(完整版)房屋建筑毕业设计 4外文翻译

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毕业设计 (论文)

外文翻译

设计（论文）题目：宁波天合家园某住宅楼

2号轴框架结构设计与建筑制图

学院名称：建筑工程学院

专业：土木工程

指导教师：马永政、陶海燕

2012 年 12 月 10 日

外文原稿1

Tension Stiffening

in Lightly Reinforced Concrete Slabs

1R. Ian Gilbert1

Abstract: The tensile capacity of concrete is usually neglected when calculating the strength of a reinforced concrete beam or slab

even though concrete continues to carry tensile stress between the cracks due to the transfer of forces from the tensile reinforcement to the concrete through bond. This contribution of the tensile concrete is known as tension stiffening and it affects the member's stiffness after cracking and or slab eventhough concrete continues to carry tensile stress between thecracks due to the transfer of forces from the tensile reinforcementto the concrete through bond. This contribution of the tensileconcrete is known as tension stiffening

and it affects the member'sstiffness after cracking and amount permittedby the relevant building code. For such members

the flexuralstiffness of a fully cracked cross section is many times smallerthan that of an uncracked cross section

and tension stiffeningcontributes greatly to the stiffness after cracking. In design

deflectionand crack control at service-load levels are usually thegoverning considerations

and accurate modeling of the stiffnessafter cracking is required.

The most commonly used approach in deflection calculationsinvolves determining an average effective moment of inertia [Ie]for a cracked member. Several different empirical equations areavailable for Ie

including the well-known equation developed byBranson [1965] and recommended in ACI 318 [ACI 2005]. Othermodels for tension stiffening are included in Eurocode 2 [CEN1992] and the [British Standard BS 8110 1985]. Recently

Bischoff [2005] demonstrated that Branson's equation grossly overestimates

thtie average sffness of reinforced concrete memberscontaining small quantities of steel reinforcement

and moment reaches the flexural tensile strength of the concrete or modulus of rupture

fr. There is a sudden change in the local stiffness at and immediately adjacent to this first crack. On the section containing the crack

the flexural stiffness drops significantly

but much of the beam remains uncracked. As load increases

more cracks form and the average flexural stiffness of the entire member decreases.

If the tensile concrete in the cracked regions of the beam carried no stress the load-deflection relationship would follow the dashed line ACD in Fig.

1. If the average extreme fiber tensile stress in the concrete remained at fr after cracking

the loaddeflection relationship would follow the dashed the actual response lies between these two extremes and is shown in Fig. 1 as the solid line AB. The difference between the actual response and the zero tension response is the tension stiffening effect （ in Fig. 1）.

As the load increases

the average tensile stress in the concrete reduces as more cracks develop and the actual response tends toward the zero tension response

at least until the crack pattern is fully developed and the number of cracks deflection calculations.

3.Models for Tension Stiffening

The instantaneous deflection of beam or slab at service loads may be calculated from elastic theory using the elastic modulus of concrete Ec and an effective moment of inertia

Ie. The value of Ie for the member is the value calculated using Eq. [1] at midspan for a simply supported member and a weighted average value calculated in the positive and negative moment regions of a continuous span

（1）

where Icr=moment of inertia of the cracked transformed section;Ig=moment of inertia of the gross cross section about the centroidal axis [but more correctly should be the moment of inertia of the uncracked transformed section Iuncr]; Ma=maximum moment in the member at the stage deflection is computed; Mcr=cracking moment =(frIg yt); fr=modulus of rupture of concrete (=7.5 fc in psi and 0.6 fc in Mpa); and yt=distance from the centroidal axis of the gross section to the extreme fiber in tension.

A modification of the ACI approach is included in the Australian Standard concrete may reduce the cracking moment significantly. The cracking moment is given by Mcr=(fr? fcs)Ig yt

where fcs is maximum shrinkage-induced tensile stress in the uncracked section at the extreme fibre at which cracking occurs（Gilbert 2003）.

