of slabs on the stiffness of beams was assumed 1.5 for one side slab and2.0 for both side slabs.

3-dimensional analysis by stiffness matrix method was used for elastic stress analysis.

(2) Elastic Dynamic Analysis
Lumped mass model with 3 degrees of freedom was used for elastic dynamic analysis. The stiffness in eachstory was determined by diminishing the full stiffness matrix used in the elastic stress analysis. Newmark's

βmethod for integration with time interval of 0.01 second was adopted for the analysis, and damping constant was
assumed to be 0.02 for critical damping, which was proportional to the stiffness. Input seismic ground motions
shown in Table 7.5, with maximum velocity level scaled to 25cm/sec and time interval of 0.01 second, wereadopted for the analysis .

(3) Pushover Analysis
The following treatments and assumptions were made for the model of pushover analysis in addition to thoseadopted in the elastic stress analysis.

1)M − θ relation assumed for beam ends was normal bilinear, which changed the stiffness at the full plastic
moment, having the second stiffness equal to 1/100 of the first.

2) Columns were assumed to remain elastic until the end of analysis since the plastic hinge formed only inbeams.

3-dimensional analysis by stiffness matrix method was used for elasto-plastic pushover analysis. Yielding inbeams was considered by using the model of rigid-plastic rotational spring at the member ends in the analysis.

(4) Elasto-Plastic Dynamic Analysis
Elasto-plastic dynamic analysis was performed by using one frame model in Y-direction extracted from the
24-story frames, to consider the hysteresis of each member directly in the analysis at every step. Shear
deformations at beam-to-column connections were considered in addition to the deformations considered in theelastic stress analysis. Input seismic ground motions shown in Table 7.5, with maximum velocity level scaled to
50cm/sec and time interval of 0.005 second, were adopted for the analysis. Hysteresis models adopted for- 53 -

columns and beams are shown in Table 7.6.

7.2.3 Results of Analysis
All results of trial design are shown in the form of comparison between CFT and structural steel.

(1) Results of Elastic Stress Analysis
The weight and stiffness of 40-story frames are shown in Table 7.7, which reveals the following characteristics.
1) The weight of CFT column is 2.1 to 3.3 times larger than that of structural steel column.
2) The cross sectional axial stiffness of CFT column is 1.5 to 2.2 times larger than that of structural steel
column.
3) The cross sectional bending stiffness of CFT column is 1.1 to 1.5 times larger than that of structural steel
column.
4) The story weight of CFT system is 1.1 to 1.16 times larger than that of structural steel system.
5) The story shear stiffness of CFT system is 1.1 to 1.3 times larger than that of structural steel system.

The story drifts of 40-story frames under the design load are shown in Table 7.8. The story drifts of CFT frame
in lower stories are larger than those of structural steel frame, while the former becomes smaller than the latterin upper stories. Figure 7.3 shows each story displacement of CFT frames and components caused by thebending and shear deformations of beams and the bending, shear and axial deformations of columns. The story
displacements caused by beam deformation δ B and column axial deformation δ C were calculated by thefollowing equations.

δ B = δ 0 − δ1
δC = δ0 −δ2where,

δ 0 : Total story displacement,δ 1 : Story displacement obtained from the analysis with making the bending and shear stiffness of beams
10000 times larger than the designed values, andδ 2 : Story displacement obtained from the analysis with making the axial stiffness of columns 10000 times
larger than the designed values.
It is observed from Figure 7.3 that 60 to 70% of the total story displacement is caused by beam deformation andthe rest is caused by column deformation in all cases of 3 frames analyzed. The proportion of axial deformationof the column to the total story displacement increases as the number of story increases, and it becomes as large
as 30% in the case of 40-story frame.

(2) Results of Elastic Dynamic Analysis
Figure 7.4 shows examples of vibration mode shapes and Table 7.9 shows the 1st natural period of vibration.

- 54 -

There's only 2% difference between CFT and structural steel.

Figure 7.5 shows the maximum response of the shear coefficients caused by El Centro. No significant difference
is observed between CFT and structural steel systems.

(3) Results of Pushover Analysis
Figure 7.6 shows Q − δ relations of 2nd and 9th floors of 40-story frames. Table 7.10 shows the energy
absorbed until the drift angle reaches 1/100, which is the area enclosed by theQ − δ curve, the horizontal

axis and the vertical line at the drift angle of 1/100. The following observations are made.

1) Yield story shear forces of CFT and structural steel systems are almost the same, because the overall framemechanism with beam hinges is adopted.
2) The energy absorbed in one story of CFT system at drift angle of 1/100 is larger by 4 to 8% than that ofstructural steel system.

(4) Results of Elasto-Plastic Dynamic Analysis
Table 7.11 and Figure 7.7 show maximum responses. From these results the following features can be observed.

1) The maximum story shear coefficient of CFT system is smaller by 5 to 8% than that of structural steel system.
2) The maximum overturning moment of CFT system is larger by 2 to 8% than that of structural steel system.
The difference may be caused mainly by the difference of mass of each system.
3) No significant difference regarding the maximum story drift can be found between CFT and structural steel
systems.

Figure 7.8 shows maximum response of ductility factors of beams and columns, and Figure 7.9 shows theplastic hinge formation in beams and columns. From these results the following features can be observed.

1) Ductility factors of beams stay within 3.0 at all stories, which are within the supposed performance of beams
of 4.0. Ductility factors of CFT and structural steel columns measured in comparison with elastic deformation
corresponding to M p stay within 0.6 at all stories.
2) Yielding occurred at the bottom of columns at the 1st story of CFT system, and no hinges are generated inany columns in structural steel system. This is derived from the difference of the definition of the firstyielding between two systems, that is, the first yielding point is defined by

M y for CFT and M p forstructural steel columns. The pattern of hinge generation in beams is very similar between two systems.
7.2.4 Amount of Steel and Cost Estimation
Figure 7.10 shows the comparison of steel amount per unit floor area used for CFT and structural steel systems,
and its ratio. Total steel amount includes steel used for columns, beams and sub-beams for entire building. Plates
and bolts for connections and reinforcing bars for floor slabs, foundation beams and footings are not included,- 55 -

which may be almost the same in both CFT and structural steel systems. The total steel amounts per unit floorarea of 3 structural steel frames are 105 kg/m2 for 10-story frame, 143 kg/m2 for 24-story frame and 189 kg/m2
for 40-story frame. These numbers are within a reasonable comparable range compared with those in theexisting buildings. The steel amount of CFT columns is less by about 25% than that of structural steel columns,
and the total steel amount of CFT system is less by about 10% than that of structural steel system.

Table 7.12 shows a cost estimation of main frames including columns, beams and sub-beams. The unit cost is
assumed to be 250,000 yen per ton for structural steel, and 35,000 yen per cubic meters for concrete. These unit
costs include materials, fabrications, transportation, and constructions. The cost of main frames for CFT system
is lower by 5 to 7% than that of structural steel system. The total building cost for CFT system would be lower
by 1% than that of structural steel system, if the cost of main frame structure is assumed to occupy 15% of thetotal building cost. As the number of stories increases, the cost merit of CFT system becomes larger.

- 56 -

CHAPTER 8: SUMMARIES AND FUTURE RESEARCH NEEDS

8.1 Summary of Each Chapter(1) Summary of Chapter 1
The overview of the research on the Japanese side on concrete-filled structural steel tube (CFT) column system
was presented in Chapter 1. It includes the definition of the composite and hybrid structures, in which CFT
column system is positioned. The status of CFT column system in Japan is next briefly summarized. As for the
research on CFT column system, various research issues on this topic are raised and they are prioritized. Finally,
the concrete research plan is summarized. The concrete research plan includes both experimental and analyticalinvestigations and design implication study through trial design is also included.

(2) Summary of Chapter 2
The ultimate strength and load vs. deformation relationships of the CFT columns with circular and squaresections are investigated based on the experimental results of 114 centrally loaded stub columns. The following
conclusions are reached on the bases of the study.
1) As for the ultimate strength of circular CFT columns, the difference between the ultimate strength andthe nominal squash load, which is provided by the confining effect on concrete strength, can be
estimated as a linear function of the tube yield strength. The biaxial stress state of the circular steel tube
at the ultimate strength is estimated based on the analytical procedure in which yield criteria for the steel
and concrete are assumed to be ones proposed by Mises and Richart, respectively.
2) Formula for a capacity reduction factor due to the local buckling of the steel tube was first empiricallyderived from the test data of centrally loaded steel column s of thin-walled hollow square tube. Then, it ismodified to the reduction factor applicable to the steel tube in a CFT column by considering restraining
effect of filled concrete on the local buckling of steel tube. This enables the estimation of ultimate
strength of square CFT columns with thin-walled steel tube, which is s maller than the nominal squash
load due to the local buckling of the steel tube.
3) Stress vs. strain models for concrete in CFT columns are formulated based on Sakino-Sun’s model
which has been proposed for concrete confined by a square steel tube acting only as the transverse
reinforcement so-called as a steel jacket.
4) Stress vs. strain model for a square steel tube is formulated based on the experimental results.

(3) Summary of Chapter 3
The ultimate strength and load vs. deformation relationships of the CFT columns with circular and squaresections were investigated based on the experimental results of 65 eccentrically loaded stub columns. The
followings are observed from the study.
1) Bending strength of eccentrically loaded circular CFT columns exceeded the superposed strength due tothe confinement effect regardless of the combination of material strength of steel tube and filling- 57 -

concrete.
2) Use of the high strength concrete generally caused the reduction of ductility of a circular CFT column.
However, it was made clear that such a non-ductile flexural behavior was improved by confining theconcrete with high strength steel tube.
3) Increase in bending strength due to the confinement effect could not be expected in the case of square
CFT columns. Moreover, the effect of local buckling must be considered.
4) Fiber analysis generally well traced the flexural behavior of eccentrically loaded CFT columns, andestimated well the ultimate strength obtained in the tests . Here, the scale effect on concrete strength,
confinement effect on concrete strength and change in nominal yield stress of a circular CFT columns,
and the reduction in nominal compressive yield stress of square steel tube due to local buckling are
included in the analytical models .

(4) Summary of Chapter 4
The ultimate strength and load vs. deformation relationships of the CFT beam-columns with circular and squaresections were investigated based on the experimental results of 33 beam-columns. The main findings are
summarized in the followings.
1) Circular CFT beam-columns show superior ductility than square ones .
2) Ductility becomes larger as the steel tube strength becomes higher, and it generally becomes smaller as
the filled concrete strength becomes higher. However, concrete strength has a little influence in the caseof the CFT beam-columns with high strength steel tube. To utilize high strength concrete, use of highstrength steel tube is effective.
3) Ductility of CFT beam-columns reduces under the variable axial load. It is because the damage
produced by the combined action of the lateral and axial loads concentrates on one of the flanges of thebeam-columns both in the positive and negative loadings.
4) Enhancement in moment resisting capacity by the confinement effect from steel tube can be expected incircular CFT beam-columns, while it cannot be expected in square CFT beam-columns. Extraconfinement from the footing and top stub of the specimens is admitted in both circular and square CFT
beam-columns.
5) Square CFT beam-columns show almost identical structural performance independent to the loading
directions.
6) The proposed analytical models in Chapters 2 and 3 are effective to represent the hysteretic behavior of
CFT beam-columns, except for the behavior after local buckling in the circular CFT beam-columns
under the high compressive axial load.

(5) Summary of Chapter 5
The ultimate strength and load vs. deformation relationships of the CFT beam-to-column subassembladges arestudied based on the experimental results of 11 beam-to-column specimens. The main findings are summarized
in the followings.

- 58 -

1) All subassemblies yielded at the shear panel, showing very good performance in elasto-plastic region
and indicated very ductile behavior. The gradual failure of panel occurred and no rapid loss of strength
was observed The intense buckling of steel tube and the crash of concrete also did not occurred Circularsubassemblies showed more ductile performance than rectangular subassemblies.
2) The shear panel contributed to share a large part of displacement in the overall displacement of .
subassemblies in plastic stage. Beams and columns preserved being almost elastic through testing.
However, when the exterior specimens subjected to tensile axial force, the contribution of shear panel
decrease the share in overall displacement as the cracks of the concrete in the panel zone had occurred3) Co mbination of strength of steel tube and filled concrete affected the maximum strength and the
behavior of specimens. Subassemblies of which the tube of panel was made of higher strength steel
showed the higher maximum strength. However, the specimens showed a little bit steeper degradation ofshear force carrying capacity after the maximum strength. This was true for both rectangular and circularsubassemblies,4) The yield and ultimate strength of specimens were greater than the shut-term allowable strength and theultimate strength of shear panel by SRC code of AIJ
5) The skeleton curves of the relationships between shear force and shear deflection angle of panels ofrectangular subassemblies were discussed on the basis of arch mechanism in the shear panel. The
confinement effect of the filled concrete in the shear panel was considered in the calculation The
obtained skeleton curves gave nearly good exp ressions with the actual relationships of specimens
6) Three dimensional FEM analysis simulated well the behavior and the ultimate strength of both
rectangular and circular subassemblies. The appropriate idealization of softening zone in concrete
stres s-strain relationship after the maximum strength gave good simulations with the exp erimental
results. Confining concrete by a steel tube and diaphragms did not contribute to increase in strength. but
did contribute to increase the ductility of panels
7) Shear force transferred by concrete and steel in panel obtained by the three dimensional FEM analysis
on actual specimens and by the numerica1 simulation. Shear force transferred by steel in elastic toplastic zone was evaluated with the bi-1inear relation Shear force transferred by concrete almost
proportionally decreased with the decrease of beam depth - colu mn depth ratios.

(6) Summary of Chapter 6
The design formulas for axial compressive load capacity, ultimate moment and deformation capacity of the CFT
columns with circular and square sections are proposed based on experimental results of specimens which wereplanned to obtain a wide range of test data usable to establish a generally applicable design methods of CFT
column systems. The proposed design formulas have following features.
1) The confinement effect in circular CFT columns, which has been ignored for a long time in AIJStandards for structural calculation of SRC structures as well as ACI Code, is taken into consideration
by assuming that a magnitude of hoop tension in steel tube at the ultimate state is− 0.19σ sy . The

ultimate moment of steel tube is estimated as the full plastic moment of the steel tube with the hoop

- 59 -

tension of − 0.19σ sy , and that of the concrete is estimated based on the ultimate strain and shapefactors of stress block of concrete confined by the steel tube with the hoop tension of − 0.19σ sy .
2) The capacity reduction due to the effect of the local buckling of steel tube is introduced to design
formulas of square CFT columns with large B/t ratio. Thus, the ultimate moment of square CFT columns
within the applicable limitations of AIJ Standards for structural calculations of SRC structures is
estimated as the full plastic moment, and that of the columns out of the applicable limitations can be
estimated by the proposed method in which the ultimate strain and shape factors of stress block for plain
concrete and the effect of the local buckling of steel tube are taken into considerations.

Ru of the CFT columns with circular and squaresections are proposed as an index of deformation capacity of the columns. The estimation of Ru makes3) The design formulas for the limit rotation angle

it possible to classify the CFT columns into four from the viewpoint of ductility demand according to theJapanese Building Standard Law.

(7) Summary of Chapter 7
From the studies of the trial design, the characteristics and merits of CFT column systems are summarized asfollows.
1) The lateral story stiffness of CFT column system is larger than that of structural steel system. The story
weight of CFT column system becomes larger than structural steel system, too. This leads to the similar
vibration characteristic of CFT and structural steel systems .
2) No major difference in natural periods and elastic responses was observed between CFT and structuralsteel systems.
3) No significant difference in the energy absorption capability and the elasto-plastic behavior wasobserved between CFT and structural steel systems, as far as the overall frame mechanism was adopted.
4) Total steel consumption of CFT column system for entire building was less by about 10% than that ofstructural steel system.
5) Within the assumption of the unit costs of steel and concrete, the cost of main frame structure became
lower by about 6% than that of structural steel system.

To sum up, cost merit was found in CFT column system compared to structural steel system. However, nosignificant merits of CFT column system could be found in terms of static and dynamic behavior within therange of small story drift of 1/100, considered in a common structural design, compared to structural steel
system. It may be said that the robustness of CFT column system beyond the range of the story drift of 1/100
supposed in a common design is higher than that of structural steel system. However, to prove the robustness of
CFT column system in terms of bearing capacity and ductility, collapse analysis into the range of large story
drift beyond the maximum bearing capacity, allowing the plastic hinge formation in columns, will be needed.

8.2 Future Research Needs
This series of research could not cover the verification by large frame tests. Therefore, further investigation of- 60 -

large frames is necessary. The concrete filling into structural steel tubes is important in securing the structural
performance of CFT column system although it cannot be inspected by the current methodology. This should
also be examined in the future. The damage to CFT columns initiates from the crash of filled concrete which is
not observable, and thus the evaluation of such damage is a significant issue to be solved.

- 61 -

ACKNOWLEDGEMENTS

The authors would like to give their sincere thanks to the members of the Technical Coordinating Committee
and the Technical Sub-Committee on CFT column system of the US-Japan Cooperative Engineering Researchon Composite and Hybrid Structures for their suggestions and comments. Prof. Emeritus H. Aoyama (Universityof Tokyo), Prof. A. Tanaka (Utsunomiya University) and Dr. H. Yamanouchi (Building Research Institute) areespecially acknowledged for serving as Chairman, Vice-chairman and Coordinator for the Technical
Coordinating Committee. Prof. S. Mahin (University of California, Berkeley), Prof. S. Goel (University ofMichigan), and Prof. C. Roeder (University of Washington) are also acknowledged for serving as Chairman,
Coordinator and CFT Sub-Committee Chair on the US-side. Ms Kumiko Hirayama, Building Research Institute,is also acknowledged for preparing figures and tables.

This study was supported by the Building Contractors Society, the Japan Structural Consultants Association, theKozai Club, and the Building Center of Japan. Kyusyu University provided testing facilities, which is especiallyacknowledged.

- 62 -

REFERENCES

1.1)"Recommendations for a U.S.-Japan Cooperative Research Program utilizing Large-Scale Testing
Facilities", Report No.UCB/EERC 79-26, September 1979

1.2)

Architectural Institute of Japan (AIJ): “Standards for Structural Calculation of Steel Reinforced Concrete
Structures”, 1967 & 1981 revision

1.3)

Building Standard Law of Japan

1.4)

Architectural Institute of Japan (AIJ): "Standards for Structural Calculation of Steel Reinforced Concrete
Structures", 1987 revision, June 20, 1987 (in Japanese)

1.5)

U.S.-Japan Planning Groups Joint Planning Workshop: “Recommendations for U.S.-Japan Cooperative
Research Program -Phase 5 Composite and Hybrid Structures”, UMCEE92-29, 1992.11.

