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NBS

PUBLICATIONS

NBSIR 81-2374

Investigation of Construction Failure

of Harbour Cay Condominium in

Cocoa Beach, Florida

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Building Technology

Washington, DC 20234

September 1981

Sponsored by

:upational Safety and Health Administration

Department of Labor

1Q0 shington, DC 20001

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31-2374

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OCT 2 2 1981

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NBSIR 81-2374

INVESTIGATION OF CONSTRUCTION FAILURE

OF HARBOUR CAY CONDOMINIUM IN

COCOA BEACH, FLORIDA

SH.)Af. Lew

N. J. Carino

S. G. Fattal M. E. Batts

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Building Technology

Washington, DC 20234

September 1981

Sponsored by

Occupational Safety and Health Administration

U.S. Department of Labor

Washington, DC 20001

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS. Ernest Ambler, Director
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ABSTRACT

The investigation of the collapse of a five-story reinforced concrete flatplate structure under construction at Cocoa Beach, Florida is presented in

this report. The investigation included onsite inspection, laboratory tests

and analytical studies.

Based on the results of this investigation, it is concluded that the most

probable cause of the failure was insufficient punching shear capacity in the

fifth-floor slab to resist the applied construction loads.

Two factors contributed to the low punching shear capacity: one in the design

and the other in the construction of the building. In the design, the omission

of a check for punching shear resulted in a smaller slab thickness than needed

to satisfy the Code requirements. In construction, the use of specified chairs

having insufficient height to support the top reinforcing steel resulted in

more than the cover specified in the structural drawings. Both factors contributed to reducing the effective depth of the slab such that it had insufficient

strength to resist the construction loads.

The analysis showed that shear stresses in the slab at many column locations on

the fifth floor exceeded the nominal shear strength. Thus, punching shear failure at one of the columns precipitated a progressive failure of the slab throughout the entire fifth floor. Collapse of the fifth floor, in turn, caused the

successive collapse of the lower floor slabs. The analysis of the structure

indicated that the failure of the fifth floor slab most likely initiated at

column G-2, an interior column which supported the last bay of freshly placed

concrete prior to the collapse.

Key Words: Building; collapse; concrete; concrete strength; construction;

failure; flat plate; shear; strength.

iii
PREFACE

An investigation of the construction failure of a five-story reinforced

concrete building (Harbour Cay Condominium) at Cocoa Beach, Florida on March 27, 1981 was carried out by the National Bureau of Standards (NBS) at

the request of the Occupational Safety and Health Administration (OSHA),

Department of Labor. This investigation was conducted pursuant to an interagency agreement between OSHA and NBS.

The NBS investigation team consisted of H. S. Lew, S. G. Fattal, N. J. Carino,

and M. E. Batts. Throughout the course of this investigation, the team

received full cooperation from the OSHA regional and area officers. Assistance provided by Mr. William P. Demery, Tampa Area Director and his staff is

gratefully acknowledged by the NBS team. Without their help, it would not

have been possible to obtain the data needed to carry out this investigation.

iv
EXECUTIVE SUMMARY

The Harbour Cay Condominium, a five-story flat-plate reinforced concrete

building, was under construction at Cocoa Beach, Florida. Shortly after 3:00

p.m., March 27, 1981, during the placement of concrete for the roof slab, the

entire structure collapsed. As a result, 11 workers died, and another 23 were

injured.

The structure had an overall length of 242 ft (74 m) and a width of 58 ft (18 m) . The slab concrete was placed on flying forms, consisting of preassembled plywood decks supported by aluminum trusses. Except for the roof, which was

planned to be cast in one day, the floor slabs were cast in two separate operations, each covering one-half of the total floor area. At the time of the

collapse, concrete had been placed on seven of the nine flying forms comprising

the entire roof area.

During the course of the investigation two site visits were made. The first

site visit was made on March 28-30, 1981, one day after the collapse, and the

second visit was made on April 7-8, 1981. During the visits, structural components were examined visually and photographic documents were made as the com- ponents were removed for rescue operations, selected key components were

identified and marked with paint for testing laboratory personnel to secure

test samples, and measurements of member dimensions were made. In addition,

construction and other documents were obtained through the OSHA area office in

Tampa

.

At the time of collapse, the flying forms were placed on the fifth floor to

receive concrete for the entire roof slab. The exact number and locations of

reshores at various floors are not certain. The first story appears to have

had only a few exterior reshores in place, and the fourth story probably had

the full number of reshores according to the reshoring plan supplied by the

general contractor after the collapse. Conflicting statements were given by

workers as to the number and locations of reshores on the second and third

stories. For the purpose of analysis, it was assumed that both the second and

third stories had the full number of reshores.

The concrete used for the structure was supplied by both the onsite batch plant

and a ready-mix plant. Approximately 1/3 of the concrete used for the floor

slabs was supplied by the ready-mix plant. The remainder was supplied by the

onsite batch plant. The concrete produced by the onsite batch plant was reported to have been nonuniform in consistency and f inishability

.

Core samples were taken from slab and column sections which were identified

during the first field visit to the site. Tests of these cores were made by an independent testing laboratory which was retained by OSHA. The compressive

strengths of cores taken from slab sections satisfied the American Concrete

Institute 318 Standard (ACI Code) strength criteria for the specified strength

of 4000 psi (27.6 MPa). The compressive strength of cores taken from column

sections were lower than the strengths of slab cores, and the average strength

of the cores from the fifth floor columns did not satisfy the ACI criteria for

the specified strength.

v
Samples of Ingredients used for the concretes produced by the onsite batch

plant and the ready-mix plant were obtained from the site. They were subjected

to chemical and physical analyses and used by NBS to make concrete mixtures

representative of what was used in the structure. The purpose of the NBS laboratory investigation was to establish the strength of the concrete at the time

of collapse. Based on the laboratory mixes and local temperature data at the

site, the estimated compressive strengths of concrete at the time of collapse

were 4000 psi (276 MPa) for the south half and 3700 psi (255 MPa) for the north

half of the fifth floor slab.

The analysis of the structure subjected to the construction loads was made with

the ICES-STRUDL-II* finite element program. The finite element model analyzed

incorporated both the structure, which is discretized by rectangular plate bending elements, and the formwork system. In determining the loading conditions

for analysis, the effect of construction load distribution between the structure

and the formwork was considered along with the effects of sequential removal of

flying forms. The results of the analysis show that maximum moments and shears

in the slab occur on the fifth floor. Combinations of shears and moments which

would produce critical conditions occurred at columns B-2 and G-2 . Description

of the collapse by workers on the site indicate that the most probable origin

of the collapse was the fifth floor slab in the vicinity just below where the

roof concrete was being placed, which supports the critical nature of column G2

.

The construction documents and field measurements revealed that the top layers

of reinforcing bars in the column strips were supported by individual high

chairs which produced concrete cover of 1 3/4 in (44 mm) instead of 3/4 in

(19 mm) as specified in the structural drawings. This discrepancy resulted in

the reduction of the effective depth of the slab which had direct effect on the

shear resisting capacity of the slab.

The chair heights were in conformance with the steel supplier's shop drawings,

which specified the number and sizes of individual high chairs required for

each floor. The record copies of these drawings show numerous check marks.

One of the six pages bore the handwritten words "approved as noted" and the

signature of the design engineer of record.

Using the moments and shears obtained from the finite element analysis, punching

shear stresses resulting from direct shear forces and unbalanced moments were

computed at slab-column connections. These shear stress values are compared

with the shear capacity of the slab based on the predicted inplace strength of

concrete at the time of collapse. The computed beam-type shear stress under

construction loading is considerably less than the beam shear strength of the

slab. On the other hand, the computed punching shear stresses at many column

locations in the fifth floor exceeded the nominal punching shear capacity of

the slab. Based on the probablistic nature of the shear strength, it is reasoned that the collapse was most likely initiated due to punching shear failure

at column G-2.

* A general purpose structural analysis program, see chapter 5 and ref. 5.5 for

further details.

vi
It is shown that high punching shear stresses resulted from use of reduced

height individual high chairs to support the top reinforcing bars in the column

strip, thereby reducing the effective depth of the slab. The omission of a check for punching shear capacity in the design calculations resulted in a slab

thickness of 8 in (203 mm). The ACI Code would have required a thickness of

11 in (279 mm) to provide adequate safety against punching shear under design

loads. It is shown that this design deficiency resulted in a reduction of the

margin of safety such that the slab had insufficient reserve strength to

accommodate construction errors.

It is concluded that two factors contributed to the collapse: one in the

design and the other in the construction of the building. In design, a check

for punching shear was omitted. In construction, the specified chairs used to

support the top reinforcing steel provided more than the cover called for in

the structural drawings. The analysis showed that the shear stresses at many

column locations in the fifth floor exceeded the nominal shear strength. Once

punching shear failure had initiated at column G-2 , it propagated throughout

the slab causing the total collapse of the fifth floor, which, in turn, caused

the successive collapse of the lower floor slabs.

vii
TABLE OF CONTENTS

Page

ABSTRACT . ill

PREFACE iv

EXECUTIVE SUMMARY v

1 . INTRODUCTION 1 1.1 Background 1

1.2 Objective and Scope of the Investigation 1

1.3 Organization of the Report 1

2. DESCRIPTION OF THE STRUCTURE AND THE COLLAPSE 3

2.1 Introduction 3

2.2 Documents Supplied by OSHA Field Office 3

2.3 Description of the Structure 5 2.4 Construction Method 8

2.5 Events at the Time of the Collapse 10

2.6 Summary 11

3. SITE INVESTIGATION 29

3.1 Introduction 29

3.2 Investigation at the Site 29

3.3 Interview Statements of Workers 30

3.3.1 Location of Reshores 30

3.3.2 Construction Deficiencies 31

3.3.3 Quality of Concrete 32

3.3.4 Description of the Collapse 33

3.4 Summary 34

4. LABORATORY INVESTIGATIONS 39

4.1 Introduction 39

4.2 Information on Field Concrete 39

4.3 Test of Samples from the Site 41

4.3.1 Concrete Ingredients 41

4.3.2 Core Strength 42

4.3.3 Petrographic Analysis 42

4.3.4 Chemical Analysis 43

4.4 NBS Laboratory Investigation 44

4.4.1 Sample Preparation 44

4.4.2 Test Results 46

4.4.3 Estimated Strength of Concrete at Time of the Collapse . 47

4.4.4 Reinforcing Steel Strength 49

4.5 Summary 49

5. STRUCTURAL ANALYSIS 71

5.1 Introduction 71

5.2 Background 71

5.3 Loads 72

viii
TABLE OF CONTENTS (Cont.)

Page

5.4 Construction Sequence 74

5.5 Method of Analysis 77

5.6 Finite Element Analysis of Entire Structure 79

5.7 Summary 80

6. INTERPRETATION OF ANALYSIS 94

6.1 Introduction 94

6.2 Analysis of Shear Failure of Slab 94

6.3 Influence of Effective Depth of Slab on Shear Capacity 96

6.4 Probable Mode of Failure 98

7. SUMMARY AND CONCLUSION 103

8. ACKNOWLEDGMENTS 105

9. REFERENCES 106

APPENDIX A. RESULTS OF TESTS ON SAMPLES FROM THE SITE A-l

ix

1.1 1 INTRODUCTION

BACKGROUND

The Harbour Cay Condominium, a five-story flat-plate reinforced concrete

building, was under construction at Cocoa Beach, Florida. Shortly after

3:00 p.m., March 27, 1981, during the placement of concrete for the roof

slab, the entire structure collapsed vertically with each floor slab breaking away from the columns and landing on the ground floor. As a result, 11

workers who were performing duties at various levels of the building died,

and another 23 were injured.

The structure had an overall length of 242 ft (74 m) , a width of 58 ft (18 m)

and contained nine bays. The slab concrete was placed on flying forms, con- sisting of preassembled plywood decks supported by aluminum trusses. The

construction system is described in detail in chapter 2. Except for the

roof, which was planned to be cast in one day, the floor slabs were cast in

two separate operations, each covering one-half of the total floor area. At

the time of the collapse, concrete had been placed on seven of the nine

flying forms comprising the entire roof area.

Within one hour of the collapse, an inspection team from the Occupational

Safety and Health Administration (OSHA), Department of Labor, arrived at the

site and began an investigation of the collapse. A team from the National

Bureau of Standards (NBS) joined the OSHA team one day later on March 28, 1981.

1.2 OBJECTIVE AND SCOPE OF THE INVESTIGATION

The National Bureau of Standards was requested to assist with the field

investigation being conducted by the Occupational Safety and Health Administration at the site of the collapse and to carry out a detailed study to

determine the most probable cause of the collapse.

In response to this request, NBS carried out field, laboratory, and analytical

studies. The NBS investigators used data obtained from onsite inspections,

OSHA case records, and other related documents such as the engineering calculations, specifications, and drawings showing details of the building

construction.

1.3 ORGANIZATION OF THE REPORT

This report is organized in nine chapters:

Chapter 2 presents the description of the structure based on the contract

drawings and the specifications and lists the various documents furnished by

OSHA. In addition, it describes the method and sequence of construction of

the structure, its appearance, and conditions before and after the collapse.

1
Chapter 3 presents observations made and information obtained from the field

investigations. It summarizes the statements made by those interviewed by

OSHA officials

.

Chapter 4 presents the results of concrete tests which were used to determine

the strength and stiffness of the structure at the time of failure. The test

results of the reinforcing steel obtained from the site are also presented.

Chapter 5 describes the rationale for defining the loads which were acting on

the structure at the time of the collapse. It also presents the analytical

approach used in the investigation of the structure and the results of the

analysis

.

Chapter 6 examines the results of the analysis of the structure, compares

them with the capacity of the structure, and presents the most probable cause

and mode of failure.

Chapter 7 summarizes the findings of the investigation and presents its

conclusions

.

Chapter 8 acknowledges those individuals . who made contributions to the various

phases of the investigation and to the preparation of this report.

Chapter 9 lists the references cited in the text.

2
2 DESCRIPTION OF THE STRUCTURE AND THE COLLAPSE

2.1 INTRODUCTION

This chapter describes the structural features of the Harbour Cay condominium

and the method used in its construction. The emphasis is on structural aspects

that are pertinent to the investigation of the collapse. The principal sources

of information were construction documents supplied by OSHA, site investigations

of the collapse, and photographs of the structure taken before and after the

collapse

.

Section 2.2 identifies the various construction and other documents supplied by

OSHA and describes briefly their uses in this investigation. The appropriate

sections under which the specific contents of these documents are discussed are identified by reference.

Section 2.3 presents selected structural data for use in subsequent analytical

investigations. It begins with a brief description of the layout of the building, the sectional configuration of its columns and floor slabs, reinforcing

plans and schedules, materials specifications, design loads and the scope of

design calculations. The latter part of. this section establishes the actual

effective depth of the slab at the columns. The effective depth is a critical

parameter in calculations of the shear and moment capacity of the slab.

Section 2.4 describes the falsework system and the procedure used in the

construction of the building. The principal topics of discussion are the con- figuration and layout of the flying form system, the construction sequence, the

state of construction at the time of the collapse, and the location of reshores

in the lower stories.

A brief description of events at the time of the collapse is given in

section 2.5. Section 2.6 provides a summary of the findings reported in this

chapter

.

2.2 DOCUMENTS SUPPLIED BY OSHA FIELD OFFICE

NBS requested certain documents in order to carry out its investigation, which

were subsequently provided by the OSHA area office in Tampa. The documents

provided information on the structural aspects of the building, the construction schedule, and the results of tests conducted by testing laboratories.

OSHA also provided copies of personal interview statements of workers, many of

whom were eyewitnesses to the collapse. In addition, OSHA supplied NBS with

photographs taken prior to and after the collapse.

Construction Documents

The following construction documents were supplied by OSHA: (1) architectural,

structural and steel supplier’s drawings and job specifications, (2) a flying

form plan, (3) a reshoring layout (prepared after the collapse), and

(4) structural design calculations.

3
The first three documents were used extensively to develop an understanding of

the structure and the construction scheme (sections 2.3 and 2.4), and to formulate the analytical models (chapter 5) employed in subsequent investigations of

the collapse (chapters 5 and 6). The flying form plan, supplemented with information from the manufacturer (section 2.4), were the principal documents used

in the evaluation of construction loads (section 5.3).

Section 2.3 includes a commentary on the structural design calculations mainly

with regard to content and certain omissions. As discussed in that section,

the question of adequacy of the design calculations for service loads is not

addressed because it is outside the scope of this investigation.

Daily Logs

Several sources were used to reconstruct the construction history. These

included the daily notes of the construction manager and of the general contractor' s field superintendent, and the engineer's daily inspection reports.

Copies of the daily reports of the onsite batch plant provided information

which resolved some discrepancies in the above sources with respect to the con- struction sequence. The schedule presented in section 2.4 is based primarily

on these documents.

Along with the batch plant reports, there were copies of mix design sheets for

concrete that the general contractor had used on a previous project. These

mix design sheets were supplemented with charts to be used by the batch plant

operator to determine the quantities of concrete for different size batches and

types of concrete. These documents are further discussed in chapter 4. In

addition, OSHA obtained copies of invoices and delivery tickets from various

materials and service suppliers to the general contractor.

Copies of daily weather reports obtained from Patrick Air Force Base contained

hourly temperature data that were used with NBS laboratory tests to predict the

inplace strength of concrete at the time of the collapse (section 4.4.3).

Testing Laboratory Reports

The general contractor employed a testing laboratory to perform quality

control testing of the concrete. The information contained in these reports

is presented in section 4.2.

OSHA retained a testing laboratory to test samples taken from the site after

the collapse. A copy of this laboratory's report was provided to NBS and the

results are summarized in section 4.3.

Interview Statements

OSHA provided copies of the statement made by workers interviewed by OSHA

Compliance Officers. These statements provided very useful information on the

construction of the building as discussed in section 2.4. In addition, they

provided eyewitness accounts of the collapse. Section 3.3 summarizes some of

the pertinent information gathered from the personal interview statements.

4
Copies of the investigation reports of the Florida Department of Professional

Regulation provided statements made by engineers, regulatory personnel, and

construction managers involved in the Harbour Cay project.

2.3 DESCRIPTION OF THE STRUCTURE

The building was a five-story flat plate structure intended for residential

use. A typical framing plan for the second through fifth floors is shown in

figure 2.1. The columns were 10 x 18 in (254 x 457 mm) in cross section with

the following exceptions: the columns on lines A and K were 10 x 12 in (254 x

305 mm) and those at the exterior ends of the stairwells were 8 x 12 in (203 x

305 mm) in cross section. The thickness of the floor slabs was 8 in (203 mm)

.

The thickness of the balcony slab along the west side varied from 7-1/4 in

(184 mm) at the column line to 6-1/2 in (165 mm) at the edge. The thickness

of the walkway slab along the east side (as scaled from the structural plans)

was 7-1/2 in (191 mm). The first story floor slab was cast on grade. The

elevator tower was structurally detached from the building. Top of slab story

heights were 8' -8" (2.64 m) . At the foundation level, the columns were

supported by pile caps containing two to nine piles each. The exterior caps were connected by continuous wall footings.

Reinforcing steel for the slabs and columns was specified as deformed bars

ASTM Grade 60. The framing plans in figures 2.1 and 2.2 show the top and

bottom reinforcement in the slab, respectively. Table 2.1 shows the reinforcing schedule for the columns. The structural drawings called for 4000 psi

(27.60 MPa) strength for concrete in the above-grade portion of the structure.

The specified live loads were 20, 40, 60 and 100 psf (960, 1920, 2870 and

4790 Pa) for the roof, floor, public balcony and stairways, respectively. The

wind design was specified to conform with the SBCC Standard Building Code,

110 mph (49 m/s) wind with a 100-ye^r return period [2.2]. Concrete block

masonry strength was specified as f = 1500 psi (10.35 MPa). Non-load bearing

concrete masonry construction was used for the exterior walls of the building.

The structural design calculations consisted of 79 consecutively numbered

pages. The first page specified the material strengths and loads assumed for

the design. Pages 3 through 9 showed the axial loads in columns, computed on

the basis of tributary areas. Pages 9 through 16 contained wind load calculations and the moments they produced in the assumed structure. Page 17 specified the preliminary column sizes based on the computed axial loads and assumed

end moments. Pages 18 through 77 contained the slab moment calculations and

sizing of the flexural reinforcement. Page 78 contained the pile cap designs

and page 79 showed the final column schedule.

