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High performance concrete (HPC) can be considered as a logical development of cement concrete in which the ingredients are proportioned and selected to contribute efficiency to various properties of cement concrete in fresh as well as hardened states.
High performance concrete (HPC) is defined as a concrete mixture which posses high workability, high strength, high modulus of elasticity, high dimensional stability, low permeability and resistance to chemical attack.
There is a little controversy between the terms “high strength concrete” and “high performance concrete”. High performance concrete is also a high strength concrete but it has a few more attributes specifically designed as mentioned above. It is there for logical to describe by the more widely embracing term “high performance concrete” (HPC).

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31-10-2010, 09:33 PM


High performance concrete(HPC) is that concrete which meet special performance and uniformity requirements that cannot always be achieved by conventional material ,normal mixing, placing and curing practices .

Architects, engineers and constructors all over the world are finding that using HPC allows them to build more durable structures at comparable cost .HPC is being used for building in aggressive environments, marine structures, highway bridges and pavements, nuclear structures, tunnels, precast units

This reports aims to discuss the application of HPC particularly for bridge structures .The use of HPC was found to have added advantages compared with normal concrete in areas of strengths , service life, construction time, economy ,etc.
An experimental study on “Behavior of instrumented prestressed high performance concrete bridge girders” by Hazim M. Dwairi, Mathew c. Wagner, Mervyn J.Kowalsky, Paul Zia is also discussed as case study

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Concrete is considered as durable and strong material. Reinforced concrete is one of the most popular material used for construction around the world. Reinforced concrete is exposed to deterioration in some regions especially in costal regions. There for researchers around the world are directing their efforts towards developing a new material to overcome this problem. Invention of large construction plants and equipments around the world added to the increased use of material .This scenario led to the use of additive materials to improve the quality of concrete. As an out come of the experiments and researches cement based concrete which meets special performance with respect to workability, strength and durability known as” High Performance Concrete” was developed


High performance concrete (HPC) is that which is designed to give optimised performance characteristics for the given set of materials, usage and exposure conditions, consistent with requirement of cost, service life and durability

The American Concrete Institute (ACI) defines HPC ‘‘as concrete which meets special performance and uniformity requirements that cannot always be achieved routinely by using only conventional materials and normal mixing, placing, and curing practices.”

High performance in a broad manner can be related to any property of concrete. It can mean excellent workability in the fresh state like self-levelling concrete or low heat of hydration in case of mass concrete, or very rigid setting and hardening of concrete in case of sprayed concrete or quick repair of roads and airfields, or very low imperviousness of storage vessels, or very low leakage rates of encapsulation containments for contaminating material.

HPC is composed of the same material as normal concrete, but it has been engineered to achieve enhanced durability or strength characteristics, or both, to meet the specified demands of a construction project and implimentation. The main ingredients of high performance concrete are cement, fine aggregate, coarse aggregate, water, mineral admixtures and chemical admixtures

If the structure of normal strength concrete (NSC) is compared with high performance concrete (HPC) one notes several differences: The matrix stiffness of HPC is larger than NSC and approaches the stiffness of the aggregate, the bond strength between matrix and aggregate is higher for HPC, matrix tensile strength is higher, Reduced internal cracking in terms of number of cracks and size of intrinsic cracks before loading. These aspects show that HPC is more elastic and more brittle than NSC.

Figure 1 shows a schematic representation of the stress-strain curve from a uniaxial test along with the simplified crack pattern

HPC has a greater Young’s modulus than NSC and the post-peak softening branch is steeper. High Strength Concrete (HPC) is more homogeneous than normal strength concrete (NSC). Initial flaws like pores, cracks and interfacial delamination in HPC are smaller and less numerous than in NSC. This makes HPC more stiff and elastic as compared to NSC

