Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, June 2009 / Copyright © 2009 Japan Concrete Institute 157

Invited paper Mechanical Properties of Concrete and Reinforcement ― State-of-the-art Report on HSC and HSS in Japan ― Minehiro Nishiyama1

Received 25 May 2009, revised 12 June 2009 Abstract Sub-working group 1 of the Japan Concrete Institute (JCI) committee on Utilization of High-Strength Concrete (HSC) and High-Performance Concrete (HPC) (JCI-TC063A) has been working on collecting information from research on HSC and HPC as well as their practical use. The committee published a report in 2006 in Japanese. The objective of the report is to present state-of-the-art information on concrete with strengths in excess of about 60 MPa and high-strength steel reinforcement such as prestressing steel, excluding concrete made with atypical materials, such as fibers, or un- common techniques. It primarily addresses the mechanical properties of high-strength concrete and high-strength steel reinforcement. A concise digest of the report has already been published as a keynote paper in the proceedings of 8th International Symposium on Utilization of HSC and HPC in October 2008 in Tokyo. This paper is based on the keynote paper with additional information on High Strength Steel, HSS, reinforcing bars in Japan.

1. Introduction joints, structural walls, piles and structures constructed of HSC, including structural performance evaluation and Sub-working group 1 of the Japan Concrete Institute design of those structures. (JCI) committee on Utilization of HSC and HPC The committee published the 560-page state-of-art (JCI-TC063A) has collected and published report on HSC/HPC in 2006 in Japanese. The state-of-the-art information on concrete and steel rein- sub-working group of the committee is planning to pub- forcement with strengths in excess of 60 MPa and 500 lish an English version within a few years. This paper is a MPa, respectively. The report excludes concrete made concise digest of the report. with atypical materials, such as fibers, or uncommon This paper starts with mechanical properties of con- techniques. crete and reinforcement in the next chapter. The chapter In this report concrete strength in excess of 60 MPa in addresses compressive strength, fatigue strength, the field of building engineering and 80 MPa in the field modulus of elasticity, stress-strain behavior, shrinkage, of civil engineering is defined as high-strength concrete. creep, and fire resistance of concrete, which is followed Concrete up to 60 MPa and 80 MPa in design compres- by a chapter of classification, grades and use of rein- sive strength has been already incorporated in the codes forcing steel in Japan. Mechanical properties and fire of the Architectural Institute of Japan (AIJ) and the Japan resistance of reinforcing steel are also included in the Society of Civil Engineers (JSCE), respectively. chapter. In Chapter 4, structural performance of members In the building engineering field, super high-rise and frames is addressed in terms of uniaxial behavior of buildings, longer span beams, and reduction of member confined high-strength concrete, flexural and shear be- sections primarily induce the use of HSC. In the field of havior of beams and columns constructed of civil engineering longer span girders, reduction of girder high-strength materials. Beam-column joints and struc- sections and weight especially in bridges and tanks, as tural walls are also stated. As an application of well as construction of highly durable structures, are high-strength materials, design of super-high-rise build- motivations for the use of HSC. ings is addressed in the last chapter. The mission of the sub-working group 1 is to compile significant information on HSC not only as a material, 2. Mechanical properties of concrete but also its practical use for structures. The aims of the group are; This chapter summarizes HSC, reinforcement and bond 1) to collect information on state-of-the research and between concrete and reinforcement based on a literature practical use of HSC and, survey of past research. 2) to compile information on structural performance of In the first section compressive strength of HSC is members such as beams, girders, columns, beam-column described considering relations with water-binder ratio, admixtures, coarse aggregate, and age. There is a ceiling

on compressive strength which can be reached with a 1 reduction of the water-binder ratio. The ceiling is con- Professor, Dept. of Architecture and Architectural sidered to be approximately 120 N/mm2. To obtain Engineering, Kyoto University, Kyoto, Japan. higher strength, enhancement of cementitious material E-mail:[email protected] 158 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 and aggregate properties and curing conditions are essential. The above sections are followed by relationships be-

) tween compressive strength and splitting-tensile and 2 flexural strengths. Fatigue and elastic properties are m

m discussed in relation to the stress-strain response. /

N Shrinkage, creep, durability and fire resistance properties (

h are also summarized. Autogenous shrinkage and explo- t

g sive spalling-off of cover concrete are important issues n

e

r that should be discussed for application of HSC to t

s

structural members. e

v Summarized in the subsequent sections are classifica- i

s tion, grades and use of high-strength reinforcement in s e 1 week

r

Japan. Mechanical properties, including those at elevated p 4 weeks temperatures, and connections for high-strength rein- m 13 weeks

o

forcing bars are presented. C Relationships between bond strength and concrete Binder-water ratio compressive strength as well as between bond stress and slip are described based on a literature survey. Experi- mental results and design equations for bond splitting Water-binder ratio (%) failure and lap splice strength are also summarized. Fig. 1 Relationship between water-binder ratio and com-

pressive strength (Tomosawa et al. 1994). 2.1 Compressive strength 2.1.1 Relationship between water-cement ratio and compressive strength In the range of water-binder ratio larger than 25%, com- 140 Standard curing, W/C=0.45

) pressive strength of concrete increases in proportion to 2 120 binder-water ratio. However, as shown in Fig. 1, in the m

m range of water-binder ratio less than 25%, compressive / 100 High-belite cement (fine powder type)

N strength peaks at about 120 N/mm2 (Tomosawa et al. ( High-belite cement (low heat type)

h Ordinary portland cement t 80

1994). Moreover, even if their water-cement ratios are g

n the same, compressive strength of concrete with large e 91-day

r

t 60

unit water content is smaller than that of concrete with s small water content. In Fig. 1, concrete was made from e v 28-day i 40

N: ordinary portland cement, SF: cement with silica s

s

e fume, and BS: cement with blast-furnace slag. r 20 7-day p

m

2.1.2 Relationship between cement type and o 0 3-day compressive strength C 20 30 40 50 60 70 For concrete with water-cement ratio larger than 30%, C2S content (%) the strength-gain rate is different depending on the ce- ment type. However, with the decrease of water-cement 140

) Standard curing, W/C=0.25

ratio to about 25%, the difference becomes smaller (To- 2 120 mosawa 1994). The relationship between the content of m 91-day

m belite (C2S) and the compressive strength are shown in /

N 100 Fig. 2. The results indicate that the compressive strength ( 28-day

h at early ages (within 28-day) decreases as C S content t

2 g 80 increases. However, at an age of 91-days, the compres- n 7-day

e

r sive strength increases with a positive increment of C S t 60 2 s content (Uchida 1997). e 3-day

v i 40

s

s

2.1.3 Relationship between admixture type and e

r compressive strength p 20

m

Shown in Fig. 3 is the relationship between compressive o 0 strength and silica fume replacement ratio with regard to C 20 30 40 50 60 70 curing conditions (Nagataki 1988). It is revealed that C2S content (%) compressive strength peaks at a different ratio depending on the curing conditions. Fig. 2 Relationship between C2S content and compres- Shown in Fig. 4 is the case of fly ash, in which the sive strength at 3, 7, 28 and 91 days (Uchida 1997). M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 159 compressive strength obtained varies depending on the 2.1.6 Modulus of elasticity curing condition (Nagataki et al. 1988). In case of steam A formula that is able to estimate modulus of elasticity or 28-day water curing, compressive strength decreases applicable for both normal and high-strength concrete is as the replacement ratio increases. In autoclave curing, shown in Eq.3 (Noguchi and Tomosawa 1995b). the maximum compressive strength is obtained at the 2 1/3 replacement ratio of 40%. 4 ⎛ γ ⎞ ⎛ σ b ⎞ E = 3.35 ×10 ⋅ k1k2 ⎜ ⎟ ⎜ ⎟ (3) ⎝ 2.4 ⎠ ⎝ 60 ⎠ 2.1.4 Relationship between splitting tensile strength and compressive strength where, E: modulus of elasticity (N/mm2), γ: mass per unit 3 Relationship between splitting tensile strength and volume (t/m ), k1: coarse aggregate coefficient, 0.95: compressive strength is approximated by Eq.1 (Noguchi quartz schist crushed stone, crushed andesite stone, and Tomosawa 1995a). crushed cobblestone, crushed basalt stone, and crushed clay slate stone, 1.2: crushed limestone and calcination 0.637 σ t = 0.291σ b (1) bauxite, 1.0: other coarse aggregate, k2: admixture coeffi-

2 cient, 0.95: silica fume, blast-furnace slag powder, and micro where, σ t : splitting tensile strength (N/mm ), σ b : 2 powder made from fly ash, 1.1: fly ash, 1.0: no admixture. compressive strength (N/mm ). The equations (1) to (3) are derived from experimental 2.1.7 Stress-strain relationship in compression results on concrete with compressive strength ranging 2 2 Figure 5 shows examples of stress-strain relationships from 27 N/mm to 160 N/mm collected from 88 tech- obtained experimentally and theoretically (Kent and Park nical papers and reports such as AIJ and JCI journals. 1972; Fafitis and Shah 1985; Muguruma and Nagai The water cement ratio ranged from 15 % to 75 %, unit 3 3 1976; Popovics 1973). For low strength concrete the water content from 108 kg/m to 232 kg/m and age of idealized curves show good agreement with the experi- one day to 21 years. The equation was obtained from mental results. However, for high-strength concrete the concrete with compressive strength of up to about 160 2 2 stress-strain curve proposed by Muguruma and Nagai is N/mm . Up to 50 N/mm , the ratio of splitting tensile the best fit to the experimental results. strength to compressive strength is about 1/12, and splitting tensile strength is approximately proportional to 2.1.8 Shrinkage compressive strength. However, the ratio becomes 2 Shrinkage of concrete can be divided into two compo- smaller as σ b increases beyond 50 N/mm . The ratio is 2 nents: autogenous shrinkage which develops during about 1/8 in the vicinity of 100 N/mm , and the scatter cement hydration process, and drying shrinkage. Auto- becomes more significant. genous shrinkage was reported in 1930s and 1940s, however it has usually been neglected because the strain 2.1.5 Relationship between flexural strength was at most 100x10-6 in concrete with a water-cement and compressive strength The relationship between compressive strength and flexural strength (modulus of rupture) is given by Eq.2 Autoclave curing (180°C - 5 hours) 80 (Noguchi and Tomosawa 1995a).