The design calculations were not examined in detail, but the following omissions

were noted with regard to compliance with the provisions of the American Concrete

Institute 318 Standard (hereafter referred to as the ACI Code) [2.1], which is

part of the applicable local Code [2.2] for this structure.

(1) There were no calculations to indicate whether the deflection or minimum

thickness provisions of the Code were met.

5
(2) There were no calculations to indicate whether the punching shear and

beam shear provisions of the Code were met.

(3) There were no checks to determine if the column reinforcement would

satisfy the necessary spacing requirements specified by the Code.

(4) The steel area calculations for the flexural reinforcement in the slabs were based on ASTM Grade 40 steel. This is at variance with the Grade 60

steel specified in the structural drawings.

(5) The effective depth of the slab in flexural reinforcement calculations

was not defined explicitly but appeared as part of a constant multiplier

of computed moments. Back calculations showed that the designer used

the value of "around" 6.3 in (160 mm) for the effective depth in steel area calculations (effective depth varies somewhat with the amount and

size of steel bars; hence the use of the word "around"). For the 8-in

(203-mm) slab, the effective depth used is consistent with the 3/4-in

(19-mm) minimum cover specified by the ACI Code (sec. 7.5 of ref. 2.1).

Explanations of "effective depth" and "cover" appear later on in this

section

.

The above statements are cited for information purposes only. Insofar as this

investigation is concerned, the structural drawings and specifications supersede the design calculations, and therefore, are the documents referenced in

subsequent comments about the "design" of the structure. Thus, questions

related to consistency of the design calculations with the contract documents

are not discussed any further.

According to the contract specifications, "protection for reinforcing steel

shall be 1 in for slabs." However, a "typical balcony section" in the structural drawings shows 3/4-in (19-mm) clear cover for top and bottom bars in the

floor slab and bottom bars in the balcony slab, and 1 in (25 mm) clear cover

for top bars in the balcony slab. According to the steel supplier's drawings

and delivery tickets, a sufficient number of 3/4- in (19-mm) slab bolsters were

supplied to obtain a 3/4-in (19-mm) clear cover for the bottom bars in the

slabs regardless of bar size.

From the same sources, the quantities and sizes of individual high chairs

(IHC) supplied for the top steel in a floor slab were identified as follows:

500 - 5-3/4" IHC, 470 - 4-1/4" IHC, 70 - 5" IHC, 50 - 4" IHC and 70 - 3-1/4"

IHC. In addition, #5 support bars were supplied for use with the high chairs

in the middle strip to support the one-way top reinforcement in these strips

(fig. 2.1). The specified spacing of high chairs was 4 ft (1.2 m) on center.

With the aid of the framing plan for top bars in the slab (fig. 2.1), and

quantities of chairs of various specified sizes supplied, it was determined

that the 5-3/4- in (146-mm) high chairs were used for the one-way top reinforcement in the middle strips and the 4-1/4- in (108-mm) high chairs were used for

the two-way top reinforcement in the column strips. Figure 2.3 shows the top

reinforcement layout in a column strip assuming bars in both directions have

the same diameter D. The relationship between chair height, bar diameter,

effective depth, cover, and slab thickness are illustrated in the same figure.

6
Chairs having specified sizes other than these two were supplied in much

smaller quantities for use with the top reinforcement of the thinner slabs in

the balcony and walkway, and those in the elevator tower and stairwells.

Noting that #4 to #6 bars sizes were used at the top of middle strips, the use

of 5-3/4- in (146-mm) chairs with #5 support bars would result in a clear cover

of 7/8 to 1-1/8 in (22-29 mm) for the top bars in the middle strips. This

exceeds the 3/4-in (19-mm) cover specified in the structural drawings but is

within the 3/8-in (9.5-mm) tolerance limit specified in section 7.5 of the ACI

Code [2.1]. According to the steel supplier’s drawings, in areas of two-way

top steel, such as the column strips, main reinforcement in one direction is

used as support bars for main reinforcement in the other direction. The use

of 4-1/4-in (108-mm) chairs in these areas where the slab is 8 in (203 mm)

thick would result in a net cover ranging from a minimum of 1-5/8 in (41 mm)

for the combination of maximum bar sizes (e.g., #8 and #9 at column C2) to a maximum of 2-1/8 in (54 mm) for the combination of minimum bar sizes (e.g., #5

and #8 at column A3). The possibility of 5-3/4-in (146-mm) chairs having been

used in the column strips is ruled out because this would have resulted in

virtually no top cover at most locations.

There is, therefore, an inconsistency in the construction documents. The

structural drawings indicated a 3 /4-in (19-mm) cover for the top slab bars,

whereas the chair height indicated on the steel supplier’s shop drawings would

result in more than 3/4- in cover for the top bars in the column strips of the

slabs. The record copies of the shop drawings contained numerous check marks.

One of the six pages bore the handwritten words "approved as noted" and the

signature of the design engineer of record.

For further verification, actual field measurements were carried out.

Measurements of top bar cover and recovery of 4-1/4- in (108-mm) chairs from

slab segments salvaged from the debris of the collapsed building, (figures

2.4 and 2.5, respectively), confirmed the use of 4-1/4- in (108-mm) chairs for

the two-way top reinforcement in the column strips.

The importance placed on actual chair heights and their location is due to

the fact that the effective depth of the slab is a critical parameter governing its punching shear and bending capacity. Figure 2.6 defines the critical

shear area used in calculations of punching shear stresses in the slab at the

periphery of a column. The punching shear area depends on the sectional dimensions c^ and C 2 of the column and the effective depth, d, which appears in all

three dimensions defining this area.

The calculated effective depth at a column where the top reinforcement in the

slab consists of #8 and #9 bars is 5.3 in (135 mm). This is obtained by adding

to the 4-1/4- in (108-mm) chair height the average diameter of the two bars. If

the effective depth is taken to the centroid of the bars, as defined in section

2.1 of the ACI Code [2.1], its value would be 5.28 in (134 mm) or 5.35 in (136

mm) depending on which size bar is assumed to be at the top. Considering all

the columns and using both methods of calculation, the effective depth is found

to fall in the range of 4.97 to 5.35 in (126 to 136 mm). These figures will be

somewhat less along the sides of columns at the edges of balcony and walkway

7
slabs which are less than 8 in (203 mm) thick. The value of d = 5.3 in (135 mm)

is assumed in the subsequent calculations of stresses in the structure (chapter

5). Generally, these stresses are somewhat on the low side because the assumed

d for the slab is high at most of the columns.

2.4 CONSTRUCTION METHOD

The falsework system consisted of flying forms in the top story to support

freshly cast concrete in the floor directly above, and wood reshores in the

lower stories.

Figure 2.7 identifies the basic components of the flying forms. The assembled

form consisted of two or three parallel trusses having a specified length

obtained by bolting together standard 30-ft (9.2-m) long truss panels and, where

necessary, 5 ft (1.5 m) and/or 10 ft (3.1 m) spacer panels, as indicated in

the figure. The assembled configuration can be seen in figure 2.8 which shows

two relatively undamaged forms that were recovered from the debris after the

collapse. The trusses were spaced at 10 ft (3.1 m) on center and were crossbraced at vertical chord panel points. The supports were swivel type screw

jacks with 4- in (102-mm) square base plates attached to the bottom chords at

the junctions of vertical and diagonal chords. These were spaced at 9'-10"

(3.0 m) on center, or at 5*-5” (1.7 m) on center where a spacer panel of that

size was used. Standard size Aluma beams (fig. 2.7) were connected to the

trusses across the top chords at 1.6 ft (0.49 m) intervals. The Aluma beams

projected beyond the exterior trusses by various amounts, depending on the

width of bay in which they were used. The plywood forms were nailed to nailer

strips within the U-shaped top flanges of the Aluma beams (fig. 2.7).

Figure 2.9 shows the overall layout of the shoring system. Circled letters

and numbers designate transverse and longitudinal column lines, respectively.

The columns are located at the junctions of these lines and are shown as solid rectangles. To form a floor required the use of nine flying forms,

all of which were laid out in the transverse (east-west) direction. The flying

form assemblies consisted of three parallel trusses each, except in the case of

bay DE, where only two trusses were used. The flying forms had four different

configurations. These are identified by the letters A, B, C and D, enclosed

in hexagons. The location of the screw jacks are designated by solid circles.

The hollow squares designate reshores used in the lower stories. These were

4 x 4 in (102 x 102 mm) nominal, 3-1/2 x 3-1/2 in (69 x 69 mm) actual, white

pine single posts. They were installed using wooden wedges at the bottom.

Ellis type wood shores were used to support narrow filler form strips along

the transverse column lines between the edges of adjacent flying forms.

Unless otherwise noted, the following sequence of construction is based on construction schedules supplied by the general contractor and on worker

statements

.

Figure 2.10 shows a typical forming schedule for the upper stories. Construction progressed at the rate of one story per week, according to the following

cycle. The south half of a particular floor slab was cast on a Tuesday. The

placement of steel bars in the north half of the slab and casting of the columns

8
over the south half were carried out the next day, or Wednesday. On Thursday,

the columns in the south half were stripped, the north half of the slab was

cast and the forms below the south half of the slab were lowered. On Friday,

the forms in the south half were replaced by reshores and flown to the same location in the story directly above, and the columns over the north half were

cast. The following Monday, the columns in the north half were stripped, the

steel bars in the south half of the slab above were placed, and the flying forms

below the north half were lowered, replaced by reshores, and flown to the story

directly above.

Figure 2.11 shows the various casting and stripping dates and the ages of the

columns and floor slabs relative to the time of the collapse. In the case of

incomplete information or conflicting witness statements regarding the date of

a particular event, a question mark appears next to the specified date in the

figure. Note that, unlike the lower story slabs, the roof slab was scheduled

to be cast in one day. At the time of the collapse, all but the two north bays

of the roof slab had been cast.

The flying forms were flown out of the west side of the building where the

mobile crane was located. The forms under the fifth floor slab were all flown

the same day proceeding consecutively from south toward north. The procedure

for removing forms is reconstructed on the basis of worker statements and verified against that recommended by the supplier. It consists of four steps as follows: (1) the form is lifted hydraulically at the four corners, (2) the screw jacks are then loosened and folded, (3) the form is lowered on rollers

and moved out gradually, and (4) the portion vacated by the form is reshored.

Figure 2.11 shows full reshores in the second through fourth stories. A fully

reshored story conforms with the reshoring layout shown in figure 2.9. There

is sufficient indication from worker statements to assume that the fourth story

was fully reshored at the time of the collapse. Likewise, it is known that

there were no rehores in the first story with the exception of some peripheral

reshores under the balconies and walkways (see sec. 3.3 and fig. 2.11).

Although there is a question as to when the first story reshores were removed,

they would have been recovered for reuse when the fourth story forms were flown.

Note that the removal of reshores from the first story marks the first time

when the forces in the total shoring system will no longer have a direct path

independent of the columns to the foundation (slab on grade, in this case).

Based on worker statements (see section 3.3) and photographic evidence, there

is some doubt as to whether the third and second stories were fully reshored

at the time of the collapse. According to one worker account, only two lines

of reshores were used in each bay of the second story. This would make the

number of reshores in that bay about two-thirds that of a fully reshored story.

Other accounts mention cracked reshores in the fourth story being replaced by

reshores removed from the lower stories, or reshores being taken out to allow

masonry work to proceed, etc.

Figure 2.12 is a partial view of the east facade of the building between the

south stairwell and the elevator tower. Because the building was viewed from

the northeast side, it is possible to identify all the columns on column lines C

9
and D, as indicated in the figure. Within the third story, it is shown, by

means of the superimposed graphics in the figure, that the projected space

between columns Cl and D3 (the distance between Cl and D3 in the figure) should

be occupied only by eight reshores within bay CD, assuming that story is fully

reshored. From figure 2.12 and the closer view shown in figure 2.13, the third

story appears to be fully reshored at the time these photographs were taken.

It also appears that the number of second story reshores is comparable to that

in the third story, although their locations appear to be different. Noting

that the flying forms within the south four bays of the fourth story are in

place, the date of the photographs is placed between March 12 to March 15, or

12 to 15 days before the collapse (see figs. 2.10 and 2.11). It is not known

whether or not any of these reshores were removed in the interim.

Figure 2.14 is an aerial photograph of the building viewed from the southeast

side. The direction of the shadows and the presence of the steel reinforcement

for the roof slab indicate the photograph to have been taken on the morning of

March 26, or the day before the collapse. The steep angle of the photograph

conceals the view of interior reshores in the fourth story and below. It is

noted that at the south end only one exterior reshore is in place in each bay

of the fourth story while no reshores are in place at the same location in the

third story. This contradicts the reshoring plans supplied by the general

contractor where two reshores are shown in each bay along the exterior edges

of the slabs. It is possible that some of the original reshores were removed

to place masonry units in the third floor, as indicated by worker statements.

Because of the uncertainties noted above, for the purpose of analysis, all

three stories below the fifth floor were assumed to be fully reshored, while

no reshores were assumed within the first story (see fig. 2.11). The effect of

this assumption would make the calculated stresses less critical than if partial

reshoring were assumed in the second and third stories.

2.5 EVENTS AT THE TIME OF THE COLLAPSE

The collapse occurred at approximately 3:00 p.m., on March 27. Based on worker

statements, it appears to have been triggered by a failure in the fifth story

slab somewhere in the center portion of the building, and propagated straight

down to the ground. This topic is discussed further in section 3.3.

Figure 2.14 provides a fairly close indication of the stage of construction at

the time of the collapse. In the south end, the exterior walls for the first

two stories were in place and stacks of masonry units were on the third floor.

In the north end, construction of masonry walls was completed in the first

three stories and masonry units were stacked on the fourth floor. Masonry

loads are taken into consideration in the analysis of the building (sec. 5.3).

Figure 2.15 shows the location of the workers in the building at the time of

the collapse. Hollow and solid circles designate respectively survivors and

workers killed in the crash. The two workers in the first story were installing the framing for the window units. The workers in the third story were mostly masons finishing up the masonry walls in that story. The two survivors

in the fourth story were making leveling adjustments in the formwork, while the

10
third worker near the south bay was cleaning up the debris in that area. The

workers on the roof were engaged in concreting operations for the roof slab.

They were mostly within the two north bays where concrete had not yet been

placed. According to worker statements, no concrete was being delivered to the

roof at the time of the collapse.

Figures 2.16 and 2.17 are photographs of the debris taken in the wake of the

collapse. Figure 2.16 gives a full view of seven bays of the collapsed building. The bays can be identified by referring to the superimposed letters

corresponding to the column lines shown in figure 2.1. Note that many of the

columns, three and four stories high, are still standing, indicating that the

second to fourth floor slabs had been punctured at the columns on their way

down. A partial view of the flying form in the second bay from the north end can be seen in the figure. This is one of the two flying forms at the north

end (fig. 2.8) that remained relatively undamaged in the collapse because it was not supporting freshly cast concrete. A partial side view of the same flying form appears in figure 2.17 which was taken from the west side of the

building after the flying form in the north bay had been removed.

2.6 SUMMARY

The following summarizes the information about the design and construction of

the building which were used in subsequent analysis of the collapse. The

documents on the collapse were supplied by the OSHA area office in Tampa.

1. Based on construction documents and the superintendent's daily log, the

construction schedule, from the casting of the first floor slab up until

the day of collapse, was established.

2. Flying forms were used to cast the slabs for the second through fifth

floors at a rate of one per week.

3. Omissions were noted in the structural design calculations. There were no calculations to indicate that the structure would satisfy the deflection

and punching shear requirements of the ACI Code. The value of effective

depth of the floor slabs used in the design was not defined explicity.

4. The top reinforcement within the column strips of the slab was supported by

4-1/4-in (108-mm) individual high chairs, and this resulted in an effective

depth about 1 in (25 mm) less than if 3/4 in (19 mm) cover were provided,

as indicated in the structural drawings. The lower effective depth reduced

the structural strength of the slab.

5. At the time of collapse, the three stories below the flying forms were

reshored. However, it could not be firmly established how many reshores

were in the second, third and fourth stories. Consequently, for the purpose of subsequent analysis, full reshoring was assumed in the three stories

below the flying forms.

11
Table

2.1

Column

Schedule

12
INDIVIDUAL HIGH CHAIR (414”)

\
Table

2.1

Column

Schedule

12
I

.

j

.

I i I . 1 > I.aJ w

Figure 2.1 Top reinforcement layout in typical floor slab

13

INDIVIDUAL HIGH CHAIR (414”)

\

9

'

. I . I 1 1 1 1 1.1 -1.J

,

Figure 2.2 Bottom reinforcement layout in typical floor slab

14

INDIVIDUAL HIGH CHAIR |4!4”)

D = Bar diameter

Figure 2.3 Configuration of chairs and top reinforcement in column strip

15
Figure 2.4 Field measurement of top bar cover in column strip

Figure 2.5 Field measurement of chair heights for top bars

in column strip

16
—COLUMN

ci +d A v =2d(ci+C2+2d)

A v = Area of Critical Section

Figure 2.6 Area of critical section for punching shear
SPECIAL

NAILER

STRIP

(1

5/8” x1

9/16”

ACTUAL

DIM )

18

TYPICAL

PANEL

ASSEMBLIES
19

Figure

2.8

Flying

forms

from

the

two

north

bays

recovered

after

the

collapse
20

Figure

2.9

Layout

of

flying

forms

and

reshores
CONSTRUCTION CYCLE

TUESDAY WEDNESDAY

THURSDAY

FRIDAY MONDAY

Cast S(l/2) Slab

Place Steel N(l/2), Cast Columns S(l/2) Strip Columns S(l/2), Cast N(l/2) Slab, Drop Forms S(l/2)

Set Forms S(l/2) Slab, Cast Columns N(l/2) Strip Columns N(l/2), Place Steel S( 1/2) , Drop and Fly Forms N( 1/2)

Figure 2.10 Typical forming schedule
LEGEND NO.

/NO. :

CASTING

DATE

(month

/day)

(NO.

/NO.)•

STRIPPING

DATE

(month/day)

(NO) :

AGE

AT

TIME

OF

COLLAPSE

(days)

00

22

Figure

2.11

Assumed

state

of

construction

at

time

of

the

col

lap
© © © © © ©

2.12 Schematics for determining amount of reshores using

partial view of east side of building
24

Figure

2.13

Partial

view

of

bay

CD

from

east

side

of

building
25

Figure

2.14

Aerial

view

of

building

taken

from

southeast
KILLED

IN

THE

COLLAPSE

SURVIVING

WORKERS

Q.°a op%

•

i \i\l\N ,

\ . 1 •

°l\ 4

o

o

o

o

o

•

N K

ri-i -i \ •

i o

• r

\

° \ \ \° \ \ \T3333TTTTa

eg

26

Figure

2.15

Location

of

workers

in

building

at

time

of

the

collapse
27

Figure

2.16

General

view

of

wreckage

after

the

collapse
28

Figure

2.17

View

of

debris

at

north

end

of

building

after

the

collapse
3 SITE INVESTIGATION

3.1 INTRODUCTION

Section 3.2 presents observations and measurements made by NBS during two site

visits to Cocoa Beach, Florida, and sec. 3.3 summarizes the statements made by

individuals who were interviewed by the OSHA officials during the course of the

investigation. The first site visit was made on March 28-30, 1981, one day

after the collapse, and the second visit was made on April 7-8, 1981. During

the visits, the structural components were examined visually as they were removed for rescue operations, selected key components were identified and

marked with paint for later removal of drilled cores, and measurements of member

dimensions were made.

3.2 INVESTIGATION AT THE SITE

A thorough inspection of the undisturbed wreckage for possible clues as to the cause of the collapse was ruled out during the rescue operation, which involved

over 200 people and was not completed until the morning of March 29. Instead,

a substantial number of photographs were taken before and during the removal

of debris. The stacked configuration of the second through fifth floor slabs

(as seen in figure 2.17) showed no evidence of sidesways or overturning. Examination of the debris and photographs did not provide any further clues than

those supplied by workers as to the exact location of the origin of the collapse.

In order to carry out rescue operations, the floor slabs were cut into

approximately 20 ft (6 m) square pieces by air hammers and removed by mobile

cranes. All slab layers (2nd, 3rd, 4th, and 5th floor) were removed in this manner. Some of the slab sections and columns were salvaged and their location

in the building was identified (figure 3.1). They were the source of core samples which were subsequently tested for ingredient content and compressive

strength (chapter 4). The slab sections were also used to determine size and

location of reinforcing bars and support chairs (sec. 2.3).

Careful examination of a number of column pieces revealed the following:

1. The slabs broke away from the columns at the slab-column interface

(figure 3.2).