2.1)

Blanks, R. F. and McNamara, C. C.: "Mass Concrete Tests in Large Cylinders", ACI Journal, Vol. 31,
January - February, 1935, pp.280-303

2.2)

Richart, F. E., Brandtzaeg, A and Brown R. L.: "The Failure of Plain and Spirally Reinforced Columns inCompression", Bulletin No.190, University of Illinois , Engineering Experimental Station, Urbana, 1929

2.3)

Architectural Institute of Japan (AIJ): "Recommendations for Design and Construction of Concrete
Filled Steel Tubular Structures", October 20, 1997 (in Japanese)

2.4)

Sakino, K. and Sun, Y. : "Stress-Strain Curve of Concrete Confined by Rectilinear Hoop”, Journal of
Structural Construction Engineering of AIJ, No.461, July, 1994, pp.95-l04 (in Japanese)

2.5)

Nakahara, H., Sakino, K., and Inai, E. : “Analytical Model for Axial Compressive Behavior of Concrete
Filled Square Steel Tubular Columns”, Proceedings of the Japan Concrete Institute, Vol.20, No.3, 1998,
pp.817-822 (in Japanese)

3.1)

Sakino, K., Ninakawa, T., and Matsumoto, M. : “US-Japan Cooperative Research Project on Composite
and Hybrid Structures (CFT-17) Axial Compressive Behavior of Concrete Filled Circular Steel Tubular
Columns”, Summaries of Technical Papers of Annual Meeting of AIJ, Structures III, 1997, pp.917-918(in Japanese)

3.2)

Nakahara, H., Sakino, K., and Inai, E. : “Analytical Model for Axial Compressive Behavior of Concrete
Filled Square Steel Tubular Columns”, Proceedings of the Japan Concrete Institute, Vol.20, No.3, 1998,
pp.171-178 (in Japanese)

3.3)

Nakahara, H., and Sakino, K.: “Behavior of Concrete Filled Square Steel Tubular Columns under aConstant Gravity Load and Cyclic Bending Moment”, Summaries of Technical Papers of Annual
Meeting of AIJ, Structures III, 1999, pp.1241-1242 (in Japanese)

4.1)

Housing Guidance Division of Ministry of Land, Infrastructure and Transport (MLIT), Japan Building
Officials Conference, Building Center of Japan: "Technical Standard Manuals for Structural Regulationsof Buildings", March 15, 2001 (in Japanese)

4.2)

Architectural Institute of Japan (AIJ): "Recommendations for Design and Construction of Concrete
Filled Steel Tubular Structures", October 20, 1997 (in Japanese)- 63 -

4.3)

Fujimoto, T., Mukai, A., Nishiyama, I., Noguchi, T., Baba, T., Sakino, K, and Morino, S. :"Eccentriccompression test of concrete filled square steel tubular stub columns using high strength materials",
Journal of Structural Construction Engineering (Transactions of AIJ), No.501, 1997, pp.173-180 (inJapanese)

4.4)

Fujii, S., Aoyama, H., and Umemura, H.: "Analytical moment-curvature relationship of reinforced
concrete sections based on the material characteristics", Summaries of Technical Papers of Annual
Meeting of AIJ, 1973, pp.1261-1262

4.5)

Meng, L., Ohi, K., and Takanashi, K.: "A simplified model of steel structural members with strength
deterioration used for earthquake response analysis", Journal of Structural Construction Engineering(Transactions of AIJ), No.437, 1992, pp.115-124

4.6)

Inai, E., Nishiyama , I., and Sakino, K.: "Simulation of shear-flexural behavior of concrete filled steel
tubular columns, Part 1 on the analytical method", Summaries of Technical Papers of Annual Meeting of
AIJ, Structures III, 1997, pp.921-922

5.1)

Architectural Institute of Japan (AIJ): "Standards for Structural Calculation of Steel Reinforced Concrete
Structures", 1987 revision, June 20, 1987 (in Japanese)

5.2)

Fukumoto, T., Sawamoto, Y., and Morita, K.: " Connection between High-Strength Concrete-Filled
Square Tubular Steel Column and Steel Beam Reinforced with Internal Diaphragm", Journal of
Structural and Construction Engineering (Transactions of A IJ), No.507, May, 1998, pp.171-178 (inJapanese)

5.3)

Naganuma, H.: " Stress-Strain Relationship For Concrete Under Triaxial Compression", Journal of
Structural and Construction Engineering (Transactions of AIJ), No.474, August, 1995, pp.163-l70 (inJapanese)

5.4)

Naganuma, H.: "Orthotropic Constitutive Relationship For Concrete Considering Nonlinear Poisson:
Effect Under Triaxial Stress", Journal of Structural and Construction Engineering (Transactions of AJJ),
No.485, July, 1996, pp.109-116 (in Japanese)

5.5)

William, K. J., and Warnke, E. P.: "Constitutive Model for the Triaxial Behavior of Concrete",
Proceedings of International Association for Bridge and Structural Engineering, Vo l.19, 1975

5.6)

Ohnuma, H. and Aoyagi, M.: "Ultimate Strength Property of Concrete under Triaxial Compressive
Stresses", Report of CRIElPI, No.381021, Dec., 1981 (in Japanese)

5.7)

Yonezawa, K., Yoshioka, K. et al.: "US-Japan Cooperative Structural Research Pro ject on Composite andHybrid Structures (CFT-22)", AIJ A nnual Convention C-1, September, 1997, pp.927-928 (in Japanese)

5.8)

Sakino, K. et al.: "US-Japan Cooperative Structural Research Project on Composite and Hybrid
Structures (CFT-17)", AIJ A nnual Convention C-1, September, 1997, pp.917-918 (in Japanese)

6.1)

Architectural Institute of Japan (AIJ): "Recommendations for Design and Construction of Concrete
Filled Steel Tubular Structures", October 20, 1997 (in Japanese)

6.2)

Harbin University of Civil Engineering and Architecture: “Design Code of Concrete Filled Steel Tubes(DL5099-1997)”

6.3)

British Standards Institution: “BS5400,Part 5, Concrete and Composite Bridges,” 1979

6.4)

European Committee for Standardization: “Eurocode 4: Design of Composite Steel and Concrete
Structures,” 1992

- 64 -

6.5)

Architectural Institute of Japan (AIJ): "Standards for Structural Calculation of Steel Reinforced Concrete
Structures", 1987 revision, June 20, 1987 (in Japanese)

6.6)

American Concrete Institute: "Building Code Requirements of Structural Concrete-ACI 318-95", 1995

6.7)

Richart, F. E., Brandtzaeg, A and Brown R. L.: "The Failure of Plain and Spirally Re inforced Columns inCompression", Bulletin No.190, University of Illinois, Engineering Experimental Station, Urbana, 1929

6.8)

Blanks, R. F. and McNamara, C. C.: "Mass Concrete Tests in Large Cylinders", ACI Journal, Vol. 31,
January - February, 1935, pp.280-303

6.9)

Sun, Y. P. and Sakino, K.: "Assessment of Ultimate Strength and Deformability of an RC Column
Retrofitted by Circular Steel Tube", Proceedings of the 2nd International RILEM/CSIRO/ACRA
Conference, September 21-23, 1998, pp.229-245

6.10)

Sun, Y. P., Sakino, K, Aklan, A . M. and Ikenono, Y. : "Ultimate strength and Deformability of RC Column
Retrofitted by Square Steel Tube", Proceedings of the 2nd International RILEM/CSIRO/ACRA
Conference, September 21-23, 1998, pp.619-632

6.11)

NEHRP Provisions (1994 Edition), NEHRP Recommended Provisions For Seismic Regulations For New
Buildings, Part 1 – Provisions, Building Seismic Safety Council, Washington, D.C. 1994

7.1)

U.S.-Japan Planning Groups Joint Planning Workshop: Recommendations for U.S.-Japan Cooperative
Research Program -Phase 5 Composite and Hybrid Structures, UMCEE92-29, 1992.11.

7.2)

Housing Guidance Division of Ministry of Land, Infrastructure and Transport (MLIT), Japan Building
Officials Conference, Building Center of Japan: "Technical Standard Manuals for Structural Regulationsof Buildings", March 15, 2001 (in Japanese)

7.3)

Building Standard Law of Japan- 65 -

Table 1.1 High Priority Research Topics Recommended in the Joint Planning Workshop [1.5]
HIGH PRIORITY RESEARCH TOPICS
Experimental and Analytical Basic Studies
Research
Topic

Interation
Materials

ofExperimental
Subassemblies

Scale effects;
Components

Materials

Rate of

2-D

3-D

Loading, etc.

Experimental Structures
Quasi-

Shaking

static

Table

Field

Analytical Studies ofStructural Bodies
Modeling

Parametricofand Design

Behavior

Studies

Design Studies
DesignDesignStudies

Guidelines

Beam-Columns:
Effect ofConfinement vs.

−

E

E

−

−

−

−

−

−

E

E

−

−

−

E

E

E

−

−

−

−

−

E

E

−

M

−

E

E

E

−

−

−

−

−

E

E

−

M

M

M

M

M

−

−

−

−

−

−

−

−

M

−

−

E

−

E

?

−

−

−

E

E

E

E

−

−

−

−

−

−

−

−

−

M

M

M

M

−

−

−

−

−

−

L

?

?

L

L

L

L

−

−

−

−

−

−

L

?

?

L

L

L

L

E

E

E

−

−

−

−

−

−

−

−

−

M

−

E or M

E or M

−

−

−

−

−

−

−

−

−

−

Composite
Method of
Evaluation
Bond and Shear
Transfer
Columns:
Creep, Shrinkage(High Rise)
Connec tions:
Force & Moment
Transfer
Panel Zone
behavior
Frames:Analytical Studies
Combining Above
Analysis ofBraced Frames
Very HighStrength Concrete
Literature Survey

E, M and L identify topics which should be done early, intermediate or late in the proposed five year research program, respectively.
Some categories are noted with a (?), and these identify areas which are desirable if funds are available.

- 66 -

Table 1.2 Medium Priority Research Topics Recommended in the Joint Planning Workshop [1.5]
MEDIUM PRIORITY RESEARCH TOPICS
Research
Topic

Experimental and Analytical Basic Studies
Materials Interation Components Scale effects;ofRate ofMaterials
Loaidng, etc.

Experimental Experimental Structures Analytial Studies ofSubassemblies
Strucrtural Bodies
2-D

3-D

Quasi- Shaking Field
static Table

Design Studies

Modeling Parametric Design
Desingofand Design Studies Guidelines
Behavior
Studies

Columns:
Buckling

M

M

M

-

-

-

-

-

-

M

M

-

M

Non-Traditional
Connection Design

*

*

*

*

*

-

-

-

-

*

*

*

-

Placement ofConcrete & As
Built vs. Design

M

-

M

-

?

-

-

-

-

-

-

-

M

Construction
Loads

-

-

-

-

-

-

-

-

-

M

-

-

M

-

-

-

-

-

-

-

-

-

L

L

L

L

Construction:

Frames:Composite,
Confinement only,or BothSome categories are noted with a (*), and these identify areas which can not be classified into H, M or L in the workshop.

- 67 -

Table 1.3 Test Program of Centrally and Eccentrically Loaded Stub Columns

Shape

Steel* 1

400

□

590

780

400

◯

590

780

Rank

B, D
(mm)

FA
FC
FD
FA
FC
FD
FA
FC
FD
FA
FC
FD
FA
FC
FD
FA
FC
FD

148
216
324
144
211
318
120
175
264
149
300
450
122
238
360
108
222
336

t
(mm)
4.38

6.36

6.47

2.96

4.54

6.47

B/t
D/t

Concrete

34
49
74
23
33
50
19
27
41
50
101
152
27
52
79
17
34
52

0 (void)
Fc20
Fc40
Fc80
0 (void)
Fc20
Fc40
Fc80
0 (void)
Fc20
Fc40
Fc80
0 (void)
Fc20
Fc40
Fc80
0 (void)
Fc20
Fc40
Fc80
0 (void)
Fc20
Fc40
Fc80

Number of
specimens
CLSC
ELSC
23*2

11

23*2

11

23*2

10

15

11

15

11

15

11

CLSC: centrally-loaded stub column
ELSC: eccentrically-loaded stub column*1 Nominal tensile strength of steel, σ u , in MPa.
*2 Eight specimens among 23 are additionally conducted specimens, so they have different B, D and t fromstandard ones. Detailed dimensions of these specimens are in Table 2.1 and Table 2.2(b).

- 68 -

Table 1.4 Test Program of CFT Beam-Columns

Shape

Steel* 1
400

□

590
780
400

◯

590
780

Rank
FA
FC
FA
FC
FA
FC
FA
FC
FA
FC
FA
FC

B, D
(mm)
210
210
180
240
240
160

t
(mm)
5.80
4.50
8.83
5.95
9.45
6.66
4.70
9.00
4.52
9.12
4.76

B/t
D/t
35
47
23
35
19
27
51
27
53
17
34

*1 Nominal tensile strength of steel, σ u , in MPa.
*2 Two specimens among eight are subjected to biaxial bending.
*3 One specimen among eight is subjected to biaxial bending.

- 69 -

Concrete
Fc40
Fc90
Fc40
Fc90
Fc40
Fc90
Fc40
Fc90
Fc40
Fc90
Fc40
Fc90

N/N0

Number of
specimens

0.4

4

0.4
variable
0.4
variable

8*2

0.4

2

0.4
variable
0.4
variable

8*3

65

Table 1.5 Test Program of Beam-to-Column Connections

Connection
type

╋
┣

Tube
shape

Diaphragm
type□

Through
Outer

◯

Ring

□
◯

Through

Panel
Steel* 1
590
780
590
590
780
590

Concrete

Section

B/t
D/t

Fc40
Fc90

□-250x4.5

55

Fc40
Fc90

◯-280x4.560□-160x3.055◯-180x3.0

60

Fc90

N/N0
0.2
0.2
0.2
0.2
0.2
variable*1 Nominal tensile strength of steel, σ u , in MPa.
*2 One specimen among three is a three-dimensional beam-to-column connection specimen.
*3 One of the two specimens was loaded constant axial force in spite of exterior joint.

- 70 -

Number of
specimens
3*2
1
1
2
1
2*3
1

Table 2.1 Results of Axial Compressive Tests of Steel Tubes Used for CFT Specimens- 71 -

Table 2.2(a) Experimental Variables and Test Results (Circular CFT)

- 72 -

Table 2.2(b) Experimental Variables and Test Results (Square CFT)

- 73 -

Table 2.3 Specific Values for the Stress vs. Strain Models for Concrete- 74 -

Table 2.4 Specific Values for the Stress vs. Strain Models for Steel Tube in Square CFT Columns- 75 -

Table 3.1(a) Summary of Eccentrically Loaded Stub Columnn Specimens (Circular Tubes)
specimen

D

t
(mm)

D/t

σ sy
fc'
(MPa) (MPa)N/NoLoading
condition(mm)
EC4-A-4-035
150
50.7
39.9 0.35
EC4-A-4-06
0.59
EC4-C-2-035
0.34
24.5
EC4-C-2-06
0.59
EC4-C-4-03
0.30
2.96
283
39.9 0.40
EC4-C-4-04
300
101.4
EC4-C-4-06
0.61
EC4-C-8-045
0.45
77.6
EC4-C-8-06
0.60
EC4-D-4-04
0.40
450
152.0
39.9
EC4-D-4-06
0.60
EC6-A-4-02
0.20
122
26.9
39.9
EC6-A-4-06
0.60
EC6-C-2-03
0.30
24.5
EC6-C-2-06
0.60
EC6-C-4-025
0.25
39.9
239
4.54
52.6
579
EC6-C-4-03
0.30 Fig. 3.2(a)EC6-C-4-06
0.60
EC6-C-8-03
0.30
77.6
EC6-C-8-06
0.60
EC6-D-4-03
0.30
360
79.3
39.9
EC6-D-4-06
0.60
EC8-A-4-015
0.15
108
16.7
39.9
EC8-A-4-06
0.60
EC8-C-2-06
0.60
24.5
EC8-C-2-08
0.80
EC8-C-4-015
0.15
6.47
835
39.9 0.30
EC8-C-4-03
222
34.3
EC8-C-4-06
0.60
EC8-C-8-06
0.60
77.6
EC8-C-8-07
0.70
EC8-D-4-015
0.15
336
51.9
39.9
EC8-D-4-045
0.45
Note:
D
= diameter of steel tubet
= wall thickness of steel tube
σ sy = yield strength of steel (yield strength was obtained by 0.2% offset.)fc'= cylinder strength of concreteN= applied axial forceNo= nominal squash load①：Eccentric axial loading②：Circular or Rectangular section③：Nominal tensile strength of steel tube: 4 - 400MPa, 6 - 590MPa, 8 - 780MPa④：Rank of D/t (B/t) ratio concerning deformation capacity of hollow steel tube: Rank-FA, FC, FD⑤：Design compressive strength of concrete: 2 - 20MPa, 4 - 40MPa, 8 - 80MPa⑥：Axial force ratio： 19 - N/No = 0.19

- 76 -

Table 3.1(b) Summary of Eccentrically Loaded Stub Column Specimens (Square Tubes)
Specimen

t
B/t
(mm)
(mm)
149
33.9
148
215
49.1
214
48.9
49.2
215
4.38
49.0

B

σ sy
(MPa)

fc'
(MPa)