0.678 σ f = 0.440σ b (2)

)

2

m where, σ f : flexural strength (modulus of rupture) 60

m 2 / Standard curing

(N/mm ). N ( for 28 days

h

t

g

n

) W/(C+Si)=0.3

e

2

150 r t 40 m Standard Steam cur- Steam cur- Steam cur- 28 s

m curing ing 65°C ing 80°C e / ing 80°C

v

N i

( 28 3 s

s h 28 28 1

t

100 e

g 28 r

n 3 3 Autoclave p Steam curing (65°C - 3 hours) e 3

r m 20 t 7 1 curing

o s 180°C

C e 3

v

i 50

s

s

e 1 Pre-curing -

r Pre curing p 4hours 24 hours

m 0 o 0 0 20 40 60

C 0503030 0500503030 05030050 Silica fume replacement (%) Fly ash replacement (%) Fig. 3 Relationship between silica fume replacement ratio Fig. 4 Relationship between fly ash replacement ratio and and compressive strength (Nagataki 1988). compressive strength (Nagataki et al. 1988). 160 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

150 150 150 150 Experiment Experiment Experiment Experiment

)

2

m 100 100 100 100

m

/

N

(

s

s

e

r t 50 50 50 50

S

0 0 0 0 0 2000 4000 6000 0 2000 4000 6000 0 2000 4000 6000 0 2000 4000 6000 Strain (x10-6) Strain (x10-6) Strain (x10-6) Strain (x10-6)

(a) Kent and Park (1972) (b) Fafitis and Shah (1985) (c) Muguruma and Nagai (1976) (d) Popovics (1973) Fig. 5 Comparison of stress-strain curves obtained experimentally and theoretically (Tomosawa et al. 1994). ratio of 50% or more. With increasing recent applications 0.56 of concrete with low water-cement ratio, autogenous ε ' t,t = ⎡1− exp −0.108 t − t ⎤ε ' cs ()o { ()o } sh (4) shrinkage has become an issue of concern. Considerable ⎣ ⎦ research has been conducted on autogenous shrinkage where, ε ' (t,t ): shrinkage strain from age of t0 to since 1990 in Japan, and state-of-the-art reports have cs o been published by technical committees organized in Japan Concrete Institute (JCI 1996; JCI 2002). 600 Shown in Figs. 6 and 7 is the effect of water-cement 40% 50% 80% 100% 500 ratio on autogenous shrinkage and drying shrinkage. )

6

- Contribution of Lower water cement ratio results in larger autogenous 0 autogenous shrinkage

1

x shrinkage. For a water-cement ratio of 0.17, autogenous ( 400

n shrinkage accounts for large percentage of total shrink- i

a

r

t age strain. Autogenous shrinkage develops rapidly at an s 300

e

early age and converges to the final value within one to g W/C=0.17, SF10%

a three days as shown in Fig. 8. k W/C=0.23, SF10%

n

i 200 r W/C=0.30 Sakata et al. (JSCE 2000) proposed the following h

s

l W/C=0.40

equation to predict shrinkage strain for concrete with a a t 100 o Autogenous shrinkage: Sealed curing water cement ratio between 40% and 65%. The equation T has been adopted by Standard Specifications for Con- Total shrinkage: 20°C, 60%R.H. 0 crete Structures of JSCE (JSCE 2002) for concrete with a 0100200 300 400 500 600 compressive strength less than 55 N/mm2. Autogenous shrinkage strain(x10-6) Fig. 7 Ratio of autogenous shrinkage to total shrinkage 1200 (Tazawa and Miyazawa 1997).

Total shrinkage

1000 Autogenous shrinkage ) 6 -200 - 0 W/C=0.50

Drying shrinkage 1

x 800 ( 0

)

n

i

6

-

a

0 r W/C=0.40

t

1 200

s

x

(

600 e

g

n

i a 400

k

a

r

n

t

i

r W/C=0.30

S 400 h 600

s

s

u

o 800 W/C=0.23 200 n

e

g

o

t 1000 W/C=0.20 0 u 0.1 0.3 0.5 0.7 A 0 20 40 60 80 100 Water-cement ratio Age (day) Fig. 6 Relationship between water cement ratio and Fig. 8 Development of autogenous shrinkage with age autogenous shrinkage and drying shrinkage (Miyazawa (Tazawa and Miyazawa 1997). 2005). M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 161

-6 -6 t (x10 ), ε 'sh : final shrinkage strain (x10 ) Figure 9 and 10 show the relationship between the coefficient β and unit water content and V/S, respectively. ε 'sh =−50 + 78{} 1− exp()RH /100 V/S is considered to have less effect on final drying 2 shrinkage strain for high-strength concrete than that for ⎡ V ⎤ (5) 38 log W 5log + e − ⎢ e ⎥ normal-strength concrete. V/S affects only the strain ⎣ 10S ⎦ development rate for drying shrinkage. Shown in Fig. 11 is the comparison between the pre- t0, t: effective age (days) at the beginning of drying and effective age during drying calculated by dicted and the measured drying shrinkages. Equation 8 can predict drying shrinkage within ±40% for drying n ⎪⎧ 4000 ⎪⎫ shrinkage data collected in Japan. A comparison of the to or t =Δti ⋅exp⎨ 13.65 − ⎬ (6) predicted drying shrinkage with the RILEM database is ∑ 273+ T Δt /T i=1 ⎩⎪ ()i o ⎭⎪ illustrated in Fig. 12. Equation 14 for prediction of autogenous shrinkage where, T = 1 °C, T: Temperature (°C), Δt : days under 0 i was developed mainly based on the research by Tazawa the temperature of T, RH: relative humidity (%), 45 % and Miyazawa (1997), and has been adopted in Standard ≤ RH≤ 80%, W: unit water content (kg/m3), 130 kg/m3 Specifications for Concrete Structures by JSCE (2002). ≤W≤230 kg/m3, V: volume of concrete (mm3), S: surface According to their research results autogenous shrinkage area of concrete exposed to the atmosphere (mm2), V/S: is strongly dependent on type of cement and wa- volume-to-surface area ratio (mm), 25 mm ≤V/S ter-cement ratio. Effect of aggregate unit volume and ≤300mm. 2 specimen size can be ignored. Figure 13 shows rela- For concrete with strength ranging from 55 N/mm to tionship between autogenous shrinkage and water ce- 80 N/mm2, drying and autogenous shrinkages are pre- dicted (JSCE 2002). The total shrinkage strain ε 'cs (t,to ) is estimated by the sum of these strain components as follows, 50 β Volume/surface ratio=25mm

:

r

o

t 45 t =14 days ε 'cs ()t,to = ε 'ds ()t,to + ε 'as (t,to ) (7) 0

c

a

f

-6 t where, ε 'ds ()t,to : drying shrinkage strain (x10 ) from n 40

e age of t0 to t, ε 'as ()t,to : autogenous shrinkage strain m t =28 days -6 p 0

(x10 ) from age of t to t. o 0 l 35

e

v

ε ' ⋅ t − t e ds∞ ()o d ε ' t,t = 30 ds ()o (8) e

β + ()t − to g a t =56 days k 25 0

n

i ε ' : final drying shrinkage, r

ds∞ h

S 20 ε ' 170 180 190 200 210 220 230 ε ' = dsρ (9) ds∞ 3 1+η ⋅to Unit water content (kg/m ) Fig. 9 Relationship between unit water content and coef- α ()1− RH /100 W ficient β (JSCE 2000). ε 'dsρ = ⎧ 500 ⎫ (10) 1+150 exp ⎨− ⎬ 70 ⎩ fc '28()⎭ β

:

r

−4 o η = 10 15exp 0.007 ⋅ f '28 + 0.25W t {}( c ( )) (11) c

a

f

t

n 50

where, fc '28(): compressive strength of concrete at e 2 2 28 days (N/mm ), f '28≤80 N/mm , α: modification m c () p

o coefficient to consider the effect of cement type on l

e

v

shrinkage, α = 11: ordinary portland cement and e low-heat cement, α = 15: high-early strength portland d 30

e

g

cement, β: modification coefficient to consider a

k

n characteristics of shrinkage development, which was i

r obtained by a regression analysis of experimental data. h S 10 04080 120 160 4WV/ S β = (12) Volume/surface ratio (mm) 100 + 0.7t o Fig. 10 Relationship between V/S and coefficient β (JSCE 2000). 162 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

1,000 1000 μ

Number of data sets= 92 ) 91-day

6

-

+20 % 0 1 800

x

% (

n 750 40 i

+ a

r % t

0 s -2 600

e

g

a

k

n

i 500 r 400 Predicted

h

0% s -4

s

u

o

n

e 200

g

o

250 t

u A 0 0.1 0.2 0.3 0.4 0.5 0.6

Experimental shrinkage strain - strain shrinkage Experimental Water-cement ratio 0 0 250 500 750 1,000 Fig. 13 Relationship between autogenous shrinkage at 91 days and water cement ratio (Tazawa and Miyazawa 1997). Calculated shrinkage strain - μ

Fig. 11 Comparison of predicted and measured drying 1.0 shrinkage for data collected in Japan (JSCE 2000). 0.8 1,000

) 0.6

Data : RILEM t μ

(

β +40% 0.4 W/C=0.20 750 W/C=0.23 W/C=0.30 0.2 W/C=0.40 W/C=0.56 500 0 0.1 1.0 10 100 1000 0% -4 Age (day) Fig. 14 Strain development rate of autogenous shrinkage 250 with various water cement ratio (Tazawa and Miyazawa 1997).

Number of data sets= 52 Experimental shrinkage strain - - strain shrinkage Experimental 0 0 250 500 750 1,000 W/C a b Calculated shrinkage strain - μ 0.20 1.2 0.4 0.23 1.5 0.4 Fig. 12 Comparison of predicted drying shrinkage with 0.30 0.6 0.5 the data base of RILEM (JSCE 2000). 0.40 0.1 0.7 0.50 ≤ 0.03 0.8

Figure 15 shows relationships between the predicted ment ratio. Figure 14 shows the strain rate β (t) of and measured autogenous shrinkages for concrete with autogenous shrinkage for various water-cement ratios. water cement ratio of 20% to 56%. The equation can predict autogenous shrinkage within ±40%.

ε 'as ()t,to = ε 'as ()t − ε 'as ()to (13) Autogenous shrinkage is significantly dependent on cement type. Figure 16 shows measured autogenous b shrinkage of cement paste of various kinds of cement. ε ' ()t = γε ' β ()t = γε ' ⎡1− exp −at− t ⎤ (14) as as∞ as∞ ⎣ { ()s }⎦ The water cement ratio is 30%. From this experimental observation, the coefficient γ for Equation 14 is given as where, γ: coefficient to account for type of cement (de- follows: scribed below), ε 'as∞ : final value of autogenous γ = 1.0 for ordinary portland cement shrinkage. = 1.3 for high early strength portland cement ε 'as∞ = 3070exp{−7.2()W / C } (15) = 0.9 for moderate heat portland cement = 0.4 for low heat portland cement where, ts: age of initial setting, a, b: coefficients listed = 1.0 for type B ground granulated blast-furnace slag in the following table for ordinary portland cement. cement M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 163

W/C=0.30

)

6

-

0

1

x

(

e

g

a

k

n

i

r

h

s

s

u

o

n

e

g

o

t

u

a

d

e

r

u

s

a

e

M

Predicted autogenous shrinkage (x10-6) Fig. 15 Predicted and measured autogenous shrinkage for concrete with water cement ratio of 20% to 56% (Ta- zawa and Miyazawa 1997).