2. One or two bottom slab bars were placed through some columns. In many

other columns no such evidence was noted.

3. In some cases, the vertical reinforcement above the floor level was bent

severely during fabrication (figure 3.2).

4. Column size measurements agreed with the structural drawings.

It was also noted from the first-story column sections still standing on the

ground level that where 8- # 10 or 8- # 11 bars were lapped, in a 10 x 18 in

(250 x 460 cm) column section, the total steel area exceeded considerably the

maximum limit of four percent of the gross area of the concrete allowed by the

29
ACI Code [2.1]. As shown in figure 3.3, the vertical reinforcement was so congested that there was hardly any space left between adjacent bars to allow

the flow of concrete during placement.

The slab thickness was measured at several locations where concrete cores were

taken. In all cases, the thickness was within +1/8 in (3 mm) of 8 in (200 mm).

Further data on slab thickness were obtained from measurements of the lengths

of the core samples (chapter 4). These measurements confirmed the agreement of

the slab thicknesses with the construction drawings.

The onsite batch plant, located southeast of the building, had a single

compartment hopper into which a front-end loader deposited both coarse and

fine aggregate. The cement was weighed by a separate hopper. The onsite batch

plant was equipped with large tanks containing the various admixtures that were

added to the concrete. Smaller admixture tanks on the west side of the building contained the superplasticizer that was added prior to discharge of the

concrete truck.

The weighed materials were deposited into a 15 cubic yard (11.5 m^) ready-mix

concrete truck for mixing and delivery to the discharge point adjacent to the

building under construction. The distance of travel of the truck between the

onsite plant to the discharge point was about 150 yd (137 m)

.

Samples of concrete-making materials were obtained from the onsite plant and

from the ready-mix plant which also supplied concrete for the project. Some

samples were sent to a testing laboratory for analysis, and others were sent to

NBS for use in making concrete mixtures (see chapter 4).

3.3 INTERVIEW STATEMENTS OF WORKERS

As part of the investigation, OSHA compliance officers interviewed workers and

others who were involved in the construction of the building. The statements

obtained from the interviews provided valuable information which served as source documents in reconstructing the events prior to the collapse. This sec- tion summarizes their contents pertinent to the investigation. These include

the location of reshores on the day of the collapse; deficiencies encountered

in the performance of the building as it was being constructed; quality of the

concrete; and eyewitness descriptions of the collapse. In the following sections, the numbers in parentheses refer to numbers assigned to the interview

statements

.

3.3.1 Location of Reshores

One of the critical elements in the determination of the cause of the collapse

is the location and number of reshores in the building at the time of the

collapse. The interview statements contained conflicting testimony. Four

employees (#7, #50, #53, #65) stated that reshores were in position on the

second, third, and fourth floors. On the other hand, others (#8, #9, #12, #22,

#25, #40) indicated that there may not have been the full amount of reshores on

the second and third floors. One of the workers (#8) claimed that reshores

were removed from the lower floors for use on upper stories, but it was also

30
stated (#39) that enough posts were put into the building to fully reshore

three floors. Other workers (#6A, #8, #37, #40) said they saw masons remove

some posts or saw posts fall down which were not replaced. These statements

were contradicted by others who said the masons did not touch the posts while

placing concrete block walls (#63) or that posts were periodically replaced

(#7).

Several workers (#7, #10, #18, #22, #24, #32, #44, #45) stated that they saw

bowed reshores and two workers (#32, #44) saw two reshores on the fourth floor

break or fall while concrete was being placed on the roof.

According to a supervisory worker (#7) there was no reshoring plan to follow,

and one of the superintendents directed him to change from two to three rows

of reshores per bay for the third floor and above. There were only two workers

who described how the reshores were installed (#45, #53); the reshores were

held in position by wedges inserted between the floor slab and the reshores.

There is no indication that the wedges were driven in so as to lift the slab

and relieve its self weight.

In summary, based on the personal interview statements, it is not possible to

determine firmly the reshoring layout at the time of collapse. The fourth

floor probably contained the full number of reshores as indicated on the

reshore drawing supplied by the general contractor after the collapse. The

first floor probably contained no interior reshores (some shores were probably

present along the balcony and walkway beneath the block walls placed on the

second floor) . The precise number and locations of reshores which existed on

the second and third floors is unknown.

3.3.2 Construction Deficiencies

There were many statements by workers of different trades concerning the

existence of cracks in all the slabs. The testimonies indicate that the cracks

were noticed shortly after the flying forms were removed and that they were

located around the columns and near the midspans. These cracks were marked

with crayon by one of the superintendents. On March 10, 1981 the structural

engineer inspected the building; the second and third floors were in place by

that time. According to one of the superintendents, the engineer suggested

that additional reinforcing bars be added in the midspan of the walkways.

Similar cracks subsequently developed in the fourth and fifth floor slabs. A

worker (#34) described them as "spider cracks," and reported that upon removing

plumbing sleeves from floor slabs, the cracks could be seen extending to a depth of 4 to 5 in (102 to 127 mm) into the slabs.

In addition to the cracks, it was reported that noticeable deflections occurred

in the slabs after the forms were removed. According to one of the superintendents, the deflections were most evident in the north and south end apartment

units; on the second floor, a deflection of 1-3/4 in (45 mm) was noted in the

end apartment on the north side of the building.

Four employees (#2, #4, #38, #63) mentioned another significant problem which

occurred during construction. During placement of the roof, a water puddle

31
formed on the east side near the elevator tower, indicating that a low point

existed. Additional concrete (an unspecified quantity) was added to the

depressed area but it was reported (#2, #4) that the water puddle reappeared

after refinishing the area. Apparently (according to #38 and #63), the same problem occurred at the same location during placement of the fourth floor.

Three workers (#40, #45, #73) mentioned that the construction joints at the

middle of the building were not of good quality. They were described as "ragged", "cracked", and "honeycombed". Two workers (#45, #14) mentioned that

they saw rebars showing on the faces of some columns. The incomplete filling

of the column forms was probably due to congestion of steel (see figure 3.3),

which, according to a supervisor (#72) was so severe in some of the second and

third floor columns that the top reinforcing steel in the slabs could not pass

through. In order to pass slab bars through some columns, the protruding column

reinforcement was "bunched together" and the slab bars were placed alongside

the column bars (figure 3.2).

3.3.3 Quality of Concrete

The concrete for the building was supplied by an onsite batch plant operated

by the general contractor and by a ready-mix company. According to the onsite

batch plant operator (#67), the ready-mix company supplied about one- third of

the concrete used in the floor slabs. All of the concrete for the columns came

from the onsite batch plant. Chapter 4 will contain more details of the mix

proportions and the results of strength tests by the testing laboratory. In

this section, workers' impressions of the concrete quality is presented.

Based on the Department of Professional Regulation's report and the diaries of

the construction manager and the general contractor's superintendent, it was determined that the developer's superintendent, who usually oversaw the batch

plant operation, was absent from February 2, 1981 until March 9, 1981. During

the week of March 2, 1981, the general contractor's superintendent was also on

vacation. A worker (#4) stated that during the placement of the second and

third floors, the concrete appeared to have erratic slump; he estimated that

the slump varied from three to nine inches. Referring back to the construction

schedule (section 2.4), it is seen that the third floor was placed while both

job superintendents were absent. The same concrete finisher stated that the

texture and slump became more consistent after the developer's superintendent

returned to the job site.

Several workers (#37, #74, #22, #34, #63) who talked to masons mentioned that

some of the concrete was difficult to finish and tended to set up quickly.

This may be attributed to the slump loss [3.1] associated with the use of

superplasticizers (which was added to the concrete in the trucks prior to

placement). Apparently on some batches with lower than normal water contents,

it was necessary to add more than the usual quantity of superplastercizer to

produce the proper consistency for placement (see section 4.2). By the time

finishing operations were performed, the workability was reduced due to slump

loss; during finishing the concrete reflected its low water content. According

to the statements of two workers (#34, #74), this problem occurred only with

the concrete prepared on site.

32
3.3.4 Description of the Collapse

Descriptions of the collapse were provided by 28 persons, 13 of whom were

located in the building. Ten of those in the building who described the collapse were on the roof, one was on the fifth floor, another on the fourth, and one on the third floor. The remainder were on the ground at various locations

with respect to the building.

According to those on the roof at the time of the collapse, concrete was being

placed at approximately column line H (see figure 2.1). Thus, the two flying

forms at the north end of the building did not have concrete on them. One

worker (#4) said he had felt the building tremble two or three times during the

day without any apparent reason. He was on the form between column lines H and

J and near the west side when the collapse occurred, and he described the plywood as tilting toward the south and east. He described the fall as "one long

slide with one thump." Another worker (#7), who was also on a dry form, stated

that his "feet were suddenly knocked out from under" him and noticed that the

finishing crew had fallen to a lower level. An employee (#18) working 50 to

60 ft (15 to 18 m) south of the finishers heard a rumble and saw the roof caving in at the position of the pour; the "wave" moved toward him as he turned

away toward the south. Other witnesses (#2, #6, #26, #38) also indicated that

the apparent origin of the roof collapse was in the vicinity of the boundary

of freshly placed concrete, i.e., column line H. A worker (#26) who rode down

on the dry form near the north end stated that he "could see that the middle of

each floor gave way as each floor was collapsing."

A person working on the fifth floor (#8) said he heard a "pop” about 5 to 10

minutes prior to the collapse. He was standing near the center of the building

near column line H when "the fifth floor fell out from under" him. He further

stated that "the floor fell straight down with slight hesitation at the fourth

and third floors."

A worker on the fourth floor (#74) also believed the collapse originated in

the middle of the building but could not say which floor fell first. However,

he stated that he didn’ t think it was the top floor otherwise he would have

been "hit on the head." A person (#3) on the third floor who was working on a block wall near column line F stated that he "heard a crack overhead, looked to

the southeast and saw a big hole and it was coming towards him." He further

stated that he tried to run but the floor fell out from under him.

A person (#22) who was working at the garage northeast of the building offered

this account of the collapse: "I heard a crack and the center of the fifth

floor started down followed by the rest of the floor, and then the fourth,

third, and second floors right on top of each other with the roof of wet con- crete following. The floors appeared to slide right down around and over the

columns." The same general description was also supplied by others on the

ground (#15, #25, #27, #29, #40). An employee (#29) who was at the northwest

corner of the building stated that "the fifth floor, about the second set of

columns from north end of floor gave out and the building came down." A worker

(#40) who was north of the building also described hearing a crack and seeing

33
the fourth and fifth floors coming down to meet the third floor; he also saw

the middle of the building "go in" first.

Different descriptions were provided by persons at other vantage points. A

worker (#39) who was on the northwest side near the second bay from the north

end he stated: "I saw the extreme northwest corner of the fourth floor near

the columns starting to fall, tilting eastward, then the rest of the fourth

floor fell, then the fifth floor and the flying forms caught up with the fourth

floor and the whole building collapsed."

A person (#42) working north of the building said he heard a loud crack like

wood splitting, looked up and saw the flying form going down while the fifth

floor was still in place. Another laborer (#43) who was 50 ft (15 m) away and

walking toward the northeast corner, heard cracking noises and witnessed the

south end of the building at the top falling first toward the north followed

by the collapse of all the columns.

3.4 SUMMARY

Based on the site visits, the following conclusions were drawn:

1) The failure mode appeared to be of the punching- shear type, because the

slabs had broken away from the columns at the slab-column interfaces.

2) There was no evidence that sidesway or overturning of the building had

occurred

.

3) Measurements of the outside dimensions of selected column and slab sections

indicated conformity with the structural drawings.

4) The reinforcing steel in some first-story columns did not meet the clear

spacing requirements of the ACI code.

The worker interview statements gave valuable information on the circumstances

prior and up to the collapse, and provided eyewitness accounts of the failure.

The following were some of the significant findings from these statements:

1) It was not possible to establish the exact layout of reshores at the time

of collapse. The first floor appears to have had only a few exterior

reshores in place, and the fourth floor probably had the full amount of

reshores. Conflicting statements were given as to the number of reshores

on the second and third floors.

2) Extensive cracking occurred in floor slabs when supporting formwork was

removed. The structural engineer was made aware of this problem.

3) The concrete produced by the onsite batch plant was non-uniform in terms

of consistency and f inishability . The greatest problems were reported

during the absence of both job superintendents.

34
4) Descriptions of the collapse by workers on the site indicate that the most

probable origin of the collapse was in the fifth floor slab in the vicinity

beneath where the roof concrete was being placed.

35
36

section

of

the

fourth

floor
37

Figure

3.2

Severely

bent

column

reinforcement
38

Figure

3.3

Congested

reinforcement

in

a

first

-story

column
4. LABORATORY INVESTIGATIONS

4.1 INTRODUCTION

The objective of the laboratory investigations was to establish the strength

of concrete in the structure at the time of the collapse. The approach used

involved gathering data from the site and preparing concrete mixtures representative of the inplace concrete. A commercial testing laboratory was contracted

by OSHA to carry out tests on concrete cores from the site. This laboratory

took drilled core samples from identified portions of the structure as directed

by NBS. The cores were used for compressive strength tests, petrographic examination, and chemical analysis. In addition, samples of the ingredients used

in the concrete were taken and subjected to physical and chemical analyses.

Based on the available mix design information, three concrete mixtures were prepared at NBS from which specimens were cast for strength testing at specified

ages. The results of these tests were compared with the core strength data.

Section 4.2 discusses the information provided by OSHA dealing with the

prescribed concrete mix proportions and the results of quality control tests

performed on the concrete.

Section 4.3 summarizes the results of tests conducted on samples of concrete

ingredients and drilled cores taken from the site after the collapse.

Section 4.4 presents the results of the NBS laboratory investigation. This

includes the strength-gaining characteristics of the concrete mixes, and the

predicted strengths of various portions of the structure at the time of collapse.

4.2 INFORMATION ON FIELD CONCRETE

OSHA interviewed the onsite batch plant operator (Univel) and the president of

the ready-mix company (Dixie) to determine the mix designs that were being

followed. Table 4.1 shows the stated mix proportions (saturated surface dry

basis) for one cubic yard of 4000 psi (27.6 MPa) concrete. The onsite batch

plant used the same proportions for the floor slab and column concrete, except

that 3/8 in (10 mm) maximum size aggregate was used for the columns, whereas

3/4 in (20 mm) maximum size was used for the slabs. The ready-mix plant only

provided concrete for the slabs and used 1 in (25 mm) maximum size aggregate.

All mixtures contained an air entraining agent (AEA) and a water reducer/set

retarder (WR/SR). In addition, a high range water reducer (HRWR) , or superplasticizer, was added by the general contractor prior to discharge from the

concrete truck mixers. According to the onsite batch plant operator, 2 gal* of

superplasticizer were added to 10 or 11 cubic yard* trucks and 1-1/2 gal were

added to 9 cubic yard trucks. A waterproofing admixture was added along with

* 1 gallon = 3.785 x 10“3 m3

1 cubic yard = 0.7645 dP

1 ounce = 2.457 x 10~^ m^

100 pounds = 45.36 kg

39
superplasticizer to concrete placed for balconies and walkways. The dosage

reported by the batch plant operator was 4-1/2 to 5 gals per truck. The manufacturer's recommended dosage for these admixtures were as follows: AEA, 3/4

to 3 oz per hundred pounds of cement; WR/SR, 5 to 8 oz per hundred pounds of

cement; HRWR, 15 to 20 oz per hundred pounds of cement.

The ready-mix company had a computer-controlled batch plant and employed a moisture meter in the sand bin. The onsite plant did not make moisture content

determinations, and based aggregate weights on assumed moisture contents of 1 percent and 3 percent for the coarse aggregate and sand, respectively.

Based on the available testimony, it appears that water was added until the

concrete reached a slump in the range of 3 to 5 in (76 to 127 mm) . The batch

plant operator stated that he performed slump tests on the first two trucks

each day, but made no record of these values. Additional water had to be

added to mixes that "tightened up" in the mixer, possibly due to absorption of

water by the highly absorptive stone when it was mixed in a dry condition.

Daily reports of the onsite batch plant indicated: the size of the batches;

the type of concrete; the truck number; the batch weights of cement, sand, coarse aggregate, water reducer and air entraining agent. The quantities of

ingredients shown on the reports agreed with the quantities called for in the

4000 psi (27.6 MPa) mix design. However, in most of the reports, the column

labeled "Gals. Water" contained no entry of the quantity of water added. When

the quantity of water was reported, it was observed that widely different quantities were reported for batches of the same size. This may reflect changes in

the moisture content of the aggregates, or that dissimilar concrete mixtures

were being produced . There were no copies of the delivery tickets of the

ready-mix concrete.

The general contractor employed a testing laboratory to inspect samples of

concrete prior to placement and prepare standard 6 x 12 in (152 x 305 mm)

cylinders for quality control testing. On the days of floor slab placement,

samples were taken from one truck from the ready-mix plant and one truck from

the onsite plant. Concrete used for the columns which was supplied by the

onsite plant was not sampled. Table 4.2 summarizes the information contained

in the test reports: the first column indicates the date and time of sampling;

the second and third columns indicate the supplier and where the concrete was

placed, respectively; the fourth column is the slump, which was measured after

the addition of the superplasticizer; the fifth and sixth columns give the test

age and compressive strength of the cylinders; and the last column indicates

the quantity of superplasticizer added. Note that only one cylinder was tested

at early ages, while two were tested at 28 days. Some 28-day strengths

apparently exceeded the capacity of the testing machine and may account for the

"+" after some strength values.

A study of table 4.2 reveals several significant facts. The typical slump was

in the range of 8 to 10 in (203 to 254 mm) except for samples of the onsite

plant concrete taken during the placement of the north half of the second

floor and both halves of the third floor. One might infer that these samples

had lower than normal water content, and this is borne out by the very high

40
early strengths of the corresponding cylinders. Both job superintendents

testified that the early age strength results were used to determine if sufficient strength had been reached for form removal. They testified that forms were removed only if the cylinder strength exceeded 2/3 to 3/4 of 4000 psi

(27.6 MPa). In general, the reported test ages agree with the formwork removal

schedule. However, it appears that the fifth floor samples were not tested

at ages corresponding to the formwork removal times. The practice of using

laboratory-cured cylinders intended for quality control testing for the purposes of scheduling formwork removal may be an unsafe approach if the inplace

maturity (see section 4.4.3) at test age is less than the maturity of the control cylinder. Finally, only three test reports give the quantity of plasticizer added to the mixes, and in each case it exceeded the quantity specified

by the batch plant operator. This may indicate that these mixes had lower than

average initial slumps and that more admixture was required to produce the

typical 8 to 10 in (203 to 254 mm) slump.

The strength and age data in table 4.2 are shown in figures 4.1 and 4.2 for the

onsite plant and ready-mix concretes, respectively. Also shown are the bestfit

curves to represent the data. In figure 4.1 two curves are shown, the solid curve includes all the data, while the dashed curve excludes those data of the

low slump samples for the north half of the second floor and both halves of the

third floor. In general, these data indicate that the 28-day strengths are con- sistent with the specified strength of 4000 psi (27.6 MPa). However, samples

of concrete from the fifth floor are weaker than samples from other floors.

4.3 TESTS OF SAMPLES FROM THE SITE

A commercial testing laboratory was retained by 0SHA to secure samples from the

site and carry out the pertinent tests. This section summarizes the information

obtained from the report of the testing labortory

4.3.1 Concrete Ingredients

Samples of cement, aggregates, water, and admixtures were secured from the

onsite plant and from the ready-mix plant. Appendix A contains the reported

results of physical and chemical tests on the samples. Table A.l shows the

physical characteristics of sand and coarse aggregate samples. The coarse aggregates had higher than typical absorption capacities for normal weight

aggregates. The sands were finer than typical concrete sands, and the coarse

aggregates did not, in general, have uniform particle size distributions which

may explain some of the finishing difficulties mentioned in section 3.4.3.

The results of chemical analyses of cement samples are presented in table A. 2.

Except for differences in setting times, there were no significant differences

between the samples from the two batch plants. The samples meet the requirements of ASTM C150 [4.1] for Type I cement. Because samples were taken only

after the collapse, it is not known whether all cement in the structure agreed

with the results in table A. 2.

41
Tables A. 3 and A. 4 show the results of tests on water and admixture samples.

Although a pond water sample was chemically analyzed, there were no indications

from the available data that pond water was used in the concrete.