HPC does not simply mean high strength concrete (HSC), but also includes other enhanced material properties such as early-age strength, increased flow ability, high modulus of elasticity (MOE),low permeability, and resistance to chemical and physical attack (increased durability). HPC is usually high strength concrete (HSC), but HSC may not always be of high performance
Architects, engineers and constructors all over the world are finding that using HPC allows them to build more durable structures at comparable cost .HPC is being used for building in aggressive environments, marine structures, highway bridges and pavements, nuclear structures, tunnels, precast units


High-performance concrete characteristics are developed for particular applications and environments; some of the properties that may be required include:
• High strength
• High early strength
• High modulus of elasticity
• High abrasion resistance
• High durability and long life in severe environments
• Low permeability and diffusion
• Resistance to chemical attack
• High resistance to frost and deicer scaling damage
• Toughness and impact resistance
• Volume stability
• Ease of placement
• Compaction without segregation
• Inhibition of bacterial and mold growth


High-performance concretes are made with carefully selected high-quality ingredients and optimized mixture designs; these are batched, mixed, placed, compacted and cured to the highest industry standards. Typically, such concretes will have a low water-cementing materials ratio of 0.20 to 0.45. Plasticizers are usually used to make these concretes fluid and workable. Table 1 lists materials often used in high-performance concrete and why they are selected.

Table 1. Materials Used in High-Performance Concrete

For many years, high-strength, high-performance concrete has been used in the columns of high-rise buildings. However, in recent years, there has been increased use of high-performance concrete (HPC) in bridges where both strength and durability are important considerations. The primary reasons for selecting HPC are to produce a more economical product, provide a feasible technical solution, or a combination of both

High performance concrete bridges include two key elements: total precast bridge systems that can dramatically improve construction speed and high performance concrete that can improve durability and structural efficiency. In HPC bridges, these improvements are achieved at no cost premium and often at a reduced initial cost.

Designing with HPC components can drastically reduce construction time because various precast components can be combined to allow a truck-to-structure systems approach without waiting for site forming and curing. Full depth precast decks are being used on both new and rehabilitated bridges. The cost for this approach can result in overall savings due to more efficient designs that permit longer spans or fewer girders and/or piers. HPC can be used effectively in virtually all bridge components to aid in minimizing construction and future maintenance. HPC components can include piles and pile caps, piers and column bents, abutments, decks, and rails and barriers. HPC uses the same materials as typical concrete but is engineered to provide higher strength and better durability. These attributes can be varied to align with the design’s needs. They will be affected by environmental and geographic conditions and the specific bridge components (that is, substructure, beams or deck).

Fig 2 Cross section of the pier elevation shows the main compmnents of a bridge system


Overall, the advantages accruing from higher durability and/or additional strength include a variety of benefits:
• Longer service life thanks to higher durability and lower chloride penetration. When needed, bridge life can extend to 100 years or even more.
• Lower maintenance and inspection requirements, especially since the bridge requires no painting or rust protection. This savings grows with the bridge’s longer service life.
• Longer spans, which can reduce costs by eliminating piers or allowing the use of concrete beams instead of steel beams.
• Wider beam spacing, reducing the number and cost of beams.
• Shallower beams due to higher concrete strength.
• Improved mechanical properties such as greater tensile strength.
• Rapid construction due to the ability to factory-cast components while site work is underway and the ability to erect pieces upon delivery. These benefits cut the time necessary for disruptions to local traffic.
• Predictable performance and close tolerances for precast members due to the high quality achieved through PCI certification and casting under controlled conditions in the plant.
In general, HPC components can produce lighter, longer precast pieces and smaller-diameter columns that creep less. This means span lengths can be lengthened and under clearances can be maximized.