S: Anti-sulfate portland cement O: Oil-well cement Expansive admixtures and shrinkage reducing agents, G: Geothermal-well cement L: Low heat cement as well as low-heat portland cement, are effective to N: Ordinary portland cement A: Alumina cement reduce autogenous shrinkage. Figure 17 (Tazawa 1994) B: Type-B blast furnace slag cement W: White cement shows the reduction of autogenous shrinkage for cement M: Medium heat portland cement paste with 30% water-cement ratio using expansive ad- H: High early strength portland cement mixtures and shrinkage reducing agents. Expansive ad- Fig. 16 Measured autogenous shrinkage of cement paste mixtures compensate for early-age shrinkage strain such for various kinds of cement (Miyazawa and Tazawa 1996). as autogenous shrinkage, while shrinkage reducing agents compensate for longer-term shrinkage such as 1000

)

6 drying shrinkage. It is also reported that simultaneous -

0

1 use of these admixtures does not impair their effects. x ( 500 n

o

i

s Age (day)

2.1.9 Creep n a 12 4 10 21 28 100 200 400 1000

p

Specific creep strain of concrete decreases its final value x 0 e 7 and variation as the compressive strength increases as -

e shown in Fig. 18. Specific creep is dependent on relative g

a

k

n humidity. High humidity reduces specific creep, which is i 500

r

h known as Picket’s effect. Creep strain measured in 100% s N30-0-0

s relative humidity is defined as basic creep, and for lower u N30-0-2.0 (D1)

o

n 1000 N30-0-2.0 (D2) relative humidity, drying creep is added to the basic e

g N30-10-0 (E1)

o creep. t N30-10-0 (E2)

u

Creep strain is assumed to be a summation of basic A N30-10-0 (E3) 1500 creep and drying creep. An equation is derived by re- gression analyses on a database of creep test results. The Fig. 17 Effect of shrinkage reducing agent and expansive following equation, which is applicable to concrete up to admixture on reduction of autogenous shrinkage (Tazawa 80 MPa in compressive strength, was proposed by JSCE 1994). 308 committee (JSCE 2000).

C t,t ' = A ⋅log t − t '+1 r () e () (16) using the following equation.

-6 where, C ()t,t ' : specific creep strain (10 /MPa), A: Cdrying ⋅W (1− h) C r A = + basic coefficient determined by the regression analyses (17) ϕdrying + fc '()t ϕbasic + fc '()t (10-6/MPa), t : age of concrete (day), t’: age of loading start (day). Regression analyses on available creep test results The coefficient A can be expressed as a linear function were conducted to determine the coefficient A. From the of relative humidity and unit water content as shown in analyses, it was revealed that ϕdrying ≅ ϕbasic , and finally Fig. 19 and 20 (JSCE 2000). The coefficient A is ex- the following approximate equation was obtained. pressed as superposition of basic creep and drying creep 164 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

4W ()1− h + 350 3. Mechanical properties of reinforcing steel A = (18) 12 + f '()t c High-strength concrete and high-strength longitudinal where, W: unit water content of concrete (kg/m3), h: relative humidity (in decimal), f’c: compressive strength of concrete (MPa).

)

Figures 21 and 22 show the comparison of the pre- a

P

dicted and measured specific creep strains. The equation M

/ μ can predict specific creep strain with a precision of ±40%. (

2.1.10 Fire Resistance of Concrete (1) Reduction of compressive strength and Young’s Modulus due to high temperatures (Morita and Ni- shida 2005) Figure 23 indicates the reduction in compressive strength and Young’s modulus as the ambient tempera- ture is increased. The notation “S20-75” indicates the test results with an initial strength ranging from 20 to 75 N/mm2, and S120，SP120，C120 indicate the test results Fig. 19 Effect of relative humidity on coefficient A (JSCE with an initial strength of 120 N/mm2. The reduction of 2000). compressive strength is small up to 200 degrees Celsius. However, the deterioration is significant at temperatures over 400 degrees Celsius. On the other hand, Young’s

modulus decreases approximately in proportion to the )

a temperature experienced. These test results reveal that P

M the initial strength of the concrete has only a small influ- / μ ence on the residual strength and Young’s modulus ratios. (

(2) Concrete explosive spalling (Morita et al. 2001) Explosive spalling can occur when concrete is exposed to high temperatures. As shown in Fig. 24, the depth of concrete explosive spalling increases as the wa- ter-to-cement ratio decreases and, therefore, as the compressive strength increases. Mixing some kinds of (kg/m3) chemical fiber such as polypropylene fiber into high-strength concrete in excess of 60 N/mm2 is ex- Fig. 20 Effect of unit water content on coefficient A (JSCE perimentally confirmed in Fig. 24 to be effective in the 2000). reduction of concrete cover spalling. )

2 150

% +40 /(N/mm μ

)

a

P

M

/ 100 μ

(

0% -4

50

Number of data sets = 146 0 0 50 100 150

(MPa) Calculated specific creep strain - μ/(N/mm 2) Experimental - creep strain specific Fig. 18 Relationship between compressive strength of Fig. 21 Comparison of predicted and measured specific concrete and ultimate creep strain (JSCE 2000). creep (Experimental data collected in Japan (JSCE 2000)). M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 165 reinforcement are often utilized for columns under high the ultimate load carrying capacities. However, under axial load in high-rise reinforced concrete buildings. service load conditions, crack widths and deflections They can also be used for beams and girders to enhance should be controlled for horizontal bending members. The stress in the longitudinal reinforcement allowed

) under service load conditions is much smaller than the 2 2 250 yield strength. It is approximately 200 N/mm , which Data : RILEM corresponds to only 30% of the yield strength, for ex- %

/(N/mm 0 ample, when reinforcing bars with a specified yield μ 200 +4 strength of 685 N/mm2 are used. Large deformation from the service load condition would be required for

150 high-strength reinforcing bars to attain the yield strength at the ultimate state. Prestress is essential for making

% good use of high-strength reinforcement, by ensuring 40 100 - crack widths and deflections can be controlled within an acceptable deformation range. High-strength reinforcement up to 1275 N/mm2 in the 50 specified yield strength is often used as shear rein- forcement in columns and beams. Design equations by Number of data sets= 140 0 which the ultimate shear strength can be calculated are 0 50 100 150 200 250 proposed by AIJ (1999) and other organizations, and have been used in practice. Calculated specific creep strain - μ/(N/mm 2) Experimental - creep strain specific 3.1 Japanese practice of classification and Fig. 22 Comparison of predicted specific creep with the utilization of reinforcing steel data base of RILEM (JSCE 2000). According to the statistics report published by The Japan

Iron and Steel Federation (JISF 2005), electric arc fur-

o

i nace steel made up 25.6% (28.85 million tons) of the

t

a

r total crude steel production (83.63 million tons) in Japan

h t in 2005. Almost all reinforcing bars in Japan are pro-

g

n

e duced by electric arc furnaces. Approximately 10 million

r

t

s tons of the total electric arc furnace steel is used as re-

e

v inforcing bars.

i

s

s High-strength reinforcement with yield strengths

e

r 2 2

p ranging from 590 N/mm to 685 N/mm are used in

m practice today. However, Japan Industrial Standard (JIS)

o

c

l G 3112-1987 for reinforcement for reinforced concrete

a

u construction only specifies reinforcing bars up to 490

d i 2 s N/mm in yield strength. The mechanical properties are

e

R summarized in Table 1. The most recent edition of Elevated temperature (°C) Japanese Architectural Standard Specification JASS 5 Reinforced Concrete Work (JASS 5) published by AIJ has regulations or requirements for construction speci-

o

i

t fications for structures using reinforcing bars up to 490

a 2 r N/mm in specified yield strength.

s

u

l AIJ published “Recommendation for Practice of

u

d High-Strength Concrete” in January 2005 (AIJ 2005),

o

m which specifies bending, connection and anchorage de-

s ' tailing of reinforcement with a yield strength greater than

g

n 2

u or equal to 490 N/mm .

o

Y In Japan, ingredient adjustment and rolling condition

l

a can produce reinforcing bars with yield strengths up to

u

d 2 i 685 N/mm . Because additives may increase the cost, the

s e production of reinforcing bars with yield strengths

R greater than 685 N/mm2 is carried out by heat treatment Elevated temperature (°C) or cold drawing. Fig. 23 Residual compressive strength and Young’s A national research project on development of tech- modulus ratios after elevated temperature (Morita and nologies for Ultra-light and Ultra-high reinforced con- Nishida 2005). crete buildings (NewRC) started in 1988. Grade 685 and

166 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

2months Mean 1year All data

)

Age: m

m

2months (

h

t

p

e

D

Age: 1year

Measured spalling depth

W/C=27.5 31.7 37.5 44.9 48.6 64.4 (%) Fig. 24 Relations between spalling depth and W/C (Morita et al. 2001).

980 reinforcement for longitudinal reinforcing bars of Table 1 Mechanical properties specified in JIS G 3112 columns and beams were developed during the project. for reinforcement. For transverse reinforcement Grade 785 and 1275 were Yield strength or developed. Recommendations for ingredients and me- Tensile 0.2% offset yield Elongation chanical properties specifications was proposed. The Grade strength strength, % mechanical properties are summarized in Table 2 for N/mm2 N/mm2 longitudinal reinforcement and in Table 3 for transverse reinforcement. A prefix USD is used for high-strength SR235 ≥235 380~520 ≥20 or ≥24* deformed bars developed in the NewRC project, while prefixes SR or SD are used for reinforcement specified in JIS G 3112, which indicates a plain or deformed bar, SR295 ≥295 440~600 ≥18 or ≥20 respectively. Table 4 summarizes the mechanical properties of re- SD295A ≥295 440~600 ≥16 or ≥18 inforcing bars currently used in practice according to a report of General Building Research Center (GBRC 2005). SD295B 295~390 ≥440 ≥16 or ≥18 A summary regardless of the diameters is shown as follows ( σ y is the specified yield strength given by the grade designation), SD345 345~440 ≥490 ≥18 or ≥20 SD295A (D10~D22): x = 1.231σ ，Standard deviation σ = 0.051σ SD390 390~510 ≥560 ≥16 or ≥18 y y SD345 (D10~D51): x = 1.125σ ，Standard deviation σ = 0.039σ SD490 490~625 ≥620 ≥12 or ≥14 y y SD390 (D25~D38): *Depends on test coupon used. x = 1.108σ ，Standard deviation σ = 0.021σ y y SD490 (D35~D38): Table 2 NewRC Specifications of mechanical properties of high strength reinforcing bars (longitudinal x = 1.067σ y ，Standard deviation σ = 0.008σ y reinforcement). The higher strength bars result in a smaller ratio of the Yield Elongation average yield strength to the specified yield strength. The Yield strength or ratio of the standard deviation to the specified yield strength/ Length of 0.2% offset Grade tensile yield Elongation strength has the same trend. However, the number of yield strength plateau (%) tests for SD390 and SD490 is much smaller than the strength (%) (%) other two grades. N/mm2 According to the report “Research on performance of gas pressure welding for SD490 reinforcement” pub- USD685A 685~785 ≥85 ≥1.4 ≥10 lished by Japan Pressure Welding Society, the following statistics can be summarized; USD980 ≥980 ≥95 － ≥7 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 167