4.3.2 Core Strength

Core samples were drilled from column and slab sections from April 1 to

April 3, 1981. Some cores were used for compressive strength determinations

and others for petrographic and chemical analyses. Table 4.3 indicates the

identification numbers and locations in the structure, as identified by NBS,

of six column sections and 14 slab sections from which cores were drilled.

Also shown are the number of cores tested in compression from each section.

An attempt was made to obtain at least six strength values from each floor

level in the structure.

The cores from each part of the structure were tested at three ages: soon after core samples were obtained and at seven and 14 days subsequent to the

first test. The samples were stored in plastic bags to preserve the inplace

moisture conditions until tested. Table A. 5 gives the dimensions, ages and

strengths of the tested cores. The cores had nominal 2, 3, and 4 in (51, 76,

and 102 mm) diameters. The reported compressive strengths include the standard

(ASTM C 42 Standard Method of Obtaining and Testing Drilled Cores and Sawed

Beams of Concrete) correction factors for cores with an aspect ratio less than

two [4.1]. Figure 4.3 shows the floor slab core strengths as a function of

their test ages. Cores from the south half are identified by a shaded symbol.

Except for two 23-days old cores from the south half of the fifth floor (and a severely damaged core from slab section 13), the strengths were greater than

the specified strength of 4000 psi (27.6 MPa). However, all cores (except the

damaged one), satisfy the "85 percent of specified strength" criterion in

section 4. 8. 4. 4 of the ACI Code [2.1].

Figure 4.4 shows the compressive strengths of cores drilled from columns. Since

it was not known whether the column sections were from the north or south halves

of the building, the age shown in Figure 4.4 is the average age of the north and

south columns on each floor. It is evident that the concrete in the columns was weaker than that in the floor slabs. The average strength of the fifth story

cores does not satisfy the ACI Code strength criterion mentioned above.

The lengths of unbroken cores from the slab sections had an average value of

8-1/16 in (205 mm) with a standard deviation of about 3/16 in (5 mm). Four cores were about 6-1/2 in (165 mm) long which indicates cores from the walkway

or balcony areas.

4.3.3 Petrographic Analysis

Nineteen 6 in (152 mm) diameter cores were used for subsequent petrographic

and chemical analyses. Eleven cores were from column sections and eight from

slab sections as indicated in table 4.3. Polished sections were examined with

a stereo-microscope according to ASTM C 856 (Standard Practice for Petrographic

Examination of Hardened Concrete [4.1]. The polished sections were also examined according to ASTM C 457 (Standard Practice for Microscopical Determination

42
of Air-Void Content and Parameters of the Air-Void System in Hardened Concrete)

[4.1], from which the volume fractions of paste, sand, coarse aggregate and air were measured. In addition, specific gravity, percent absorption and percent

voids were measured according to ASTM C 642 (Standard Test Method for Specific

Gravity, Absorption and Voids in Hardened Concrete) [4.1].

Table A. 6 in appendix A gives the results of the above analyses for each core;

these results are summarized in table 4.4. For quantitative determinations

the numbers in table 4.4 represent the average values and the numbers within

parentheses are the standard deviations of the averages. The similarity of

the slab and column cores in terms of percent volumes of the ingredients

confirms the statements that the same mix proportions were used in the columns

and in the slabs. An unexplained finding is the color of the cement paste:

most of the fifth floor cores were gray, while the cores from the other floors were buff colored. The columns had slightly higher air content than the slabs,

but in each case the air bubbles were characteristic of entrapped air (large

size, large spacing) rather than entrained air. This phenomenon is consistent

with a reported effect of superplasticizers [3.1].

4.3.4 Chemical Analysis

In addition to petrographic analysis, the core samples were used for chemical

analyses to determine cement and admixture content. Analysis for cement was carried out in accordance with ASTM C 85 (Standard Test Method for Cement Content of Hardened Portland Cement Concrete) [4.1]. However, it is recognized

that the method is imprecise when the concrete contains significant amounts of

calcareous aggregates, which is the situation in this case since the coarse aggregate was calcareous and the sands also contained shell fragments. Separate

tests were performed on samples of fine and coarse aggregates so that an allowance for the lime in the aggregates could be made when computing the cement content. Cement analyses were performed on seven column cores and six slab cores. The calculated cement contents, expressed as percentages of the unit dry weights

of the concrete samples, are given on the last line of table A. 6. Using the

weight percent of cement and the corresponding dry specific gravity of each

sample, the cement content was expressed in pounds per cubic yard of concrete.

These quantitites are given in table A. 7 of appendix A, and the average values

are given in table 4.5. Using the volume percentage of paste (table A. 6) and

cement content, the approximate water content was calculated and the values are shown in table A. 7, while the average values are given in table 4.5. Likewise,

using the percentage volumes of sand and coarse aggregate, and their corresponding specific gravities (saturated surface dry), the weights of sand and coarse aggregate per cubic yard of concrete were calculated. Table 4.5 also shows the

average aggregate contents for the column and slab cores.

Study of table 4.5 revealed the following points: (1) The calculated weights

of sand and coarse aggregate are in excellent agreement with the quantities of

the mix designs shown in table 4.1. (2) The calculated cement contents are significantly lower than the mix design values. (3) The calculated average

water cement ratios are 0.65 and 0.72 for the columns and slabs, respectively.

The low cement contents and corresponding high water-cement ratios are inconsistent with the core strengths. It is highly improbable that these high water

43
cement ratios would result in the strengths shown in figures 4.3 and 4.4. This

leads one to suspect that the calculated cement contents are not precise.

In addition to cement content analyses, the cores were examined for their

admixture contents. The general procedure involved making reference samples of

concretes with different admixture combinations; then, a solution of calcium

carbonate was used to extract the admixtures from the samples; finally, the

extractions were analyzed by infrared spectrometry and a comparison was made

between the reference samples and samples from the cores. The procedure is

best suited for sulfonated admixtures which include the water reducer/ set

retarder and superplasticizer that were used on the job. Table 4.6 summarizes

the results of the analyses. The cores are identified in terms of their location as described in table 4.3. The quantities of admixture are expressed as

a percent of the weight of concrete. According to the infrared analyses, the core samples had signature spectra that corresponded with reference samples

made with the water reducer/set retarder or with a combination of superplasticizers (HRWR) and accelerating admixture. At the bottom of the table are shown

the quantities of the admixtures used in the laboratory mixes (to be discussed

in the next section) representing the ideal mix designs, and the manufacturer's

recommended dosages for a cement content of 530 lbs per cubic yard (315 kg per

m3) .

There are several contradictions between table 4.6 and the information obtained

from the site. First is the appearance of the accelerating admixture, which

according to the onsite batch plant operator was not used in the columns or floor slabs. For the columns, the measured quantities of water reducer/ set

retarder are greater than the mix design values for the onsite plant. For the

floor slabs, the measured quantities of superplasticizer and accelerator are in

general lower than the mix design values. i Some cores from the fifth and fourth

floors had higher than expected quantities of water reducer/set retarder.

In summary, some of the chemical analysis results contradict or are inconsistent with other information. It is difficult to conclude whether such

differences are real, or whether they are due to the difficult nature of the

chemical analyses and the inherent uncertainties in the results.

4.4 NBS LABORATORY INVESTIGATION

The purpose of the NBS laboratory investigation was to establish the strength

of the concrete in the structure at the time of the collapse. This was to be

accomplished by developing the strength-maturity relations of the column and

slab concretes. In addition, it was necessary to determine the elastic moduli

of the concretes for inclusion in the structural analyses of the structure.

Finally, it had to be established whether the concretes had typical tensile

strength properties.

4.4.1 Sample Preparation

Samples of aggregate, cement, and admixtures were obtained from the onsite

batch plant and from the ready-mix plant. These materials were used to prepare

44
the concrete mixtures to be discussed. The original plan called for the

preparation of concrete batches according to the stated mix designs as indicated in table 4.1, and if necessary, for the preparation of mixes according

to the results of the petrographic and chemical analyses of the core samples.

Thus, three mixes were initially prepared: Mix I represented the concrete

supplied by the ready-mix plant; Mix II, the onsite batch plant concrete for

the floor slabs; and Mix III, the onsite plant concrete for the columns. It was assumed that the column concrete contained superplasticizer in the same proportion as the slab mixes.

In order to minimize the quantity of concrete that had to be mixed, 4 x 8 in

(102 x 203 mm) cylindrical specimens were used. However, some 6 x 12 in

(152 x 305 mm) cylinders were made for checking the results of the smaller

specimens. Forty-five, 4 x 8 in (102 x 203 mm) cylinders and seven 6 x 12 in

(152 x 305 mm) cylinders were molded from a 4-1/2 cu ft (0.13 m^) batch of

each mix. A 9 cu ft (0.25 m^) capacity, revolving drum mixer was used.

Table 4.7 shows the batch quantities used for each mix. Coarse aggregate, sand

and cement were placed in the mixer in that order. Water was added in two portions, one portion contained the air entraining agent, the other contained the

water reducer/set retarder. The initial quantity of water was less than shown

in table 4.7. After a period of initial mixing, a slump test was performed and

additional water was added as needed to achieve a slump of three to five inches.

When the desired slump was attained, superplasticizer was mixed into the con- crete and slump, air content and unit weight determinations were made. As

shown in table 4.7, addition of superplasticizer at the dosage rate of 2 gal

per 10 cubic yards (9.9 x 10”^ m^ per m^) of concrete produced an increase in

slump of about 3 in (76 mm) Referring back to table 4.2, which shows the slump

values for samples from the job, it may be concluded that in most cases either

more than the above dosage of superplasticizer was used or the initial slumps

were greater than about 5 in (127 mm)

The air contents of the three mixes are in good agreement with the results of

the petrographic analyses of core samples (table 4.4). Note that Mix I (representing the ready-mix concrete) contained more air than Mix II (the onsite plant

concrete), which is consistent with the higher quantity of air entraining agent

in Mix I. Mix III (containing small aggregate) had the highest air content,

which is consistent with the known effect of aggregate size on air content:

small aggregates tend to entrap more air in the mix than large aggregates.

When the aggregates were being weighed, samples were taken for moisture content

determination. The moisture content values along with the batch weights and

specific gravities, were used to calculate the volume percentages of paste and

aggregates in each mix. Table 4.8 shows the results of these calculations, and

for comparison the values from the petrographic analyses of the cores are also

shown. There is good agreement between the laboratory mixes and the concrete

in the structure. Thus, it was concluded that concrete used in the building

agreed fairly well with the stated mix designs in terms of the relative volumes

of paste, air and aggregates.

45
After completing the fresh concrete tests, the cylinder molds were filled as follows: the 6 x 12 in (152 x 305 mm) molds were filled in the standard manner

using three layers and 25 strokes of the tamping rod; the 4 x 8 in (102 x

203 mm) plastic molds were filled in two layers, and each layer was compacted

by striking the outside of the mold 25 times with a rubber mallet. It was felt

that rodding the small specimens would tend to cause segregation of the coarse aggregate toward the bottom of the molds. Specimen preparation was completed

about two hours after mixing was begun.

Since the available information suggests that no special curing procedures were

employed during construction, the molded specimens were kept in their molds

with the top surfaces uncovered and allowed to cure in the laboratory at an ambient temperature of about 70°F (21°C). For comparison, seven 4 x 8 in

(102 x 203 mm) specimens and the seven 6 x 12 in (152 x 305 mm) specimens were

covered with plastic and subsequently cured under water at 70°F (21 °C). For

each batch, thermocouples were placed in two uncovered small cylinders, in one covered small cylinder and in one 6 x 12 in (152 x 305 mm) cylinder. Temperatures were monitored by a data logger at regular time intervals, and the resulting data was used to calculate maturity of the specimens when they were tested.

4.4.2 Test Results

The cylinders were used for compression tests according to ASTM C 39

(Standard Test for Compressive Strength of Cylindrical Concrete Specimens) and

for splitting tensile tests according to ASTM C 496 (Standard Test for Splitting

Tensile Strength of Cylindrical Concrete Specimens) [4.1]. In addition stressstrain curves were recorded during the compression tests of all 4 x 8 in (102 x

203 mm) cylinders. The moist-cured cylinders were not tested for splitting

tensile strength. Specimens were tested at the following ages: 0.5 days, 1.3

days, 3 days, 7 days, 14 days, and 28 days. Figure 4.5 shows the compressive

strength as a function of age. Each point represents the average of three

cylinders. The shaded data points represent the average strengths of the moistcured specimens which were tested at 7 and 28 days. It appears that the method

of curing plays a more significant role on strength development in the weaker

concrete. The results of Mix III were significant. Apparently, the increased

air content and increased water demand of the small aggregate mix resulted in

significantly lower compressive strength than the other mixes.

Table 4.9 compares the average strengths and standard deviations of moist-cured

4 x 8 in (102 x 203 mm) and 6 x 12 in (152 x 305 mm) cylinders. The results

indicate that there were no statistically significant differences between the

two sizes, and confirm that using the smaller specimens is an acceptable

alternative

.

In order to check whether the strength-gaining properties of the laboratory

mixes were consistent with the core strengths, the best-fit curves in figure

4.5 were superimposed on the core strength versus age data. Figure 4.6 shows

the curves for Mixes I and II compared with the core strengths of the slab

concretes. The agreement is reasonable. The third floor core strengths agree

with the strength development curve of Mix I. The generally greater strength

for the third floor cores is consistent with the results of the inspection

46
laboratory reports (table 4.2), which showed that the slumps for the onsite

plant concrete used on the third floor were lower than normal. Mix II has a strength development curve consistent with the core strengths of the second,

fourth and fifth floor slabs.

Figure 4.7 compares Mix III with the column core strengths. The laboratory mix

is consistent with these core tests.

The overall agreements noted above led to the conclusions that, in general, the

laboratory mixes were representative of the concrete in the structure. Thus,

the strength development characteristics of these mixes were used for estimating the inplace strength of various parts of the structure at the time of the

collapse

.

Figure 4.8 shows the increase with age of the secant modulus of elasticity at

50 percent of the ultimate strength for the laboratory mixes. Mixes I and II

gave nearly identical results, while the weaker Mix III was significantly less

stiff. Note that the development of stiffness occurs at a faster rate than

strength, and the stiffness reaches its limiting value sooner than strength.

The results of the splitting tensile tests were used to determine whether the

tensile strength characteristic of concrete made with the job-site materials

was similar to typical normal weight concrete. Figure 4.9 shows the indirect

tensile strength (average of three specimens) as a function of the compressive

strength of companion specimens tested at the same age. Also shown is the

function

f

t = (f

c )0 * 73 (*-D

which is known to be a good representation of the relation between splitting

tensile strength (f t ) and compressive strength (

f

c

)

.

The data indicate that

concrete made with the materials from the job site would have tensile strength

properties that are typical of normal weight concrete.

In summary, the three mixes prepared and tested by NBS are in agreement with

the strength of cores from the structure, and it may be concluded that they

are representative of concrete in the structure. As a result, it became

unnecessary to make additional mixes to correspond with the results of the

chemical analyses of the cores.

The next section describes the development of the strength-maturity functions

for the laboratory mixes, which were used in making in place strength

predictions at the time of collapse.

4.4.3 Estimated Strength of Concrete at Time of the Collapse

In order to make the best possible estimate of the strength of concrete in the

structure at the time of collapse, the maturity concept was employed [4.2].

The maturity concept attempts to represent the combined effects of temperature

and time on the strength development of concrete by means of a single quantity

47
called "maturity", which is calculated from the temperature- time history of

the concrete.

The premise of the maturity concept is that strength of concrete is a singlevalued function of maturity, independent of the actual temperature- time history. Thus, a given concrete possesses a unique strength-maturity relation;

and once this relation is established under any particular (but measured)

curing history, it can be used to predict strength due to any other curing

condition. For this investigation, the strength-maturity functions of the

laboratory mixes were developed and used to predict the inplace strength.

Reference [4.3] provides additional information on the nature of the strengthmaturity function of concrete.

As discussed in section 4.4.1, thermocouples were installed in two 4 x 8 in

(102 x 203 mm) cylinders of each laboratory mix. Temperatures were recorded

automatically every hour, from the filling of the molds until the last cylinder

was tested. The average temperature of the two cylinders was used to represent

the thermal history of each mix. Maturity was calculated for each test age,

and the resulting compressive strength versus maturity data are shown in

figure 4.10. For each mix, least squares curve fitting was used to determine

the curves shown in figure 4.10. Each curve is an hyperbola with the general

equation

Su K (M-Mq )

1 + K (M-M0 ) (4.2)

where, S = compressive strength for maturity M

Su = limiting strength as maturity approaches infinity

M0 = maturity when strength development is assumed to begin

K = slope of curve at Mq divided by Su

The results of the least-squares fit gave 6125, 4870 and 3870 psi (42.2, 33.6,

and 26.7 MPa) as the best fit values of the limiting strengths for Mixes I,

II, and III, respectively. The strengths for each mix were divided by their

corresponding limiting strengths, and the normalized strengths were plotted as function of maturity as shown in figure 4.11. The equation of the best-fit

function was found to be as follows:

S = 0.0209 (M-11.4) f4.3

Su 1 + (M-11.4)

where maturity is expressed in °F-days. Equation (4.3) was used to estimate

the fraction of ultimate strength in various portions of the structure at the

time of the collapse.

In order to use eq. (4.3), the maturity of the concrete in the structure had

to be established from the inplace temperature history. Since such a history

was not available, an approximation was used. It was assumed that the thermal

history of the structure was equal to the air-temperature history obtained

from Patrick Air Force Base, located about five miles south of the structure.

48
Maturity at the time of the collapse was calculated for each portion of the

structure and the results are given in table 4.10.

Based on the computed maturities, the corresponding values of concrete strength

as a fraction of ultimate strength were calculated according to eq. (4.3).

These values are also given in table 4.10. Finally, using the NBS laboratory

data and the core strength information, the relative strengths were converted

to absolute values. The limiting strength of Mix III was assumed to be representative of the column concrete. While for the slab concrete, the limiting

strength of Mix I was assumed to be representative of the third floor, and the

limiting strength of Mix II was used for the remaining floor slabs. The last

three columns in table 4.10, give the estimated strengths, rounded off to the

nearest 100 psi (0.7 MPa), for the various components of the structure at the

time of the collapse.

4.4.4 Reinforcing Steel Strength

Samples of reinforcing steel were obtained from the site and had the identification markings of Grade 60 bars. One bar of each size (#4 to #11) was tested in

uniaxial tension and the measured yield strengths varied from 61 to 67.5 ksi

(420 to 465 MPa). Thus, tests confirmed that Grade 60 steel was used in the

structure as required by the contract documents.

4.5 SUMMARY

Based on the information supplied by OSHA and the results of the NBS laboratory

study, the following conclusions were drawn relative to the construction

materials used on the project:

1) There was insufficient documentation of the quantities of ingredients

used in each batch of concrete.

2) The compressive strength of cores taken from slab sections satisfied

the ACI Code strength criteria for a specified strength of 4000 psi

(27.6 MPa)

.

3) The compressive strength of cores taken from column sections were

lower than the strength of slab cores, and the average strength of the cores from the fifth floor columns did not satisfy the ACI Code

criteria for the specified strength.

4) Core strength values had high variability.

5) Petrographic analyses of core samples revealed that the volumetric

proportions of paste and aggregate were in agreement with the stated

mix designs. The proportions, however, did show high variability.

6) The computed cement contents based on the chemical analysis of core samples were lower than stated in the mix designs. Since the low

values are inconsistent with the core strengths, it was concluded

that the calculated cement content values were imprecise.

49
7) The strength development curves of the three mixes prepared by NBS were consistent with the average value of the core strengths.

8) The tensile strength characteristics of the NBS mixes were typical of

normal weight concrete.

9) The reinforcing steel was Grade 60.

In summary, there are no indications that a gross error was committed in

concrete proportioning that would have resulted in seriously deficient concrete.

The high variability of the concrete, as indicated by core strengths and other

data, is significant and its importance is discussed in chapter 6.