Behavior of instrumented high performance concrete bridge girders byHazim M. Dwairi, Matthew C. Wagner, Mervyn J. Kowalsky, Paul Zia
In this study a comprehensive monitoring of the behavior of four prestressed high performance concrete (HPC) bridge girders, with higher compressive strength, during construction and while in-service, is presented. The monitoring program covered instrumentation and monitoring of a series of four girders during the casting operation, after construction, under the effects of traffic and thermal loads, as well as under controlled
load conditions. Information regarding transfer length, prestress loss, heat of hydration, compressive strength, modulus of elasticity (MOE), modulus of rupture (MOR), creep, shrinkage, coefficient of thermal expansion, and chloride permeability of the concrete used is obtained and presented. Furthermore, the in-service monitoring and controlled load tests and details regarding thermal expansion, bridge stiffness, and load distribution factors are also presented. This paper provides details of testing of the concrete properties and field instrumentation of the bridge girders as well as a discussion of service level monitoring and controlled load testing. Comparisons are made between experimental and theoretical results

Fig.3 shows the bridge for three southbound lanes under construction, which forms the basis for the work described in this paper. Figs.4 and5 show the plan of the bridge and a typical cross section, respectively.

Fig. 3. US 401 Southbound Bridge over the Neuse River, Raleigh, NC.

Fig. 4. US 401 Southbound Bridge over the Neuse River plan view.

Fig.5. Typical cross-section of Southbound Bridge.

The objective of the research presented in this paper is to develop an understanding of the behavior of HPC in bridge structures. This objective was achieved through a comprehensive monitoring program including: (1) characterization of the actual HPC utilized in the construction of the bridge, (2) extensive instrumentation and monitoring of a series of four girders in the bridge during the casting operation, and (3) monitoring of the bridge structure after construction under the effects of traffic and thermal loads, as well as controlled load conditions. Information regarding transfer length, prestress loss, heat of hydration, as well as compressive strength, modulus of elasticity (MOE), modulus of rupture (MOR), creep, shrinkage, coefficient of thermal expansion, and chloride permeability of the concrete was obtained and presented. In addition, the in-service monitoring and controlled load tests, details regarding thermal expansion, bridge stiffness, and load distribution factors were obtained.

For the HPC used in the US 401 Bridge, performance criteria were selected in accordance with FHWA definition of HPC as shown in Table 2 Tests were conducted on the material to evaluate its compressive strength, flexural strength, modulus of elasticity (MOE), modulus of rupture (MOR), creep, shrinkage, thermal properties, and chloride permeability. In all cases, the concrete samples were taken from batches of material used in four instrumented bridge girders. Two were AASHTO Type IV girders (designated as A4 and B4) and two were AASHTO Type III girders (designated as C4 and D4) as shown in Fig 4
Table .2

Numerous 102 * 203 mm cylinders, six 76 * 76 *286 mm prisms, and six 152 * 152 *508 mm (6 6 20 in.) prisms were cast for the material testing. The specimens were cured along side with the girders to keep the curing temperatures for the specimens as close as possible to those of the actual girders. One concern in the use of HPC is that standard cylinder tests may not provide an accurate measure of the in-place strength of the concrete. High temperatures generated during hydration can affect the in-place strength of HPC. For this reason, it was desirable to match-cure the cylinders that were to be used for compression tests and the determination of the modulus of elasticity. The match curing was performed with the help of the FHWA Mobile Concrete Laboratory. Four cylinders were matchcured for each girder, for a total of sixteen specimens. The match-cured cylinders were prepared in the same manner as standard non-match-cured cylinders.
The performance requirements of the HPC and the laboratory test results are listed in Table 2. Table 3 shows the mix proportion of the concrete that was used for the girders. The slump and the air content of the concrete are shown in Table 4. Both the air content and the slump met the requirements of the specifications .Table4 summarizes the compressive strength of match cured and non-match-cured cylinders. The average compressive strength of the concrete for both Type III and Type IV girders met specifications as shown in Table 2. It is noted that by using silica fume the
concrete gained strength rapidly at early age and after 56 days there was virtually no increase in strength.
Table. 3



The measured and predicted results for the modulus of elasticity (MOE)are shown in table.6
MOE of concrete calculated by the equations
E_(c=) w^1.5 0.043√(〖f'〗_c ) Mpa (ACI Code) (1)
E_c=3321.4√(〖f'〗_(c ) ) + 6895 Mpa (ACI 363) (2)
By using obtained compression strength and the measured unit weight of concrete,24kN/m3 ,the predicted values of MOE are shown in table.6

The modulus of rupture (MOR) for the HPC was determined according to ASTM C 78–84. It was found to be 1.25 MPa for the Type III girder and 1.11 MPa for the Type IV girder. According to the ACI Code [5], the modulus of rupture may be calculated by
fr=.62√(f'c) Mpa (3)
Using an average value of 72.4 MPa for f’c, Eq. (3) gives a value of 5.28 MPa ,which compares very well with the test results. Note that MOR was not specified as a performance criterion.