less than or equal to 12.6mm) are classified into three Table 3 NewRC Specifications of mechanical properties categories of B, C, and D according to their tensile of high strength reinforcing bars (transverse reinforcement). strengths as shown in Table 6. They are further divided Yield strength or Tensile into two categories normal (N) and low (L) according to 0.2% offset Elongation Grade strength their relaxation ratio. yield strength 2 (%) 2 N/mm Mechanical properties of seven-wire and 19-wire N/mm strands specified in JIS G3536 are summarized in Table USD785 ≥785 ≥930 ≥8 7 and 8, respectively. The relaxation ratios for normal- and low-relaxation steels are less than or equal to 8.0% USD1275 ≥1275 ≥1420 ≥7 and 2.5%, respectively. The production processes for prestressing steel bars are summarized in Table 9. SD345: number of tests n = 111, average x = 1.108σ y , Table 10 indicates statistics of the mechanical prop- Standard deviation σ = 0.038σ erties of smaller diameter prestressing steel bars, pri- y marily used for shear reinforcement in high-strength 1.107 SD390: number of tests n = 70 , average x = σ y , concrete members. The ratio of the average yield strength to the specified Standard deviation σ = 0.045σ y yield strength 1275 N/mm2 and the standard deviations are summarized as follows, SD490: number of tests n = 176 , average x = 1.077σ y ,

Standard deviation σ = 0.023σ SBPDN1275/1420 (9.0mm): x = 1.115σ y , standard y deviation σ = 0.013σ y 3.2 Prestressing steel SBPDN1275/1420 (10.7mm): x = 1.104σ , standard JIS has three categories for prestressing steel: JIS G 3109 y for prestressing steel bars; JIS G 3137 for small diameter deviation σ = 0.011σ y deformed prestressing steel bars; and JIS G 3536 for SBPDN1275/1420 (12.6mm): x = 1.067σ , standard prestressing steel wires and strands. y Prestressing steel bars are classified into three cate- deviation σ = 0.008σ y gories of A, B, and C according to their yield strengths. The elongation of the prestressing steel bars shown in The yield strength is defined as the stress corresponding Table 10 (at most 9.4%) is much smaller than that ob- to a 0.2% offset. They are further classified into two served for ordinary strength deformed bars in Table 4 categories of No.1 and No.2 according to the ratio of the (approximately 25%). yield strength to the tensile strength. In fact, bars pro- duced through heat treatment have a large yield strength 3.3 Mechanical properties of reinforcement to tensile strength ratio and consequently are classified as Typical stress-strain curves for reinforcing bars are No.1, while those produced using cold drawing or hot shown in Fig.25. The distinctive features can be sum- drawing are classified as No.2 due to their lower yield marized as follows, strength to tensile strength ratio. The mechanical prop- 1) Young’s modulus ranges from 2.0 x105 N/mm2 to erties are summarized in Table 5. 2.05x105 N/mm2 regardless of the reinforcement Smaller diameter deformed bars (nominal diameter type.

Table 4 Mechanical properties of reinforcing bars tested in GBRC (GBRC 2005). Yield strength (N/mm2) Tensile strength (N/mm2) Elongation at fracture (%) Number of Standard Standard Standard Grade Diameter Average, Average, Average, tests deviation, deviation, deviation,

x σ x σ x σ D10~D13 241 365 15.5 519 18.6 27 1.9 SD295A D16~D22 93 358 11.7 518 16.2 26 2.4 Sub total 334 363 14.9 519 18.0 27 2.1 D10~D13 129 386 15.4 552 23.4 25 1.9 D16~D22 567 387 13.1 570 22.3 24 2.2 D25~D32 309 388 12.4 576 20.8 27 2.1 SD345 D35~D38 60 396 9.9 591 11.1 26 2.2 D41~D51 84 392 14.6 586 11.5 25 1.6 Sub total 1149 388 13.4 572 22.9 25 2.5 D25~D32 29 431 8.4 624 17.5 23 2.0 SD390 D35~D38 3 437 0.6 647 0.6 22 0.6 Sub total 32 432 8.2 626 17.9 23 1.9 SD490 D35~D38 3 523 3.8 711 0.9 21 0.0 168 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

Table 5 Mechanical properties of prestressing steel bars. 0.2% offset Tensile strength Elongation Relaxation* Grade yield strength N/mm2 % % N/mm2 A No.2 SBPR 785/1030 ≥785 ≥1030 No.1 SBPR 930/1080 ≥930 ≥1080 B ≥5 ≤4.0 No.2 SBPR 930/1180 ≥930 ≥1180 C No.1 SBPR 1080/1230 ≥1080 ≥1230 * Relaxation after 1,000 hours under the tensile stress of 60% of yield strength.

Table 6 Mechanical properties of smaller diameter deformed prestressing steel bars (φ≤12.6mm). 0.2% offset Tensile strength Elongation Relaxation* Grade yield strength N/mm2 % % N/mm2 SBPDN 930/1080 ≤4.0 B ≥930 ≥1080 SBPDL 930/1080 ≤2.5 SBPDN 1080/1230 ≤4.0 C No.1 ≥1080 ≥1230 ≥5 SBPDL 1080/1230 ≤2.5 SBPDN 1275/1420 ≤4.0 D ≥1275 ≥1420 SBPDL 1275/1420 ≤2.5 * Relaxation after 1,000 hours under the tensile stress of 60% of yield strength.

Table 7 Mechanical properties of seven-wire strands. Table 8 Mechanical properties of 19-wire strands. 0.2% offset 0.2% offset Tensile Tensile Diameter Elongation Diameter yield Elongation Grade yield strength strength strength mm % mm strength % kN kN kN SWPR 9.3 ≥75.5 ≥88.8 kN 7AN 10.8 ≥102 ≥120 17.8 ≥330 ≥387 A SWPR SWPR 19.3 ≥387 ≥451 12.4 ≥136 ≥160 19N 7AL 20.3 ≥422 ≥495 ≥3.5 15.2 ≥204 ≥240 SWPR ≥3.5 21.8 ≥495 ≥573 SWPR 9.3 ≥86.8 ≥102 19L 7BN 10.8 ≥118 ≥138 28.6 ≥807 ≥949 B SWPR 12.4 ≥156 ≥183 * Relaxation after 1,000 hours under the tensile stress of 7BL 15.2 ≥222 ≥261 60% of yield strength. * Relaxation after 1,000 hours under the tensile stress of 60% yield of strength Table 9 Production processes of prestressing steel bars. 2) Reinforcing bars lower than Grade 785 show clear Production processes yield points, while higher grades do not show a Products (Base material: round plain hot-rolled, definite yield point. The NewRC specifications re- non-alloyed, high-carbon steel rod） quire Grade USD685 to have a yield plateau of more Hot-rolled Stretching - bluing than 1.4%. Quenching - tempering 3) The higher the yield strength, the larger the ratio of Heat-treatment Cold drawing- quenching - tempering yield strength to tensile strength. SD785 has the ratio of 0.85 and some prestressing steels have a ratio Cold-drawing Cold drawing - bluing greater than 0.9. We should bear in mind that the rotational capacity of a member reinforced with high-strength longitudinal steel becomes smaller. 3.4 Mechanical properties of reinforcing steel at 4) The higher the yield strength, the smaller the elon- elevated temperatures gation at fracture. Reinforcement higher than Grade JIS G 0567 specifies the test procedure for reinforcement 490 conforming to JIS G 3112 should have an at elevated temperatures. Most tests reported in the lit- elongation strain at fracture larger than 12~24%. In erature were conducted on structural steel, not reinforc- the NewRC specifications, USD685 and USD980 ing bars or prestressing steel bars. Only limited research should have elongation strains larger than or equal to is found on reinforcing steels. 10% and 7%, respectively. Bending workability dete- Japanese Society of Steel Construction published a riorates as the grade of reinforcement becomes higher. report on mechanical properties of reinforcing bars and M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 169

Table 10 Statistics of mechanical properties for small diameter prestressing steel bars. Yield strength (N/mm2) Tensile strength (N/mm2) Elongation (%) Diameter, Number of Standard Standard Standard Grade Average, Average, Average, mm tests, n deviation, deviation, deviation, x x x σ σ σ 9.0 63 1421 17.2 1493 17.2 9.4 0.5 SBPDN 10.7 100 1408 14.0 1484 10.4 9.4 0.5 1275/1420 12.6 100 1361 10.6 1473 9.5 8.9 0.4

Table 11 Tensile tests of steel at elevated temperatures. Diameter Strain rate（%/min） Temperature Literature Steel or grade (mm) Up to yielding After yielding (°C) SR235 φ22.0 0.16~0.29 7.5 SR295 φ22.0 0.19~0.22 7.5 20, 100, 200, 300, 350, 400, JSSC 1969 SD345 D25 0.07~0.30 7.0~10.0 450, 500, 550, 600 SD390 D25 0.20~0.30 6.7~10.0 Nagao et al. SD390 D16 JIS G0567 20, 200, 300, 400, 600, 800, 2000 SD490 D16 JIS G0567 1000 Baba et al. Normal temperature, 100, 200 USD685 D22 0.5 2002 300, 400, 500, 600, 800, 1000 Matsudo et USD685 D19 Normal temperature, 200, 300, 0.3 al. 2003 SBPD1275 U9.0 350, 400, 500, 600, 800, 1000 SBPR785/1030 φ23.0 0.22 74 SBPR930/1080 φ24.0 0.2 10 20, 100, 200, 300, 350, 400, JSSC 1969 SBPR930/1180 φ23.0 0.22 74 450, 500, 550, 600 SBPR1080/1230 φ24.0 0.2 10 prestressing steel bars at elevated temperatures in 1969 SBPR930/1080, SBPR930/1180, SBPR1080/1230. (JSSC 1969). The report also includes properties after Since then, several studies have been done on returning a specimen to room temperature. The following stress-strain idealization of reinforcement at elevated reinforcing and prestressing bars are included in the temperatures and experimental data has been accumu- report: SR235, SR295, SD345, SD390, SBPR785/1030, lated. However, the data are still not enough. For prestressing steel, Michikoshi et al. (2003) provided only the residual strength after specimens were cooled down Prestressing bar Elastic 2000 (0.92 twisted) Plastic to room temperature. region region Nagao et al. (2000) conducted tensile tests on SD390 Prestressing bar (0.90) Plastic flow Hardening and SD490 D16 reinforcing bars at room temperature, region 200, 300, 400, 600, 800, and 1000 degrees Celsius. The

s

s

Prestressing e r figures below show yield strengths and tensile strengths

1500 t bar (0.88) S plotted against the temperature. Figure 26 below shows

) 2 yield strengths and tensile strengths plotted against

m

m temperature. Figure 27 depicts ratios of yield strength / Prestressing

N ( 1000 bar type B and tensile strength at elevated temperatures to those at s Strain s room temperature. Nagao et al. concluded that there is no

e

r SD785 (0.85) t SD590 (0.75) difference between SD390 and SD490 reinforcement.