50
Table 4.1. Mix Designs (S.S.D. Basis)

Quantities Per Cubic Yard

Ingredient UNIVEL* DIXIE*

Cement* (lbs) 526 545

Sand (lbs) 1178 1230

Coarse

Aggregate (lbs) 1772 1600

Water (lbs) __ 250

WR/SR (oz) 5.3 28

AEA (oz) 1 .6 4

* Type I Univel = Onsite batch plant

Dixie = Ready-mix plant

1 lb = 0.4536 kg

1 oz = 2.957 x 10“5 m3

51
Table 4.2. Summary of Testing Laboratory Results

Sampling Date (hour) Supplier Floor (N/S)Slump

(in) Test Age(days)

Strength

(psi)

Plasticizer

(gal)

2/13 - (800) Univel* 1 (S) 10 7 5260 (1)* * 28 5660+

2/13 - (930) Dixie* 1 (S) 9 1/2 7 4500 (1) —

28 5650

2/19 - (930) Univel TTn) 10 5 3940 (1) —

28 5380

2/19 - (900) Dixie l 00 8 5 4060 (1) —

28 5660+

2/26 - (900) Univel 2 (S) 9 2 3680 (1) 8

7 5310 (1) 28 6010+

2/26 - (815) Dixie 2 (S) 7 1/2 2 2020 (1) 3 1/2

7 4240 (1)

28 5860

2/27 - (920) Univel 2 (N) 4 3 4560 (1) —

7 5250 (1) 28 5850

2/27 - (945) Dixie 2 (N) 9 3 3250 (1) —

7 4360 (1) 28 4780

3/3 - (830) Univel 3 (S) 3 3/4 2 4300 —

-

7 5060

28 6260

3/3 - (840) Dixie 3 (S) 9 l/4 2 3050 —

7 3580

28 5870

3/5 - (1025) Univel 3 (N) 5 l/2 1 4190 (1) —

7 5660+(l)

28 5840+

3/5 - (1000) Dixie 3 (N) • 9 1 1450 (1)

7 3760 (1) 28 5200

3/10 - (1035) Univel 4 (S) 9 l/2 2 2940 (1) 5

7 4040 (1) 28 5050

3/12 - (900) Univel 4 (N) 8 l/2 4 4000 (1) —

7 4780 (1) 28 5360 \ 3/12 - (845) Dixie 4 (N) 8 1/2 4 3470 (1) —

7 4260 (1) 28 5290

3/17 - (900) Univel 5 (S) 9 1/2 2 2790 (1) —

7 3650 (1) 28 4710

3/17 - (945) Dixie 5 (S) 10 2 2640 (1)

7 4260 (1) 28 5510

3/19 - (850) Univel 5 (N) 8 1/2 2 2310 (1) —

7 3750 (1) 28 4740

3/19 - (840) Dixie 5 (N) 8 2 1490 (1) —

7 3180 (1) 28 4320

3/27 - (230) Dixie R 6 4 3360 (1) Hydrocel*

7 4390

28 5050

3/27 - (215) Univel R 8 4 2940 (1) Hydrocel

7 3710

28 4600

* (1) Indicates one cylinder tested — Indicates quantity not given in the report

Hydrocel is a waterproofing admixture

Univel is the onsite batch plant

Dixie is the ready-mix plant

1 in = 25.4 mm

100 psi «* 0.69 MPa

1 gal - 3.785 x 10-3 m3

52
Table 4.3. Location and Number of Cores Tested

Number of Cores

Section

Number Location

Comp.

Strength

Petr/ Chem

Analysis

c0 1C 4th-5th floor 1 2

L 2C 4th-5th floor 2 2 u 3C 5th floor 2 1 M 4C 4th-5th floor 4 2 N 5C 4th-5th floor 2 2

S 6C 3rd-4th floor 8 2

1 5th floor, N.W. Tower, E. side 1 2 5th floor, N.W. Tower, @ 4th Col. 7 1 3 4th floor, N.W. of Tower, E. side

between 2 and 3

3

4 4th floor N.W. 2 1 s 5 3rd floor N.W. of Tower, E. side 3 1 L 6 3rd floor 2 A 7 2nd floor East side. North end 3 1 B 8 2nd floor N.E. side of Tower 4

S 9 4th floor 1 10 3rd floor N.W. 1 11 4th floor North

12 3rd floor South 6 1 13 4th floor South 5 1 14 5th floor South 6 1

53
Table 4.4. Petrographic Analysis Average Results

Columns Slabs

ASTM C856 (Comments)

Paste

color buff (4 gray - 7 buff) buff (2 gray - 6 buff)

texture medium medium

uniform ok variable

hydration ok ok

Voids

entrained slight- some slight

entrapped mostly mostly

water gain some instances few instances

Re tempering some instances some instances

Cracks none none Aggregate

fine ok ok

coarse ok ok

Other - - ASTM C457 (percent by volume]

Paste 27 (1.9) 28 (2.5)

Fine Aggregate 27 (2.3) 27 (1.9)

Coarse Aggregate 41 (2.9) 41 (3.6)

Air Voids 4.2 (1.1) 3.2 (1.6)

Spacing Factor, in. 0.025 (0.006) 0.024 (0.007)

ASTM C642 (after immersion

and boiling)

Absorption, % 9.1 (0.6) 8.6 (0.7)

Specific Gravity, dry 2.02 (0.02) 2.07 (0.04)

Specific Gravity, SSD 2.20 (0.02) 2.25 (0.03)

Voids, permeable, % 18.4 (1.0) 17.8 (1.2)

1 in = 25.4 mm

54
Table 4.5. Average of Approximate Material Proportions in Concrete

Column Slab

Cement* (lbs/yd3

) 483 (29)+ 442 (41)

Fine Aggregate (lbs/yd3

) 1208 (103) 1213 (83)

Coarse Aggregate (lbs/yd3

) 1750 (123) 1743 (149)

Water (lbs/ yd3

) 315 (38) 320 (42)

Unit Weight, SSD,

lbs/ft3** 137.5 (1.4) 139.6 (1.4)

Air Voids, % 4.2 (1.1) 3.2 (1.6)

Water Cement Ratio 0.65 (0.11) 0.72 (0.15)

Notes

+ Standard deviation.

* Cement content determined from ASTM C 85, Modified

ASTM C85 and dehydration to allow for calcareous

coarse aggregate (limerock and calcareous material

(shell) present in fine aggregate.

** Unit weight SSD estimate by using bulk SSD specific

gravity (ASTM C 642) after immersion and boiling.

1 lb/yd3 = 0.5933 kg/m3

1 lb/ft 3 = 16.02 kg/m3

55
Table 4.6. Results of Admixture Analysis

Core I.D.

Location

Section (Floor)

Admixture

(% concrete wt.) WR/SR HRWR ACC.

P01 1C (5) 0.02 X*

P02 2C (5) 0.01 X

c P03 3C (5) 0.02 X

0 P04 4C (5) 0.01 X X

L P05 5C (5) 0.06 X

u P06 1C (4) 0.05

M P07 6C (4) 0.08 X

N P08 2C (4) 0.06 X

S P09 5C (4) 0.09 X

P10 4C (4) 0.02 X

Pll 6C (3) 0.02 X X

P12 1 (5) 0.02 X X

S P13 2 (5) 0.10 X

L P14 14 (5) 0.08 X

A P15 4 (4) 0.08 X

B P16 13 (4) 0.09 X X

S P17 5 (3) 0.01 X X

P18 12 (3) 0.01 X X

P19 7 (2) 0.02 X

Mix Design:

On-site 0.01 0.05 0

Ready Mix 0.06 0.05 0

Manufacture r+ 0.05-

0.09

0.11-

0.27

0.12-

0.37

* An "X" indicates presence of this admixture.

+ Quantities based on manufacturer's recommended dosages and assuming

530 lbs of cement per cubic yard (315 kg per m^) of concrete.
Table 4.7. Summary of Laboratory Mixes

Mix I Mix II Mix III

Cement (lbs) 93.6 86.5 86.5

Water (lbs) 40.6 48.4 53.1

Sand (lbs) 216 198 198

C. Aggregate (lbs) 278 291 284

Moisture Content (%)

Sand 1.5 1.1 1.7

C. Aggregate 2.1 -1.3 -1 .7 Slump (in)

Before HRWR 4 1/2 4 1/2 4 3/4

After HRWR 7 7 1/2 7 3/4

Air Content (%) 3.8 2.9 4.9

Unit Weight (pcf) 140.2 142.9 136.0

1 lb = 0.4536 kg

1 in = 25.4 mm

1 pcf = 16.02 kg/m^

Table 4.8. Calculated Volume Percentages of Ingredients in Laboratory

Mixes Compared with Measured Values from Cores

Mix I Mix II

Measured

Slab* Mix III

Measured

Column*

Paste 28.5 27.1 28 (2.5) 27.9 27 (1.9)

Sand 28.7 26.9 27 (1.9) 26.1 27 (2.3)

C. Aggregate 39.1 43.1 41 (3.6) 41 .1 41 (2.9)

Air 3 .8 2.9 3.2 (1.6) 4.9 4.2 (1.1)

* From petrographic analysis of cores from structure. Quantities in parentheses are standard deviations of average values.

57
Table 4.9. Comparison of Compressive Strength of

4 x 8 in and 6 x 12 in Cylinders

Average

Compressive Strength*

Mix

Age

(days) 4 x 8 in 6 x 12 (in)

I 7 5060 (68) 5125 (69)

28 5815 (42) 6010 (97)

II 7 4185 (113) 4115 (144)

28 5350 (124) 5490 (221)

III 7 3270 (80) 3250 (54)

28 4345 (112) 4250 (84)

* Average of three specimens, number

in parentheses are standard

deviations

.

1 in = 25.4 mm

1000 psi =6.9 MPa

58
Table 4.10 Estimated Maturities and Strengths at Time of Collapse

Floor Portion

Approx

.

Age (Days)

Calculated

Maturity

(F-Days)

Predicted

S/Su

Strength (psi)

Column

Su = 3900

Slab

Su - 4900 Su * 6100

C-N* 7 148 0.74 2900

5 S-N* 8 163 0.76 3700

C-S 9 192 0.79 3100

S-S 10 213 0.81 4000

C-N 14 305 0.86 3400

4 S-N 15 329 0.87 4300

C-S 16 351 0.88 3400

S-S 17 375 0.88 4300

C-N 21 467 0.90 3500

3 S-N 22 492 0.91 4500 5600

C-S 23 520 0.91 3500

S-S 24 544 0.92 4500 5600

C-N 27 621 0.93 3600

2 S-N 28 643 0.93 4600

C-S 28 643 0.93 3600

S-S 29 662 0.93 4600

C-N 35 791 0.94 3700

1 S-N 36 815 0.94 4600

C-S 39 895 0.95 3700

S-S « 969 0.95 4700

* C-N denotes the columns in the north-half while S-N denotes the slab in the north-half;

C-S and S-S are the columns and slab in the south half.

F-days = 0.555 C-days

1000 psi * 6.9 MPa

59
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drilled

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64

Figure

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of

laboratory

mixes
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66

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of

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laboratory

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MODULUS

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67

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elasticity

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70

MATURITY

(F-DAYS)
5. STRUCTURAL ANALYSIS

5.1 INTRODUCTION

This chapter presents the analytical investigation of the Harbour Cay structure

under conditions existing at the time of the collapse. Section 5.2 provides

background information with the intent to acquaint the reader with the complex

force transfer mechanisms in shored building construction. The definition of

loads and the bases used in their calculations are the topics discussed in

section 5.3. The possible effect of the reshoring sequence, including the

sequence of removal of flying forms from a story, is examined in section 5.4.

The basis for selecting the method of analysis is discussed in section 5.5.

The results of the analysis are presented and explained in section 5.6. A

summary of major findings in this chapter is presented in section 5.7.

5.2 BACKGROUND

In multistory flat plate building construction, it is common practice to

support the freshly cast slab by a shoring system which distributes the load to

lower floors. The number of stories requiring shores* and reshores* depends

primarily on the magnitude of imposed construction loads, curing characteristics of the concrete, and the rate of construction. The convention ns +nr is

sometimes used to designate the number of shored and reshored stories, respectively. Shoring combinations in which 1 ns _< 3 and 0 nr 5 (other than

n s+nr= l) have been used in the past.

Shoring systems commonly utilize metal frames and/or single posts. In systems

without reshores (nr=0) , a newly placed slab remains shored until the n s _i

floors above it have been cast. In mixed systems, an intermediate step occurs when the story under a particular floor is reshored. Depending on the method

used, the forces on the reshores may vary from almost zero to levels comparable

to those existing in the shores they replace.

Regardless of the shoring scheme used, as construction progresses upward, the

building is subjected to unidentical cycles of loading and unloading which produce residual forces in the structure. At a certain stage of construction,

when a new floor is cast, the forces introduced combine with the residual forces

in the structure to produce the highest load on a particular slab, which is the

maximum load any slab will experience during the construction of the building.

The magnitude of this load does depend on the shoring scheme (ns+nr ) used, as well as on the stiffness properties of the partially matured concrete structure

and, where nr >0, on the amount of precompression in the reshores when they are first installed. It should be pointed out, however, that insofar as the most

critical condition is governed by available capacity as well as loads, it may

not necessarily occur in the slab which develops the peak load.

* Shores are vertical members which support formwork and freshly cast concrete.

Reshores are vertical members used to support new structural members after

shores and formwork are removed.

71
In the case of the Harbour Cay building, flying forms were used in lieu of

conventional shores (frames and/or posts) in one story (ns =l). At the time of

the collapse, three stories were reshored with wood posts. This identifies

the shoring scheme as ns+nr=l+3. A flying form is an integrated system which

incorporates shores and formwork into a large unit capable of forming several

bays at a time. The method of replacing flying forms with reshores (described

in section 2.3) results in unloading a tier of bays at a time (2 or 3 bays in

this case), allowing the floor above that area to deflect under its own weight,

before reshores are installed. Consequently, the forces on the reshores may be

minimal; that is, they are installed snug enough to keep them from falling down.

The supposition that reshores were not subjected to appreciable precompression

is corroborated by interview statements (section 3.4) of several instances of

workers restoring fallen reshores.

5.3 LOADS

The following list identifies possible sources of loading on the structure at

the time of the collapse.

(1) The self weight of the structure

(2) The weight of fresh concrete in the roof slab

(3) The weight of the shoring system

(4) The weight of construction workers, equipment and materials

(5) Wind

(6) Impact

Impact loads due to the placement of concrete could be significant only if

concrete were being discharged from the bucket at the time of the collapse. It

has been established that concrete was not being discharged onto the forms at

the time of the collapse (chapter 2) and no other materials were being delivered to the roof. The prevailing wind was estimated at 9 mph (4 m/s) from the

northwest direction based on weather reports from nearby Partick Air Force

Base. This corresponds to wind pressures which are less than one percent of

the design wind pressures stipulated by the applicable Code [2]. The ratio of

actual-to-design wind forces on the structure may be even smaller due to the

absence of the upper story walls and roof parapet. Thus, compared to gravity

loads, the effect of wind on the structure is not significant.

Figure 2.13 shows the approximate location of the workers who were in the

building at the time of the collapse. Very few workers were in the stories

below the roof and they appeared to be dispersed. There was a concentration of

13 workers near the edge of the completed portion of the roof. If they are all

assumed to be in bay HJ (fig. 2.1), the load in that bay due to their combined

weight (assuming 200 lb or 90 kg per worker) would be slightly over 4 psf

(190 Pa). No evidence was discovered of any heavy equipment (such as compressors) in the building. Miscellaneous other equipment in the building, such as hand-operated concrete finishing tools, vibrators, etc., are light and too few

to make a significant contribution to existing gravity loads.

Loads attributed to the self-weight of the structure were calculated as a uniform load of 105 psf (5030 Pa) for the second through fifth floor slabs, of

72
which 100 psf (4790 Pa) represents the weight of concrete in the 8-in (203-mm)

slab based on a unit weight of 150 pcf (2400 kg/m^) for the concrete. The

additional 5 psf (240 Pa) load accounts for the weight of reinforcing bars in

the slab, determined in accordance with the reinforcing plans shown in the

structural drawings.

Loads attributed to the cast portion of the roof slab (between column lines A

to H, figure 2.1) were likewise calculated as 105 psf on the same basis as above. Loads attributed to the weight of flying forms were estimated as a

5 psf (240 Pa) uniform load based on information supplied by the manufacturer.

For the purpose of analysis, the flying form loads were lumped with the roof

loads so that a total load of 110 psf (5270 Pa) was assumed acting on the flying form at the roof level where concrete was in place. A load of 10 psf

(480 Pa), representing the combined weight of reinforcement and flying forms, was applied to the portion of the roof which was void of concrete (the area between column lines H and K, figure 2.1).

The weight of the masonry walls on the periphery of the structure plus stacked

masonry pallets on the north bay of the fourth floor (bay JK, fig. 2.1) and the

south bay of the third floor (fig. 2.12 and bay AB, fig. 2.1) were included in

the analysis of the structure. The walls were in place in the first, second

and most of the third stories (see figure 2.1 for the location of masonry walls

in a completed story). The wall loads for the second and third stories were estimated at 0.3 kips per lineal ft (4.4 kN/m) based on 35 lb (15.9 kg) unit

weight for the masonry. The pallet loads were estimated at about 2 kips (9 kN)

per end bay (south or north). Both wall loads and pallet loads were converted

to equivalent concentrated loads and were placed at the appropriate nodes in

the finite element model of the structure.

Based on the foregoing comments, only the first three sources of loading

defined above were considered in the analysis of the structure. It should be

noted that according to independent laboratory investigations subsequent to

the collapse (see table A. 7), the unit weight of concrete was closer to 140 pcf

(2240 kg/m^) than the 150-pcf (2400-kg/m^) assumed in load calculations.

However, this difference is compensated by other load effects not considered

in the analysis, such as items (4) and (5) cited above, the additional weight

of concrete placed in the slab where ponding occurred (see section 3.3), and

the weight of reshores in the second through fifth stories.

Figure 5.1 shows a partial drawing of flying form types A and B within the two

bays at the north end of the building to assist in vizualizing the layout.

Note that in the case of form type A, instead of a continuous beam, two shorter

beams are used at each location that overlap at the middle truss.

Referring to figures 2.7 and 2.9, it is noted that roof loads are transmitted

to the fifth floor as concentrated loads through the flying form reaction

points. The widths of the flying forms were 25.0, 26.0, 24.5 and 21.0 ft

(7.63, 7.93, 7.47, 6.41 m) for types A to D, respectively (the widths of forms

A and B are shown in fig. 5.1). Since the widths of the transverse bays were

also different, the filler panels (figures 2.9 and 5.1) had, in general, different widths as well. The lengths of the Aluma beams for types A, C and D

73
are the same as the corresponding widths of the forms. The Aluma beams for

types B and C are therefore continuous over the center truss. No continuity

occurs in the case of type D which has only two trusses.

Each flying form was assumed to support the entire load within its bay.

Therefore, the portion of loads on the filler strips between the edge of the

form and column centerline was assumed as a uniform line load applied to the

top of the exterior truss of the form. The Aluma beam reactions at the trusses were calculated taking into consideration the effect of continuity where such

was the case. These reactions were also applied at the top of the trusses as

a uniform line load. The truss reactions were then determined on the basis of

tributary spans for simplicity. In the case of form B, the tributary reactions

were compared with those obtained by analysis using the sectional and material

properties of the trusses supplied by the manufacturer and assuming continous

top and bottom chords and pinned verticals and diagonals. This comparison

showed that the differences between corresponding reactions were within 12

percent. In the analysis of the structure, forces opposite to the tributary

reactions were applied to the fifth floor as point loads.

5.4 CONSTRUCTION SEQUENCE

The forming and reshoring sequence at various stages of construction was examined in the analysis of the Harbour Cay building because it affects the

load paths in the system. An analytical approach to evaluating the effect of

loading history on the interaction between slabs and falsework was first formulated by Neilsen [5.1], taking into consideration the deformation characteristics of the slabs and the shores. Grundy and Kabaila [5.2] developed a simplified method for analyzing the distribution of loads between slabs and falsework

based on the assumption of rigid shores. For a single bay structure, without

reshores (nr = 0, sec. 5.1) this approach yielded a maximum load ratio of 2.36

(i.e., the maximum load on a slab divided by its dead load). This was close to

the more rigorous analysis by Nielsen which yielded a maximum load ratio of

2.56.

Both of these analyses assume a shored system (no reshores). To reduce the

load ratios, Taylor [5.3] proposed a scheme which involved physically relieving

reshore loads prior to the placement of a new concrete floor. This reduced the

maximum load ratio to 1.44. The method of construction in the Harbour Cay

building (sections 2.4 and 5.2) allows the slab to deflect under its own weight

in the reshoring operation. This results in a reduction of the maximum load

ratio in a manner similar to that described by Taylor.