The rapid chloride permeability testsaccording to AASHTO T 277 were carried out in order to get permiability value. Standard 102 * 203 mm cylinders were sliced into thirds and the top and middle thirds were used for testing. The results are given in table .7

2.4. Girder behavior during casting
A single line of girders was instrumented in order to monitor temperature and strains within the girders, as shown in Fig. 4. There are five girder lines, each with four spans: 28.0, 28.0, 17.5, and 17.5 m (91.9, 91.9, 57.4, and 57.4 ft). The longer spans use AASHTO Type IV girders and the shorter spans use AASHTO Type III girders. The instrumented girders are designated A4, B4, C4, And D4.
The use of the HPC mix eliminated one line of girders and increased transverse girder spacing from the original design using the conventional concrete. The strength requirement for the girders was 69.0 MPa (10,000 psi) to 76.0 MPa (11,000 psi) at 28 days, and the average strength of the tested cylinders met the requirements as shown in Table 2.
In order to monitor the temperature gradients within the girders both during the curing period and the long-term testing, a total of 22 thermocouples (Omega FF-K-24) were placed at five cross sections of each of the four girders. Ten thermocouples were placed at mid-span, three at 1/4 L, three at 3/4 L, and three at a distance of L/ 50 from either end (where L is the girder span. To measure strains in the concrete, a series of Vibrating Wire ages (VWGs) (Roctest EM-5) were placed at the center of gravity of the prestressing strands at mid-span and 1.524 m (5 ft) for girders C and D, and 1.219 m (4 ft) for girders A and B, in either side from mid-span to measure long-term strains. Prestressing force was measured with load cells and transfer length was determined from strains measured by using an embedded steel bar with attached strain gages.

Load cells (Strainsert model PC-50 with 220 kN (50,000 lb) capacity) were placed at the dead end of the casting bed to measure the prestressing force after tensioning, as well as after curing, and immediately prior to detensioning of the strands. The load cells were placed on four strands on each of the casting beds (Type III and Type IV). Tables 8 and 9 summarize the load cell readings at various times after tensioning.



The transfer length of prestressing strands at the ends of the girders was determined by measuring the strains in a ‘‘strain gage bar” (SGB) embedded in each girder. Bars used in girders C4, D4, and B4 had eight strain gages, while the bar used in girder A4
had nine strain gages. The gages were read with a standard strain gage indicator after the concrete had cured and just before detensioning, in order to obtain the initial readings. The gages were then read immediately after detensioning. The change in strain for each gage is plotted in relation to its distance from the end of the girder in Figs. 10 and 11. The average strain is also plotted in these figures. To determine the transfer length from the measured strains, a method similar to that proposed by Oh and Kim [10] was followed to establish the strain plateau, which was obtained by drawing a horizontal line at 95% of the maximum value of the plotted average strains. The transfer length is taken as the horizontal distance from the origin (i.e. the end of the girder) to where the horizontal line intersects the plotted average strain profile (see Figs. 10 and 11). For the Type III girder, a transfer length of 0.711 m (28 in.) was estimated, while 0.660 m (26 in.) was estimated for the Type IV girder.

Fig 6. Transfer length for Type III girder.

Fig 7. Transfer length for Type IV girder

Figs. 12 and 13 represent a sample of the girder curing temperatures along with the ambient temperature.

Fig. 8. Thermocouples 1–5 for girder C4 at mid-span.

Fig 9. Thermocouples 6–10 for girder C4 at mid-span.