S SD490 (0.70) Baba et al. (2002) conducted tensile tests on USD685 SD390 (0.67) at room temperature, 100, 200, 300, 400, 500, 600, 800, 500 SD345 (0.67) and 1000 degrees Celsius. The results are indicated in SD295 (0.65) Fig.28 (a) in terms of stress at 1% strain and tensile Yield ratio in parenthesis strength. Residual strength ratios, which are defined as the ratios of stress at 1% strain and tensile strength after 0 10 20 30 specimens are cooled down to those at room temperature, Strain (x10-3 ) are also shown in Fig.28 (b). Fig. 25 Stress-strain relations for reinforcement used in Figures 26 and 28 indicate that the strength decay of Japan (Usami 2000). SD390 and SD490 at elevated temperatures is almost the 170 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 same as that of USD685. Up to 300 degrees Celsius no 720 − T k = for reinforcing steel reduction can be observed. Between 300 and 800 degrees y,T 470 Celsius an approximately linear relation can be found. These characteristics can be idealized by the following 700 − T k = for prestressing steel equations, which are specified for structural steel SS400 y,T 550 and SM490 in Recommendation for Fire Resistant De- sign of Steel Structures published by AIJ (2008). Ito et al. (2005) conducted tensile tests on prestressing steel bars of 13 mm and 9.2 mm in diameter. The bars F ()T = F ×κ ()T were Grade C No.1 and produced by heat-treatment. The s s loading rates were 0.5 kN/sec for 13mm bars, and 0.3

⎧ 1:TR ≤ T ≤ 300 kN/sec for 9.2 mm bars. The temperature of the bars was ⎪ κ ()T = ⎨ 750 − T elevated by 5 degrees Celsius per minute. The :300≤ T ≤ 750 stress-strain curves were obtained at room temperature, ⎩⎪ 450 100, 200, 300, 400, 500, and 600 degrees Celsius. Yield strength was not measured at the room temperature, where, Fs ()T : Effective yield strength of steel at T de- because the bars fractured within the threaded portion, gree Celsius, Fs : Design specified strength of steel, κ ()T : Strength reduction ratio at T degree Celsius, T : which was outside the gauge length for measuring strains. Temperature of steel, TR : Room temperature (20 degrees Celsius). Figure 29 shows the stress-strain curves obtained BS8110, AS4100 and NZ3404 give the following experimentally. As shown in this figure, yield and tensile equations for strength reduction ratio ky,T at T degree strengths, in addition to the modulus of elasticity, re- Celsius. ky,T is a ratio of yield strength at T degree Cel- duced significantly under high temperature. The sius to that at room temperature. ky,T should be smaller strengths before heating are reported to almost recover than or equal to unity. after cooling down from the elevated temperature up to 450 degrees Celsius. Prestressing steel bars are often

600 800

700 500 SD390 ) SD390

)

2 2 SD490 m 600 SD490

m

m

m 400 /

/

N

N 500

(

(

h

h

t

t

300 g

g 400

n

n

e

e

r

r

t

t

s 300 s 200

e

l

d

i

l

s

e

i 200

n

Y 100 e T 100 0 0 200 400 600 800 1000 0 200 400600 800 1000

Elevated temperature (°C) Elevated temperature (°C) Fig. 26 Yield strengths and tensile strengths under elevated temperatures (Nagao et al. 2000).

)

)

1.2 . 1.2

.

p

p

m

m

e SD390

t e SD390 t 1.0 1.0 SD490

m

m SD490

o

o

o

o

r

r

/

/ 0.8 0.8

d

d

e

e

t

t

a

a

v

v

0.6 e 0.6

l

e

l

e

e

(

(

o

i

o

i 0.4 t 0.4

t

a

a

r

r

h

h

t

t

g

g 0.2 0.2

n

n

e

e

r

r

t

t

s s 0 0

e

l

d

i l 0 200 400 600 800 1000 0 200 400600 800 1000

s

e

i

n

Y Elevated temperature (°C) e T Elevated temperature (°C) Fig. 27 Ratio of yield and tensile strengths at elevated temperatures to those under room temperature (Nagao et al. 2000). M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 171 used as shear reinforcement, which may be susceptible to increases. These properties can enhance the stiffness and elevated temperature during a fire. However, shear re- elastic limit of reinforced concrete structural members. inforcement is assumed not to be heated higher than 350 However, brittle failure beyond the peak of the degrees Celsius if a specified concrete cover is main- stress-strain curve has been pointed out as a drawback, tained. Explosive spalling of high-strength concrete at which can lead to lower ductility in structural members. elevated temperatures, however, should be taken into Large autogenous shrinkage in early age is observed in consideration. concrete in excess of 100 N/mm2 in compressive strength. This results in modulus of elasticity and flexural and 4. Structural performance of HSC members tensile strengths leveling off at approximately 100 N/mm2 although the compressive strength still increases High-strength concrete in excess of 60 N/mm2 in stan- beyond 100 N/mm2. dard cylinder compressive strength is used for bridge The surface of cracks caused by applied shear will girders and underground tanks. Columns and structural tend to be smoother in HSC than in NSC because coarse walls in the lower stories of super high-rise condomin- aggregates are sheared off along the crack surface rather ium buildings are constructed of high-strength concrete. than cracking through the cement paste. Strain concen- Spun concrete piles are also produced using HSC. trations are observed when concrete is about to fail in Modulus of elasticity and the strain at the peak of the compression, which can be another reason for the brittle stress-strain curve increase as the compressive strength failure of HSC. To overcome the above drawbacks, and to fully utilize 1.0 HSC for structural members, appropriate reinforcement

) . is needed. p 0.9 Strength at 1% strain

m The current Standard Specifications for Concrete

e Tensile strength t 0.8 Structures published by JSCE (2007) can be used for m 2 o concrete with compressive strength up to 80 N/mm . It

o

r / 0.7 enables performance evaluation of structural members,

d

e t including the effects of autogenous shrinkage. AIJ pub- a 0.6

v

e lishes a design standard called Standard for Structural

l

e ( 0.5 Calculation of Reinforced Concrete Structures (AIJ

o

i t 1999), which can be applied to concrete with a design a 0.4 r 2

h compressive strength up to 60 N/mm . AIJ has another

t

g 0.3 guideline Recommendation for Practice of

n

e

r

t High-Strength Concrete (AIJ 2005) published in 2005,

s 0.2 l which deals with concrete with compressive strengths up

a u 0.1 2 d to 100 N/mm . The recommendation places a focus on

i

s

e production and construction of HSC. R 0 200 400 600 800 1000 Chapter 4 of the JCI-TC063A report describes the Elevated temperature (°C) structural performance of beams, columns, beam-column joints, frames, structural walls and piles. Uniaxial com- 900 pression of columns confined by transverse reinforce- ment, flexure, shear and deformation capacity are sum- 800 Strength at 1% strain marized. Experimental investigations are summarized in Tensile strength which flexural-shear loading tests have been conducted 700 on simply supported bridge girders constructed using 2

) concrete strengths up to approximately 130 N/mm with 2 600

m or without shear reinforcement. Cyclic flexural-shear

m / 500 loading tests on columns using concrete strengths up to

N ( about 170 N/mm2 to simulate earthquake loading in

s s 400 e building structures are also summarized.

r

t

S 300 4.1 Beams and columns 200 4.1.1 Uniaxial compression behavior of con- fined high-strength concrete 100 (1) Triaxial compression tests on HSC The following formula is proposed to express the com- 0 0 200 400 600 800 1000 pressive strength σ cc of laterally confined concrete based on the unconfined compressive strength σ ' : Elevated temperature (°C) co

Fig. 28 Strength at 1% strain and tensile strength, and σ cc = σ co '+ k '⋅ p (19) residual strength after cool-down (Baba et al. 2002). 172 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 where, p :lateral pressure, k ' :constant. Richart et al. σ k ⋅ p (1928) proposed 4.1 for k ' . cc = 1 + (20) σ ' σ ' Experiments on HSC in excess of 100 N/mm2 have co co been conducted and reported that a similar equation to k = 21.2 − 0.05σ ' (21) Eq.19 can be used for HSC. The studies reported that the co ratio of longitudinal strain to lateral strain at the peak stress is approximately 2.2 regardless of the compressive (2) Uniaxial compression tests on reinforced con- strength and lateral pressure. crete columns Xie et al. (1995), for example, reported tri-axial Lateral confinement by transverse reinforcement is not as compression tests on concrete cylinders confined by effective as hydrostatic lateral pressure on concrete. lateral pressures up to 50% of the compressive strength. When transverse reinforcement is used for lateral con- The standard cylinder strengths were 60, 90, and 120 finement, k ' in Eq.19 should be smaller than 4.1 as 2 2 N/mm . Figure 30 shows the results for 120 N/mm proposed by Richart et al. (Saatcioglu and Razvi 1992). concrete. As the lateral pressure increases, the compres- Lateral pressure from transverse reinforcement is not sive strength and the strain at the peak in the stress-strain uniform along the longitudinal axis of the members as curve increase and the descending branch becomes well as in the member section (Mander et al. 1988). High milder without abrupt reduction in load carrying capacity. strength confining reinforcement does not always attain Xie et al. proposed the following equations, developed its yield strength although the confining effect is usually by regression analyses of the experimental data. calculated using its yield strength. φ =13mm Past research (Cusson and Paultre1994; Razvi and Staatcioglu 1999; Saatcioglu and Razvi 1998; Naga- shima et al. 1992; Itakura and Yagenji 1992; Li et al. 2001; Sakino and Sun 1993; Sun and Sakino 1993; Na- kazawa et al. 2003; Zhang and Mori 1997; Akiyama et al. 2004; Akiyama et al. 2005) surveyed includes columns whose water-to-binder ratio and concrete compressive strength are within the following ranges 25~35% and 80~120 N/mm2, respectively. Silica fume was used for concrete for which the compressive strength was higher than 80 N/mm2. The maximum aggregate size was 10~13mm. The test results can be summarized as fol- lows, a) In HSC columns reinforced appropriately with high-strength lateral confining steel, yielding of lateral reinforcement can be observed at the maxi- mum load. The shape and volume of lateral rein- forcement are of importance for successful use of high-strength lateral reinforcement. b) Failure was concentrated to a smaller region in HSC φ =9.2mm columns compared with NSC columns. HSC col- umns confined by a smaller amount of lateral rein- forcement experienced a further concentration of the failure region, and c) A stiff loading machine is needed to test HSC. If a flexible loading machine is used sudden and explo- sive failure immediately after peak stress and cover concrete spalling is inevitable.