N/No

Loading
condition

ER4-A-4-19
0.19
41.1
ER4-A-4-57
0.57
ER4-C-2-25
0.25
25.4
ER4-C-2-56
0.56
ER4-C-4-21
0.21
262
41.1
ER4-C-4-38
0.38
ER4-C-4-51
0.51 Fig. 3.2(b)ER4-C-8-33
214
48.9
0.33
80.3
ER4-C-8-46
215
49.0
0.46
ER4-D-4-27
0.27
323
73.7
41.1
ER4-D-4-60
0.60
ER6-A-4-22
0.22
144
22.7
41.1
ER6-A-4-61
0.61
ER6-C-2-25
211
33.2
0.25 Fig. 3.2(c)25.4
ER6-C-2-58
33.1
0.58
210
ER6-C-4-18
33.0
0.18
41.1
6.36
618
ER6-C-4-44
33.1
0.44
ER6-C-4-57
209
32.9
0.57 Fig. 3.2(b)ER6-C-8-24
33.1
0.24
210
80.3
ER6-C-8-54
33.0
0.54
ER6-D-4-23
50.2
0.23
319
41.1
ER6-D-4-47
50.1
0.47
ER8-A-4-08
121
18.7
40.5
0.08
ER8-C-2-38
27.0
0.38
25.4
ER8-C-2-57
27.1
0.57
175
ER8-C-4-24
27.0
0.24
40.5
ER8-C-4-38
27.1
0.38 Fig. 3.2(c)6.47
835
ER8-C-4-57
27.2
0.57
176
ER8-C-8-39
27.3
0.39
77.0
ER8-C-8-58
175
27.1
0.58
ER8-D-4-40
0.40
265
40.9
40.5
ER8-D-4-60
0.60
Note: B= width of steel tubet
= wall thickness of steel tube
σ sy = yield strength of steel (yield strength was obtained by 0.2% offset.)
fc’ = cylinder strength of concreteN= applied axial forceNo= nominal squash load①：Eccentric axial loading②：Circular or Rectangular section③：Nominal tensile strength of steel tube: 4 - 400MPa, 6 - 590MPa, 8 - 780MPa④：Rank of D/t (B/t) ratio concerning deformation capacity of hollow steel tube: Rank-FA, FC, FD⑤：Design compressive strength of concrete: 2 - 20MPa, 4 - 40MPa, 8 - 80MPa⑥：Axial force ratio： 19 - N/No = 0.19

- 77 -

Table 3.2(a) Material Properties of Structural Steel Tubes (Circular Tubes)
t
Es
σ sy
σ st
ε
(mm) (MPa) (MPa) σ sy /σ st (GPa)(%)
400MPa 2.96
283
408
0.693
224
29.1
590MPa 4.54
579
646
0.895
228
15.2
780MPa 6.47
834
879
0.949
218
10.1
Note: σ st = tensile strength, Es = Modulus of elasticity, ε =Elongation
Table 3.2(b) Material Properties of Structural Steel Tubes (Square Tubes)
t
(mm)
4.38
6.36
6.47

400MPa590MPa780MPa

Es
σ sy
σ st
σ sy /σ st (GPa)(MPa) (MPa)
262
411
0.637
214
618
673
0.918
219
834
879
0.949
218

ε
(%)
32.5
15.1
10.1

Table 3.3(a) Material Properties of Concrete (Circular Tubes)
Fc
fc' (MPa)
E c (GPa)εB (%)
20MPa
25.4
25.4
0.206
40MPa
40.7
32.1
0.220
80MPa
78.1 - 85.1 36.3 - 38.5 0.281 - 0.288
Note: fc' = compressive Strength, E c = modulus of elasticity
εB = strain at the compressive strength
Table 3.3(b) Material Properties of Concrete (Square Tubes)
Fc
20MPa
40MPa
80MPa

fc' (MPa)

E c (GPa)εB (%)

25.4
25.1 - 26.0 0.205 - 0.207
40.5 - 41.1 29.3 - 33.4 0.219 - 0.223
77.0 - 80.3 35.9 - 37.3 0.280 - 0.304- 78 -

Table 3.4(a) Experimental and Analytical Results (Circular Tubes)

Specimen

Mu of test

Mcal.1

Mcal.2

Mcal.3(kN･m)(kN･m)(kN･m)(kN･m)Mu
M cal.1

Mu
M cal.2

Mu
M cal .3

EC4-A-4-035
32.0
28.0
27.6
29.9
1.14
1.16
1.07
EC4-A-4-06
36.5
23.4
22.7
25.1
1.14
1.18
1.06
EC4-C-2-035
128.3
125.9 119.0 128.9
1.02
1.08
1.00
EC4-C-2-06
109.7
108.2 93.8 104.2
1.01
1.17
1.05
EC4-C-4-03
150.0
155.4 147.4 157.0
0.97
1.02
0.96
EC4-C-4-04
156.6
158.2 146.6 154.9
0.99
1.07
1.01
EC4-C-4-06
130.5
138.8 116.4 122.5
0.94
1.12
1.07
EC4-C-8-045
194.1
237.8 215.1 218.3
0.82
0.90
0.89
EC4-C-8-06
160.8
218.8 181.6 173.7
0.74
0.89
0.93
EC4-D-4-04
409.7
457.5 406.1 425.2
0.90
1.01
0.96
EC4-D-4-06
346.8
414.1 312.1 316.5
0.84
1.11
1.10
EC6-A-4-02(46.9)
40.8
40.8
46.4
1.15
1.15
1.01
EC6-A-4-06
42.2
28.6
28.4
37.3
1.48
1.49
1.13
EC6-C-2-03(156.9)164.8 161.4 185.1
0.95
0.97
0.85
EC6-C-2-06
132.3
121.1 114.2 149.1
1.09
1.16
0.89
EC6-C-4-025(173.8)184.4 180.4 197.3
EC6-C-4-03(177.6)183.2 178.2 195.2
0.97
1.00
0.91
EC6-C-4-06
155.4
140.3 129.6 153.7
1.11
1.20
1.01
EC6-C-8-03
217.3
222.6 215.6 229.3
0.98
1.01
0.95
EC6-C-8-06
178.2
182.6 165.1 175.8
0.98
1.08
1.01
EC6-D-4-03(460.5)474.5 451.8 488.8
EC6-D-4-06
399.6
379.3 322.6 369.0
1.05
1.24
1.08
EC8-A-4-015(48.2)
58.0
58.0
65.1
EC8-A-4-06
53.8
37.3
37.4
50.4
1.44
1.44
1.07
EC8-C-2-06(219.8)176.8 171.9 236.3
1.24
1.28
0.93
EC8-C-2-08
149.3
96.3
90.0 170.9
1.55
1.66
0.87
EC8-C-4-015
301.7
280.8 278.2 314.9
1.07
1.08
0.96
EC8-C-4-03
282.3
269.5 265.5 306.5
1.05
1.06
0.92
EC8-C-4-06
261.3
192.2 184.2 247.2
1.36
1.42
1.06
EC8-C-8-06(267.0)230.2 216.3 254.3
EC8-C-8-07
262.0
184.9 168.0 207.5
1.42
1.56
1.26
EC8-D-4-015(604.1)
695.9 682.9 761.5
EC8-D-4-045
648.1
618.3 582.9 685.3
1.05
1.11
0.95
The values in ( ) and ( ) mean the experimental ultimate moment of specimens failed
in cracking in welded portion, before and after the value of φD reaching 2.5%,
respectively.
M cal.1 = theoretical ultimate moment (cylinder strength was used as concrete strength)
M cal.2 = theoretical ultimate moment (scale effect on concrete strength was considered)
M cal.3 = analytical ultimate moment by the fiber analysis- 79 -

Errata
Table 3.4(b) Experimental and Analytical Results (Square Tubes)

Specimen

Mu of test

Mcal.1

Mcal.2

Mcal.3(kN･m)(kN･m)(kN･m)(kN･m)Mu
M cal .1

Mu
M cal .2

Mu
M cal .3

ER4-A-4-19
53.9
48.2
47.7
50.3
1.12
1.13
1.07
ER4-A-4-57
38.9
40.5
38.6
38.6
0.96
1.01
1.01
ER4-C-2-25
101.4
101.9 99.4
99.4
1.00
1.02
1.02
ER4-C-2-56
69.0
83.3
76.4
67.9
0.83
0.90
1.02
ER4-C-4-21
115.4
115.7 113.0 114.5
1.00
1.02
1.01
ER4-C-4-38
103.2
118.0 112.4 108.2
0.87
0.92
0.95
ER4-C-4-51
83.0
109.3 101.0 87.4
0.76
0.82
0.95
ER4-C-8-33
146.3
160.0 153.1 151.0
0.91
0.96
0.97
ER4-C-8-46
121.4
159.0 147.6 133.7
0.76
0.82
0.91
ER4-D-4-27
296.8
322.2 305.0 285.4
0.92
0.97
1.04
ER4-D-4-60
200.9
283.8 226.6 184.2
0.71
0.89
1.09
ER6-A-4-22
124.0
119.9 119.3 124.8
1.03
1.04
0.99
ER6-A-4-61
83.7
75.5
73.5
79.6
1.11
1.14
1.05
ER6-C-2-25
253.0
256.9 253.5 240.5
0.99
1.00
1.06
ER6-C-2-58
145.7
175.5 166.5 142.9
0.83
0.88
1.02
ER6-C-4-18(257.4)279.1 275.4 268.1
ER6-C-4-44
214.2
244.1 235.4 207.3
0.88
0.91
1.03
ER6-C-4-57
164.3
203.9 192.1 161.6
0.81
0.86
1.02
ER6-C-8-24(299.4)316.6 309.8 301.9
0.95
0.97
0.99
ER6-C-8-54
206.3
263.5 245.3 206.2
0.78
0.84
1.00
ER6-D-4-23(592.9)711.9 688.6 587.2
ER6-D-4-47
407.7
632.7 581.8 400.3
0.64
0.70
1.02
ER8-A-4-08(103.0)102.6 110.0 113.0
ER8-C-2-38(203.0)205.1 203.0 172.0
0.99
1.00
1.18
ER8-C-2-57
157.9
143.5 139.7 114.4
1.10
1.13
1.38
ER8-C-4-24(211.8)240.4 238.3 222.2
ER8-C-4-38(212.8)214.9 210.7 181.9
0.99
1.01
1.17
ER8-C-4-57
153.0
157.7 151.5 121.4
0.97
1.01
1.26
ER8-C-8-39(235.4)245.2 240.2 210.2
0.96
0.98
1.12
ER8-C-8-58
182.4
190.0 180.6 140.3
0.96
1.01
1.30
ER8-D-4-40
446.2
544.1 531.2 362.8
0.82
0.84
1.23
ER8-D-4-60
319.7
415.2 394.7 236.8
0.77
0.81
1.35
The values in () and ( ) mean the experimental ultimate moment of specimens failed
in cracking in welded portion, before and after the maximum moment, respectively.
M cal.1 = theoretical ultimate moment (cylinder strength was used as concrete strength)
M cal.2 = theoretical ultimate moment (scale effect on concrete strength was considered)
M cal.3 = analytical ultimate moment by the fiber analysis
The values in red ink are corrected on Oct. 7, 2004.

- 80 -

Table 4.1 Summary of Beam-Column Specimens
Steel
Concrete
Depth Thickness Strength
Strength
SectionSpecimenDt
Axial Load
Shape
σu
Fc
(mm)
(mm)
(MPa)
(MPa)
SC4-A-4-C
40
400
240
4.5
SC4-A-9-C
90
0.4No
SC6-A-4-C
40
SC6-A-9-C
9.0
90
SC6-A-9-V
0.7No 〜-0.3Nso
590
240
SC6-C-4-C
40
0.4No
CircularSC6-C-9-C
4.5
90
SC6-C-9-V
0.7No 〜-0.3Nso
SC8-A-4-C
40
0.4No
SC8-A-9-C
9.0
SC8-A-9-V
780
160
0.7No 〜-0.3Nso
90
SC8-C-9-C
0.4No
4.5
SC8-C-9-V
0.7No 〜-0.3Nso
SR4-A-4-C
40
6.0
SR4-A-9-C
90
400
210
SR4-C-4-C
40
4.5
0.4No
SR4-C-9-C
90
SR6-A-4-C
40
SR6-A-9-C
9.0
90
SR6-A-9-V
0.7No 〜-0.3Nso
590
210
SR6-C-4-C
40
Square0.4No
SR6-C-9-C
6.0
90
SR6-C-9-V
0.7No 〜-0.3Nso
SR8-A-4-C
40
0.4No
SR8-A-9-C
9.0
90
SR8-A-9-V
0.7No 〜-0.3Nso
780
180
SR8-C-4-C
40
0.4No
SR8-C-9-C
6.0
90
SR8-C-9-V
0.7No 〜-0.3Nso
Steel
Concrete
Depth Thickness Strength
Section Strength
Axial Loading
Specimen
D
t
Shape
σu
Fc
Load Angle
(mm)
(mm)
(MPa)
(MPa)
SR6-A-9-C-45
9.0
590
210
45°
SR6-C-9-C-45
6.0
Square90
0.4No
SR8-A-9-C-45
780
180
9.0
SR8-A-9-C-22.5
22.5°
S R 6-A-9-C-45
D：Column Depth (Diameter)① ② ③④⑤ ⑥ ⑦
t：Thickness of Steel Tube①：Beam-Column test②：Shape of Section, R-Square section, C-Circular section③：Steel Strength, 4-400, 6-590, 8-780(MPa)④：Width (Diameter)-to-thickness ratio, A-FA Class, C-FC Class⑤：Concrete Strength, 4-Fc40, 9-Fc90⑥：Axial Load Ratio：C-Constant(0.4No), V-Variable(0.7No〜-0.3Nso）⑦：Loading Angle, 45-45degree, 22.5-22.5degree- 81 -

Table 4.2 Material Properties of Structural Steel Tubes

Coupon Test
Comp.
Steel Thickness Yield
Yield
TensileSectionYield Ｅｌｏｎｇａｔｉｏｎ Strength
Strength
Strength Strength
Shape
Ratio
εu
ｔ
σsyc
σu
σsy
σst
σsy/σst
（％）
（
ｍｍ）
（
Ｍ
Ｐａ）
（
Ｍ
Ｐ
ａ
）
（
Ｍ
Ｐ
ａ
）
（ＭＰａ）
400
Circular

590
780
400

Square

590
780

4.71
4.52
9.00
4.76
9.12
4.50
5.84
5.95
8.83
6.66
9.45

284
504
482
771
820
276
295
540
537
824
825

449
662
618
788
833
412
434
669
673
851
865

0.633
0.761
0.780
0.979
0.984
0.669
0.679
0.808
0.797
0.968
0.954

24.7
23.6
24.0
10.0
11.7
29.0
29.0
12.9
14.1
11.0
11.9

338
530
508
785
806
326
323
609
588
805
837

Table 4.3 Material Properties of Concrete

DesignStrength
Fc
（ＭＰａ）
40
90

CylinderStrengthfc'（ＭＰａ）
35.5〜42.4
84.5〜94.5

Young's
ModulusＥc（ＧＰａ）
31.3〜33.8
36.2〜38.6- 82 -

Poisson'sRatioν
0.18〜0.23
0.21〜0.24

TensileStrengthσＴ（ＭＰａ）
3.12〜3.26
4.85〜5.01

Table 4.4 Summary of Test Results

Specimen
SC4-A-4-C
SC4-A-9-C
SC6-A-4-C
SC6-A-9-C

Ｄ or B
(ｍｍ)

ｔ
(ｍｍ)

241
238

4.70

241

9.00

SC6-A-9-V
SC6-C-4-C
SC6-C-9-C

238
240

SC6-C-9-V

241

SC8-A-4-C
SC8-A-9-C

4.52

Ｄ/ｔor
B/t
51.3
50.6
26.7
26.8

α

σscy
(ＭＰａ)

0.08

338

0.07

508

SC8-C-9-V
SR4-A-4-C
SR4-A-9-C
SR4-C-4-C
SR4-C-9-C
SR6-A-4-C
SR6-A-9-C

N
N
or
No
N so

Ｎ
(ｋＮ)

39.2
88.2
35.5
84.4

0.37
0.38

1034
1784
1809
2567
4682
-947
1462
2086
3874
-491
1612
1988
3457
-1026
1347
2359
-513
1162
1895
1021
1791
1959
2649
4880
-1070
1545
2368
4326
-769
2576
3077
5294
-1569
2003
2540
4437
1142
2644
2358
2965
2782

26.7

91.7

52.6
53.1

35.5
84.4

0.14

530

53.2

91.7
35.5

161

9.12

17.7

0.07

806

0.12

785

SC8-A-9-V
SC8-C-9-C

fc'
(ＭＰａ)

160
159

93.9

33.6
4.76

33.4

210

5.80

36.2

1.37

323

210
209

4.50

46.7
46.4

1.71
1.70

326

23.9

1.22

211
8.83

588

93.9
39.2
88.2
39.2
88.2
39.3
88.3

0.38
0.67
(-0.28)
0.45
0.39
0.68
(-0.27)0.41
0.71(-0.30)0.39
0.69(-0.29)
0.40
0.39
0.40
0.38

SR6-A-9-V

210

23.8

1.21

91.7

0.70
(-0.29)SR6-C-4-C
SR6-C-9-C

211
210

35.5
35.3

1.82
1.81

39.3
93.7

0.38

SR6-C-9-V

212

35.6

1.82

91.7

SR8-A-4-C
SR8-A-9-C

178
179

1.19
1.20

42.3

SR8-A-9-V

178

SR8-C-4-C
SR8-C-9-C

5.95

9.45

18.9

837

1.19

180

6.66

27.0

805
1.71

210
211
181
180

8.83
5.95
9.45

23.8
35.5
19.2
19.0

1.21
1.82
1.21
1.20

94.5
42.3

1.70

SR8-C-9-V
SR6-A-9-C-45
SR6-C-9-C-45
SR8-A-9-C-45
SR8-A-9-C-22.5

609

588
609
837
837

94.5
88.3
88.3
91.7
84.4

α = B / t σ scｙ/ E s , α = (D / t )⋅ (σ scｙ/ E s ) , Es: Young's Modulus of Steel- 83 -

0.69(-0.30)
0.43
0.42
0.72
(-0.32)
0.42
0.41
0.71
(-0.31)
0.38
0.39
0.40
0.39

Ｍue
Ｍuc
(ｋＮ・ｍ) (ｋＮ・ｍ)
173
202
371
422
274
-323
220
254
153
-185
245
261
173
-147
151
110
-82
187
225
151
202
373
402
259
-302
263
295
163
-226
345
377
217
-280
240
264
146
-210
374
270
371
390

128
174
255
313
223
-234
151
218
168
-132
153
175
109
-156
112
75
-84
145
197
128
177
306
363
243
-276
244
303
217
-209
275
322
182
-278
217
262
149
-202
343
284
326
324

M ue
M uc

Ｍua
(ｋＮ・ｍ)