Concrete strains were recorded by using embedded VWGs. A typical result is shown in Fig. 14 for a series of vibrating wire gages. It is noticed that the strain values change as the heat of hydration develops.

Fig.10. VWG strains for girder D4.

Table 10 shows that the prestress loss due to elastic shortening based on the strain measurement is less than predicted, especially for the Type III Girder. It is suspected that the gages failed to record the entire compressive strain during detensioning; this could be due to some reasons such as inadequate consolidation of concrete around the embedded gages or failure of the gage itself. By using the predicted loss due to elastic shortening, it appears that the total prestress loss for the Type IV Girder given in Table 11 is slightly overestimated.




As previously noted, only girder-line 4 shown in Fig 10 was instrumented.

Fig10.Plan view
In this phase of reserch three types of instruments were utilized in this phase, twelve previously installed thermocouples (Omega FF-K-24) were retained and an additional two thermocouples were placed in the deck at a distance of L/4 from the supports in spans A and D, as shown in Fig. 12. Twelve of the previously installed EM-5 Vibrating Wire Gauges (VWGs), were retained and additional VWG’s were placed in the deck at supports. Finally, one additional LVDT was used at each abutment and two extra LVDT’s were used at the expansion joint to measure the longitudinal movement of the girder. All instruments were connected to CR23X Campbell scientific data-loggers, placed at bent diaphragms under the bridge and powered by solar panels. Two data-loggers were used one for spans A and B, and the other for spans C and D. The data was recorded every 4 h over a period of four month under normal traffic loading, at each period the data-logger recorded the instruments’ readings for five minutes. The data-logger records up to 1500 readings per second. The monitoring of the bridge started two months after it was opened for traffic. Fig. 11shows span C and D end displacements due to thermal effects in addition to differential displacements measured by LVDTs during the four month period.

Fig. 11. End displacements due to thermal effect and traffic loading (a) Girder C, and (b) Girder D.

Thermal effects were calculated in reference to the lowest temperature, these thermal effects were calculated based upon the temperature gradients obtained from thermocouples number 3, 4, and 5 for Girder D4 and thermocouples number 6, 7, and 8 for Girder C4 as shown in Fig. 12

Fig. 12. Retained thermocouples of girder cross-sections along girder-line 4 and locations in each girder cross-section.

The differential displacements represent the difference between the LVDT’s reading at anytime and its reading when the lowest temperature was recorded. The displacements measured by the LVDTs placed at the end of the girders show additional end displacements due to traffic loading; however, maximum girder end than 0.0064 m (1/4 in.). End displacement of girders caused by end rotations due to temperature gradient along the depth of the girder cross-section was found to be minimal and it has insignificant effect on total end displacements.displacement due to thermal effects and traffic loading was less
Two static live load tests were conducted on this bridge, the first test took place before the bridge was opened for traffic and the second test was eight months after the bridge was put in-service. In both tests, a five-axle truck was used (Type 3S2 AASHTO designation), loaded roughly to full capacity in one run and to half capacity in the second; the truck and it is total weight are shown in Fig. 13

Fig. 13. Configuration and weights of the truck used in live load tests.
The truck was positioned on 10 different locations as shown in Fig. 14 to maximize moment at mid-spans and supports, and to estimate load distribution factors by positioning the driver’s wheel on Girder-4 in one run and the passenger’s wheel on the adjacent girder in another.

Fig. 14. Truck loading positions.
In addition to the internal gages alreadyembedded in the girder (VWG’s and LVDT’s), two temporary string potentiometers were placed under the bridge at mid-span D and mid-span A to record maximum deflections. The live load tests were performed by placing the loaded five-axle truck on the desired location, when the truck and trailer came to rest at the designated loading position, instruments readings were recorded for a period of 30 s. The truck was then moved to next position without
unloading the bridge, and then readings were recorded in the previous manner.

Fig.15. Strain distribution versus cross-section depth, for girder Type-IV at the support between span A and B due to different loading positions.