4.1.2 Flexural properties (1) Flexural cracking strength Figure 31(a) compares the flexural cracking strengths observed in reinforced concrete members constructed of concrete with standard cylinder compressive strength in excess of 60 N/mm2 with theoretical values calculated using Eq.22. The experimental and calculated data gen- erally agree well, but considerable variation is observed. Fig. 29 Stress-strain relations of prestressing steel bars No correlation between the concrete compressive at elevated temperatures (Ito et al. 2005). strength and the ratio of the experimental results to the

M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 173

(2) Flexural crack width It has been reported that higher strength concrete can decrease flexural crack width and interval as a result of higher bond strength to reinforcement (Gokaku et al. 2001). Figure 32 shows flexural crack widths measured in axial tensile tests of reinforced concrete prisms in which concrete compressive strength was an experi- mental parameter. The measured crack widths at a rein- forcement stress of 300 N/mm2 (normalized by the crack width for specimen with a concrete compressive strength of 30 N/mm2) are plotted against concrete compressive strength. For low concrete stresses, it is known that crack width is suppressed by tension stiffening. Concrete strength is considered to considerably affect tension stiffening. In the above experiments it was difficult to evaluate how Fig. 30 Longitudinal stress versus longitudinal strain for large an effect concrete strength has on tension stiffening tri-axial compression tests for 120 N/mm2 concrete (Xie because the number of cracks was not yet at a steady state. 1995). For practical reasons, tensile strength of concrete is disre- garded in the crack width calculation proposed by JSCE. cracking strength from Eq.22 is observed, as seen in Fig.31 (b). However, the ratio measured to calculated )

m

•

cracking strength decreases as axial load ratio becomes N

k

(

large, as shown in Fig. 31 (c) (Sato et al. 2004). t

n

e

M = σ +σ Z m cr ()t o e (22) i

r

e

2 p where, σ : flexural tensile strength of concrete (N/mm ), x t e

, σ : compressive strength of concrete y σ t = 0.56 σ B B b

2 2 t

(N/mm ), σ o : axial stress (N/mm ), Z e : equivalent sec- n

e

tion modulus taking steel reinforcement into account m

3 o

(mm ) m

The JSCE Standard Specifications for Concrete g

n

i Negative loading

k

Structures-2002 “Structural Performance Verification” c

a Positive loading

r

(JSCE 2002) gives a member analysis method for ob- c

l

a

taining flexural cracking strength, which utilizes tension r

u softening and splitting tensile strength of concrete. x

e

l

However, because of the complexity of the JSCE ap- F proach, Eq.23 taking tension softening properties and Flexural cracking moment calculated with Eq.22(kN•m) member height into account as parameters, is proposed for obtaining a theoretical flexural cracking strength Fig. 31 (a) Comparison of flexural cracking moment ob- (Uchida et al.1992; Kato et al. 1992). tained experimentally and from Eq.22 (Sato et al. 2004).

⎧ ⎫

1 o

t fb = k0b ⋅ ft = ⎨1 + ⎬ ⋅ ft (23)

l

⎩ 0.85+ 4.5()h /lch ⎭ a

t

n where, f : flexural cracking strength, f : tensile e b t m

i

s

t

r strength of concrete, k : coefficient expressing the l 0b e

u

p s

relationship between flexural strength of concrete and its x e

r

e

tensile strength attributable to tension softening proper- d

f

e

t ties, h: height of member cross-section (m) (>0.2), l : o ch a

l

2 o

u

i characteristic length (m) ( = G E / f ), G : fracture t f c t f c

l

a energy (N/m) (area under the tension softening curve: in a

R 1/3 c general, G = 10 d f ' d : maximum dimen- f ()max c , max sion of coarse aggregate (mm), f c ' : compressive strength of concrete (N/mm2). Concrete compressive stress (N/mm2)

In the calculation of lch , the unit for elastic modulus Ec Fig. 31 (b) Effect of compressive strength on the ratio of 2 and f t is N/m . experimental to theoretical values (Sato et al. 2004).

174 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

1.4

o

0

t

3 l 1.2 15

c a k2 = +0.7

t F f' +20

n c

o

t e 1.0

d

m

i

e

s

t

r

z

l

i

l 0.8

e

u

a

p s

x e

m

r

r

e 0.6

o

d

f n

e

t

o

h

t a 0.4

l

d

o

i

u

i

t

c

w

l

a Self-compacting concrete

a k 0.2

c

R c

a r 0 C 0 20 40 60 80 100 120 Axial load ratio Concrete compressive stress (N/mm2) Fig. 31 (c) Effect of axial load on the ratio of experimental Fig. 32 Relationship between average crack width at 2 to theoretical values (Sato et al. 2004). stress in reinforcement of 300 N/mm and concrete compressive strength (Gokaku et al. 2001).

(3) Flexural strength compressive strength, N: axial load, b and D: width and Three design equations for calculation of flexural depth, g1: ratio of distance between tension and com- strength are compared: Equations 24-26 from Guidelines pression longitudinal reinforcements (sectional area of published by the Building Center of Japan (BCJ 1991), each is ag /2) to D. Eq. 27 from ACI318-08 (ACI 2008) and Eq.28 from NZS3101:2006 (NZS 2006). Equations 24-26 give an ACI318 rectangular stress block: (27) approximate estimation of flexural strength for a member Uniformly distributed concrete stress of 0.85 fc ' section with multi-layer longitudinal reinforcing bars The height a = β1c , c: neutral axis depth based on the plane-section assumption. Equation 27 and 28 give concrete stress-block coefficients for an extreme β1 = 0.85 when fc ' ≤ 27.5MPa compression fiber strain of 0.003. The plane-section f '− 27.5 β = 0.85 − 0.2 c when 27.5MPa N troduced for reducing compressive strength. In NZS3101, max b 0.85 of 0.85 f ' depends on concrete compressive c M u = {0.5 ⋅ ag ⋅σ y ⋅ g1 ⋅ D strength as shown below. ⎛ N − N ⎞ (24) 2 max α1 = 0.85 when fc ' ≤ 55MPa +0.024 ⋅()1+ g1 ⋅()3.6 − g1 ⋅b ⋅ D ⋅σ B }⎜ ⎟ ⎝ Nmax − Nb ⎠ α1 = 0.85 − 0.004(fc '− 55)≥ 0.75 when fc ' > 55MPa

Nb ≥ N ≥ 0 Figure 33 (Ghosh and Ahmad 2006) illustrates the

M u = 0.5 ⋅ ag ⋅σ y ⋅ g1 ⋅ D comparison of the stress intensity factor α1 between (25) ACI318, NZS3101, ACI Innovation Task Group 4 and ⎡ ⎤ +0.5 ⋅ N ⋅ D ⋅ ⎣1− {}N / ()b ⋅ D ⋅σ B ⎦ CSA A23.3. ACI Innovation Task Group 4 has proposed the reduction of α for concrete exceeding 55 N/mm2 in 0 > N ≥ N 1 min compressive strength. Flexural strengths were collected from 225 specimens M = 0.5 ⋅ a ⋅σ ⋅ g ⋅ D + 0.5 ⋅ N ⋅ g ⋅ D u g y 1 1 (26) appearing in literature in the past. Symbols in the legend → N = b ⋅ D ⋅σ + a ⋅σ of Fig. 34 to Fig. 39 indicate that F: flexural failure, F max B g y B: bond failure after flexural yielding, F→S: shear fail- ure after flexural yielding, F→Sc: brittle failure due to Nb = 0.22 ⋅()1+ g1 ⋅b ⋅ D ⋅σ B axial compression load after flexural yielding, and S/F: N =−a ⋅σ flexural and shear failures occurring simultaneously. min g y Equations 24-26 by BCJ give a good estimation with where, ag and σ y : gross sectional area and yield the average ratio of 1.03, but the standard deviation of strength of longitudinal reinforcement, σ B : concrete 0.21 indicates a slightly large scatter. No significant M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 175 effect of concrete strength on the ratio is observed. The Shear reinforcement ratio (pw): 0.2 – 1.8 % ACI concrete stress block expressed by Eq.27 gives the Degree of shear reinforcement (pw σwy): best estimation among three design equations. NZS 3101 4.0 – 20.0 N/mm2 is primarily the same as the ACI except α1 for taking the effect of high-strength concrete into account. However, as seen in the average ratio of 1.24, NZS 3101 underes- timates the experimental results.

4.1.3 Shear behavior Shear loading tests on reinforced concrete members constructed of concrete in excess of 60 N/mm2 and shear reinforcement of more than 1000 N/mm2 have been car- ried out and reported in Japan as well as abroad. The experimental parameters are shear reinforcement ratio, axial load ratio, shear span ratio, effective depth and so on. In the civil engineering field, specimens with large section and low reinforcement ratio under low axial loads have been tested. In building engineering field members with large shear reinforcement ratio and axial load ratio with small shear span ratio have been tested. Design equations were derived from each set of test results. The design equation can be applied to members with a similar specification of reinforcement and dimensions. The literature survey indicates that specimens tested can be divided into three groups: 1) HSC beams without shear reinforcement - specimens failing in diagonal tension Concrete strength: 60 – 130 N/mm2 Shear span to effective depth ratio: 2.5 – 6.0 Effective depth: 150 – 1200 mm Axial load ratio: 0 - specimens failing in diagonal compression Concrete strength: 60 – 130 N/mm2 Shear span ratio: 1.0 – 3.0 Effective depth: 150 – 700 mm Fig. 33 Comparison of α1 in various design codes Axial load ratio: 0 (Ghosh and Ahmad 2006). 2) HSC beams with limited shear reinforcement Concrete strength: 60 – 130 N/mm2 Shear reinforcement yield strength (f ): wy BCJ design equation 250 – 1200 N/mm2 Shear span to effective depth ratio: 2.5 – 5.0 Mean SD Effective depth: 150 – 1200 mm )