M ue
M ua

1.35
1.16
1.45
1.35
1.23
1.38
1.45
1.16
0.91
1.40
1.60
1.49
1.59
0.94
1.35
1.48
0.97
1.29
1.14
1.18
1.14
1.22
1.11
1.07
1.09
1.08
0.97
0.75
1.08
1.25
1.17
1.19
1.01
1.11
1.01
0.98
1.04
1.09
0.95
1.14
1.20

176
201
387
408
310
-239
225
266
203
-131
223
233
145
-163
143
96
-83
183
222
146
185
387
423
246
-306
271
314
151
-230
344
375
205
-291
252
287
139
-220
418
295
392
390

0.98
1.00
0.96
1.03
0.88
1.35
0.98
0.95
0.75
1.41
1.10
1.12
1.19
0.90
1.06
1.15
0.99
1.02
1.01
1.03
1.09
0.96
0.95
1.05
0.99
0.97
0.94
1.08
0.98
1.00
1.01
1.06
0.96
0.95
0.92
1.05
0.95
0.89
0.92
0.95
1.00

Table 5.1 Summary of Beam-to-Column Connection Specimens
Rectangular Tube

Specimen
R1

R2

R3

R4

R6

╋ plain

Shape of Specimen

Circular Tube
R5

R5’

C1

┣ plain

C2

solid

C3

C4

╋ plain┣ plainNominal Tensile
Strength of Steel
Tube

590

590

780

590

590

590

590

590

590

780

590

σ u (MPa)
Design ConcreteStrength

90

40

90

90

90

90

90

90

40

90

90

H-250x250x9x12

H-160x160
x12x161000

Fc (MPa)
Beam section(mm)
Beam length(mm)Columnsection(mm)Column length(mm)H-250x250x9x12
3000

3500

H-160x160
x12x161000□-250x12
3000

3000

○-280x12○-280
x9□-160x9

2500

2000

3000

2000

Panel section(mm)□-250x4.5□-160x3○-280x4.5○-180x3

Panel p Ds / p t

55.6

53.3

62.2

60.0

Panel p Ds / B d

1.05

1.11

1.18

1.25

Diaphragm

3050○-180x9

continuing to beam flanges
Com.
0.67 p N oCom. 0.2 p N o
Tens.
0.3 p N so

Axial force ratio

continuing to beam flanges
Com.
0.2 p N oCom. 0.2 p N o

Com.
0.67 p N oTens. 0.3 p N soexterior column : ┣ interior column : ╋Table 5.2 Material Properties of Panel Zone Steel and Filled Concrete
R1

R2

Rectangular Tube
R3
R4
R6

R5,R5’

C1

Circular Tube
C2
C3

C4

Thickness p t (mm)

4.58

4.58

4.72

4.58

4.58

3.08

4.64

4.64

3.09

Diameter p Ds (mm)

248.3

249.5

251.4

237.3

251.5

161.5

280.5

280.0 280.0 179.8

207

207

204

203

207

207

204

204

204

203

Specimen

Steel
Panel

FilledConcrete

Young’s modulus E s(kN/mm2)
Yield point pσ sy

4.78

492

492

756

442

492

513

439

439

730

448

(MPa)
Max. Strength pσ st

656

656

809

616

656

658

641

641

805

640

(MPa)
Elongation (%)

23.4

23.4

18.6

34.4

23.4

23.7

24.0

24.0

16.8

23.9

43.90

35.80

42.22

42.22

44.6

41.36

42.88

34.28

41.1

41.36

0.253

0.217

0.242

0.242

0.255

0.229

0.238

0.207 0.244 0.229

109.7

54.4

102.5

102.5

97.7

99.6

98.4

49.1

Young’s Modulus Ec(kN/mm2)
Poisson’s ratio υComp. Strength fc '(MPa)

- 84 -

94.2

99.6

Table 5.3 (a) Test Results of Rectangular Tube Specimens
R1

R2

R3

R4

R5

R5’

R6

16.90

14.12

15.42

11.70

5.80

5.13

10.04

eQpyo (kN )

113

90.6

186

68.8

40.2

37.5

138

eQpy (kN )

140

106

213

102

53.6

49.5

167

SpecimeneKfexperiment3( x10 kN / rad )eQm + (kN)

226

175

251

188

85.8

90.9

224.6

eQm − (kN)

-216

-171

-251

-186

-88.4

-95.5

-223.3

15.3

14.72

15.03

14.50

5.62
-3.05

5.60
-3.05

12.29

cQpa (kN )

129

105

153

121

46.5

46.5

120

cQpu (kN )

155

126

186

144

55.8

55.8

144

cQca (kN)

326

294

748

504

80.0

142

232

cQby (kN )

243

243

390.2

241

195

194

201

cQdy (kN )

395

395

662

259

222

222

339

1.09

1.01

1.39

0.84

1.15

1.06

1.39

cKf
calculation3( x10 kN / rad )eQpy + / cQpa
eQm + / cQpu

1.45

1.39

1.35

1.31

1.53

1.62

1.56

eQm − / cQpu

-1.39

-1.36

-1.35

-1.29

-1.58

-1.71

-1.55

Table 5.3 (b) Test Results of Circular Tube Specimens
C1

C2

C3

C4

15.96

16.16

13.98

5.45

eQpyo(kN)eQpy(kN)

191

136

170

36.7

208

171

194

46.6

eQm + (kN)

284

228

277

92.6

eQm − (kN)

-281

-224

-272

-94.0

19.88

19.05

14.71

5.73
-3.18

cQpa (kN )

130

107

139

41.9

cQpu (kN )

156

128

167

50.5

cQca (kN)

268

240

358

62.0

cQby (kN )

295

295

406

99.0

SpecimeneKfexperiment3( x10 kN / rad )cKf
calculation3( x10 kN / rad )

238

238

343

??

1.60

1.60

1.40

1.11

eQm + / cQpu

1.82

1.78

1.66

1.83

eQm − / cQpu

-1.80

1.75

1.63

-1.86

cQdy (kN )eQpy + / cQpa

e: experiment c: calculation

K f :elastic stiffness

eQpyo : equivalent to first yield of shear panel eQpy : equivalent to yield of shear panel
cQpa : equivalent to short term allowable shear strength of shear panel by AIJ-SRC Standards
cQpu : equivalent to ultimate shear strength of panel by AIJ-SRC Standards
eQm + : maximum strength in positive loading cycleeQm − : maximum strength in negative loading cycle
cQca : equivalent to short term allowable strength of column
cQby : equivalent to yield strength of steel beam
cQdy : equivalent to yield strength of diaphragm- 85 -

Table 5.4 Virtual Specimens for Numerical Simulation

Rectangular
ColumnCircular
ColumnImaginarySpecimen

ColumnDepth Dc
(mm)

Beam Depth
H
(mm)

H/Dc

Concrete
Strength fc '(MPa)

R3-150
R3-075
R3-125
R3-C300
C3-150C3-075
C3-125
C3-C300

238
238
238
238
268
268
268
268

357
179
298
238
402
201
335
238

1.5
0.75
1.25
1
1.5
0.75
1.25
0.89

102.5
102.5
102.5
29.4
94.2
94.2
94.2
29.4

Steel Yield
Point σ sy
(MPa)
756
756
756
756
730
730
730
730

D’*1
1.42
1.42
1.42
2.67
1.86
1.86
1.86
2.97

*1 D’: Coefficient of Declivity in Concrete Stress-Strain Softening Region in Sakino’s Model

Table 5.5 Results of Numerical Simulation
Specimen

Rectangular
ColumnCircular
ColumnR3-150
R3-075
R3-125
R3-C300
R1
R2
R3
C3-150C3-075
C3-125
C3-C300
C1
C2
C3

H/Dc
1.5
0.75
1.25
1
1
1
1
1.5
0.75
1.25
0.89
0.89
0.89
0.89

D’
1.42
1.42
1.42
2.67
1.04
1.99
1.42
1.86
1.86
1.86
2.97
1.39
2.23
1.86

fc '

σ sy

Qp − max

cQp − maxsQp − max

(MPa)
102.5
102.5
102.5
29.4
109.7
54.4
102.5
94.2
94.2
94.2
29.4
98.4
49.1
94.2

(MPa)
756.0
756.0
756.0
756.0
492
492
756.0
730
730
730
730
439
439
730

(kN)
2510
2839
2611
1780
2344
1721
2670
2648
3049
2743
1995
2363
1855
2950

(kN)
1574
1869
1676
639
1757
1073
1727
1713
2003
1774
774
1910
1214
1974

(kN)
986
1037
1012
1192
697
685
1107
942
1112
1019
1253
545
671
1024

- 86 -

cQp − max/( Ac ⋅ fc ' )
0.27
0.32
0.29
0.38
0.28
0.35
0.30
0.32
0.38
0.33
0.47
0.34
0.44
0.37

Table 6.1 Characteristic Parameters of Concrete Stress Block for Circular CFT Section

Table 6.2 Characteristic Parameters of Concrete Stress Blockfor Square CFT Section

Table 6.3 Relationships Between Ductility Grades and Limit Rotation Angles of Beam-Columns- 87 -

Table 6.4 Summary of Database on Stiffness Degrading Ratio ay

Circular CFT

Square CFT

US-Japan

Others

US-Japan

OthersNumber of Date

13

45

33

108

D / t ( B / t)

17.7~53.2

20.4~77.0

18.9~46.7

15.6~70.0

σ sy [MPa]284~819283~549276~824194~642fc' [MPa]35.5~93.9

28.0~84.9

39.2~94.5

20.0~101.7

N / N00.39~0.70

0.0~0.71

0.40~0.72

0.0~0.83

a / D (a / B )3.0

1.75~5.2

3.0

1.5~5.67- 88 -

Table 7.1 Investigation Items and Material Properties of Each Theme Structure

Table 7.2 Design Loads

Table 7.3 Gravity and Seismic Loads of 40-Story Frames

Table 7.4 List of Members for 40-Story Frames- 89 -

Table 7.5 Input Seismic Ground Motions for 24-Story Frames

Table 7.6 Hysteresis Models for 24-Story Frames

Table 7.7 Weight and Stiffness of 40-Story Frames

Table 7.8 Story Drift of 40-Story Frames- 90 -

Table 7.9 First Natural Period

Table 7.10 Absorbed Energy

Table 7.11 Results of Response Analyses

Table 7.12 Cost Estimation of CFT and Steel Frames- 91 -

Member Level

System Level

Steel Tube + Concrete

Cement + Fiber

Core RC Wall + Exterior Steel Frame

Frame Level

1-D to 2-D

2-D to 3-D

RC Column + Steel Beam

Figure 1.1 Classification of Composite and Hybrid Structures- 92 -

Figure 1.2 Typical CFT Column System and Concrete Filling- 93 -

Japan side

U.S. side

Ministry ofConstruction
Building
Contractors
Society

Building
Center ofJapan
Japan Structural
Consultants
Association

National
ScienceFoundation

Universities
research cooperation
Domestic
Cooperative
Research

Building
Research
Institute

Joint Technical
Coordinating Committee(JTCC)
Joint Steering Committee(JSC)

Kozai ClubJapan TCC
Japan Technical
Coordinating Committee

U.S. TCC
U.S. Technical
Coordinating Committee

Japan TSC

U.S. TSC

Joint Technical
Sub-Committees(JTSC)

Figure 1.3 Overall Research Operation System of the US -Japan Cooperative Research Programon Composite and Hybrid Structures- 94 -

３Ｄ

３Ｄ

Ｄ

(a)

(b)

(c)

(d)
Figure 1.4 Element Tests on CFT Column System- 95 -

CFT-Column and Steel Beam

6.400
Steel Beam
6.400

CFT-Column

9.600

35.200

Steel Beam

6.400
6.400
6.400

6.400

6.400

6.400

6.400

38.400Figure 1.5 Key Plan of CFT Theme Structure- 96 -

6.400

Figure 2.1 Test Setup Used for Centrally Loaded Stub Columns- 97 -

Figure 2.2 Sizes of Specimens

Figure 2.3 Scale Effect on Compressive Strengths of Circular Plain Concrete Columns- 98 -

Figure 2.4 Stress States of Steel Tube and Concrete

Figure 2.5 Experimental Axial Compressive Strengths of Circular CFT Columns- 99 -

Figure 2.6 Comparisons Between Experimental and Calculated Strengths (Circular CFT
Columns)

Figure 2.7 Compressive Strength Factors of Hollow Square Steel Tubes- 100 -

Figure 2.8 Comparisons Between Experimental and Calculated Strengths (Square CFT Columns)

Figure 2.9 Stress vs. Strain Models for Concrete in CFT Column- 101 -

Figure 2.10 Stress vs. Strain Model for Steel Tube in Circular CFT Column

Figure 2.11 Stress vs. Strain Models for Steel Tube in Square CFT Column- 102 -

(a) CC4-A, C, D-4

Circular CFT Columns(b) CC8-A, C, D-8

Circular CFT Columns

Figure 2.12 Comparisons Between Experimental Results and Proposed Models- 103 -

(a) CR4-A, C, D-4

Square CFT Columns(b) CR8-A, C, D-8

Square CFT Columns

Figure 2.13 Comparisons Between Experimental Results and Proposed Models- 104 -

40

40
160 80
400
80

900

50 40
3 00
4 00

300
420

60

132

216
480

132

unit:mm

1 32
240

40

216
480

13 2 40

100

64 8

200
4 80

100

40

60

40

50

40

40

380
420

100100 40
480

Figure 3.1 Dimensions of Eccentrically Loaded Stub Columns

」
(a)

(b)

(c)

Figure 3.2 Loading Condition of Eccentrically Loaded Stub Columns- 105 -

40

M (kN･m)50

N/No =0.35 (EC4-A-4-035)

M (kN･m)N/No=0.20 (EC6-A-4-02) 60

M (kN･m) N/No =0.15 (EC8-A-4-015)

40

30
20

σsy =579MPa
fc' =39.9MPa
D/t =2710
φ･D (%)

0
0

2

4

M (kN･m)6

φ･D (%)

0

2

200

4

6

8

M(kN･m)200

2

4

φ･D (%)
802N/No =0.80 (EC8-C-2-08)
σsy =835MPa
fc' =24.5MPa
D/t =3450
φ･D (%)

φ･D (%)

4

6

8

0

400
N/No =0.25 (EC6-C-4-025)
N/No=0.30 (EC6-C-4-03)

M (kN･m)N/No =0.40 (EC4-C-4-04)
150

8

0

200

M (kN･m)6

N/No =0.60 (EC8-C-2-06)

100

σsy =579MPa
fc' =24.5MPa
D/t =53

50

6

M (kN･m)N/No=0.60 (EC6-C-24-06)

0
0

4

150

N/No =0.60 (EC4-C-2-06) 100

0

2

200

150

σsy =283MPa
fc' =24.5MPa
D/t =1010250

N/No =0.30 (EC6-C-2-03)

100

50

φ･D (%)

08

N/No =0.35 (EC4-C-2-035)

σsy =835MPa
fc' =39.9MPa
D/t =17

20

0

150

N/No =0.60 (EC8-A-4-06)
×:Cracking

20

σsy =283MPa
fc' =39.9MPa
D/t =51

10

N/No=0.60 (EC6-A-4-06) 40

30

N/No=0.60 (EC4-A-4-06)

150

2

4

6

8

M (kN･m)N/No=0.15 (EC8-C-4-015)

300

N/No =0.30 (EC4-C-4-03)
100

N/No =0.60 (EC6-C-4-06)

100
σsy =283MPa
fc' =39.9MPa
D/t =101

50

N/No =0.60 (EC4-C-4-06)

σsy =579MPa
fc' =39.9MPa
D/t =53

50
φ･D (%)

0200

2
M (kN･m)4

68

N/No=0.45 (EC4-C-8-045)

φ･D (%)

0250

2
M (kN･m)4

6

8

N/No =0.30 (EC6-C-8-03)

10050

σsy =579MPa
fc' =77.6MPa
D/t =53

50
φ･D (%)

0
500

2

4

6

8

M (kN･m)400

M (kN･m)0

2
M (kN･m)φ･D (%)

50

0

4

6

8

0
700

N/No =0.30 (EC6-D-4-03)

N/No =0.60 (EC6-D-4-06)

2

4

6

8

2

4

M(kN･m)6

8

N/No =0.45 (EC8-D-4-045)

N/No =0.15 (EC8-D-4-015)

200

400
300

σsy =579MPa
fc' =39.9MPa
D/t =79100

σsy =835MPa
fc' =39.9MPa
D/t =52

200
φ･D (%)00

φ･D (%)

600

300

100

N/No =0.60 (EC8-C-8-06)

σsy =835MPa
fc' =77.6MPa
D/t =34

500

200

8

N/No =0.70 (EC8-C-8-07)

100

N/No =0.40 (EC4-D-4-04) 400

N/No=0.60 (EC4-D-4-06)
σsy =283MPa
fc' =39.9MPa
D/t =152
φ･D (%)

6

0

500

300

4

150

0
0

300

2

200
N/No =0.60 (EC6-C-8-06)

100

σsy =283MPa
fc' =77.6MPa
D/t =101

0

250

150

N/No =0.60 (EC4-C-8-06)

φ･D (%)

0

200

150

σsy =835MPa
fc' =39.9MPa
D/t =34100

00N/No =0.30 (EC8-C-4-03)
N/No =0.60 (EC8-C-4-06)

200

100

φ･D (%)

0
0

2

4

6

8

0

2

Figure 3.3(a) Moment vs. Curvature Relations (Circular Specimens)- 106 -

4

6

8

60

M (kN･m)N/No=0.19 (ER4-A-4-19)

150

M(kN･m)N/No =0.22 (ER6-A-4-22)

120

M (kN･m)N/No=0.08 (ER8-A-4-08)

100

N/No=0.57 (ER4-A-4-57)
40

N/No =0.61 (ER6-A-4-61)

100

80
×:cracking

60
σsy =262MPa
fc' =41.1MPa
B/t =34

20

σsy =618MPa
fc' =41.1MPa
B/t =2350
φ･Ｄ(%)

040
φ･Ｄ(%)

0
0

2

120

4

M(kN･m)6

8

N/No =0.25 (ER4-C-2-25)

100

0
300

2

60

φ･Ｄ(%)

4

6

8

0

2

4

M (kN･m)N/No=0.25 (ER6-C-2-25)

40

N/No =0.58 (ER6-C-2-58)

20

100
φ･Ｄ(%)