Vertical Construction Sequence

To examine the effect of the construction history on the Harbour Cay building, an analysis was performed according to the method proposed by Grundy and Kabaila

[5.1] assuming no loads are taken by the reshores when they replace the flying

forms. The results of this analysis are schematically summarized in figure

5.2. The numbers on the right side of each sketch are load ratios (defined

above). In this analysis, reshores are assumed to be weightless and axially

rigid and the stiffness of the slabs to be equal. A step-by-step explanation

74
(steps a through n) of the shoring sequence and corresponding load ratios,

(designated by r) follows. In the case of reshores, r designates the load on reshores in a story divided by the dead load of the slab.

(a) Second floor placed. Flying form ("form," in brief) takes entire

load.

(b) Form removed. First floor assumes self weight (r=l).

(c) Reshores are installed snug in first story (r=0). First floor still

carries its own weight (r=l).

(d) Flying forms are placed in the second story and the third floor is cast.

Flying form carries third floor's weight (r=l) which in turn is transferred directly to the first story reshores (r=l). The first floor slab

still carries its own weight (r=l).

(e) Flying form removed causing third floor to carry its own weight (r=l).

First story reshores are relieved (r=0).

(f) Reshores are installed in a snug fit.

(g) to (j) is the same as the cycle from (d) to (f).

(k) Reshores in the first story are removed. (For the Harbour Cay building,

it was assumed that the reshores in the first story were removed before

the flying forms in the fourth story were flown since the reshores would

be needed on hand to replace the flying forms as they were removed.) The

load relieved from the vacated reshores (r=l) is equally shared by floors

2 to 4 (r=l/3 each). The load in the second story reshores are relieved

by two-thirds (r=-2/3) since one-third of the vacated reshore load is

carried by the second floor slab. The third story reshores are unloaded

by one-third (r=-l/3) since an additional one-third of the vacated reshore

load is carried by the third floor slab.

(l) Flying forms in the fourth story are removed. The fifth floor assumes

its own self weight (r=l). Floors 2 to 4 are unloaded equally (r=-l/3

each). The loads in the reshores in the third story are relieved by

two-thirds (r=-2/3) since one-third of the vacated flying form load is

taken by the fourth floor slab. The loads in the reshores in the second

story are relieved by one-third (r=-l/3) since another one-third of the

vacated flying form load is taken by the third floor slab.

(m) The result of (1), i.e., the reshores are unloaded (r=0).

(n) Finally the reshores are installed snug in the fourth story.

To summarize, this analysis shows that prior to placement of the roof, all

the floor slabs are supporting their own weight and the reshores are unloaded;

i.e., the reshores are not involved in transmitting the slab weights of the

second through fourth floors.

75
The values of the load ratios would be somewhat different if actual reshore and

slab stiffnesses were considered in the analysis.

It should be kept in mind that this is a simplified analysis that highlights the

response of a single bay plane structure to a particular shoring scheme and

sequence of construction. In a multibay space structure there are other con- struction effects that need to be examined. This topic is discussed below.

Effect of Lateral Form Removal

Because the Harbour Cay building was a multibay structure, the possibility

exists that the sequence of form removal from a particular story will have an

effect on load distribution. Specifically, the question to resolve is whether

or not the replacement of flying forms with reshores in a multibay story

causes the entire slab above to carry its own weight.

As explained in chapter 2, the sequence of form removal was from the south

side bays to the north side bays. All indications are that the forms were

removed sequentially, beginning with bay AB and ending with bay JK. With

this sequence, an analysis of the effect of the lateral removal of the flying

forms is depicted in figure 5.3. The steps of the analysis are explained

below.

Figure 5.3 (a) shows a cast slab being supported by the flying forms. In

(b) the flying forms are removed from the first bay. This loads the frame by an amount equal to the floor dead weight in that slab (2.35 k/ft or 34.3 kN/m).

A frame analysis of the structure in (b) shows that the flying form in the

adjacent bay BC is unloaded by 13.5 kips (60.1 kN) , 6.5 kips (29.0 kN), and

0 kips as shown in (c). The effect of unloading the first bay causes negligible

forces in all bays but the adjacent bay.

Next, reshores are inserted in a snug fit in the first bay and the flying

forms in the second bay are removed. This loads the frame by the slab weight

of that bay. However, that load is reduced by the amount caused by the removal

of the flying forms from the first bay, i.e., the 13.5 kips (60.1 kN), 6.5 kips

(29.0 kN), and 0 kips. This net loading of the frame results in an unloading

of the flying forms in the adjacent bays by 9.3 kips (41.4 kN), 4.4 kips

(19.6 kN), and 0 kips. Again the effect of removing the flying forms induces

negligible forces in all bays but the adjacent bays. Since in (c) the reshores

in the first bay were simply inserted in a snug fit, i.e. zero load, a reduction

in the load in these reshores caused by removing the flying form in the second

bay would cause them to go into tension.

Since reshores cannot develop tension, it is reasonable to assume that they were installed just tight enough to keep them from falling down when forms in

adjacent bays are lowered. Thus, in (d), with the removal of the flying form

in the third bay, the load in the reshores in the first bay would still be zero. The removal of this flying form again loads the frame by the slab dead

weight of that bay less the amount caused by the removal of the flying form in

the second bay, i.e., 9.3 kips (41.4 kN), 4.4 kips (19.6 kN), and 0 kips.

76
This process is repeated until the north bay is reshored, at which time, the

forces in the reshores are all zero or nearly so, and the slab is supporting

its own dead weight as assumed in the analysis shown in figure 5.1.

Definition of Analysis Models

The above analyses indicate that when the flying form system is used, the

reshores are effective in transferring only the newly placed concrete loads to

the lower floors. When the flying forms are removed, the forces in the reshores

revert to zero and the slabs carry their own dead weight. As a result, it is

reasonable to assume that before the roof is cast the fifth, fourth, third,

and second floor slabs are supporting their own weight and the reshores are unloaded (as indicated in fig. 5.2, step n) . When the roof is cast, the load

will be shared by all the floor slabs through the action of the reshores. Thus,

to determine the state of stress at the time of collapse, two separate cases should be analyzed. In the first case, the structure without reshores is

loaded by the dead weight of floors 2 through 5 as shown in figure 5.4a. In

the second case, the reshored structure is loaded with the dead weight of the

roof as shown in figure 5.4b. The final result is obtained by adding the

results from the two analyses.

5.5 METHOD OF ANALYSIS

The design of a flat plate structure such as the Harbour Cay Condominium

typically involves either a simplified "direct design" analysis in which the

structure is designed for assumed upper and lower bounds of bending and shear

stresses based on tributary areas, or, the structure is designed to resist

moments and shears determined by analysis using the equivalent frame method

(EFM). The equivalent frame method involves modeling the structure as a series

of plane frames along the column lines, where the slabs between the centerlines

of adjacent spans are modeled as wide beams. As specified in the ACI Code

[2.1], the actual column stiffnesses in the EFM are modified after Corley and

Jirsa [5.4] to better represent the transfer of bending moments at the

slab-column connections (table 5.1).

The EFM is a simple method to use. However, since the Harbour Cay condominium

was reshored irregularly in addition to having irregularly spaced flying form

loads, the EFM analysis was compared to results of a finite element method

(FEM) of analysis which can model the three-dimensional aspect of the loading.

Equivalent Frame Analysis of Fifth Floor

The comparison was made using only the fifth story of the structure and

assuming the columns were fixed at the base and the reshores pinned at both

ends, as shown in figure 5.5a (see section 2.4 for the number and location of

reshores). The EFM analysis summarized in figure 5.5 was the frame along

column line 2. The loading including the distributed dead weight of the fifth

floor slab (106 psf or 5.08 kPa) plus the weight of the flying forms and the

newly cast roof up to column line H (figure 2.1), applied as joint loads. The

roof loads were calculated as explained in section 5.3. The pertinent structural data for the EFM is summarized in table 5.1. The results of the analysis

77
as summarized in figure 5.5(a) are for the frame with reshores and in figure

5.5(b) for the frame without reshores. These results were then compared to a FEM analysis of the fifth floor alone using the same loading conditions as above.

Table 5.1. Structural Data for Fifth Floor only EFM

A

reshores = 302.25) = 36.75 in2 (237 cm

2

) Ereshores

= 1.600,000 psl

(= 11 GPa)

bs iab = 266 in. (6.76 m)

’slab “ H350 in Interior col * 566 i"4

lexterior col “ ^4 *n

E c = 4,280,000 psi (29.5 GPa)

WD = (0.106) (22 .2

' ) = 2.35 klf (34,300 N/m)

Finite Element Analysis of Fifth Floor

The finite element software used in the analysis was the Sperry Univac

implementation of ICES-STRUDL-II [5.5]. Rectangular plate bending elements

(type BPR) were used. This element is described in reference 5.6. The finite

element mesh used for the fifth floor is shown in figure 5.6. The rationale

behind the fairly dense mesh can be seen upon a careful examination of a plan

view of the fifth floor (figure 5.7). Nodal points were necessary at the locations of the flying form truss jacks and at the locations of the reshores underneath the slab. Nodal points were added between the flying form trusses to

reduce the maximum aspect ratio of the plate bending elements from 5.0 to 2.5.

The FEM model of the fifth floor consisted of the plate bending mesh shown in

figure 5.6 supported by columns fixed at their base. As before, two different

analyses were made: one for the structure with reshores and one for the structure without reshores.

The results for the column forces for the FEM fifth-floor model with reshores

are shown in figure 5.8. Due to almost symmetric loading and closeness of span

lengths, the column moments are all relatively small. A comparison of axial

forces in columns on column line 2 with those obtained from the EFM analysis

(fig. 5.5a) shows poor agreement between the two sets of results.

The results for the column forces for the FEM model without the reshores are shown in figure 5.9. The largest column force is 146 kips (650 kN) occurring

in column B-2. This value compares well with the value of 144 kips (641 kN)

for the EFM results summarized in figure 5.5(b). The other column forces also

agree reasonably well.

The results for column axial forces from the EFM and FEM analyses agree fairly

well for the case without reshores since the axial force is primarily due to

tributary area loads. With the addition of reshores, the load transmission

78
becomes three-dimensional and the EFM column axial forces do not match the

FEM results even reasonably well; for this reason, a three-dimensional FEM

analysis is required.

5.6 FINITE ELEMENT ANALYSIS OF ENTIRE STRUCTURE

Due to the absence of flying forms in the second through fourth floors a coarse FEM mesh, as shown in figure 5.10, could be used for modeling these

levels. The fifth floor mesh (figure 5.6) was retained for this analysis.

The columns were modeled as three-dimensional beam-column elements and the

reshores were modeled as three-dimensional truss elements.

The structural geometry is defined in chapter 2. The values for the modulus

of elasticity of concrete, E c , (as described in chapter 4) are summarized in

table 5.2. Poisson's ratio was assumed as 0.2. The modulus of elasticity of

the reshores was taken as 1,600 ksi (11 GPa).

Table 5-2. Assumed Concrete Modulus of Elasticity

Ec

5th Floor Slab 4,150 ksi (28.6 GPa)

4th Story Columns 3,590 ksi (24.7 GPa)

4th Floor Slab 4,270 ksi (29.4 GPa)

3rd Story Columns 3,630 ksi (25.0 GPa)

3rd Floor Slab 4,310 ksi (29.7 GPa)

2nd Story Columns 3,650 ksi (25.2 GPa)

2nd Floor Slab 4,340 ksi (29.9 GPa)

1st Story Columns 3,660 ksi (25.7 GPa)

As described in section 5.4, two cases are analyzed. In the first case, the

structure without reshores is loaded with the dead weight of the floors, as shown in figure 5.4(a). In the other case, the structure with reshores in the

second, third, and fourth stories, as shown in figure 5.4 (b), is loaded with

the flying forms and the roof slab, which was cast from column line A to column

line H. Superposition of the results of these two analyses yield the final

results which are discussed below.

The FEM analysis of a large structure such as this (5328 members, 2543 nodes,

12679 degrees of freedom) causes problems as to the most efficient manner of

presenting the results. Since it was suspected that failure was due to punching

shear, shear forces at the columns and top column moments are presented. These

quantities are used to calculate punching shear stresses in the slab.

Although the typical failure mode of a flat plate structure of this type is by

punching shear, in the interest of completeness, various failure modes were checked prior to the complete analysis of the structure.

Bending stresses were checked according to the ACI Code [2.1] direct design

method. The steel ratios in the slabs were also checked and found to be less

79
than the balanced ratio as required. Column slenderness effects were also

investigated; however due to the nature of the collapse, i.e., vertical with

no sidesway component, this mode of failure was not suspected. Axial stresses

in the columns were checked for tributary area loads and found to be well below

allowable. Initial estimates of punching shear stresses (from tributary areas)

showed the degree of safety to be marginal for this mode of failure.

Figure 5.11 shows the combined results of punching shear forces (difference

between column axial forces at the slab-column connections) for the slab-column

connections along column line B. Figure 5.12 shows the same results for column

line 2. These figures indicate that the fifth floor was the most heavily

loaded floor (in punching shear) as suspected from the load path analysis in

section 5.4. The punching shear and unbalanced column moments for all the

fifth floor slab column connections are tabulated in figure 5.13. The most

heavily loaded connections are on the fifth floor at column B-2 and G-2; 102

kips and 100 kips (454 kN and 445 kN), respectively. The analysis of punching

shear stresses corresponding to the axial forces and moments shown in figure

5.13 is presented in the next chapter.

5.7 SUMMARY

The following are the major points resulting from the analysis of the structure

under construction loads;

1) At the time of collapse, gravity loads were the only significant kinds

of loading on the structure.

2) Using a simplified and approximate analysis of the load distribution

during construction, it is shown that at the stage just before the roof

loads are applied each slab carries its own weight and there are no loads in the reshores. When the roof is cast, the reshores are effective in distributing the roof load to the slabs of the reshored stories.

3) The final analysis of the construction load effects was obtained by

superposition of the results of two analyses. In one case, the unshored

structure was loaded by the self weight of floor slabs 2 through 5.

In the second case, the shored structure was loaded with the roof load.

4) In order to account for the three-dimensional response of the shored

structure, the finite element method was used for the final analysis.

5) The results of the analysis indicate that the most heavily loaded

slab-column connections are in the fifth floor at columns B-2 and G-2.

80
//^ / ^//

^

/~^^7

ASSUMED SUPPORT CONDITIONS OF ALUMA BEAMS

Figure 5.1 Configurations of flying forms A and B in north bays

81
rvs, i

a) b)

SHORED STORY (TYP)

r-t

flp

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ft 0

1

1

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Figure 5.2 Effect of vertical construction sequence

82
83

Figure

5.3

Lateral

sequence

of

flying

form

removal
1 v j

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Figure 5.4 Two cases analyzed to obtain final solution

84
85

Figure

5.5

Results

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FEM

analysis

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fifth

floor

only
86

Figure

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87

Figure

5.7

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88

Figure

5.8

Axial

loads

and

bending

moments

in

columns

from

FEM

analysis

of

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floor

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(with

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Figure

5.9

Axial

loads

and

bending

moments

in

column

from

FEM

analysis

of

fifth

floor

only

(without

reshores)
90

Figure

5.10

FEM

mesh

for

second

through

fourth

floors
91
CN

92

Figure

5.12

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shear

forces

along

column

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93

Figure

5.13

Fifth

floor

punching

shear

and

column

moments
6. INTERPRETATION OF ANALYSIS

6.1 INTRODUCTION

This chapter evaluates the results of the analysis of the structure and draws

conclusions therefrom on the most probable cause of the collapse. First, in

section 6.2 shear stresses at the critical sections are computed and compared

with the capacity of the structural elements. Second, in section 6.3 the

thickness of the slab is examined to see whether the slab thickness of 8 in

(203 mm) was adequate to support the design loads and if not, to what extent

the design inadequacy contributed to the collapse. In section 6.4 the cause

of the collapse is established based on this analysis, and the most probable

mode of failure and the reason for the occurrence of progressive collapse are

given.

6.2 ANALYSIS OF SHEAR FAILURE OF SLAB

It has been shown in a number of studies [6.1] that flat plate structures are inherently weak in shear at the slab-column connection and, when punching

shear failure occurs, the structure is highly susceptible to progressive

collapse unless it is adequately designed and carefully constructed [6.2, 6.3].

Judging from the manner in which the Harbour Cay structure failed and in the

absence of any evidence of ductile failure, it was suspected that the structure

had experienced shear failure at slab-column connections.

To verify that such a failure had occurred, punching shear stresses resulting

from direct shear forces and unbalanced moments were computed at slab-column

connections using the shear and moment values for construction loads given in

figure 5.13. These shear stress values were compared with the shear capacity

of the slab based on the predicted inplace strength of concrete at the time of

collapse. Because the punching shear stress values are maximum at the fifth

floor level (see figures 5.12 and 5.13), the shear stresses are evaluated only

for this level. The guidelines given in the ACI Code [2.1] are followed in

computing the punching shear stress. Figure 6.1(a) shows the critical section

in the slab around an interior column. The critical section where shear

stresses are evaluated is taken perpendicular to the plane of the slab and a distance a half of the effective depth away from the periphery of the column.

The assumed distribution of shear stresses at the critical section is shown in

figure 6.1(b). The critical sections for edge and corner columns and corresponding assumed shear stress distributions at these critical sections are shown

in figure 6.2(a) and (b) and figure 6.3(a) and (b), respectively.

The results of the punching shear stress calculations are listed in figure 6.4.

In this figure, the numbers given in parentheses are the shear stresses resulting from the direct shear only and the other numbers are the shear stresses

resulting from the combined effect of the direct shear and the unbalanced

biaxial bending moments. The negative values shown at the edge and corner

columns are upward shear stresses resulting from the combined effect. Negative

stresses occur on a side of the critical section when the absolute value of the

moment-induced shear stress exceeds the absolute value of the direct shear

stress

.

94
It is seen from figure 6.4 that maximum shear stresses resulting from direct

shear occur at columns B-2 and G-2 with shear stress values of 250 psi

(1.72 MPa) and 244 psi (1.70 MPa), respectively. Except for the edge columns

in line A and corner column B-4, maximum combined shear stresses occur also

at columns B-2 and G-2. It needs to be pointed out that the large shear

stresses at the edge and corner columns may not develop as a result of moment

redistribution to the interior columns due to cracking of the slab in the

vicinity of exterior columns caused by high moments [6.4]. If this redistribution of moments takes place, the moment-induced stress would increase at

the columns in line B and J and would decrease at the edge columns in line A

and K. Because combined shear stresses at these edge and corner columns

would be limited by cracking in the slab and because shear stresses due to

vertical loads are relatively small, it is less likely that punching shear

failure would develop at these column locations, and they are not considered

as the critical locations in the failure analysis.

For the purpose of design, the nominal shear strength specified by the ACI Code

[2.1] is 2 / f Q for beam-type shear and 4 / f£ for punching shear where

is the compressive strength of concrete.

For the south half of the fifth floor, the beam-type shear strength given by

the ACI Code would be 126 psi (0.87 MPa) while the punching shear strength

would be 253 psi (1.74 MPa) for an estimated compressive strength ( f £) of

4000 psi (27.6 MPa) at the time of collapse. The corresponding strengths for

the nor^h half would be 121 psi (0.83 MPa) and 243 psi (1.66 MPa) for an estimated f c of 3700 psi (25.5 MPa). It should be recalled that the south half

of the slab was two days older than the north half.

The computed beam-type shear stress under construction loading is 51 psi

(352 kPa) and 49 psi (338 kPa) , respectively, at the critical sections corre- sponding to columns B-2 and G-2. These values are considerably less than the

beam shear strength of the slab. Therefore, it is concluded that beam-type

shear failure did not precipitate the collapse.

When the ACI Code punching shear strengths are compared with the computed

shear stresses, it appears that the punching shear stresses at many columns in

the fifth floor exceeded the nominal punching shear capacity of the slab. This

would suggest that failure should have occurred prior to the placement of concrete to column line H. To examine why failure did not occur earlier requires

a review of the ACI provision of 4 / f ^. F igure 6.5 shows experimental punching shear strengths, normalized by bd / f^, versus c/d where b is the perimeter of critical section, c is the side length of column and d is the effective

depth of the slab [6.1]. It is clearly seen in this figure that the punching

shear strength is a probabilistic parameter which exhibits a high degree of

variability. It is evident that 4 / f

'c as used in the ACI Code is a lower

bound strength, that would be justified for the purpose of design. For failure

analysis, however, t he actual shear strength should be used, and it would be

greater than 4 / f £ . According to the data shown in figure 6.5, the mean

punching shear strength could be as high as 7 / f£ and the probability of a punching shear strength of 4 / f£ occurring at any given location in the

structure is relatively low. However, the probability of failure initiating at

95
some location in the entire structure increases rapidly with the number of

locations where the shear stresses are in the neighborhood of 4 / f^,.