Fig.15.represents the strain distribution at the support between span A and B due to the different loading positions. It is clear that strains due to half-loaded truck are approximately half of the strains due to the fully loaded truck, which indicates that the girder behaved elastically as expected. Load versus strain for both tests is shown in Fig. 16. Strain is measured at mid-span D due to loading position 1 and at the mid-depth of deck-slab at support between spans C and D due to loading position 2. The
strains recorded in the second live load test are larger than those recorded in the first live load test, possibly indicating some minor softening of the system due to micro cracking in the tension zone.

Fig. 16 Load versus strain for both live load tests; (a) strain measured at mid-span D. (b) strain measured at support between span C and D.

A comparison between experimental and calculated strains is shown in Figs.17 and 18 .A simplified model was used for strain calculations; every two spans were assumed to be a continuous beam, although, joints constructed between spans do not guarantee full rigidity. Load distribution factors were obtained according to AASHTO provisions and axle loads were distributed accordingly.

Fig.17. Experimental and calculated strain values due to half-full and full trucks.

Fig. 17 a and b show measured and calculated strains at mid-spans D and A, respectively. Calculated strains were higher than measured strains for both spans and for different loading values and positions. As a result, one can conclude that the load distribution factors given by AASHTO are higher than the actual load distributions.

Fig.18. Experimental and calculated strain distribution along girder Type IV at middle support between spans A and B due to loading position 5.

Fig. 18 shows the strain distribution along the cross-section depth at the middle support between spans A and B. Note that the calculated neutral axis depth was found to be smaller than the actual one. The slight drop in the neutral axis depth between live load test one and two could be attributed to minor cracking in the bridge deck and diaphragm. Another comparison can be made in terms of middle span deflection. Fig. 19 a and b shows measured and calculated deflections at middle spans D and A, respectively. Again, the calculated deflections were found to be higher than the actual recorded values.

Fig. 19. Experimental and calculated deflections due to half-full truck (250 kN).
This research examined the material properties and behavior of four prestressed HPC girders during casting and initial curing as well as during service. Based on this research, the following conclusions can be drawn:
1. In general, the HPC used in this research was of good quality.
. The concrete performed well and met the expected results in compressive
strength, modulus of rupture, and creep.
2. The modulus of elasticity of the concrete was lower than expected,
3. The results of the rapid chloride permeability test were higher than expected. The higher values of permeability are most likely the results of higher cement content of the HPC mix and the use of heat and air cure of the test specimens. Both of these two factors could cause more and larger pore structures of the concrete paste, which would in turn increase the permeability of the concrete.
4. During concrete curing, the temperature measured by the embedded thermocouples showed that peak temperatures occurring 7–8 h after casting never reached more than
80°cTherefore, there was no danger of thermal cracking.
5. Based upon the load cell readings (Tables8and 9), practically there were no changes of the initial prestressing force up to the time of detensioning. Therefore the measurement suggested that there was no loss of prestress due to strand relaxation prior to detensioning.
6. Upon detensioning, the transfer lengths for the 0.015 m (0.6 in.) strand were found to be 0.711 m and 0.660 m respectively, for Type III and Type IV girders. These values are slightly less than the standard design value of 50 times the strand diameter or 0.762 m
7. The calculated prestress loss due to elastic shortening was 82.7 MPa for the Type III girders and 124.8 MPa for the Type IV girders. Total prestress loss was 179.3 MPa), i.e. 12.9%, for the Type III girders and 262.7 MPa, i.e. 19.1%, for the Type IV girders.
8. The predicted camber compared closely with the measured camber. The close prediction was possible because the use of load cells at the anchoring end of the prestressing bed provided a more accurate value of the prestressing force at
transfer than the normally assumed prestressing force based on estimated loss of prestress.
9. Girder end displacements were caused mainly by thermal effects with small effect due to traffic loading, while displacements due to end rotations could be neglected, however, maximum total girder end displacement was less than a quarter an inch.
10. The calculated strains and deflections based on AASHTO load distribution factors were found to be higher than actual recorded data.

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