0

σsy =618MPa
fc' =25.4MPa
B/t =3350

150

N/No =0.57 (ER8-C-2-57)

100

σsy =835MPa
fc' =25.4MPa
B/t =2750

2120 M(kN･m)

4

6

8

N/No =0.21 (ER4-C-4-21)

100

0
0

300

2

N/No =0.38 (ER4-C-4-38)
N/No=0.51 (ER4-C-4-51)

60

σsy =262MPa
fc' =41.1MPa
B/t =49

40
20

4

6

8

N/No =0.44 (ER6-C-4-44)

200

σsy =618MPa N/No =0.57 (ER6-C-4-57)
fc' =41.1MPa
B/t =33
φ･Ｄ(%)

φ･Ｄ(%)

50

160 M(kN･m)

4

6

8

0

2
M(kN･m)4

6

N/No=0.38 (ER8-C-4-38)
N/No =0.57 (ER8-C-4-57)

100

σsy =835MPa
fc' =40.5MPa
B/t =2750

8N/No =0.24 (ER6-C-8-24)

300

120

0

2

N/No =0.46 (ER4-C-8-46)

80
40
0

N/No =0.54 (ER6-C-8-54)
σsy =618MPa
fc' =80.3MPa
B/t =33.2
φ･Ｄ(%)

50
0

0

2

4

6

8

350 M (kN･m)

0

2

200
150

4

6

50

σsy =618MPa
fc' =41.1MPa
B/t =50

100

2

4

6

8

4

6

8

M(kN･m)N/No=0.40 (ER8-D-4-40)

300
200

σsy =835MPa
fc' =40.5MPa
B/t =41100
φ･Ｄ(%)

002

N/No=0.60 (ER8-D-4-60)

N/No =0.47 (ER6-D-4-47)

200

0

0

400

300

φ･Ｄ(%)

φ･Ｄ(%)

8400

σsy =262MPa
fc' =41.1MPa
B/t =74

100

50

500

500
N/No=0.60 (ER4-D-4-60)

σsy =835MPa
fc' =77.0MPa
B/t =27

N/No =0.23 (ER6-D-4-23)

600

250

100

0

700 M (kN･m)
N/No=0.27 (ER4-D-4-27)300

8N/No=0.58 (ER8-C-8-58)

100
φ･Ｄ(%)

6

150

150

20

4

N/No =0.39 (ER8-C-8-39)

200

200

σsy =262MPa
fc' =80.3MPa
B/t =4960

φ･Ｄ(%)

250 M (kN･m)

250

100

8

0

350

N/No =0.33 (ER4-C-8-33)

140

6

200
150

100

4

N/No =0.24 (ER8-C-4-24)

N/No=0.18 (ER6-C-4-18)

0
2

2

150

0
0

0
250 M (kN･m)

M(kN･m)250

80

φ･Ｄ(%)

φ･Ｄ(%)

008N/No =0.38 (ER8-C-2-38)

200

150
σsy =262MPa
fc' =25.4MPa
B/t =496250 M (kN･m)

200

N/No =0.56 (ER4-C-2-56)

20
0

250

80

σsy =835MPa
fc' =40.5MPa
B/t =19

φ･Ｄ(%)
0

0

2

4

6

8

0

2

Figure 3.3(b) Moment vs. Curvature Relations (Square Specimens)- 107 -

4

6

8

Co mp ressive StressConfined Concrete

Plane Concrete

StrainC ompressive Stress(a) Circular Section

Confined Concrete

Plane Concrete

Strain(b)Square Section

Figure 3.4 Stress vs. Strain Relations of Filled Concrete- 108 -

Compressive Stress

0.91 σ
sy

Et

Es
Es

Tensile Strain

Et

Compressive Strain

1.08 σsy

Tensile Stress(a) Circular Tube

Compressive Stress
S･σ
σ
σ

sy
FA Rank
FC Rank
FD RanksysT
Es

Tensile Strain

Es

SεCompressive Strainεε

sy sy sB

1.1 σ

ε

sT

sy

Tensile Stress(b) Square Tube

Figure 3.5 Stress vs. Strain Relations of Steel Tube

- 109 -

40

M (kN･m)50

M (kN･m)N/No =0.20

40

30

60

30

N/No =0.60
40N/No =0.60
20
σsy =283MPa
fc' =39.9MPa
D/t =51

10

Exp.
Ana.

σsy =579MPa
fc' =39.9MPa
D/t =2710
φ･D (%)

0

Exp.
Ana.

0

2

4

M (kN･m)

6

8

φ･D (%)
0

2

4

M(kN･m)6

8

0N/No =0.30

Exp.
Ana.

Exp.
Ana.

σsy =579MPa
fc' =24.5MPa
D/t =53

50

φ･D (%)
0
200

2

4

6

8

N/No=0.30

8N/No =0.60
N/No =0.80

Exp.
Ana.

σsy =835MPa
fc' =24.5MPa
D/t =3450
φ･D (%)

02
M (kN･m)

4

6

8

φ･D (%)

0

2

4

6

8

N/No=0.15

N/No =0.30

400 M (kN･m)

N/No =0.25
N/No =0.30

200

150

6

0

250

M (kN･m)M (kN･m)

100

0
0

4

150

N/No=0.60

100

σsy =283MPa
fc' =24.5MPa
D/t =101

φ･D (%)

200

150

50

2

250

100
N/No =0.60

Exp.
Ana.

0

200

N/No =0.35

N/No=0.60
σsy =835MPa
fc' =39.9MPa
D/t =17

20

0

150

M (kN･m)N/No =0.15

N/No =0.35

20

80

300

150
N/No =0.40
N/No =0.60

100
σsy =283MPa
fc' =39.9MPa
D/t =10150

200

N/No =0.60
100

Exp.
Ana.

σsy =579MPa
fc' =39.9MPa
D/t =53

50

Exp.
Ana.

φ･D (%)

0100
φ･D (%)

0
0

250

2

4

6

8

M (kN･m)

200

0

2

4

M (kN･m)

150

100

100

6

8

N/No =0.30

0

2

4

300 M (kN･m)

Exp.
Ana.

50

200

σsy =579MPa
fc' =77.6MPa
D/t =53

N/No =0.70

50

500

2

4

6

8

M (kN･m)50
φ･D (%)

N/No =0.60

100

0

2

4

6

8

M (kN･m)800

300

σsy =283MPa
fc' =39.9MPa
D/t =152

4

6

8

M (kN･m)6

σsy =579MPa
fc' =39.9MPa
D/t =79200

8N/No =0.15
N/No=0.45

400

100

Exp.
Ana.

σsy =835MPa
fc' =39.9MPa
D/t =52

200

Exp.
Ana.

φ･D (%)

0
4

2

600

N/No =0.60

400

0
2

0

500

Exp.
Ana.

φ･D (%)
0

φ･D (%)

N/No =0.30

300
200

Exp.
Ana.

0

600
N/No=0.40400

σsy =835MPa
fc' =77.6MPa
D/t =34

100

Exp.
Ana.

0
0

8

N/No=0.60

φ･D (%)
06

N/No =0.60

150
N/No=0.60

φ･D (%)

250

200

150

Exp.
Ana.

0

250
σsy =283MPa
fc' =77.6MPa
D/t =101

N/No=0.45

N/No =0.60
σsy =835MPa
fc' =39.9MPa
D/t =34

φ･D (%)
0

0

2

4

6

8

0

2

Figure 3.6(a) Comparisons Between Experimental and Analytical Results (Circular Tubes)- 110 -

4

6

8

60

M(kN･m)N/No=0.19

150 M (kN･m)

M (kN･m)120

N/No =0.22

100
N/No =0.0940100N/No =0.61

8060

N/No=0.57
σsy =262MPa
fc' =41.1MPa
B/t =34

20

50

Exp.
Ana.

σsy =618MPa
fc' =41.1MPa
B/t =23φ･Ｄ(%)0φ･Ｄ(%)

002

120 M(kN･m)

4

6

8

N/No=0.25

0

250

80

200

N/No =0.56

20

2
M (kN･m)

4

6

8

0

σsy =262MPa
fc' =25.4MPa
B/t =4940

250

N/No =0.25

2

200

4

20

φ･Ｄ(%)

50

0

2

4

120 M(kN･m)

6

0
0

8
300

N/No =0.21

100

2
M (kN･m)

4

6

8

250

N/No =0.51
σsy =262MPa
fc' =41.1MPa
B/t =49

40
20

100

Exp.
Ana.

φ･Ｄ(%)

0160

2
M (kN･m)4

6

8

N/No =0.33

Exp.
Ana.

0

2
M (kN･m)4

N/No =0.46

6

8

40
20

Exp.
Ana.

0

2

4

6

Exp.
Ana.

σsy =618MPa
fc' =80.3MPa
B/t =33.2

50

8

M (kN･m)

350

0

2

N/No =0.60

150

σsy =262MPa
fc' =41.1MPa
B/t =74

100
50

50

4

6

8

N/No =0.47

2

4

6

8

φ･Ｄ(%)
6

8

M (kN･m)N/No =0.40
σsy =835MPa
fc' =40.5MPa
B/t =41

200

Exp.
Ana.

σsy =618MPa
fc' =41.1MPa
B/t =50

100

4

300

N/No =0.60
100

Exp.
Ana. φ･Ｄ(%)

φ･Ｄ(%)

0
0

2

400

200

0

0
500

300

φ･Ｄ(%)

8

Exp.
Ana.

σsy =835MPa
fc' =77.0MPa
B/t =27

φ･Ｄ(%)

400
Exp.
Ana.

6

N/No =0.39

N/No =0.23

500

200

M (kN･m)100

M (kN･m)

600

250

4

0

700
N/No=0.27

300

2

150

0

0

0

N/No=0.58

N/No

100
φ･Ｄ(%)

φ･Ｄ(%)

200

150

σsy =262MPa
fc' =80.3MPa
B/t =4960

σsy =835MPa
fc' =40.5MPa
B/t =27

250

200

80

Exp.
Ana.

N/No=0.5750

N/No=0.24

250

100

N/No=0.24

φ･Ｄ(%)

300

120

8

0

350

140

6

N/No =0.38

0
0

4

100

σsy =618MPa
fc' =41.1MPa
B/t =33

50

M (kN･m)150

N/No=0.57

150

2

200

N/No=0.44

200

60

0
250

N/No =0.18

N/No=0.3880

σsy =835MPa
fc' =25.4MPa
B/t =27
φ･Ｄ(%)

Exp.
Ana.

50

φ･Ｄ(%)

008N/No =0.57

100

Exp.
Ana.

σsy =618MPa
fc' =25.4MPa
B/t =33

6

N/No =0.38

150

N/No=0.58

100

Exp.
Ana.

φ･Ｄ(%)

M (kN･m)

150

60

Exp.
Ana.

0

300

100

σsy =835MPa
fc' =40.5MPa
B/t =19

40

Exp.
Ana.

0
0

2

4

6

8

0

2

Figure 3.6(b) Comparisons Between Experimental and Analytical Results (Square Tubes)- 111 -

4

6

8

σsy

σcp

σsy
Cross-section

Steel

Concrete

1.0○：400MPa◇：590MPa□：780MPaM u /M cal.1

M u /M cal.1

Figure 3.7 Assumed Stress Blocks

0.0

1.0
○：
400MPa◇：590MPa□：780MPa
0.0

0

50

100
D/t Ratio

150

0

20

40
B/t Ratio

60

80

1.0
○：
400MPa◇：590MPa□：780MPaM u /M cal.2

M u /M cal.2(a) Circular Tube(b) Square Tube
Figure 3.8 Values of M u/M cal.1

0.0

1.0
○：400MPa◇：590MPa□：780MPa
0.0

0

50

100

150

0

20

D/t Ratio

40

60

80

B/t Ratio

1.0○：400MPa◇：590MPa□：780MPa

M u /M cal.3

M u /M cal.3(a) Circular Tube(b) Square Tube
Figure 3.9 Values of M u/M cal.2

1.0○：400MPa◇：590MPa□：780MPa
0.0

0.0
0

50

100
D/t Ratio

0

150

20

40
B/t Ratio(a) Circular Tube(b) Square Tube
Figure 3.10 Values of M u/M cal.3- 112 -

60

80

500〜780
Hole

D

t 2t

CFT
ColumnStiffener
6*Dt

500〜780

Footing

780〜1200
Seam
Seam
CircularSpecimen
Seam-+
Loading Direction(unit : mm)

SquareSpecimen

Figure 4.1 Details of Beam-Column Specimens- 113 -

Actuator

Specimen

Actuator

Reaction Wall

Actuator

Reaction Floor(a) Test Setupθθ（ｂ） Ｌｏａｄｉｎｇ Condition

Figure 4.2 Test Setup and Loading Condition for Beam-Column Specimens

0.06

θ

0.05
0.04
0.03
0.02
0.01
0
-0.01
-0.02
-0.03
-0.04

Cycle

-0.05

Figure 4.3 Loading History for Beam-Column Specimens- 114 -

N

No

Q
Variable
Nc=0.7No

N(0.4No)

Nt→Nc

Nc
Comp.

ConstantR

Q Tens.
Nt

Nt(-0.3Nso)-Nso

Figure 4.4 Rules for Axial Force Loading in Beam-Column Tests- 115 -Nc→Nt

ＳＣ４Ａ４Ｃ
400N/mm2
FA Class
Fc40

250

M(kN･m)

θ
0.06

-0.06

250

-0.002
θ

-0.06

0.08
□：L.B.
○：Max.
-250
500-0.006
ＳＣ４Ａ４Ｃ
400N/mm2
FA Class
Fc40

θ

-0.06

0.08

θ
0.08

ε

-250
θ
0.06

-0.06
ＳＣ６Ａ４Ｃ
2
590N/mm
FA Class
Fc40-0.002

□：
L.B.
○：
Max.
-0.010500-0.0060.08

εＳＣ６Ｃ４Ｃ
590N/mm2
FC Class
Fc40

θ
0.06

-0.06
-0.002
θ

-0.06

0.08

ＳＣ６Ｃ４Ｃ
590N/mm2
Fc Class
Fc40-500
300 M(kN･m)
ＳＣ６Ｃ９Ｃ
590N/mm2
FC Class
Fc90

300-0.0100.08

ＳＣ８Ａ４Ｃ
780N/mm2
FA Class
Fc40

θ
0.06

-0.06
ＳＣ８Ａ４Ｃ

θ

-0.06

0.08

780N/mm
FA Class
Fc40-0.0022-0.006□：L.B.
△：Crack○：Max.
-300-0.010ＳＣ６Ｃ９Ｃ
590N/mm2
FC Class
Fc90

‑300

ε
ＳＣ８Ａ９Ｃ
780N/mm2
FA Class
Fc90

300-0.006-0.010

ε
θ
0.06

-0.06

θ
0.08

-300
200

-0.006-0.010εM(kN･m)

θ
0.06

-0.06
-0.002
θ

‑0.06

0.08

‑200

ＳＣ８Ａ９Ｃ -0.002
780N/mm2
FA Class
Fc90□：L.B.
△：Crack○：Max.

ＳＣ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

□：
L.B.
△：
Crack
○：
Max.

ＳＣ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

Figure 4.5 Test Results of Circular Interior Beam-Column Specimens- 116 -

θ
0.06

M(kN･m)

-0.06

ε

ε

θ

□：L.B.
○：Max.

M(kN･m)-0.006ＳＣ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90-0.010-0.002

□：
L.B.
○：
Max.
-300

θ
0.06

-0.06

‑0.06

-0.006

ε

θ

-0.06

-0.010-0.010-0.002

□：
L.B.
○：
Max.

-500

-0.006-0.06□：L.B.
○：Max.

300 M(kN･m)

ＳＣ４Ａ９Ｃ
400N/mm2
FA Class
Fc90

M(kN･m)ＳＣ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90-0.002

θ
0.06

-0.06

-0.06

M(kN･m)ＳＣ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

M(kN･m)ＳＣ４Ａ９Ｃ
400N/mm2
FA Class
Fc90-0.006-0.010ε

ＳＣ６Ａ９Ｖ
590N/mm
FA Class
Fc90

400

M(kN･m)

θ

-0.04

0.06

-400
250

□：L.B.
△：Crack○：Max.

-0.04-0.002

θ
0.06
zＳＣ６Ａ９Ｖ
2
590N/mm
-0.006FA Class
Fc90-0.010M(kN･m)

-0.04
θ
0.06

ε

θ

150

0.06

ＳＣ８Ａ９Ｖ
2
780N/mm
FA Class
Fc90-0.006

□：
L.B.
○：
Max.

□：L.B.
△：Crack○：Max.

θ

□：
L.B.
△：
Crack
○：
Max.

Figure 4.6 Test Results of Circular Exterior Beam-Column Specimens- 117 -

ε

θ

0.06

-150

ＳＣ６Ｃ９Ｖ
2
590N/mm
FC Class
Fc90

0.002

-0.04

-0.04

ε

-0.002-0.010M(kN･m)ＳＣ８Ｃ９Ｖ
2
780N/mm
FC Class
Fc90

0.06

-0.0060.06-200-0.002-0.010

θ
-0.04

-0.04

θ

-0.04

-250

ＳＣ６Ｃ９Ｖ 200
2590N/mm
FC Class
Fc90

0.002

ＳＣ８Ａ９Ｖ
2
780N/mm
FA Class
Fc90

0.002

M(kN･m)

0.002

2

0.06

-0.002-0.006-0.010ε

ＳＣ８Ｃ９Ｖ
2
780N/mm
FC Class
Fc90

ＳＲ４Ａ４Ｃ
2
400N/mm
FA Class
Fc40

300

θ
0.06

-300
300

ＳＲ４Ａ４Ｃ
400N/mm2
FA Class
Fc40‑0.010‑0.030 ε

M(kN･m)

θ
0.06

0.06

-300
500

□：
L.B.
○：
Max.

‑0.010

□：
L.B.
○：
Max.

0.000

-0.06

0.06

ＳＲ６Ａ４Ｃ
2
590N/mm
FA Class
Fc40

0.06

300

‑0.010

ＳＲ６Ａ９Ｃ
590N/mm2
FA Class
Fc90

ＳＲ４Ａ９Ｃ
2
400N/mm
FA Class
Fc90

□：
L.B.
○：
Max.

□：L.B.
○：Max.