According to figure 6.4, there are at least 30 such locations in the fifth

floor slab.

If, for the sake of illustrat ion, the probability of failure at any given

location where a stress of 4 f £ occurs is assumed to be two percent (P^ =

.02), which is quite plausible in light of the data shown in figure 6.5, the

probability of failure, ?2 , of an element in a system consisting of 30 such

locations would be (assuming indepdence)

P

2 = 1 - (1-P

1

) 30 = 0.45,

or 45 percent. If a value of 0.03 (or 3 percent) is assumed for P]_, P2 would

increase to 0.60 (or 60 percent), and so on. This means that even though the

probability of failure at each column location is relatively low, the probability

of the entire structure having a failure is relatively high.

The likelihood of failure having initiated at B2 or G2, where the shear stresses

are maximum, will be somewhat greater than at other locations (except those

excluded along column line A for reasons stated above). In fact, in terms of

the pa ramet er / f

'Ci the shear stresses at B2 and G2 are virtually equal

(4.57 / f£ and 4.52 /" f respectively); an indication that failure could

have occurred at either location. The supposition that failure initiated at

column G2 rather than B2 is borne out by witness statements (as discussed in

sec. 3.3) indicating that the collapse started somewhere in the region where

concrete was being cast. However, from a probabilistic view, this is not an absolute certainty (failure could have been initiated at location other than

B2 and G2).

It should be pointed out that the data shown in figure 6.5 was obtained from

laboratory experiments where, in general, much better quality control can be

exercised than in the field. Therefore, it is reasonable to expect that if

actual field conditions were simulated, the data would have shown greater

scatter and a lower expected value than indicated in figure 6.5. Both factors

imply a greater probability of strength below 4 / f There is evidence,

as discussed in chapter 4, that the concrete used in the structure had high

variability in terms of composition and strength.

6.3 INFLUENCE OF EFFECTIVE DEPTH OF SLAB ON SHEAR CAPACITY

Without special reinforcement at the slab-column connection, the effective

depth of slab at the critical section must be sufficient to resist punching

shear failure around the column. Both the thickness of slab and the location

of tension reinforcement govern the effective depth. It was discussed in

section 2.3 that the 4-1/4 inch (108 mm) individual high chairs used to support

the top reinforcing bars in the column strips produced more than the 3/4-in

(19 mm) minimum coverage required by the structural drawing. This resulted in an effective depth of the slab of 5.3 in (135 mm). Had higher size individual

high chairs been used, to just meet the 3/4-in coverage requirements, the slab

could have had a 6.3 in (160 mm) effective depth.

96
The increase in effective depth by 1 inch (25.4 mm) increases the critical

cross sectional area to resist the punching shear by 25 percent, which, in

turn, decreases the shear stress acting on the critical section.

Because under construction loads the computed shear stresses in the slab at

many of the column locations exceeded the shear strength of 4 / f

'c , the slab

thickness of 8 inch (203 mm) was checked to see whether it was adequately

designed to resist punching shear stresses due to the design loads. The

required thickness for the slab was computed following the equivalent frame

method described in the ACI Code [2.1] in accordance with the following three

load combinations in the code:

(1) U = 1.4 D + 1.7 L

(2) U = 0.75(1.4 D + 1.7 L + 1.7 W) , and

(3) U = 0.9 D + 1.3 W

where U is the required strength to resist factored loads,

D is dead load,

L is live load , and

W is wind load.

The strength reduction factor of 0.85 is applied to 4 / f£ in computing the

design shear strength.

It was mentioned in section 2.3 that the original design calculations were

examined and that no computation was noted for checking the punching shear

capacity of the slab. For the dead load of the required thickness slab, the same live loads used in the design (see section 2.3) and the wind load specified in the Southern Standard Building Code [2.2], the required slab thickness

should have been 11 in (279 mm) instead of 8 in (203 mm) for the column size

of 10 x 18 in (254 mm x 457 mm) . The slab thickness is governed by the load

combination of (1.4 D + 1.7 L) . This clearly indicates that the design of

the slab was inadequate to support the design loads with the margin of safety

implied by the ACI Code. Table 6.1 lists the punching shear stresses at column

G-2 for the 11-in slab with a 3/4-in (19 mm) cover and with a 1 3/4-in cover

(44 mm) for the same construction condition which existed in the structure.

In addition, the punching shear stresses at column G-2 for the 8-in slab with

a 3/4-in cover and with a 1 3/4-in cover are also listed.

It is seen in table 6.1 that for the 11-in slab with the tension reinforcement

placed such that it has a 3/4- in cover (d = 9.3 in), the punching shear stress

under the construction loads is 174 psi (1.19 MPa) which is considerably less

than 4 / f£ = 243 psi (1.66 MPa). With 1 3/4-in cover for the tension reinforcement (d = 8.3 in), the punching shear stress would still be lower than

the nominal strength, 205 psi vs. 243 psi. It is also clear that had proper

size high chairs been used for the 8-in slab such that the tension reinforcement

had a 3/4 in cover (d = 6.3), the punching shear stress would have been less

than the nominal strength, 218 psi vs. 243 psi. Only when the design error (which resulted in an 8- in slab) and the use of improper size chairs (which

resulted in 1 3/4-in cover) occur simultaneously, the punching shear stress

exceeds the nominal strength, 275 psi vs. 243 psi at column G-2. Thus,

97
Table 6.1 Punching Shear Stresses at Column G-2 for 11 and

8 in Slab Under Actual Construction Loads

8-in Slab 11-in Slab

d = 6.3 in (3/4 in cover)

v = 218 psi

v/(4 / f£) = 0.90

d = 9.3 in (3/4 in cover)

v = 174 psi

v/(4 /~f£) = 0.72

d = 5.3 in (1 3/4 in cover)

v = 275 psi

v/(4 /"fj) = 1.13

d = 8.3 in (1 3/4 in cover)

v = 205 psi

v/(4 /~fj) = 0.84

1 in = 25.4 mm

100 psi = 0.69 MPa

the lack of consideration of punching shear in the design of the slab resulted

in a reduction of the margin of safety such that the slab had little reserve strength to accommodate construction errors. Based on this analysis, it is

clear that both the design error and the use of chairs having insufficient

height contributed to the collapse.

6.4 PROBABLE MODE OF FAILURE

The analysis of the structure has shown that failure of the fifth floor was most likely initiated at column G-2. This conclusion was corroborated by the

witness accounts. The loss of support at column G-2 was due to punching

shear failure of the slab rather than failure of the column below the

slab. The computed strength of the column at the time of failure was more

than three times the stress produced by the loading.

When the slab failed, the portion of loads originally resisted in shear at the

four sides of column G-2 was transferred to four adjacent columns (columns F-2

,

G-l , G-3 , and H-2). This transfer action resulted in a substantial increase

in the total shear in the slab-column connection at these four columns. Likewise, moments transferred to these four columns would also have increased the

moment- induced shear stresses. Because shear stresses in the slab-column

connection at these columns were already near the limiting shear strength of

the slab, this transfer action would have precipitated punching shear failure

at these four locations, which in turn would have triggered failures at other

column locations. The propagation of punching shear failure continued until

it swept through the entire fifth floor slab.

Once the fifth floor had failed and landed on the fourth floor, punching

shear failure of that slab was inevitable. The total weight of the roof and

the fifth floor slab was more than enough to produce shear stresses greater

98
than the shear strength of the fourth floor slab. The third and second floor

slabs failed by the same process. Thus, collapse of the entire structure was initiated by punching shear failure of the slab at the fifth floor level.

It is concluded that the most probable cause of the failure was the lack of

adequate punching shear capacity in the fifth-floor slab at several of the

columns to resist the applied construction loads. This deficiency was brought

about by the design inadequacy which resulted in the slab thickness of 8 in

instead of 11-in and by the placement of the top reinforcement at the column

strips with significantly greater cover than that which would have met the cover requirement specified by the structural drawings. This resulted in an insufficient effective depth of the fifth floor slab to resist punching shear

under construction loads.

99
(a) (b)

Figure 6.1' Shear stress distribution at interior column

Figure 6.2 Shear stress distribution at edge column

(a) (b)

Figure 6.3 Shear stress distribution at corner column

100
4000

psi f

'c=

3700

psi

253

psi 4

/Pc=

243

psi

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60

•H

pH

O CM 0 cn LO ^ UO CO MCM CO un co O cn co 00 CM CM CM CM CM CM CM CM —

1 • a a B ^ 

^ x NO O *3- CTiOuf *3- r—

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O CM uo *— co un 0 uo

O 00 CO 1— CO r— r— CO 1— .— CO 1— 1— CM CM '— 1 — 1 —- 1 w 1 a a m

6) ' (S) (^) (Z>

101
16

14

a

<9

10

VTEST

bd/F7

C = column width

d = effective slab depth

h = perimeter of critical section

08 m

* gfsj 7. ^

< fO

Q> §

o

for illustration purposes only

i . i 10

c/d

Figure 6.5 Results of punching shear tests of slab-column

connections (ref. 6.1)

102
7 SUMMARY AND CONCLUSION

As a result of the collapse of a five story reinforced concrete building during

construction at Cocoa Beach, Florida, the National Bureau of Standards (NBS) was requested by the Occupational Safety and Health Administration (OSHA) to

carry out an in-depth study aiming at the determination of the most probable

cause of the collapse. In response to this request, NBS has carried out com- prehensive field and laboratory investigations, including tests of concrete

and reinforcing steel. NBS also performed analytical evaluation of the

response of the structure under the loading condition which existed at the

time of the collapse. The findings based on the results of the NBS study

are presented in this report, together with other information such as structural drawings and calculations and OSHA case records. The conclusions drawn

from this study are summarized in the following statements:

(1) The structure as built had a number of deviations from standard practice

of construction.

(a) A number of columns examined disclosed that the inclined portion of

offset bars above the floor level were bent more than 1 in 6 as required by the ACI Code. Furthermore, the bars were not placed

parallel to the axis of the column above the offset bend.

(b) The placement of a large number of large-size reinforcing bars

(12 - #8 through 16 - #11) for lap splices in 10 in x 18 in (254 mm x 457 mm) columns created a congestion of bars which violated the ACI

Code requirement for clear spacing. Examination of columns showed

that such a congestion prevented flow of concrete into the space

between reinforcing bars. This lack of concrete prevented the

development of adequate bond between reinforcing bars.

(c) Both the reinforcing steel shop drawings and field measurements

revealed that 4-1/4 in (108 mm) individual high chairs were used to

support the two-way top layer reinforcing steel for the slabs at the

column strips. This resulted in a net concrete cover ranging from

1-5/8 in (41 mm) to 2-1/8 in (54 mm) with an average effective depth

of the slab of 5.3 in (135 mm). This was about 1 in (25 mm) less

than the effective depth corresponding to the 3/4 in (9.5 mm) cover specified in the structural drawings. Excessive concrete cover

reduced the effective depth of the slab, thereby reducing the punching

shear capacity of the slab.

(2) The results of tests on core samples and laboratory prepared specimens

were used to predict the strength of concrete at the time of failure.

(a) The results of tests on cores obtained from the floor slabs indicated

that the compressive strength of the slab concrete met the specification value. That is, the compressive strength of cores exceeded 85

percent of the specified value of 4000 psi (27.6 MPa). However, the

average compressive strength of cores obtained from the fifth-story

columns fell below the 85 percent requirement. The strengths of the

103
cores obtained from the third- and fourth-story columns meet this

requirement

.

(b) There are no indications that a gross error was committed in concrete

proportioning. However, the petrographic analysis of core samples

revealed high variability in volumetric proportions of paste and

aggregate. High variability was also noted in core strengths.

(c) The result of the NBS laboratory tests and the results of core tests

agreed well. Based on the laboratory tests, the estimated compressive

strength of concrete at the time of collapse was 3700 psi (25.5 MPa)

and 4000 psi (27.6 MPa), respectively, for the north- and south-halves

of the fifth floor slab, and 2900 psi (20.0 MPa) and 3100 psi

(21.4 MPa), respectively, for the north- and south-halves of the

fifth-story columns.

(3) The analysis of the structure showed there are two major contributing

factors which caused the collapse.

(a) The use of lower individual chairs for supporting the top layer

reinforcing bars in the column strips than required to provide the

3/4 in (19 mm) cover called for in the structural drawings reduced

the effective depth and punching shear capacity of the slab.

(b) Omission in the slab design of a check for punching shear resulted in an 8-in (203-mm) slab whih was shallower than needed to safely support

the design loads for the completed structure. This reduced the margin

of safety such that the slab had insufficient reserve strength to

accommodate the use of 4-1/4-in (108-mm) individual high chairs within

the column strips.

(4) The analysis showed that failure of the slab most likely initiated at

Column G-2. The analysis also showed that the shear stresses at many

column locations on the fifth floor exceeded the nominal shear strength.

Once punching shear failure had initiated at column G-2, it propagated

throughout the slab causing the total collapse of the fifth floor, which,

in turn, caused the successive collapse of the lower floor slabs.

104
8 ACKNOWLEDGMENTS

The authors acknowledge the contributions made by the following persons:

Dr. E. 0. Pfrang provided technical guidance and constructive criticism;

Dr. E. V. Leyendecker and Dr. W. C. Stone critically reviewed the report;

and the staff of the CBT Word Processing Center typed the manuscript.

105
9. REFERENCES

2.1 ACI Standard 318-77, Building Code Requirements for Reinforced Concrete,

American Concrete Institute, Detroit, Michigan, 1977.

2.2 Standard Building Code, Southern Building Code Congress International,

Birmingham, Alabama, 1979.

3.1 Malhotra, V. M., Superplasticizers: Their Effect on Fresh and Hardened

Concrete, Concrete International, V3, No. 5, pp. 66-81, May 1981.

4.1 Annual ASTM Standards , Part 14, American Society for Testing and

Materials, Philadelphia, PA, 1981.

4.2 Plowman, J. M., Maturity and Strength of Concrete, Magazine of Concrete

Research 8_, No. 22, pp. 13-22, March 1956.

4.3 Carino, N. J., Temperature Effects on the Strength-Maturity Relation of

Mortar, National Bureau of Standards (U.S.), NBSIR 81-224, 98 pp, March

1981.

5.1 Nielson, K. E. C., Loads on Reinforced Concrete Floor Slabs and their

Deflections During Construction, Bulletin No. 15, Swedish Cement and

Concrete Research Institute, Royal Institute of Technology, Stockholm,

1952.

5.2 Grundy, P., and Kabaila, A., Construction Loads on Slabs with Shored

Formwork in Multistory Buildings, ACI Proceedings, V. 60 , No. 12,

pp. 1729-1738, December 1963.

5.3 Taylor, P. J., Effect of Formwork Stripping Time on Deflections of Flat

Slabs and Plates, Australian Civil Engineering and Construction (Melbourne), V. 8 , No. 2., pp. 31-35, February 1967.

5.4 Corley, W. G. and Jirsa, J., Equivalent Frame Analysis for Slab Design,

J. American Concrete Institute, V. 67, No. 11, pp. 875-885, November

1970.

5.5 ICES STRUDL-II Engineering User’s Manual, MIT, Report R68-91, November

1968.

5.6 Przemieniecki, J. J., Theory of Matrix Structural Analysis , McGrawHill, N.Y., 1968.

6.1 ACI-ASCE Committee 426, The Shear Strength of Reinforced Concrete

Members - Slabs, Journal of the Structural Division, ASCE, V. 100, August

1974.

6.2 Leyendecker, E. V. and Fattal, S. G., Investigation of the Skyline Plaza

Collapse in Fairfax County, Virginia, Nat. Bur. Stand. (U.S.), Bldg. Sci.

Ser. 94, June 1973.

106
6.3 Litle, W. A., The Boston Collapse at 2000 Commonwealth Avenue, Boston,

Mass., Symposium on Progressive Collapse, ACI Annual Convention, Boston,

April 1975.

6.4 Neth, U. W., DePaiva, H. A. R., and Long, A. E., Behavior of Models of

Reinforced Concrete Flat Plate Edge-Column Connection, J. American

Concrete Institute, V. 78

,

No. 4, pp. 269-275, July-August 1981.

107

APPENDIX A

RESULTS OF TESTS ON SAMPLES FROM

THE SITE

A-l
TABLE

A.

1

AGGREGATE

TEST

RESULTS

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A2
TABLE A.

2

CEMENT ANALYSIS RESULTS

SAMPLE

PHYSICAL DATA:

Time of setting, Gillmore

Initial, hrrmin

Final, hr:min

Autoclave Expansion, Z

Fineness Blaine, m2/kg

Compressive Strength, psi

1 day

3 days

7 days

28 days

Air Content, Z vol

CHEMICAL ANALYSIS : Silicon Dioxide, X

Aluminum Oxide, X

Ferric Oxide, X

Calcium Oxide, X

Magnesium Oxide, X

Sulfur Trioxide, X

Alkalies, X

Loss on Ignition, X

Insoluble Residue, X

Tricalcium Silicate, C3S, X

Dicalcium Silicate, C2S, X

Tricalcium Aluminate, C3A, X

Tetracalcium Aluminoferrite, C4AF,

AOl A12 ASTM

UNIVEL DIXIE C150,TYPE

3:00 1:42 1:00 min.

4:46 2:54 10:00 max

+0.04 +0.06 0.80 max. 400 392 • 280 min.

1490 1640 —

3030 2810 1800 min.

3890 3870 2800 min.

5460 5370 3500 min.

7.7 6.5 12.0 max.

20.1 19.2 - 3.7 4.5 - 3.0 3.3 - 60.0 61.3 - 3.9 0.7 6.0 max. 3.0 2.9 3.0 max.

0.64 0.42 0.60 max.*

0.90 1.43 3.0 max. 0.42 0.04 0.75 max. 56 60 - 15 10 - 4.7 6.5 - 9 10 -

For low alkali cement
TABLE A.

3

WATER TEST RESULTS

SAMPLE AO A AOS A15

MIX

UNIVEL

POND

DIXIE

MIX

pH, electrometric 7.2

ANALYSIS

7.6 7.9

Alkalinity as ppm CaC03

Phenolphthalein 0.0 0.0 0.0

Total 70 204 46

Total Hardness as ppm

CaC03 60 612 74

Calcium, ppm 20.1 142.0 16.0

Magnesium, ppm 2.

A

62.7 8.3

Sodium, ppm 80 370 75

Potassium, ppm 5 25 4

Manganese, ppm * * * Phosphorus, ppm 0.1 0.1 0.1

Iron, ppm A.O 0.4 0.6

Carbonate, ppm 0.0 0.0 0.0

Bicarbonate, ppm 85.4 248.9 56.1

Chloride, ppm 300 2080 320

Sulfate, ppm 60.0 82.8 58.0

Conductivity, micromhos

per cm 580 2400 520

*Less than 0.05 ppm

A-4
TABLE A.

A

ADMIXTURE TEST RESULTS

SAMPLE AO6 AO 7 AO 8 AO 9

UNIVEL

DAREX 79 19 TYPE 3

Specific Gravity

@ 75 C 1.016 1.229 1.175

Total Solids, % 5.9 A1.8 34.