N

Tension reinforcement ratio: 1.5 – 7.0 % k

(

s

Shear reinforcement ratio (r=A /bs, A : sec- t

v v l

u

tional area, b: member width and s: spacing): s

e

r

0.1 – 2.0 % l

a Degree of shear reinforcement (r fwy): t

n 0.2 – 8.0 N/mm2 e

m

i Axial load ratio: 0 r

e

p

3) HSC columns with high shear reinforcement ratio and x axial load E Width and depth: 200x200 mm – 400 x 400 mm Length: 400 – 1200 mm Concrete strength: 60 – 170 N/mm2 Shear reinforcement yield strength (σwy): 350 – 1500 N/mm2 Calculation results (kN) Shear span to effective depth ratio: 1.0 – 2.0 Fig. 34 Comparison of experimental and calculated re- Tension reinforcement ratio: 0.8 – 2.2 % sults in flexural strength (BCJ). 176 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

Axial load ratio (=P/(Agf'c), P: axial load, Ag: 4.1.4 Deformation Sectional area, and f'c: compressive strength): (1) Initial stiffness and secant stiffness at yielding 0 – 0.7 Figure 40 illustrates the comparison of initial stiffness between experimental and theoretical results. The ex- perimental data were collected from reinforced concrete BCJ design equation columns using high-strength concrete in the range of 36 2

o

t to 120 N/mm (Ishikawa and Kimura 2006), while the

l

a

t theoretical results were based on theory of elasticity. The

n

e s initial stiffness obtained experimentally was not signifi-

t

l

m

i

u

r cantly influenced by the concrete compressive strength,

s

e

e

p r and corresponds to about 70% of the calculated values.

x

d

e

e Figure 41 shows a comparison of experimental and

f

t

o

a l theoretical results for the secant stiffness at yielding. The

o u

i

t

c

l

a test results were from reinforced concrete beams with a 2

R c concrete compressive strengths from 30 to 120 N/mm and high-strength longitudinal reinforcement with yield strengths from 390 to 685 N/mm2. The theoretical results Concrete compressive stress (N/mm2) were calculated by Eq.29 (Kumagai et al. 2006). Figure Fig. 35 Ratio of experimental to calculated flexural strengths in terms of concrete compressive strength (BCJ).

NZ design equation ACI design equation Mean

Mean ) SD

SD N

k

(

)

s

N

t

l

k

u

(

s

s

t

e

l

r

u

l

s

a

t

e

r

n

l

e

a

t

m

n

i

r

e

e

m

p

i

r

x

e

E

p

x

E

Calculation results (kN) Calculation results (kN) Fig. 36 Comparison of experimental and calculated re- Fig. 38 Comparison of experimental and calculated re- sults in flexural strength (ACI318). sults in flexural strength (NZS3101).

NZ design equation

ACI design equation o

t

l

o

a

t

t

l

n

a

e t

s

t

l

n

m

i

e

u

s

r

t

s

l

e m

e

i

u

r

p

r

s

x

e

e

d

e

r p

e

f

x

t

d

o

e

a

l

e

f

t

u

o

i

o

a

c

t

l

l

a

u o

a

i

t

c

c

R

l

a

a

c R

Concrete compressive stress (N/mm2) Concrete compressive stress (N/mm2) Fig. 37 Ratio of experimental to calculated flexural strengths Fig. 39 Ratio of experimental to calculated flexural strengths in terms of concrete compressive strength (ACI318). in terms of concrete compressive strength (NZS3101).

M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 177

41 shows that Eq.29 can be applied to estimate the secant lished in AIJ and JCI (including JCI annual meetings), modulus at yielding for reinforced concrete beams con- Hayashi et al. (2006) found the following; structed using high-strength materials. - Approximately a quarter of cruciform specimens and one-half of external joint specimens have lower strength αy 3EI Qy /Ry = 2 (29) concrete in beams than in columns and joints. In almost 1+ β 1+ β L ()b ()s a all specimens high-strength reinforcement is used for

−2 joint shear reinforcement and degree of reinforcement ⎛ d ⎞ σ d ⎛ L ⎞ 2 2 α = 3 np ⎜ − 0.5⎟ , β = 1.5 y b , β = 0.864⎜ a ⎟ (expressed as p ⋅ σ ) ranges from 3 N/mm to 8 N/mm . y t ⎝ D ⎠ b σ 2/3L s ⎝ D ⎠ w w y B a In this section of the JCI-TC063A report, beam-column joint specimens constructed of concrete (2) Ultimate deformation with compressive strength higher than 60 N/mm2 are Figure 42 shows the comparison of ultimate deformation investigated in terms of shear strength, shear stress-shear distortion relation, damage evaluation, bond and anchor ( Ru , member rotation angle) between test results and estimated values. The test results were from columns properties of beam longitudinal reinforcement, and with concrete strengths ranging from 36 to 160 N/mm2. load-deformation relation of a whole moment-frame The ultimate deformation is empirically defined as the structure. deformation when the load carrying capacity reduces to 80% of the maximum shear force, which includes P-δ effect. The estimation was carried out using Eq.30 (Ishikawa and Kimura 2005). The ultimate deformation

)

is influenced by factors such as axial force, amount of m

m Note: specimens lateral reinforcement, yield strength of longitudinal re- /

N failed in flexure

inforcement and concrete strength. These factors are k

(

] taken into consideration in Eq.30. t

n

e ⎡ 0.25 ⎤

m

⎛ ⎞ i pw ⋅ f yh f y −3

r Ru = φb ⋅φs ⎢ 22.5⎜ ⎟ λ tu + 2.5⎥ ×10 ≤ 0.044 ' e f c ⎣⎢ ⎝ ⎠ ηc ⎦⎥ p

x

2 e ⎡ ⎛ ' ⎞ ⎤ [ η ⋅ f c ⎛ s ⎞ φ = 2.0⎢ 0.2 − ⎜ c ⎟ ⎥ +1.0 ≤ 1.0 s b ⎜ ⎟ ⎜ ⎟ s

n p f e ⎣⎢ ⎝ ⋅ g ⋅ y ⎠ ⎝ db ⎠ ⎦⎥

n

f

f

i ⎛ ⎞ ⎛ ⎞ t nh − 2 ns − 2 s φs = 0.30pw ⎜ ⎟ ⎜ ⎟ + 0.95 ≤ 1.1 l

a ⎝ nh ⎠ ⎝ ns ⎠ (30) i

t

i

n

I λ tu = 10 εy ⋅ηt +1.1

4.2 Beam-column joints and frame structures Initial stiffness [calculated] (kN/mm) In columns in the lower stories of high-rise reinforced Fig. 40 Initial stiffness (Ishikawa and Kimura 2006). concrete buildings taller than, for example, 50 stories, concrete with standard cylinder compressive strength in excess of 100 N/mm2 is used. Those columns are usually

g

n

confined by transverse reinforcement of Grade SD780 or i

d higher. In beams of such high-rise buildings concrete up l

e

2 i to 60 N/mm is used with Grade SD490 or lower longi- y

t

tudinal reinforcement. a

e

There are two construction methods; cast-in-situ re- l

g

inforced concrete and precast concrete. Several types of n

)

a precast systems are used in practice. d

a

n

r

( Stress intensity in beam-column joints during earth- o

i

]

t

t

a quake excitation is higher in HSC buildings than in NSC n

t

e

o buildings because HSC member sections are smaller and r

m i Normal strength

r

m therefore, the volume of the joints is relatively small. e SD685 class

a p

x

Shear strength of the joints and anchorage of beam lon- e

e

B gitudinal reinforcement should be secured. It should be [ noted that structural performance of joints has a large effect on structural performance of the whole frame. Not many experiments can be found in literature for Beam rotation angle at yielding [calculated] (rad) beam-column joints constructed of concrete in excess of Fig. 41 Beam rotaion angle at yielding (Kumagai et al. 60 N/mm2. From a literature survey into research pub- 2006). 178 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009

4.3 Structural walls )

d

4.3.1 Why high strength concrete is needed for a

r

3

structural walls? -

0

High-strength concrete is used in structural walls to 1

x provide higher horizontal load resistance for buildings. (

]

t

Most of the lateral load induced by earthquakes or wind n

e

is resisted by structural walls due to their higher stiffness m

i

r and strength compared with other structural lateral force e

p resistant members such as columns. Use of structural x

e walls can result in lower construction cost. [

n

o

i

Three-dimensional shear walls are useful for satisfying t

a

not only structural demands but also architects' demands. m

r

o

Structural walls used for high-rise buildings are de- f

e signed to fail in flexure, which also requires ductility. d

e That is why high-strength concrete is used for such walls. t

a

m

i

t

l

4.3.2 Summary of characteristics U Shear strength of shear walls depends on concrete com- Ultimate deformation [calculated] ( x10-3 rad) pressive strength. Concrete compressive strength along shear cracks is lower than that before shear cracks occur. Fig. 42 Ultimate rotation angle (Ishikawa and Kimura The ratio of compressive strengths before and after shear 2005). cracking decreases as the concrete compressive strength obtained from standard cylinder specimens increases. Nielsen’s proposal as average: This ratio is called effective compressive strength coef- ficient, ν . An equation for ν was proposed by Nielsen σ o o ν = 0.8 − B (32) (1984), CEB (1978), and Naganuma (1991). The results o 200 calculated by these equations are shown in Fig. 43 with experimental results. In Design Guideline for Earthquake CEB: Resistant Reinforced Concrete Buildings Based on Ul- ν = 1.698σ −0.333 (33) timate Strength Concept (AIJ 1990) and Design Guide- o B lines for Earthquake Resistant Reinforced Concrete Naganuma’s proposal: Buildings Based on Inelastic Displacement Concept (AIJ ν = 1.907σ −0.34 (34) 1999), Nielsen’s equation, Eq.31, for the lower limit for o B ν is recommended for shear design of members con- where, νo : effective compressive strength ratio, σ B : o 2 structed of concrete with strength up to 60 N/mm2. For standard cylinder compressive strength (N/mm ). concrete with strength in excess of 60 N/mm2, the guidelines adopt the CEB equation in their commentaries. 5. Design of building structure This is because the equation for the lower bound of νo by Nielsen does not agree well with the experimental results 5.1 Seismic design of super-high-rise RC in the range of concrete compressive strength greater buildings than 60 N/mm2 as shown in Fig. 43. Figure 43 indicates In structural performance evaluation of super high-rise that Eq.31 can evaluate well the lower bound of the ex- RC buildings, which are taller than 60 m, focus is placed 2 on design for earthquake loading. Described below is an perimental results for σB smaller than 60 N/mm , and outline of the seismic design practice in Japan for su- Eq.33 for σB larger than that. For their average values per-high-rise buildings. Eq.32 gives a good estimation for σB smaller than 60 2 N/mm , and Eq.34 for σB larger than that. For evaluating the ductility after flexural yielding, 5.1.1 Design earthquake ground motion accurate modeling of the concrete stress-strain curves is Seismic design is conducted considering two kinds of needed. Kumagai et al. (2005) idealized the concrete earthquake motions: earthquake motions occurring rarely compressive behavior by observing the experimental and those occurring extremely rarely. results, and showed that the flexural inelastic behavior of The design earthquake motions consist of seismic three-dimensional structural walls could be evaluated by wave forms having the acceleration response spectrum of a section analysis with this idealization of concrete be- the exposed engineering bedrock, considering appropri- havior (Kumagai et al. 2005). ate wave-amplification in the subsurface of the building Nielsen’s proposal for lower bound: (hereinafter referred to as “the notification waves”). At least six earthquake wave forms shall be used. σ B Three wave froms are based on an appropriate consid- ν = 0.7 − (31) o 200 eration of phase-difference distribution and should meet the requirements in terms of duration, etc. Seismic wave M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 179

neousness with the horizontal earthquake motions.