‑0.030 ε

M(kN･m)

0.000

θ
0.06

500

ＳＲ４Ｃ９Ｃ
2
400N/mm
FC Class
Fc90

□：
L.B.
○：
Max.
0.000

M(kN･m)‑0.06
ＳＲ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90

0.06

ＳＲ６Ｃ４Ｃ
2
590N/mm
FC Class
Fc40

400‑0.030M(kN･m)

0.000

θ

-0.06

0.06

ＳＲ６Ｃ９Ｃ
590N/mm2
FC Class
Fc90

θ
0.06

‑0.06
ＳＲ６Ｃ４Ｃ
590N/mm2
FC Class
Fc40

-500

‑0.010400‑0.010v‑0.0300.000

M(kN･m)‑0.06

θＳＲ６Ｃ９Ｃ
590N/mm2
FC Class
Fc90

0.06

-400
ＳＲ８Ａ４Ｃ
780N/mm2
FA Class
Fc40

500‑0.030M(kN･m)

□：
L.B.
○：
Max.

0.000

θ
0.06

‑0.06

θ
-0.06

0.06

ＳＲ８Ａ４Ｃ
2
780N/mm
FA Class
Fc40-400
ＳＲ８Ａ９Ｃ
2
780N/mm
FA Class
Fc90‑0.010

500

□：
L.B.
○：
Max.
‑0.0300.000

M(kN･m)‑0.06

θＳＲ８Ａ９Ｃ
2
780N/mm
FA Class
Fc90

0.06

-500
ＳＲ８Ｃ４Ｃ 400
2780N/mm
FC Class
Fc40‑0.030M(kN･m)

0.000

-500
θ

ＳＲ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

‑0.06

θ
-0.06

0.06

-400

□：L.B.
○：Max.

ＳＲ８Ｃ４Ｃ
2
780N/mm ‑0.010
FC Class
Fc40‑0.020

400

θ
0.06

-400

‑0.0300.000

M(kN･m)-0.06

ＳＲ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

□：
L.B.
○：
Max.

Figure 4.7 Test Results of Square Interior Beam-Column Specimens- 118 -□：L.B.
○：Max.

‑0.030

‑0.06

□：
L.B.
○：
Max.

□：
L.B.
○：
Max.

□：L.B.
○：Max.

θ
0.06‑0.010‑0.020‑0.020□：L.B.
○：Max.

θ
0.06‑0.020

□：
L.B.
○：
Max.

-0.06

□：L.B.
○：Max.

‑0.010‑0.020□：L.B.
○：Max.

θ
0.06‑0.020

□：
L.B.
○：
Max.

-0.06

‑0.010‑0.030 εθ□：L.B.
○：Max.

θ
0.06‑0.020

□：
L.B.
○：
Max.

-0.06

□：L.B.
○：Max.

‑0.010‑0.020‑0.020

-500

θ
0.06

‑0.06

-300
θ
0.06

‑0.06

θ

θ

-0.06

‑0.020‑0.030 ε

M(kN･m)

ＳＲ４Ｃ９Ｃ
400N/mm2
FC Class
Fc90

0.000

M(kN･m)
‑0.06

-300

0.000

θ

300

□：
L.B.
○：
Max.

‑0.06
ＳＲ４Ｃ４Ｃ
400N/mm2
FC Class
Fc40

ＳＲ４Ａ９Ｃ
2
400N/mm
FA Class
Fc90

-0.06

‑0.020□：L.B.
○：Max.

-0.06

ＳＲ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

θ
0.06

‑0.06

-0.06

ＳＲ４Ｃ４Ｃ
2
400N/mm
FC Class
Fc40

0.000

M(kN･m)θ
0.06‑0.010‑0.020‑0.030□：L.B.
○：Max.

500
SR6A9C45
45 DegreeM(kN･m)0.08‑0.010

0.08

‑500

500

‑0.030ε‑0.030ε

0.000

M(kN･m)θ

‑0.06

SR8A9C22.5
22.5 Degree‑0.0100.08‑0.010θ

0.08

SR8A9C45
45 Degree‑0.020□：L.B.
△：Crack○：Max.

‑500‑0.020□：L.B.
△：Crack○：Max.

θ‑0.06

SR6C9C45
45 Degree

0.08

500

0.08

SR8A9C45
45 Degree‑0.010

‑400

θ

‑0.06

θ
0.08

θ
‑0.06

0.000

M(kN･m)

0.000
‑0.06

SR6C9C45
45 Degree‑0.020□：L.B.
△：Crack○：Max.

M(kN･m)SR6A9C45
45 Degree

θ
‑0.06

400

θ

0.000
‑0.06

‑0.030

‑0.06

ε

0.08

‑500

SR8A9C22.5
22.5 Degree‑0.020□：L.B.
△：Crack○：Max.

‑0.030εFigure 4.8 Test Results of Square Interior Beam-Column Specimens Subjected to Biaxial Bending

ＳＲ６Ａ９Ｖ
590N/mm2
FA Class
Fc90

400

0.002

M(kN･m)
-0.04

θ
-0.04

0.04

-400
ＳＲ８Ａ９Ｖ
2
780N/mm
FA Class
Fc90

400

□：
L.B.
△：
Crack
○：
Max.

θ
0.04

-0.002

ＳＲ６Ａ９Ｖ
2
590N/mm
FA Class-0.006Fc90-0.010M(kN･m)ε-0.04

ＳＲ８Ｃ９Ｖ
2
780N/mm
FC Class
Fc90

0.04

-0.002

□：
L.B.
△：
Crack
○：
Max.

-0.04

-400

0.04
□：
L.B.
△：
Crack
○：
Max.

-0.010-0.04-0.010ε□：L.B.
△：Crack○：Max.

0.04-300□：L.B.
△：Crack○：Max.

-0.010□：L.B.
△：Crack○：Max.

θ
0.04-0.002
ＳＲ８Ｃ９Ｖ
2
780N/mm
FC Class-0.006
Fc90

Figure 4.9 Test Results of Square Exterior Beam-Column Specimens- 119 -

ε

-0.04

θ
-0.006

0.04

0.002

300 M(kN･m)

2

780N/mm
FA Class
Fc90

θ

-0.04
-0.002
ＳＲ６Ｃ９Ｖθ2
590N/mm0.04 FC Class-0.006
Fc90

-300
θ

0.002

300 M(kN･m)-0.04□：L.B.
△：Crack○：Max.

0.002

θ ＳＲ８Ａ９Ｖ

ＳＲ６Ｃ９Ｖ
2
590N/mm
FC Class
Fc90

ε□：L.B.
△：Crack○：Max.

Figure 4.10 Effect of Material Strength on Ductility

Figure 4.11 Effect of Material Strength on Enhancement in Flexural Strength

My

My

Mx

Mx

(a) 590MPa Steel Tube(b) 780MPa Steel Tube

Figure 4.12 Effect of Loading Direction on Flexural Strength- 120 -

N
Mx(i)
My(i)
Hinging Portion

Hinging Portion

Elastic Portion

My(j)

N

Mx(j)

(a) Column Model

M=0(b) Curvature(c) Axial Strain

Figure 4.13 Model of Beam-Column, and Assumed Curvature and Axial Strain Distributions- 121 -

(a) Square CFT(b) Circular CFT

Figure 4.14 Analytical Fiber Elements at the Critical Sections of Square and Circular CFT

StressStress
(MPa)

(MPa)
Comp.

Tens.
(%)

(%)
Comp.
StrainStrain(a) Concrete(b) Steel Tube

Figure 4.15 Typical Stress vs. Strain Relationships of Concrete and Steel Tube

- 122 -

250

ＳＣ４Ａ４Ｃ
2
400N/mm
FA Class
Fc40

M(kN･m)-0.002θ-0.06-0.006

0.08

-250
500

-0.010

θ

-0.06

0.08

θ
0.06

ＳＣ６Ｃ４Ｃ
590N/mm 2
FC Class
Fc40

300-0.002-0.010-0.002θ
ＳＣ６Ｃ４Ｃ
590N/mm2
Fc Class
Fc40

-300
300

M(kN･m)ＳＣ８Ａ４Ｃ
780N/mm2
FA Class
Fc40

θ

-300

-0.006-0.010θ
0.06-0.002

θ-0.0060.08

300-0.010ε

ＳＣ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90

θ
0.06

M(kN･m)-0.06-0.002
θ-0.006

0.08

ＳＣ８Ａ９Ｃ
2
780N/mm
FA Class
Fc90-0.010300

ＳＣ６Ｃ９Ｃ
590N/mm2
FC Class
ε Fc90M(kN･m)

θ
0.06

-0.06

-0.06

θ

ＳＣ８Ａ９Ｃ -0.002
780N/mm2
FA Class
Fc90

0.08

-0.006-300
200
ＳＣ８Ｃ９Ｃ
2
780N/mm
FC Class
Fc90-0.010εM(kN･m)

θ
0.06

-0.06
-0.002
θ

‑0.06

0.08

‑200

Figure 4.16 Analytical Results of Circular Interior Beam-Column Specimens- 123 -

ε

-0.06

ＳＣ６Ｃ９Ｃ
2
590N/mm
FC Class
Fc90

εＳＣ４Ａ９Ｃ
2
400N/mm
FA Class
Fc90

M(kN･m)

‑300

θ
0.06
-0.002

0.08

500

ε

-0.06

-0.06

-0.010‑0.06-0.006-0.010-0.006

-500
θ
0.06

-0.06

0.08

θ

-0.06

M(kN･m)-0.06-0.002

ＳＣ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90

ε

θ
0.06

-0.06

-250

-0.006-500

M(kN･m)

0.08

ε

-0.06
ＳＣ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

250
ＳＣ４Ａ９Ｃ
400N/mm2
FA Class
Fc90

-0.06

ＳＣ４Ａ４Ｃ
400N/mm2
FA Class
Fc40

M(kN･m)ＳＣ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

ＳＣ８Ａ４Ｃ
2
780N/mm
FA Class
Fc40

θ
0.06

-0.06

-0.006-0.010ＳＣ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

εＳＣ６Ａ９Ｖ

400

M(kN･m)M(kN･m)
0.002

θ

2

590N/mm
FA Class
Fc90

0.002

ＳＣ６Ｃ９Ｖ 200

θ

2

θ

-0.04

0.06

-0.04

ＳＣ８Ａ９Ｖ
780N/mm 2
FA Class
Fc90

-0.04

-0.010M(kN･m)εθ-0.002-0.006-0.010ε

ＳＣ８Ａ９Ｖ
780N/mm 2
FA Class
Fc90

ＳＣ８Ｃ９Ｖ
780N/mm 2
FC Class
Fc90

-0.04

-0.010M(kN･m)ＳＣ６Ｃ９Ｖ
2
590N/mm
FC Class
Fc90

ε

0.002
θ
-0.04
θ
0.06

-150

Figure 4.17 Analytical Results of Circular Exterior Beam-Column Specimens- 124 -

-0.002-0.006

0.06

150

0.06

0.06

θ

-200
θ

-0.04

-0.04

-0.04

0.002

0.06

-250

0.06
z

ＳＣ６Ａ９Ｖ
590N/mm2-0.006FA Class
Fc90

-400
250

-0.002

590N/mm
FC Class
Fc90

0.06

-0.002-0.006-0.010ε

ＳＣ８Ｃ９Ｖ
780N/mm2
FC Class
Fc90

ＳＲ４Ａ４Ｃ
2
400N/mm
FA Class
Fc40

300

0.000

M(kN･m)
‑0.06

θ

-0.06

0.06

ＳＲ４Ａ４Ｃ
2
400N/mm
FA Class
Fc40

M(kN･m)

0.000
‑0.06

θ

-0.06

0.06

ＳＲ４Ｃ４Ｃ
400N/mm2
FC Class
Fc40

-300
500

θ

-0.06

0.06

‑0.06
ＳＲ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

400

M(kN･m)

ＳＲ６Ｃ４Ｃ
590N/mm2
FC Class
Fc40-400
ＳＲ８Ａ４Ｃ
780N/mm2
FA Class
Fc40

500

ＳＲ６Ａ９Ｃ
590N/mm2
FA Class
Fc90

θ
0.06

M(kN･m)ＳＲ６Ｃ９Ｃ
590N/mm 2
FC Class
Fc90

θ
0.06

ＳＲ８Ａ４Ｃ
780N/mm2
FA Class
Fc40‑0.031 ε
0.000

M(kN･m)

ＳＲ８Ｃ４Ｃ 400
2780N/mm
FC Class
Fc40

-0.06

-400

v

0.000

M(kN･m)

‑0.06

θ
0.06

ＳＲ６Ｃ９Ｃ
590N/mm2
FC Class
Fc90

θ
0.06

‑0.010‑0.020‑0.030θ
0.06

ＳＲ８Ａ９Ｃ
2
780N/mm
FA Class
Fc90

500

0.000

M(kN･m)

‑0.06

θ

-0.06

0.000

0.06

‑0.010‑0.020-400‑0.010‑0.06

θＳＲ６Ａ９Ｃ
2
590N/mm
FA Class
Fc90‑0.030

-0.06

0.06

ＳＲ８Ｃ４Ｃ ‑0.010
2780N/mm
FC Class
Fc40‑0.020θ
0.06

ＳＲ８Ｃ９Ｃ
780N/mm2
FC Class
Fc90

400‑0.0300.000

M(kN･m)
‑0.06

θ

-0.06

0.06

ＳＲ８Ｃ９Ｃ
2
780N/mm
FC Class
Fc90

-400

‑0.030Figure 4.18 Analytical Results of Square Interior Beam-Column Specimens- 125 -

θ
0.06

ＳＲ８Ａ９Ｃ ‑0.010
2780N/mm
FA Class‑0.020Fc90-500‑0.030M(kN･m)

θ
0.06

‑0.06

‑0.020-500‑0.011‑0.021

0.06

400

θ
0.06

‑0.001‑0.020

0.000

0.06

0.06

ＳＲ４Ｃ９Ｃ
400N/mm2
FC Class
Fc90

-500

‑0.010

‑0.06

-0.06

θ

θ

‑0.030

θ

‑0.06

500

‑0.010‑0.030 ε

M(kN･m)

-0.06

0.000

0.06

300

θ
0.06

‑0.020-300‑0.010

‑0.06

-0.06

0.06

-0.06

‑0.020‑0.030θθ‑0.06
ＳＲ４Ａ９Ｃ
400N/mm2
FA Class
Fc90‑0.020-500
ＳＲ６Ｃ４Ｃ
2
590N/mm
FC Class
Fc40

ＳＲ４Ｃ９Ｃ
400N/mm2
FC Class
Fc90‑0.010

0.000

0.000

M(kN･m)

-300
θ
0.06

‑0.030 ε

M(kN･m)

300

-0.06

‑0.030 ε

ＳＲ４Ｃ４Ｃ
400N/mm2
FC Class
Fc40

ＳＲ６Ａ４Ｃ
590N/mm2
FA Class
Fc40

ＳＲ４Ａ９Ｃ
400N/mm2
FA Class
Fc90‑0.010‑0.020

-300
300

θ
0.06

‑0.010‑0.020‑0.030

θ
0.06

500
SR6A9C45
45 DegreeM(kN･m)

0.000

400

θ

‑0.06

0.08

M(kN･m)SR6C9C45
45 Degree‑0.010θ

0.08

SR6A9C45
45 Degree‑0.020‑500

500‑0.0300.08‑0.020SR6C9C45
45 Degree

‑400

θ

‑0.06
SR8A9C45
45 Degree

‑0.06

ε

0.000

M(kN･m)500

0.08‑0.030θ‑0.06

SR8A9C22.5
22.5 Degree‑0.010ε

0.000

M(kN･m)0.08‑0.010

θ
‑0.06

θ
0.08

‑0.010

θ
‑0.06

0.000
‑0.06

θ

0.08

‑0.06

‑0.020‑0.030‑500

0.08‑0.020SR8A9C45
45 Degreeε‑0.030‑500

εSR8A9C22.5
22.5 Degree

Figure 4.19 Analytical Results of Square Interior Beam-Column Specimens Subjected to Biaxial Bending

ＳＲ６Ａ９Ｖ
590N/mm2
FA Class
Fc90

400

0.002

M(kN･m)
-0.04

θ
0.04

-0.002

θ ＳＲ６Ａ９Ｖ

ＳＲ６Ｃ９Ｖ
2
590N/mm
FC Class
Fc90

0.002

300 M(kN･m)

θ

2

-0.04

0.04

590N/mm
FA Class
Fc90-0.010-400
ＳＲ８Ａ９Ｖ
2
780N/mm
FA Class
Fc90

400

M(kN･m)-0.04-0.006

ε

0.002
-0.04

θ ＳＲ８Ａ９Ｖ

0.04

θ
0.04

-0.002

-0.04

0.04

-400

-0.0100.04

2590N/mm
FC Class
Fc90-0.006-0.010ＳＲ８Ｃ９Ｖ 300 M(kN･m)

0.002

ε
θ

2

780N/mm
FC Class
Fc90

-0.04

θ
-0.04

-0.006-0.04-0.002
ＳＲ６Ｃ９Ｖ-300

2780N/mm
FA Class
Fc90

θ

ε

0.04

-300

Figure 4.20 Analytical Results of Square Exterior Beam-Column Specimens- 126 -

0.04

-0.002
ＳＲ８Ｃ９Ｖ
2
780N/mm
FC Class-0.006Fc90-0.010εFigure 5.1 Shapes of Specimens- 127 -

Figure 5.2 Details of Panel Zone

- 128 -

Figure 5.3 Test Setup for Specimens R1 and R2

No.8
No.6
No.5
No.7
―

+
No.1

No.3

No.4

No.2
z

y
x

Figure 5.4 Test Setup for Specimen R6- 129 -

Figure 5.5 Loading History for Beam-to-Column Specimens

Q

N
pNo
VariableNc(0.7pNo)

Nc

Nt→Nc
R

Nc(0.2pNo)

ConstantQ

Nt(-0.3pNso)

Nt

Nc→Nt

-pNso

Figure 5.6 Rules for Axial Force Loading for Beam-to-Column Specimens- 130 -

Figure 5.7 Instruments for Measurements of Story Driftand Displacement of Panel Zone

- 131 -

Figure 5.8 Crack Patterns of Filled Concrete in Panel Zone

- 132 -

250

cQ(kN)

200

cQ(kN)

pQu

150

pQu
100

pQy
50

pQy

R(rad.)