5

1.157

31.1

A16 A17

DIXIE

DAREX 79

1.013 1.182

34.5 5.7 39.7
Table A. 5. Core Compressive Strength Test Results

r

i Section

No. Core No. T

1 Dia.

(in) Length (in) Drilled I Test

Age at Test

(days)

Compressive

Strength (psi)

r Comments

1 3rd Floor I 6C 2 T 2.95 10 5 7/8 34 36 4360 | |

1 r 6C 3 T 2.95 10 5 7/8 34 36 4190 | |

1 r 6C 4 1 2.95 10 5 1/2 41 43 4650 | | c 1 i 6C 5 1 2.95 6 1/4* 5 3/4 41 43 3810 | |

1 r 6C 6 T 2.95 10 5 7/8 55 57 4460 | | 0 1 i 6C 7 T 2.95 10 5 7/8 55 - 57 4560 | |

1 4th Floor | 6C 4 1 3.95 10 7 7/8 27 30 3270 | | L 1 1 4C 7 T 3.95 10 7 7/8 27 30 3390 steel embedment at 2 loca1 1 1 tlons, 1 ” in from each end

U 1 1 4C 8 1 3.95 10 7 7/8 27 - 30 4170 | |

1 r 5C 2 1 3.95 10 7 7/8 34 37 3860 | | M 1 i 6C 5 T 3.95 10 7 7/8 34 37 3630 | |

1 r 4C 3 T 3.95 10 7 7/8 48 51 3750 | | N 1 i 2C 6 T 3.95 10 7 7/8 48 51 4230 | |

1 5th Floor | 1C 1 T 3.95 10 7 7/8 20 22 2900 | | S 1 1 3C 7 T 3.95 10 7 7/8 20 22 3080 | |

1 f 2C 3 1 3.95 10 7 7/8 27 - 29 3620

1 1 4C 5 1 3.95 10 7 7/8 27 “ 29 3190

1 1 3C 4 ! 3.95 7 3/4* 5 7/8 41 - 43 3040

1 1 5C 6 1 3.95 10 7 7/8 41 43 3610

1 2nd Floor I 7 2A (top) T 2.18 8 1/2 4 1/2 41 4870

1 (north) | 2B ( bottom) I 2.18 8 1/2 4 1/4 41 5050 wire mesh embedment

1 1 8 6 1 2.95 8 5 7/8 41 4930 crack in tested core with initial

1 1 1 fracture at crack location

1 1 7 3 1 2.18 7 3/4* 4 1/4 48 5320 crack in one face, extending 3"

1 1 1 into core

1 1 8 4 1 2.95 5 1/4* 4 1/4 48 5420 core has thin spall (approx. 2

"

1 1 1 length) on one end; tested with

1 1 1 this shaved section

1 1 8 7 1 2.95 6 l/2* 5 l/2 62 4710

1 1 8 8 1 2.18 5 1/2* 4 1/4 62 5130

1 3rd Floor I 5 3 1 2.95 6 1/2* 5 5/8 35 5110

1 (north) | 10 7 1 2.95 8 1/4 6 35 5490

1 1 5 2 1 2.95 7 1/2* 5 7/8 42 6090 horizontal crack around 50 percent

1 1 1 of perimeter

1 1 6 5 1 2.95 6 1/2 5 7/8 42 5790

1 1 5 4 1 2.95 8 5 5/8 56 4640

1 f 6 6 1 2.95 6 1/2 5 56 6640

i 3rd Floor I 12 2 T 2.95 8 6 37 6020

1 (south) | 12 6 1 2.95 8 5 7/8 37 4540 | |

1 1 12 4 T 2.95 7 3/4 5 1/2 44 6180

1 1 12 7 I 2.95 8 1/4 5 7/8 44 5490

1 1 12 3 1 2.95 7* 5 5/8 58 5640

1 1 12 5 1 2.95 8 5 7/8 58 5730

S 1 4th Floor | 4 3 1 2.95 6 1/4 3 7/8 28 4100

1 (north) | 3 5 1 2.95 8 5 7/8 28 4270 L 1 1 4 2 ! 2.18 6 1/2 4 1/4 35 4360 crack in one face, extending

1 1 1 4" into core A 1 1 3 4 1 2.95 8 5 7/8 35 4900 horizontal crack around 50

1 1 1 percent of perimeter, did not

B 1 I 1 break at crack

1 1 3 6 1 2.95 6* 5 1/4 49 5200

S 1 ! 9 7 1 2.95 8 5 1/4 49 5500

1 4th Floor | 13 2 1 2.95 8 5 7/8 30 4960

1 (south) | 13 5 1 2.95 8 4 3/4 30 2640 2 intersecting cracks on one face.

1 1 1 3 cracks extending 6" into core

1 1 13 3 T 2.95 5 1/2* 4 3/8 37 5220 ;

1 1 13 1 4 1 2.95 6* 4 51 4900 I 1

1 1 13 6 1 2.95 8 5 l/2 51 4300 horizontal, semicircular crack

1 1

1 around 50 percent of perimeter.

1 1 1 not intersecting faces; crack

1 1 1 within tested section

1 5th Floor | 2 3 T 2.95 8 5 7/8 21 5010

1 (north) | 2 4 T 2.95 5 1/2* 4 5/8 21 4050

1 1 2 2 T 2.95 5 l/2* 4 1/2 28 5190 hairline crack extending length of

1 1 1 core; crack across one face

1 f 2 6A ( top) T 2.18 8 3 1/2 28 6260

1 1 6b ( bottom) 1 2.18 8 4 28 4890

1 1 2 5 1 2.95 8 5 7/8 42 5790

1 1 2 7 T 2.18 6* 3 1/4 42 5150

1 5th Floor | 14 3 1 2.95 8 1/8 5 7/8 23 3690

1 (south) | 14 4 1 2.95 6 1/4* 5 7/8 23 3740

1 f 14 2 1 2.95 8 1/8 5 30 4140 cracks 2 1/2" into core at 2 loca1 1

1 tions after trimming

1 1 14 6 T 2.95 7 7/8* 5 7/8 30 4680 i!

1 1 14 5 T 2.95 8 5 1/2 44 4710 | 1

1 1 14 7 1 2.95 | 8 5 1/4 44 5020 1 1

* Broken core not extending through structure.

1 In “ 25.4 mm 1000 psi « 6.9 MPa

A6
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 5th Floor, Columns

SAMPLE: P01 P02 P03 P04 P05

TEST RESULTS

C856 (Comments)

Paste

color gray gray buff gray gray

texture medium medium medium medium medium

uniform no ok ok ok ok

hydration ok ok ok ok ok

Voids

entrained slight slight slight slight slight

entrapped mostly mostly mostly mostly mostly

water gain slight slight slight slight slight

Re tempering some some none slight slight

Cracks

Aggregate

none none none none none

fine ok ok ok ok ok coarse ok ok ok ok ok

Other wood cement - - -

piece ball

C457 (percent by volume)

Paste 26 24 29 28 29

Fine Aggregate 27 29 24 25 30

Coarse Aggregate 44 44 43 44 37

Air Voids 2.6 2.7 3.5 3.8 3.5

Spacing Factor, in.

C642 (after immersion

0.030 0.029 0.026 0.034 0.029

and boiling)

Absorption, Z 8.8 9.3 9.4 9.3 9.5

Specific Gravity, dry 2.05 2.04 2.00 2.03 2.02

Specific Gravity, SSD 2.23 2.23 2.19 2.22 2.21

Voids, permeable, Z

C85 (modified, percent

18.0 19.0 18.8 18.9 19.1

of dehydrated sample) - - - 13.3 13.7
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 4th Floor, Columns

SAMPLE: P06 P07 P08 P09 P10

C856 (Comments)

Paste

color buff

TEST RESULTS

buff buff buff buff

' texture medium medium medium medium course

uniform ok ok ok ok ok

hydration ok ok ok ok ok

Voids

entrained some slight some some some entrapped large mostly mostly mostly mostly

water gain none slight none slight some Retempering slight none slight slight some

Cracks none none none none none

Aggregate

f ine ok ok ok ok ok coarse ok ok ok ok ok

Other

C457 (percent by volume)

Paste 27 25 27 28 31

Fine Aggregate 26 31 27 26 27

Coarse Aggregate 42 40 40 40 36

Air Voids 5.3 3.9 5.6 5.3 5.7

Spacing Factor, in. 0.020 0.022 0.029 0.015 0.018

C642 (after immersion

and boiling)

Absorption, %

Specific Gravity, dry 2.02 1.99 2.05 1.98 1.99

Specific Gravity, SSD 2.20 2.20 2.22 2.16 2.17

Voids, permeable, % 17.7 20.6 17.0 18.1 18.3

C85 (modified, percent

of dehydrated sample) 14.5 15.3 15.2 14.2 14.0

A-8
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 3rd Floor, Columns

SAMPLE: Pll

TEST RESULTS

C856 (Comments)

Paste

color buff

texture medium

uniform ok

hydration ok

Voids

entrained some entrapped mostly

water gain none

Re tempering some Cracks none Aggregate

f ine ok coarse ok

Other -

C457 (percent by volume)

Paste 27

Fine Aggregate 24

Coarse Aggregate 45

Air Voids 4.3

Spacing Factor, in. 0.019

C642 (after immersion

and boiling)

Absorption, Z 8.5

Specific Gravity, dry 2.03

Specific Gravity, SSD 2.20

Voids, permeable, Z 17.2

C85 (modified, percent

of dehydrated sample) —
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 5th Floor, Slab

SAMPLE: P12 P13 P14

C856 (Comments)

Paste

TEST RESULTS

color gray gray buff

texture medium fine coarse

uniform ok ok ok

hydration

Voids

ok ok ok

entrained slight some slight

entrapped mostly mostly mostly

water gain none slight none

Re tempering some some some Cracks

Aggregate

none none none

fine ok ok ok

coarse "dry" ok ok

Other -

C457 (percent by volume)

«

Paste 30 28 27

Fine Aggregate 30 30 26

Coarse Aggregate 36 39 44

Air Voids 4.3 2.6 3.2

Spacing Factor, in.

C642 (after immersion

and boiling)

0.026 0.017 0.023

Absorption, I 9.7 9.5 8.7

Specific Gravity, dry 2.02 2.04 2.08

Specific Gravity, SSD 2.22 2.23 2.26

Voids, permeable, Z

C85 (modified, percent

19.7 19.4 18.2

of dehydrated sample) — 11.0 12.41

A-10
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 4th Floor, Slab

SAMPLE: P15 P16

C856 (Comments)

Paste

TEST RESULTS

color buff buff

texture coarse coarse

uniform ok ok

hydration

Voids

ok ok

entrained some slight

entrapped mostly mostly

water gain some none

Re tempering some some Cracks

Aggregate

none none

fine ok ok

coarse ok ok

Other - -

C457 (percent by volume)

Paste 27 24

Fine Aggregate 28 25

Coarse Aggregate 39 47

Air Voids 6.3 3.3

Spacing Factor, in.

C642 (after immersion

and boiling)

0.012 0.022

Absorption, X 8.3 7.9

Specific Gravity, dry 2.04 2.10

Specific Gravity, SSD 2.21 2.26

Voids, permeable, %

C85 (modified, percent

16.9 16.5

of dehydrated sample) 14.4 12.6

A-ll
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 3rd Floor, Slab

SAMPLE: P17 PI 8

C856 (Comments)

Paste

TEST RESULTS

color buff buff

texture medium medium

uniform no no hydration

Voids

ok ok

entrained slight slight

entrapped mostly mostly

water gain none none

Re tempering slight some Cracks

Aggregate

none none

f ine ok ok coarse ok ok

Other - -

C457 (percent by volume)

Paste 27 31

Fine Aggregate 27 25

Coarse Aggregate 44 43

Air Voids 2*4 1.3

Spacing Factor, in®

C642 (after immersion

and boiling)

0*034 0.027

Absorption, % 7.9 8.1

Specific Gravity, dry 2.12 2.12

Specific Gravity, SSD 2.29 2.29

Voids, permeable, Z

C85 (modified, percent

16.7 17.2

of dehydrated sample) 13.3 12.2

A-12
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: 2nd Floor, Slab

SAMPLE: PI 9

TEST RESULTS

C856 (Comments)

Paste

color buff

texture medium

uniform no

hydration ok

Voids

entrained slight

entrapped mostly

water gain none

Re tempering slight

Cracks none Aggregate

fine ok

coarse ok

Other - C457 (percent by volume)

Paste 31

Fine Aggregate 27

Coarse Aggregate 40

Air Voids 1.8

Spacing Factor, in.

C642 (after immersion

and boiling)

Absorption, Z 8.7

Specific Gravity, dry 2.07

Specific Gravity, SSD 2.23

Voids, permeable. Z 18.0

C85 (modified, percent

of dehydrated sample)

A-13
TABLE A.

6

PETROGRAPHIC ANALYSIS RESULTS

Concrete: Average Results

SAMPLE: COLUMNS SLABS .

C856 (Comments)

Paste

color

TEST RESULTS

buff (4 gray - 7 buff) buff (2 gray - 6 buff)

texture medium medium

uniform ok variable

hydration ok ok

Voids

entrained slight- some slight

entrapped mostly mostly

water gain some instances few instances

Re tempering some instances some instances

Cracks none none Aggregate

fine ok ok

coarse ok ok

Other - -

C457 (percent by volume)

Paste 27 (1.9) 28 (2.5)

Fine Aggregate 27 (2.3) 27 (1.9)

Coarse Aggregate 41 (2.9) 41 (3.6)

Air Voids 4.2 (1.1) 3.2 (1.6)

Spacing Factor, in. 0.025 (0.006) 0.024 (0.007)

C642 (after immersion

and boiling)

Absorption, Z 9.1 (0.6) 8.6 (0.7)

Specific Gravity, dry 2.02 (0.02) 2.07 (0.04)

Specific Gravity, SSD 2.20 (0.02) 2.25 (0.03)

Voids, permeable, Z 18.4 (1.0) 17.8 (1.2)

C85 (modified, percent

of dehydrated sample) 14.3 (0.8) 12.7 (1.2)

A-14
\

TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 5th Floor, Columns

SAMPLE: P01 P02 P03 P04 P05

Material Proportions,

approximate, lbs./yd^

Cement* 454 451

Fine Aggregate 1210 1299 1099 1112 1353

Coarse Aggregate 1866 1874 1814 1844 1582

Water** - - - 327 345

Admixture, X *** 0.02 0.01 0.02 0.01 0.06

WRDA 79 X X X - X

WRDA 19 - - - X - Type III — — X —

Unit Weight, SSD,

lbs. /ft. 3**** 139.2 139.2 136.7 138.5 137.9

Air Voids , X 2.6 2.7 3.5 3.8 3.5

Spacing Factor, in. 0.030 0.029 0.026 0.034 0.029

Absorption, X 8.8 9.3 9.4 9.3 9.5

Permeable Voids, X 18.0 19.0 18.8 18.9 19.1

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

**Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity (C642)

after immersion and boiling.

A-15
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 4th Floor, Columns

SAMPLE: P06 P07 P08 P09 P10

Material Proportions,

approximate, lbs./yd3

Cement* 494 513 526 475 470

Fine Aggregate 1166 1380 1210 1180 1223

Coarse Aggregate 1771 1690 1696 1696 1523

Water** 297 258 287 320 372

Admixture, % *** 0.05 0.08 0.06 0.09 0.02

WRDA 79 - X X X X

WRDA 19 - - - - - Type III — — — — —

Unit Weight, SSD,

lbs. /ft. 3 **** 137.3 137.3 138.5 134. 8 135.4

Air Voids, % 5.3 3.9 5.6 5.3 5.7

Spacing Factor, in. - 0.020 0.022 0.029 0.015 0.018

Absorption, Z 8.7 10.3 8.3 9.1 9.2

Permeable Voids, % 17.7 20.6 17.0 18.1 18.3

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

**Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity- (C642)

after immersion and boiling.

A-16
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 3rd floor. Columns

SAMPLE: Pll

Material Proportions,

approximate, lbs./yd^

Cement* -

Fine Aggregate - 1058

Coarse Aggregate 1890

Water** -

Admixture, Z *** 0.02

WRDA 79

WRDA 19 X

Type III X

Unit Weight, SSD,

lbs. /ft. 3**** 137.3

Air Voids, Z 4.3

Spacing Factor, in., 0.019

Absorption, ' % 8.5

Permeable Voids, Z 17.2

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

** Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity-(C642)

after immersion and boiling.

A-17
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 5th Floor, Slab

SAMPLE: P12 P13 . P14

Material Proportions,

approximate, lbs./yd3

Cement* - 375 427

Fine Aggregate 1323 1326 1148

Coarse Aggregate 1517 1660 1852

Water** — 352 319

Admixture, Z *** 0.02 0.10 0.08

WRDA 79 - X X

WRDA 19 X - -

Type III X — —

Unit Weight, SSD,

lbs. /ft. 3**** 138.5 139.2 141.0

Air Voids, % 4.3 2.6 3.2

Spacing Factor, in. 0.026 0.017 0.023

Absorption, Z 9.7 9.6 8.7

Permeable Voids, Z 19.7 19.4 18.2

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

** Water estimate determined by using percentage of paste volume and

calculated cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity_ (C642)

after immersion and boiling.

A-18
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 4th Floor, Slab

SAMPLE: P15 P16

Material Proportions,

approximate, lbs./yd^

Cement*

Fine Aggregate

Coarse Aggregate

Water**

Admixture, Z *** WRDA 79

WRDA 19

Type III

Unit Weight, SSD,

lbs. /ft. 3****

Air Voids, Z

Spacing Factor, in.

Absorption, Z

Permeable Voids, Z

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

** Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity- (C642)

after immersion and boiling.

494 443

1245 1139

1634 1984

297 263

0.08 0.09

X

X

X

137.9 141.0

6.3 3.3

0.012 0.022

8.3 7.9

16.9 16.5

A-19
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 3rd Floor, Slab

SAMPLE: ' P17 P18

Material Proportions,

approximate, lbs./yd3

Cement* 475 435

Fine Aggregate 1196 1102

Coarse Aggregate 1833 1798

Water** 303 383

Admixture, Z *** 0.01 0.01

WRDA 79 - - WRDA 19 X X

Type III X X

Unit Weight, SSD,

lbs. /ft. 3**** 137.9 141.0

Air Voids, Z 2.4 '1.3

Spacing Factor, In. 0.034 0.027

Absorption, Z 7.9 8.1

Permeable Voids, Z 16.7 17.2

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

** Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity- (C642)

after immersion and boiling.

A-20
TABLE A.

7

SUMMARY OF CONCRETE TESTS

Concrete 2nd Floor, Slab

SAMPLE: P19

Material Proportions,

approximate, lbs. /yd3

Cement* - Fine Aggregate 1223

Coarse Aggregate 1669

Water** -

Admixture, Z *** 0.02

WRDA 79 X

WRDA 19

Type III ~

Unit Weight, SSD,

lbs. /ft. 3 **** 140.4

Air Voids, Z 1.8

Spacing Factor, in. 0.030

Absorption, Z 8.7

Permeable Voids, Z 18.0

Notes:

*Cement content determined from C85, Modified C85 and dehydration to allow

for calcareous coarse aggregate (limerock) and calcareous material (shell)

present in fine aggregate.

** Water estimate determined by using percentage of paste volume and calculated

cement content.

***Admixtures extracted from concrete using carbonate method for

sulfonates. (X) indicates presence of admixture.

****Unit weight SSD estimate by using bulk SSD specific gravity. (C642)

after immersion and boiling.

A-21
NBS-I14A (rev. 2-bc) U.S. DEPT. OF COMM.

4.

BIBLIOGRAPHIC DATA

SHEET (See instructions) TITLE AND SUBTITLE

PUBLICATION OR

REPORT NO.

2. Performing Organ. Report No. 3. Publication Date

81-2374

INVESTIGATION OF CONSTRUCTION FAILURE OF HARBOUR CAY CONDOMINIUM IN COCOA BEACH,

FLORIDA

5. AUTHOR(S)

H. S. Lew, N. J. Carino, S. G. Fattal, and M. E. Batts

5. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City . State, ZIP)

Occupational Safety and Health Administration

U. S. Department of Labor

Washington, D.C. 20001

10. SUPPLEMENTARY NOTES

Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it here)

The investigation of the collapse of a five-story reinforced concrete flat -plate

structure under construction at Cocoa Beach, Florida is presented in this report. The

investigation included onsite inspection, laboratory tests and analytical studies.

Based on the results of this investigation, it is concluded that the most probable

cause of the failure was insufficient punching shear capacity in the fifth-floor slab

to resist the applied construction loads.

Two factors contributed to the low punching shear capacity: one in the design and

the other in the construction of the building. In the design, the omission of a check

for punching shear resulted in a smaller slab thickness than needed to satisfy the

Code, requirements. In construction, the use of specified chairs having insufficient

height to support the top reinforcing steel resulted in more than the cover specified

in the structural drawings. Both factors contributed to reducing the effective depth

of the slab such that it had insufficient strength to resist the construction loads.

The analysis showed that shear stresses in the slab at many column locations on the

fifth floor exceeded the nominal shear strength. Thus, punching shear failure at one

of the columns precipitated a progressive failure of the slab throughout the entire

fifth floor. Collapse of the fifth floor, in turn, caused the successive collapse of

the lower floor slabs. The failure of the fifth floor slab most likely initiated at

column G-2 , an interior column which supported the last bay of freshly placed roof

concrete prior to the collapse.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) Building; collapse; concrete; concrete strength; construction; failure; flat plate;

shear; strength

13. AVAILABILITY 14. NO. OF

Unlimited

1 .for Official Distribution. Do Not Release to NTIS

H Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

PRINTED PAGES

20402. 15. Price

[22 Order From National Technical Information Service (NTIS), Springfield, VA. 22161

USCOMM-DC 6043-P80

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