5.1.4 Evaluation criteria It shall be confirmed that the inter-story drift response does not exceed 1/200 rad. for earthquake motions oc- curring rarely. It shall also be confirmed that stress in- duced in the primary elements be smaller than the specified allowable stress, and that no remarkable re- sidual cracks or deformation after the earthquakes be expected. For earthquake motions occurring extremely rarely it shall be confirmed that the inter-story drift angle does not exceed approximately 1/100 rad. The ductility factor of each story shall not exceed approximately 2.0. The duc- tility factor of the structural members shall not exceed the limit value assigned according to the structural sys- tem and the mechanical properties of each member. “Standard for Structural Calculation of Reinforced Fig. 43 Effective compressive strength coefficients ex- Concrete Structure” (RC standard) of AIJ is referred to perimentally obtained and proposals of their design for the structural design of RC buildings. The maximum equations. specified design strength is 60 N/mm2. AIJ publishes another design guideline “Design Guideline for Earth- forms simulating those which would extremely rarely quake Resistant Reinforced Concrete Buildings Based on occur at the construction site (hereinafter referred to as Inelastic Displacement Concept,” which is often referred “the site waves”) based on the active fault distribution, a to as an ultimate strength design procedure. Again, the fault rupture model, the historical earthquake activity, the specified design strength of concrete must be less than 60 bedrock structure and other factors can be used as the N/mm2. Therefore, in the design of reinforced concrete input earthquake motions. Three other wave forms re- structural members using high-strength concrete ex- corded during the past major earthquakes are used. Their ceeding 60 N/mm2, there is no authorized reference in maximum velocities are scaled to 25 cm/sec for the rare Japan. In order to verify their structural performance, earthquakes and 50 cm/sec for extremely rare earth- experiments on structural members and subassemblies, quakes. Some additional wave forms may be used. and construction trials, if needed, are often conducted. The research reports on the New RC Project and tech- 5.1.2 Idealization of a building for dynamic nical papers on structural experiments and analyses analyses published in journals are often referred to. A lumped mass system model is usually used. Internal forces and deformations of each member of the building 6. Applications should be appropriately idealized. Dynamic interaction between the building and the ground (soil-structure in- Almost all HSC applications are for high-rise condo- teraction) should be appropriately taken into considera- minium buildings. Specifications of high-rise reinforced tion. The story shear force - story drift relationships are concrete buildings such as number of stories, height, idealized based on static nonlinear analysis results taking foundation systems, strengths of concrete and rein- the plastic behavior of each member into account. forcement have been collected and analyzed. As of June 2005, there are 458 buildings over 60 m in height in 5.1.3 Seismic response analysis Japan (summarized in Building Letter, November 1972 – Dynamic response analyses on the lumped mass system June 2005) published by Building Center of Japan, are used to obtain responses of the building. The design GBRC by General Building Research Corporation of earthquake motions are applied separately along each of Japan and so on). These institutes have a board of review the two principal axes of the building. The earthquake for structural design of buildings higher than 60 m and motions are also input along the two principal axes si- publish concise specifications of buildings reviewed. The multaneously or along the axis 45 degrees to the princi- highest design strength used in high-rise buildings in pal axis. Japan was 150 N/mm2 when the survey was conducted A static non-linear push-over analysis is carried out to (June 2005). Higher design strength is planned to be a roof displacement in excess of the peak roof dis- applied to taller buildings. placement obtained from the dynamic analyses. This is Concrete strength specified in bridge structures having done to check if brittle shear failures occur in any of the been constructed in Japan is up to 80 N/mm2 although columns. there are some examples in which 120 N/mm2 concrete The effect of vertical earthquake motions shall be was used for pedestrian decks and ultra-high-strength evaluated appropriately, taking into account its simulta- 180 M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 fiber-reinforced concrete up to 180 N/mm2 for bridges. concrete columns under concentric compression and High-strength concrete is used in bridges to make the its modeling of stress-strain relationship.” girders height lower and, therefore, to make them lighter. Proceedings of JSCE, 753(V-62), 137-151. (in High durability is another advantage because civil en- Japanese) gineering construction should be in operation as long as Akiyama, M., Hong, Kee-Nam, Sasaki, T., Maeda, N. possible with the smallest maintenance costs and work. and Suzuki, M. (2005). “Concentric loading test of rc columns with normal- and high-strength materials and 7. Concluding remarks stress-strain model for confined concrete considering compressive fracture energy.” Proceedings of JSCE, Research and practice on HSC were concisely described 788(V-67), 81-98. (in Japanese) in this paper. The paper is based on a research report American Concrete Institute (2008). “Building code prepared by Research Committee on High-Strength requirements for structural concrete (ACI318-08) and Concrete Structures (JCI/TC-042) organized in JCI. The commentary (ACI318R-08).” report was written in Japanese. A full report in English Architectural Institute of Japan (1990). “Design will be published within a few years. More information guideline for earthquake resistant reinforced concrete will be included and it should be of great value for re- buildings based on ultimate strength concept.” (in searchers as well as practitioners. Japanese) The Journal of Advanced Concrete Technology has Architectural Institute of Japan (1999). “Design released a special issue of high-strength and guidelines for earthquake resistant reinforced high-performance concrete including an invited paper concrete buildings based on inelastic displacement on ultra-high-strength fiber-reinforced concrete (UFC) concept.” (in Japanese) written by Prof. Sugano, who chaired the JCI committee Architectural Institute of Japan (1999). “Standard for on Utilization of HSC and HPC (JCI-TC063A). The structural calculation of reinforced concrete issue includes another five papers, providing more in- structures based on allowable stress concept.” (in formation on HSC and HPC (Jinnai et al. 2007; Kimura Japanese) et al. 2007a; Kimura et al. 2007b; Kuroiwa et al. 2007; Architectural Institute of Japan (2005). Tanimura et al. 2007). “Recommendation for practice of high-strength concrete.” Acknowledgement Architectural Institute of Japan (2008). The author would like to express his sincere thanks to the “Recommendation for fire resistant design of steel members of sub-working group 1 of the JCI-TC063A. structures.” Members of Sub-working group 1: Manabu Fujita Baba, S., Kobayashi, Y. and Michikoshi, S. (2002). “Fire (Sumitomo Mitsui Construction Co.), Kazuya Hayashi resistance of high-strength reinforced concrete bars.” (Fujita Corporation), Tsutomu Komuro (Taisei Corpora- Summaries of Technical Papers of AIJ Annual Meeting tion), Hitoshi Kumagai (Shimizu Corporation), Hiroshi (Hokuriku), 25-26. (in Japanese) Mutsuyoshi (Saitama University), Hiroshi Okamoto Building Center of Japan (1991). “Guidelines for (Railway Technical Research Institute), Soji Oshiro structural calculation of buildings.” (in Japanese) (West Nippon Expressway Company Limited), Mo- Comité Euro-International du Béton (1978). “CEB-FIP toyuki Suzuki (Tohoku University), Masaomi Teshiga- model code for concrete structure.” wara (Nagoya University), Kazuaki Tsuda (Obayashi Cusson, D. and Paultre, P. (1994). “High-strength Corporation), Hiroshi Watanabe (Public Works Research concrete columns confined by rectangular ties.” Institute). Journal of Structural Engineering, ASCE, 120(3), Thanks are extended to the former committee mem- 783-804. bers (April 2005 - March 2007) below who significantly Fafitis, A. and Shah, S. P. (1985). “Lateral reinforcement contributed to preparation of the original report in for high-strength concrete columns.” ACI, SP-87, Japanese; Mitsuyoshi Akiyama, Yukihiro Tanimura, 213-232. Kazukiyo Tamaki, Kazuaki Tsuda, Masaru Teraoka, General Building Research Corporation of Japan (2005). Natsuki Hara, Hiroshi Fukuyama, Makoto Maruta, Hideo “Test result statistics for construction materials in the Murakami, Gakuho Watanabe, Yuji Ishikawa, Nobuyuki fiscal year of 2004, Part 2 Deformed reinforcing bars Izumi, Toshimichi, Ichinomiya, Eiichi Inai, Norio Inoue, and their connections.” GBRC, July. Kazuyoshi Izume, Kazuto Kamisakoda, Hiroaki Oyama, Ghosh, S. K. and Ahmad, S. H. (2006). “High-strength Toshiki Okada, Hideki Kimura, Takumi Shimomura, concrete for seismic applications.” Concrete Kumiko Suda, Kazuhiro Watanabe, and Hiroyoshi Watanabe. International, 28, (9), 47-49. Gokaku, W., Shimomura, T. and Maruyama, K. (2001). References “Influence of quality of concrete on crack width in RC Akiyama, M., Hong, Kee-Nam, Sato, M., Suzuki, M., member.” Proceedings of the Japan Concrete Institute, Maeda, N. and Suzuki, M. (2004). “Effect of 23(3), 1333-1338. (in Japanese) transverse reinforcement on behavior of high strength Hayashi, K., Takamori, N. and Teraoka, M. (2006). M. Nishiyama / Journal of Advanced Concrete Technology Vol. 7, No. 2, 157-182, 2009 181

“Research and examination on shear strength of R/C Journal of Advanced Concrete Technology, 5(2), beam-column joints using high-strength concrete.” 181-191. Proceedings of the Japan Concrete Institute, 28(2), Kimura, H., Ishikawa, Y., Kanbayashi, A. and Takatsu, 295-300. (in Japanese) H. (2007b). “Seismic behavior of 200 MPa ultra-high Ishikawa, Y. and Kimura, H. (2005). “Limit strength steel-fiber reinforced concrete columns displacement under compressive axial force of R/C under varying axial load.” Journal of Advanced columns using high strength concrete and high Concrete Technology, 5(2), 193-200. strength steel.” Concrete Research and Technology, 16 Kumagai, H., Tozawa, M. and Kurose, Y. (2005). (1), 55-66. (in Japanese) “Compression failure in reinforced concrete shear wall Ishikawa, Y. and Kimura, H. (2006). “Initial stiffness of with ultra-high-strength materials.” Concrete R/C columns with high strength concrete and steel Research and Technology, 16(3), 59-68. 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