R(rad.)
0-0.05-500.00

0.05

0.10

-150

Specimen : R1

-0.05

Specimen : R2

cQ(kN)

250
pQu

100

pQy

0.00

0.05

pQu
pQy

50

0.10

-0.05

-100
-200

cQ(kN)

150
R(rad.)

0

R(rad.)-500.00-150

Specimen : R3

-300
100

0.10

-200

200

-0.05

0.05

-100

-250
300

0.00

0.05
Specimen : R4-250
cQ(kN)

100

60

pQu

cQ(kN)

60

pQu

20

pQy
R(rad.)

pQy
20
-0.05

-200.00-60

0.10

R(rad.)
0.05

0.10

-0.05

-200.00-60

Specimen : R5

-100

0.05
Specimen : R5'-100

Figure 5.9(a) Story Shear vs. Story Drift Angle Relations of Rectangular Specimens- 133 -

0.10

300

cQ(kN)

300

200

pQu
100

pQy
R(rad.)

0
-0.05

200

pQu

100

0.00

0.05

0.10

-0.05

Specimen : C1

0.10

-300
cQ(kN)

100

200

pQu

100

pQy

0.00

0.05

cQ(kN)
pQu

50

pQy
R(rad.)

0

-200

0.05
Specimen : C2

-200

-300

-100

0.00

R(rad.)

-100

-200

-0.05

pQy

0

-100

300

cQ(kN)

0.10

R(rad.)

0
-0.05

0.00

0.05

-50

Specimen : C3

Specimen : C4

-300

-100

Figure 5.9(b) Story Shear vs. Story Drift Angle Relations of Circular Specimens- 134 -

0.10

2000

3000

1500

pQu
pQy

1000
500
0
-500

-2000
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
642

γp (rad.)cQ (kN)

500
400
cQu
300
cQy
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2
θc (rad.)300bQybQ (kN)bQ (kN)

100
0

-100

-100

-200

-200

-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

θb (rad.)SPECIMEN: R1

θb (rad.)SPECIMEN: R2

1000
cQu
800
cQy
600
400
200
0
-200
-400
-600
-800
-1000
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2
θc (rad.)

bQu
bQy

200

100
bQ (kN)-3000-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
642

γp (rad.)

bQu

0

0

-2000

-1500

cQ (kN)cQ (kN)

200

1000

-1000

θc (rad.)
300

pQu
pQy

-1000

γp (rad.)
500
cQu
400
cQy
300
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

2000
pQ (kN)pQ (kN)pQ (kN)

2500
2000
pQu
1500
pQy
1000
500
0
-500
-1000
-1500
-2000
-2500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

500
bQu
400
bQy
300
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2
θb (rad.)SPECIMEN: R3

Figure 5.10(a) Hysteresis Curves of Panel Zone, Column and Beam of Rectangular Specimens(continue)- 135 -

1500

500

-500

-1000

-1000

-1500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

-1500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

150

-400
-600
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2
θc (rad.)200bQ (kN)bQ (kN)

100

-100
-200
-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2
θb (rad.)SPECIMEN: R4

cQy

50
0
-50

150

cQu

100

cQy

50
0
-50

-100

-100

-150

-150

-200
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

-200
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

θc (rad.)

bQu
bQy

0

200
cQu

cQ (kN)cQ (kN)cQ (kN)-200

300

γp (rad.)

200
100

0

pQu
pQy

0

γp (rad.)

cQu
cQy

200

500

-500

250
bQu
200
bQy
150
100
50
0
-50
v
-100
-150
-200
-250
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2
θb (rad.)SPECIMEN: R5

θc (rad.)bQ (kN)

400

1000

pQu
pQy

0

γp (rad.)
600

1500

pQ (kN)

1000
pQ (kN)pQ (kN)

2500
2000
pQu
1500
pQy
1000
500
0
-500
-1000
-1500
-2000
-2500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

250
bQu
200
bQy
150
100
50
0
-50
-100
-150
-200
-250
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2
θb (rad.)SPECIMEN: R5'

Figure 5.10(a) Hysteresis Curves of Panel Zone, Column and Beam of Rectangular Specimens- 136 -

2000

500

pQy

1000

pQy

-500

1000
500

0

pQu
pQy

0

-1000

-500

-2000

-1000

-1000
-1500
-2000

-3000

-1500

-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
642

γp (rad.)γp (rad.)γp (rad.)γp (rad.)

300
200

600

cQu

400

cQy

0
-100

-200

-400
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

0
-50

-200
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

θc (rad.)bQu
bQy

bQ (kN)

100
0

-100

-100

-200

-200

-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

SPECIMEN: C1

cQy

-150

-600
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

500
bQu
400
bQy
300
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

θb (rad.)θc (rad.)300
200
bQ (kN)

200

50

cQu

-100
-400

-300

bQu
bQy

θb (rad.)0θc (rad.)

0

100

-200

300

100

150

cQy

200

cQ (kN)

100

200
cQu

cQ (kN)400bQ (kN)bQ (kN)

20000θc (rad.)200

pQu
pQ (kN)pQ (kN)

1000

1500
pQu

pQ (kN)

pQu
pQy

500
400
cQu
300
cQy
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

300

3000

1500

cQ (kN)pQ (kN)cQ (kN)

2500
2000
1500
1000
500
0
-500
-1000
-1500
-2000
-2500

0
-100
-200
-300
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

SPECIMEN: C3

Figure 5.10 (b) Hysteresis Curves of Panel Zone, Column and Beam of Circular Specimens- 137 -bQy100

θb (rad.)SPECIMEN: C2bQuθb (rad.)SPECIMEN: C4

1.0

1.0

0.8

0.8
Panel

0.6
0.4

0.4
Beam

Beam

0.2

0.2
Column

0.0
0.00

0.0
0.02
0.04
Story Drift Angle R(rad.)

0.06

0.00

SPECIMEN: R1
1.0

0.8

0.8
Panel

0.6

Column0.02
0.04
Story Drift Angle R(rad.)

Panel

0.6

0.4

0.4
Beam

Beam

0.2

0.2
Column
0.02

0.04

0.06

0.0
0.00

Story Drift Angle R(rad.)Column
0.02

0.04

0.06

Story Drift Angle R(rad.)SPECIMEN: R3

SPECIMEN: R4

1.0

1.0

0.8

0.8
Panel

Panel

0.6

0.6

0.4

0.4
Beam

Beam

0.2

0.2
Column

0.0
0.00

0.06

SPECIMEN: R2

1.0

0.0
0.00

Panel

0.6

0.02

Column
0.04

0.06

0.0
0.00

0.02

0.04

Story Drift Angle R(rad.)Story Drift Angle R(rad.)SPECIMEN: R5(comp.)

SPECIMEN: R5(tens.)

0.06

1.0
0.8
Panel

0.6
0.4

Beam
0.2
0.0
0.00

Column
0.02

0.04

0.06

Story Drift Angle R(rad.)SPECIMEN: R5'

Figure 5.11(a) Contribution of Deformation of Panel Zone, Beam and Column
to Overall Story Drift in Rectangular Specimens- 138 -

1.0

1.0

0.8

0.8
Panel

0.6
0.4

0.4
Beam

Beam

0.2

0.2
Column

0.0
0.00

Panel

0.6

0.02
0.04
Story Drift Angle R(rad.)

0.06

0.0
0.00

SPECIMEN: C1

Column0.02
0.04
Story Drift Angle R(rad.)0.06

SPECIMEN: C2

1.0

1.0

0.8

0.8
Panel
Panel

0.6

0.6

0.4

0.4
Beam

Beam

0.2

0.2
Column

0.0
0.00

0.02

Column
0.04

0.06

0.0
0.00

Story Drift Angle R(rad.)

0.02

0.04

0.06

Story Drift Angle R(rad.)SPECIMEN: C3

SPECIMEN: C4(comp.)
1.0
0.8
Panel
0.6
0.4
Beam
0.2
Column
0.0
0.00

0.02
0.04
Story Drift Angle R(rad.)0.06

SPECIMEN: C4(tens.)

Figure 5.11(b) Contribution of Deformation of Panel Zone, Beam and Column
to Overall Story Drift in Circular Specimens- 139 -

250

cQ(kN)
pQu

150

pQy
50
-0.05

R(rad.)-500.00

0.05

-150

0.10

Specimen : R6

-250
Figure 5.12 Story Shear vs. Story Drift Angle Relationof Three Dimensional Specimen R6
2000
1500

pQu

1500

pQy

pQy

1000

500

500

pQ (kN)pQ (kN)

1000

2000
pQu

0
-500

0
-500

-1000

-1000

-1500

-1500

-2000
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

-2000
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
642
γp (rad.)
500
cQu
400
300
200
cQy
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

cQ (kN)cQ (kN)γp (rad.)
500
cQu
400
300
cQy
200
100
0
-100
-200
-300
-400
-500
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

θc (rad.)

bQu
250
200
bQy
150
100
50
0
-50
-100
-150
-200
-250
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
64
2

bQ (kN)bQ (kN)θc (rad.)

250
bQu
200
bQy
150
100
50
0
-50
-100
-150
-200
-250
-0.0 -0.0 -0.0 0.00 0.02 0.04 0.06 0.08
6
4
2

θb (rad.)θb (rad.)X-direction

Y-direction

Figure 5.13 Hysteresis Curves of Panel Zone, Column and Beamof Three Dimensional Specimen R6- 140 -

1.0

1.0

0.8

0.8
Panel

0.6

Panel

0.6

0.4

0.4
Beam

Beam

0.2

0.2
Column

0.0
0.00

0.02
0.04
Story Drift Angle R(rad.)

0.0
0.00

0.06

Column0.02
0.04
Story Drift Angle R(rad.)X-direction

0.06

Y-direction

Figure 5.14 Contribution of Deformation of Panel Zone, Beam and Column
to Overall Story Drift in Rectangular Specimens
2.0

pQ/pQy

2.5
2.0

1.5

Steel:590MPa
Concrete:90MPa 1.0

Steel:590MPa
Concrete:90MPa

R/Ry

1.5
1.0
0.5

10

0.0
-2 -0.5

Concrete:40MPa

0.5

Steel:780MPa

0.0
-6

-2

-0.5

2

6

-6

-1.5

Concrete:40MPa
Steel:780MPa

R/Ry

6

10

2

-1.0
-1.5
-2.0

R1
R2
R3

-1.0

pQ/pQy

-2.0

C1
C2
C3

-2.5

(a) Effect of Material Strength
2.5
2.0

pQ/pQy2.5
2.0

Exterior Joint

1.5
1.0
Variable Axial Load
0.5
-0.3pNso
0.0-6-2 -0.5

pQ/pQyInterior Joint

1.5
1.0

Interior Joint

0.5

Variable Axial Load, 0.7pNo R/Ry

0.0
2

6

-1.0

10

-6

-2.0

2

R/Ry

6

10

-1.0

R1
R5
R5'

-1.5

-2 -0.5

Exterior Joint
Variable Axial Load
0.7pNo

-1.5
-2.0

-2.5

Variable Axial Load
-0.3pNso

C1
C4

-2.5

(b) Effect of Axial Load
2.0

2.5

pQ/pQy1.5

1.0

Through Diapragm

0.5

1.0

-2

-0.5

0 Degree

0.5

R/Ry

Outer Diapragm

0.0-6

pQ/pQy2.0

1.5

45 Degree

R/Ry

0.0
2

6

10

-6

-2 -0.5

2

6

10

-1.0

-1.0

R1
R4

-1.5

-1.5
-2.0

-2.0

-2.5

(c) Effect of Diaphragm Type(d) Effect of Loading Direction

Figure 5.15 Normalized Shear Force of Panel Zone vs.
Normalized Story Drift A ngle Relations- 141 -

R1
R6

300

250

250

200

pQy
(AIJ-SRC)

150
100●:Circular Specimen■:Square Specimen

50

Maximum Shear Strength of Test
pQmax(kN)

Yield Shear Strength of Test
pQ(kN)

300

0

200

pQu=1.2 × pQy(AIJ-SRC)

150

pQy(AIJ-SRC)

100●:Circular Specimen■:Square Specimen

50
0

0

50

100

150

200

250

300

0

50

100

150

200

250

Design Yield Shear Strength pQy(kN)

Design Yield Shear Strength pQy(kN)(a) Yield Strength(b) Ultimate Strength

Figure 5.16 Yield and Ultimate Shear Strengths of Panel Zoneof Two Dimensional Specimens

Figure 5.17 Yield and Ultimate Shear Strengths of Panel Zoneof Three Dimensional Specimen- 142 -

300

Figure 5.18 Assumed Tri-Linear Skelton Curves for Steel and Filled Concrete in Panel Zone

Figure 5.19 Arch Mechanisms in Panel Zone

- 143 -

(a) Specimen R1(b) Specimen R2(c) Specimen R3(d) Specimen R4(e) Specimen R5 (Compressive Column Axial Force) (f) Specimen R5 (Tensile Column Axial Force)

Figure 5.20 Comparisons of Calculated and Experimental Skeleton Curves- 144 -

Figure 5.21 Analytical Meshes for Finite Element Method

Figure 5.22 Stress vs. Strain Relationships of Concrete- 145 -

Figure 5.23 Story Shear Force vs. Story Drift Angle Relationshipsof Beam-to-Column Connection Specime ns- 146 -

Figure 5.24

pQc / Ac / fc'−γ Relationships

Figure 5.25

pQs − γ Relationships- 147 -

Figure 5.26 Contour and Flow of Principal Compressive Concrete Stress in Panel Zones

Figure 5.27

pQc / Ac / fc'− H / Dc Relationships

Figure 5.28

pQc / Ac / fc'−H / D ' Relationships- 148 -

Figure 6.1 Biaxial Stress State of Steel Tube under Yielding- 149 -

(a) Specimens of US -Japan Research Program(b) Specimens in Database Collected from the Proceedings of AIJ Annual Meetings

Figure 6.2 Comparisons Between Calculated Ultimate Axial Loads and Experimental Onesof Circular Columns- 150 -

Figure 6.3 Local Bucking Modes of Square Steel Tube

- 151 -

(a) Specimens of US -Japan Research Program(b) Specimens in Database Collected from the Proceedings of AIJ Annual Meetings

Figure 6.4 Comparisons Between Calculated Ultimate Axial Loads
and Experimental Ones of Square Columns- 152 -

MPa%
Figure 6.5 Stress vs. Strain Curves for Plain and Confi ned Concrete

Figure 6.6 Stress Blocks for Concrete and Steel in Circular CFT Section(Prescribed in the AIJ -CFT Recommendations [6.1])

Figure 6.7 Stress Blocks for Concrete in Circular CFT Section(Proposed by Sun and Sakino [6.9])- 153 -

Figure 6.8 Comparisons Between Calculated Ultimate Flexural Strengths andExperimental Ones of Circular Columns- 154 -

Figure 6.9 Stress Blocks for Concrete and Steel in Square CFT Section(Prescribed in the AIJ -CFT Recommendations [6.1])

Figure 6.10 Proposed Stress Blocks for Concrete and Steel in Square CFT Section with HighStrength Concrete and / or Thin Steel Tube Wall

- 155 -

Figure 6.11 Comparisons Between Calculated Ultimate Flexural Strengths andExperimental Ones of Square Columns- 156 -

Figure 6.12 Beam-Column Under Combined Forces

Figure 6.13 Definition of Limit Rotation Angle- 157 -

(a) Circular Columns(b) Square Columns
Figure 6.14 Comparisons Between Calculated Ultimate Flexural Strengths and Experimental Ones

- 158 -

(a) Circular Columns(b) Square Columns
Figure 6.15 Comparisons Between Calculated Ultimate Flexural Strengths and Experimental Ones
Attained within Rotation Angle Limitation of 0.01 radian- 159 -

Figure 6.16 Comparison Between Predicted Limit Rotation and Experimental One(Circular Columns)

Figure 6.17 Comparison Between Predicted Limit Rotation and Experimental One(Square Columns)- 160 -

Moment

MomentM max

Mu

0.85 M max

My

αy :Stiffness
Degrading Ratio

1/3M max

α y・Ke

R

Ke

K1

K2

αy =K 2/K 1

R

Ru

Figure 6.18 Tri-linear Skeleton Model

Figure 6.19 Definition of Experimental ay

1.0

αy

αy

1.0
0.8
0.6

0.6

0.4

0.4
● U.S.‑Japan○ Others

0.2
0.0
0.0

0.2

0.4

0.6

0.8

0.2

1.0

N/N0

M(kN・m)M(kN・m)

500

M(kN・m)300

0.2

M(kN・m)

150
200

300
100

0.4

0.6

N/N 0

0.8

1.0

Figure 6.21 Effect of N/No on ay(Square CFT)

200

400

150

● U.S.‑Japan○ Others

0.0
0.0

Figure 6.20 Effect of N/No on ay(Circular CFT)

200

0.8

M(kN・m)

500

500

400

400

300

300

M(kN・m)100
200
SC4A4C

50

Model
R

0
0

0.02

0.04

SC6A9C
Model
R

100
0
0

0.02

0.04

200

100

SC8A9C
Model
R

0
0

0.02

0.04

50
0
0

200

SR4A4C

SR6A9C

Model 100R
00.02 0.04
0Model

SR8A9C100

0.02 0.04

00.00 0.02 0.04(b) Square CFT(a) Circular CFT

Figure 6.22 Comparisons Between Tri-linear Skeleton Model and Experimental Results- 161 -

Model

R

R

250

500

M(kN･m)300

M(kN･m)SC8‑A‑9‑C

R
‑0.04

0
0.00

‑0.02

M(kN･m)SC6‑A‑9‑C

SC4‑A‑4‑C

0.02

0.04

0.06

R

R
‑0.04

0
0.00

‑0.02

0.02

0.04

0.06 ‑0.04

0
0.00

‑0.02

0.02

▲：Ru

▲：Ru
‑250

0.04

0.06

▲：Ru

‑500

‑300

(a) Circular CFT
200

M(kN･m)500

SR4‑A‑4‑C

‑0.04

‑0.02

M(kN･m)500

SR6‑A‑9‑C

0
0.00

R
0.02

0.04

‑0.04

‑0.02

0
0.00

R
0.02

0.04

0.06

‑0.04

‑0.02

0
0.00

▲：Ru

▲：Ru
‑200

M(kN･m)SR8‑A‑9‑C

R
0.02

0.04

0.06

▲：Ru
‑500

‑500

(b) Square CFT

Figure 6.23 Comparisons Between Normal Tri-linear Hysteretic Model and Experimental Results- 162 -

Figure 7.1 Elevations of Theme Structures

Figure 7.2 Floor Plan of Theme Structures- 163 -

Figure 7.3 Story Displacements Caused by Beam and Column Deformations

Figure 7.4 Vibrational Mode Shapes- 164 -

Figure 7.5 Maximum Responses of Shear Coefficients(El Centro NS : 25cm/sec)

Figure 7.6 Q − δ Relations of 40-Story Frames- 165 -

Figure 7.7 Maximum Responses of Shear Coefficients, Overturning Moments and Story Drifts

Figure 7.8 Maximum Responses of Ductility Factors of Beams and Columns(HACHINOHE : 50cm/sec)- 166 -

CFT system

Steel system

Figure 7.9 Formation of Plastic Hinges in Beams and Columns

Figure 7.10 Comparison of Steel Amount Between CFT and Steel Frames- 167 -

APPENDIX

The list of the technical papers written in English is shown below, which includes the papers on concrete-filled
structural steel tube column system together with those on new materials, elements and systems, reinforced
concrete column and structural steel beam systems, and reinforced concrete and steel reinforced concrete wall
systems .

***** Fiscal Year of 1993 *****[1]-1993
Hiroyuki Yamanouchi, Stephen A. Mahin, Subhash C. Goel and Isao Nishiyama
U.S.-Japan Cooperative Earthquake Research Program on Composite and Hybrid Structures
Proceedings of the 25th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Gaithersburg,
USA, pp.585-591, May 17-20, 1993***** Fiscal Year of 1995 *****[1]-1995
Isao Nishiyama and Hiroyuki Yamanouchi
U.S.-Japan Cooperative Earthquake Engineering Research Program on Composite and Hybrid Structures Research Plan and its Progress Proceedings of the 27th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Tsukuba, Japan,
pp.333-343, May 16-19, 1995
[2]-1995
Kabeyasawa, T. , Ohkubo, T. and Nakamura, Y.
Tests and Analyses of Reinforced Concrete Shear Walls in Hybrid Structures
Pacific Conference on Earthquake Engineering, Australia, Vol.2, pp.145-154, November 20-22, 1995***** Fiscal Year of 1996 *****[1]-1996
Shosuke Morino, Kenji Sakino, Akiyoshi Mukai and Kenzo Yoshioka
U.S.-Japan Cooperative Earthquake Research Program on CFT Column Systems
Proceedings of 5th International Colloquium on Stability of Metal Structures, SSRC, April 1996
[2]-1996
Subhash C. Goel and Isao Nishiyama
U.S.-Japan Cooperative Earthquake Engineering Research Program on Composite and Hybrid Structures Research Progress and Current Status Proceedings of the 28th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Gaithersburg,
USA, pp.319-325, May 14-17, 1996

- 168 -

[3]-1996
Chen Lian, Sanada Y. and Kabeyasawa T.
3-Dimensional Analysis of Hybrid Wall System -Static and Earthquake Response Analysis
Transactions of the Japan Concrete Institute, Vol. 18, pp. 213-220, 1996
[4]-1996
Hiroshi Kuramoto
Seismic Resistance of Through Column Type Connections for Composite RCS Systems
11th World Conference on Earthquake Engineering (11WCEE), Acapulco, Mexico, June 23-28, 1996
[5]-1996
Makoto Kato, Ken-ichi Sugaya and Norikazu Nagatsuka
Optimum Moment Distribution between Shear Walls and Boundary Beams of Coupled Shear Wall with Flange
Walls
11th World Conference on Earthquake Engineering (11WCEE), Acapulco, Mexico, June 23-28, 1996
[6]-1996
T. Kabeyasawa, T. Ohkubo and Y. Nakamura
Tests and Analyses of Hybrid Wall Systems
11th World Conference on Earthquake Engineering (11WCEE), Acapulco, Mexico, June 23-28, 1996
[7]-1996
S. Morino
U.S.-Japan Cooperative Earthquake Research Program on Composite and Hybrid Structures
The 2nd International Symposium on Civil Infrastructure Systems, Hong Kong/China, December 9-12, 1996***** Fiscal Year of 1997 *****[1]-1997
H. Kuramoto and H. Noguchi
An Overview of Japanese Research on RCS Systems
ASCE Structures Congress XV, Portland, Oregon, pp.716-720, April 13-16, 1997[2]-1997
H. Noguchi and K. Kim
Analysis of Beam-Column Joints in Hybrid Structures
ASCE Structures Congress XV, Portland, Oregon, pp.726-730, April 13-16, 1997[3]-1997
M. Teshigawara
An Overview of Japanese Research on Hybrid Wall Systems
ASCE Structures Congress XV, Portland, Oregon, pp.1096-1100, April 13-16, 1997[4]-1997
S. Morino, K. Sakino, A. Mukai and K. Yoshioka
Experimental Studies on CFT Column Systems - U.S.-Japan Cooperative Earthquake Research Program- 169 -

ASCE Structures Congress XV, Portland, Oregon, pp.1106-1110, April 13-16, 1997[5]-1997
Isao Nishiyama and Hiroyuki Yamanouchi
U.S.-Japan Cooperative Earthquake Engineering Research Program on Composite and Hybrid Structures - JapanSide Progress Proceedings of the 29th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Tsukuba, Japan,
pp.473-478, May 13-16, 1997[6]-1997
Nobuhiro Araki, Yasuhiro Matsuzaki, Katsuhiko Nakano, Takahiro Kataoka and Hiroshi Fukuyama
Shear Capacity of Retrofitted RC Members with Continuous Fiber Sheets
Proceedings of the Third International Symposium on Non-Metallic (FRP) Reinforcement for Concrete
Structures (FRPRCS-3), Sapporo, Japan, Vol.1, pp515-522, 1997-10[7]-1997
Takahiro Kataoka, Nobuhiro Araki, Katsuhiko Nakano, Yasuhiro Matsuzaki and Hiroshi Fukuyama
Ductility of Retrofitted RC Columns with Continuous Fiber Sheets
Proceedings of the Third International Symposium on Non-Metallic (FRP) Reinforcement for Concrete
Structures (FRPRCS-3), Sapporo, Japan, Vol.1, pp547〜554, 1997-10***** Fiscal Year of 1998 *****[1]-1998
Isao Nishiyama, Hiroyuki Yamanouchi and Hisahiro Hiraishi
U.S.-Japan Cooperative Earthquake Engineering Research Program on Composite and Hybrid Structures Current Status of Japan Side Research Proceedings of the 30th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Gaithersburg,
USA, pp.443-451, May 12-15, 1998 (Wind and Seismic Effects, NIST SP 931, Issued August 1998)[2]-1998
S. Morino
An Overview of U.S.-Japan Cooperative Earthquake Research Program on CFT Column Systems
Structural Engineers World Congress, Paper Reference T169-1, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[3]-1998
K. Sakino, T. Ninakawa, H. Nakahara and S. Morino
Experimental Studies and Design Recommendations on Concrete Filled Steel Tubular Columns -US-Japan
Cooperative Earthquake Research ProgramStructural Engineers World Congress, Paper Reference T169-3, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[4]-1998
I. Nishiyama, H. Itadani and K. Sugihiro- 170 -

Bi-directional Seismic Response of Reinforced Concrete Column and Structural Steel Beam Subassemblages
Structural Engineers World Congress, Paper Reference T177-2, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[5]-1998
H. Noguchi and K. Kim
Shear Strength of Beam-to-column Connections in RCS System
Structural Engineers World Congress, Paper Reference T177-3, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[6]-1998
V. C. Li, H. Fukuyama and A. Mikame
Development of Ductile Engineered Cementitious Composite Elements for Seismic Structural Applications
Structural Engineers World Congress, Paper Reference T177-5, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[7]-1998
T. Kabeyasawa and Y. Nakamura
Displacement-based Design of Hybrid Core Wall System
Structural Engineers World Congress, Paper Reference T186-1, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[8]-1998
M. Teshigawara, K. Sugaya, M. Kato and Y. Matsushima
Seismic Test on 12-story Coupled Shear Wall with Flange Walls
Structural Engineers World Congress, Paper Reference T186-4, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[9]-1998
M. Teshigawara, K. Sugaya, M. Kato and Y. Matsushima
Energy Absorption Mechanism and the Fluctuation of Shear Force in the Coupled Shear Walls
Structural Engineers World Congress, Paper Reference T186-5, San Francisco, USA July 19-23, 1998
(Structural Engineering World Wide 1998)[10]-1998
K. Sakino, E. Inai and H. Nakahara
Tests and Analysis on Elasto-Plastic Behavior of CFT Beam-Columns - U.S.-Japan Cooperative earthquake
Research Program Proceedings of Fifth Pacific Structural Steel Conference, Edited by Dong-Il Chang, Hyo-Nam Cho, Chung-Bang
Yun and Sang-Dae Kim, Seoul, Korea, Vol. 2, pp. 901-906, October 13-16, 1998[11]-1998
R. Kanno and G. G. Deierlein
Bearing Strength of Joints between Steel Beams and Reinforced Concrete Columns
Proceedings of Fifth Pacific Structural Steel Conference, Edited by Dong-Il Chang, Hyo-Nam Cho, Chung-Bang- 171 -

Yun and Sang-Dae Kim, Seoul, Korea, Vo l. 2, pp. 919-924, October 13-16, 1998[12]-1998
Isao Nishiyama
An Overview of U.S.-Japan Cooperative Earthquake Research Program on Composite and Hybrid Structures
Proceedings of Fifth Pacific Structural Steel Conference, Edited by Dong-Il Chang, Hyo-Nam Cho, Chung-Bang
Yun and Sang-Dae Kim, Seoul, Korea, Vol. 2, pp. 925-930, October 13-16, 1998[13]-1998
H. Nakahara and K. Sakino
Axial Compressive and Uniform Bending Tests of High Strength Concrete Filled Square Steel Tubular Columns
Proceedings of Fifth Pacific Structural Steel Conference, Edited by Dong-Il Chang, Hyo-Nam Cho, Chung-Bang
Yun and Sang-Dae Kim, Seoul, Korea, Vol. 2, pp. 943-948, October 13-16, 1998[14]-1998
J. Kawaguchi, S. Morino, J. Shirai and E. Tatsuta
Database and Structural Characteristics of CFT Beam-Columns
Proceedings of Fifth Pacific Structural Steel Conference, Edited by Dong-Il Chang, Hyo-Nam Cho, Chung-Bang
Yun and Sang-Dae Kim, Seoul, Korea, Vol. 2, pp. 955-960, October 13-16, 1998[15]-1998
Hiroshi Kuramoto and Isao Nishiyama
Equivalent Damping Factor of Composite RCS Frames
1998 ACI Fall Convention, Westin Century Plaza, Los Angeles, USA, October 25-30, 1998[16]-1998
Hiroyuki Yamanouchi, Isao Nishiyama and Jun Kobayashi
Development and Usage of Composite and Hybrid Structures (CHS) Based on Performance
ACI SP-174 - Hybrid and Composite Structures, 1998***** Fiscal Year of 1999 *****[1]-1999
Isao Nishiyama, Hiroyuki Yamanouchi and Hisahiro Hiraishi
U.S.-Japan Cooperative Earthquake Engineering Research Program on Composite and Hybrid Structures Japanese Side Research AccomplishmentsProceedings of the 31th Joint Meeting of U.S.-Japan Panel on Wind and Seismic Effects, UJNR, Tsukuba, Japan,
pp.422-430, May 11-14, 1999
[2]-1999
H. Fukuyama, Y. Matsuzaki, K. Nakano and Y. Sato
Structural Performance of Beam Elements with PVA -ECC
Proceedings of the Third International RILEM Workshop on High Performance Fiber Reinforced Cement
Composites (HPFRCCS-3), Mainz, Germany, Edited by H. W. Reinhardt and A. E. Naaman, RILEM
Proceedings PRO6, RILEM Publications S.A.R.L., pp.531-541, 1999-5- 172 -

[3]-1999
Hiroshi Fukuyama, Yukihiro Sato, Victor C. Li, Yasuhiro Matsuzaki and Hirozo Mihashi
Ductile Engineered Cementitious Composite Elements for Seismic Structural Application
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[4]-1999
Yasuhiro Matsuzaki, Shigeru Fujii, Hiroshi Fukuyama and Katsuhiko Nakano
Seismic Retrofit using Continuous Fiber Sheets
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[5]-1999
Nozomu Baba and Yasushi Nishimura
Seismic Behavior of RC Column - S beam Moment Frames
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[6]-1999
Subhash C. Goel, Atsuo Tanaka, Hiroyuki Yamanouchi and Hiroshi Fukuyama
Experimental Study on the Performance of the RC Research on New Materials, Elements and Systems
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[7]-1999
Charles W. Roeder and Shosuke Morino
Research on CFT Column Systems
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[8]-1999
John W. Wallace and Akira Wada
Hybrid Wall Systems: US-Japan Research
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[9]-1999
Gregory G. Deierlein and Noguchi Hiroshi
Research on RC/SRC Column Systems
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[10]-1999
Yoshiyuki Matsushima, Masaomi Teshigawara, Ken-ichi Sugaya and Makoto Kato
Seismic Performance Evaluation Method for a Building with Centre Core Reinforced Concrete Walls and- 173 -

Exterior Steel Frame
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[11]-1999
Isao Nishiyama, Hidehiko Itadani, Kunio Sugihiro and Hiroshi Kuramoto
Bi-Directional Behavior of Interior-. Exterior-, and Corner-Joints of RCS System
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[12]-1999
Shao-Hua Chen and Toshimi Kabeyasawa
Modelling of Re inforced Concrete Shear Wall for Nonlinear Analysis
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[13]-1999
Hiroyuki Nakahara and Kenji Sakino
Flexural Behavior of Concrete Filled Square Steel Tubular Beam-Columns
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[14]-1999
Hiroyoshi Tokinoya, Akiyoshi Mukai, Kenzo Yoshioka, Toshiyuki Fukumoto, Takashi Noguchi, Yoshiyuki
Murata and Yo shinari Tanaka
Earthquake Resistance Behavior of CFT Columns
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[15]-1999
Ken-ichi Sugaya, Makoto Kato, Yoshiyuki Matsushima and Masaomi Teshigawara
Experimental Study on Carrying Shear Force Ratio of 12-Story Coupled Shear Walls
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[16]-1999
Toshiaki Fujimoto, Eiichi Inai, Makoto Kai, Koji Mori, Osamu Mori and Isao Nishiyama
Behavior of Beam-To-Column Connection of CFT Column System
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February
4, 2000[17]-1999
Makoto Kato, Yoshiyuki Matsushima, Ken-ichi Sugaya and Masaomi Teshigawara
Seismic Energy Dissipation System of 12-Story Coupled Shear Walls
12th World Conference on Earthquake Engineering (12WCEE), Auckland, New Zealand, January 30 - February- 174 -

4, 2000[18]-1999
Subhash C. Goel, Hiroyuki Yamanouchi and Isao Nishiyama
U.S.-Japan Cooperative Earthquake Research Program on Composite and Hybrid Structures: An Overview
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 23-34, March 22-24, 2000
[19]-1999
Mizuaki Uchikoshi, Yukio Hayashi and Shosuke Morino
Merits of CFT Column System - Results of Trial Design of Theme Structures
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 135-142, March 22-24, 2000
[20]-1999
Hiroyuki Nakahara and Kenji Sakino
Practical Analysis for High-Strength CFT Columns Under Eccentric Compression
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 441-448, March 22-24, 2000
[21]-1999
Kenji Sakino and Hiroyuki Nakahara
Flexu ral Capacities of Concrete Filled Square Steel Tubular Beam-Columns with High Strength Concrete
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 473-480, March 22-24, 2000[22]-1999
Shosuke Morino
Recent Development on CFT Column Systems - US-Japan Cooperative Earthquake Research Program
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 531-538, March 22-24, 2000[23]-1999
Toshiaki Fujimoto, Eiichi Inai, Hiroyoshi Tokinoya, Makoto Kai, Koji Mori, Osamu Mori and Isao Nishiyama
Behavior of Beam-to-Column Connection of CFT Column System under Seismic Force
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 557-564, March 22-24, 2000[24]-1999
Toshiyuki Fukumoto and Koji Morita
Elasto-Plastic Behavior of Steel Beam to Square Concrete Filled Steel Tube (CFT) Column Connections
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 565-572, March 22-24, 2000[25]-1999
Eiichi Inai, Takashi Noguchi, Osamu Mori and Toshiaki Fujimoto- 175 -

Deformation Capacity and Hysteresis Model of Concrete Filled Steel Tubular Beam-Columns
Proceedings of the Sixt h ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 1, pp. 605-612, March 22-24, 2000
[26]-1999
Nozomu Baba and Yasushi Nishimura
Stress Transfer on Through Beam Type Steel Beam - Reinforced Concrete Column Joints
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 753-760, March 22-24, 2000[27]-1999
Subhash Goel, Atsuo Tanaka, Hiroyuki Yamanouchi and Hiroshi Fukuyama
Research on New Materials, Elements and Systems: An Overview
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A . Mahin, Los Angeles, California, USA, Vol. 2, pp. 961-968, March 22-24, 2000
[28]-1999
H. Fukuyama, Y. Matsuzaki, Y. Sato and H. Suwada
Structural Performance of Engineered Cementitious Composite Elements
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 969-976, March 22-24, 2000[29]-1999
K. Kobayashi, Y. Matsuzaki, H. Fukuyama, S. Hakuto, T. Toritani and H. Kumagai
Performance Evaluation of RC Elements with Ultra Light Weight Concrete
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 977-984, March 22-24, 2000[30]-1999
Hiroshi Noguchi and Kazuhiro Uchida
FEM Analysis of Hybrid Structural Frames with R/C Columns and Steel Beams
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 1099-1106, March 22-24, 2000[31]-1999
Kenji Yonezawa and Kenzo Yoshioka
Three Dimensional Finite Element Analysis on CFT Column - Steel Beam Joints
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 1123-1130, March 22-24, 2000[32]-1999
Hiroshi Kuramoto, Isao Nishiyama and Hiroshi Noguchi
Outline of Japanese Design Guidelines for Composite RCS Buildings
Proceedings of the Sixth ASCCS International Conference on Steel-Concrete Composite Structures, Edited by Y.
Xiao and S. A. Mahin, Los Angeles, California, USA, Vol. 2, pp. 1183-1190, March 22-24, 2000

- 176 -