（2）

where distribution coefficient accounting for moment level and degree of cracking and is given by

（3）

and 1=1.0 for deformed bars and 0.5 for plain bars; 2=1.0 for a single

short-term load and 0.5 for repeated or sustained loading; sr=stress in the

tensile reinforcement at the loading causing first cracking (i.e.

when the moment equals Mcr)

calculated while ignoring concrete in tension; s is reinforcement stress

at loading under consideration (i.e.

when the in-service moment Ms is acting)

calculated while ignoring concrete in tension; cr=curvature at the section

while ignoring concrete in tension; and uncr=curvature on the uncracked transformed section.

For slabs in pure flexure

if the compressive concrete and the reinforcement are both linear and elastic

the ratio sr s in Eq.(3) is equal to the ratio Mcr Ms. Using the notation

of Eq.（1）

Eq.(2) can be reexpressed as

（4）

For a flexural member containing deformed bars under shortterm loading

Eq. (3) becomes =1?（Mcr Ms）2 and Eq.（4）can be rearranged to give the

following alternative expression for Ie for short-term deflection calculations [recently proposed by Bischoff (2005)]: (5)

This approach

which .

4parison with Experimental Data

To test the applicability of the ACI 318

Eurocode 2

and BS 8110 approaches for lightly reinforced concrete members

the measured moment versus deflection response for 11 simply supported singly reinforced one-way slabs containing tensile steel quantities in the range 0.0018<<0.01 are compared with the calculated responses. The slabs (designated S1 to S3

S8

SS2 to SS4

and Z1 to Z4) were all prismatic

of rectangular section

850 mm wide

and contained a single layer of longitudinal tensile steel reinforcement at an effective depth d (with Es=200

000 MPa and the nominal yield stress fsy=500 Mpa). Details of each slab are given in Table 1

including relevant geometric and material properties.

The predicted and measured deflections at midspan for each slab when the moment at midspan equals 1.1

1.2

and 1.3 Mcr are presented in Table 2. The measured moment versus instantaneousdeflection response at midspan of two of the slabs (SS2 and Z3) are compared with the calculated responses obtained using the three code approaches in Fig. 2. Also shown are the responses if cracking did not occur and if tension stiffening was ignored.

5.Discussion of Results

It is evident that for these lightly reinforced slabs

tension stiffening is very significant

providing a large proportion of the postcracking stiffness. From Table 2 the ratio of the midspan deflection obtained by ignoring tension stiffening to the measured midspan deflection (over the moment range Mcr to 1.3 Mcr)is in the range 1.38-3.69 with a mean value of 2.12. That is

on average

tension stiffening contributes more than 50% of the instantaneous stiffness of a lightly reinforced slab after cracking at service load.

For every slab

the ACI 318 approach underestimates the instantaneous deflection after cracking

particularly so for lightly reinforced slabs. In addition

ACI 318 does not model the abrupt change in direction of the moment-deflection response at first cracking

nor does it predict the correct shape of the postcracking moment-deflection curve.

The underestimation of short-term deflection using the ACI318 model is considerably greater in practice than that indicated by the laboratory tests reported nature of cracking

the agreement between the Eurocode 2 predictions and the test results over such a wide range of tensile reinforcement ratios is quite remarkable. With the ratio of () in Table 2 varying between 0.80 and 1.39 with a mean value of 1.07

the Eurocode 2 approach certainly provides a better estimate of short-term behavior than either ACI 318 or BS8110.

6.Conclusions

Although tension stiffening 11 laboratory tests on slabs containing varying quantities of steel reinforcement. The Eurocode 2 approach （Eq.（5）

并在一个有效深度载有纵向拉伸单层钢筋d(Es=200000MPa和屈服应力=500MPa)

每个板块的详细情况见表1

包括有关的几何和材料特性

在每个板跨中挠度的预测结果与实测时

在跨中力矩等于1.1

1.2和1.3Mcr列出在表2

与瞬时变形响应的测量力矩的两跨中的板

（SS2 and Z3）进行比较和计算结果获得图2

使用三个代码方式同时显示的结果

如果没有出现开裂

如果张力加劲被忽略

5.讨论结果

很明显

这些轻型钢筋板

张力加劲非常显著

提供一个大比例的开裂后刚度

从表2

跨中挠度的比例得到了加劲

对测量张力跨中挠度忽视（在Mcr和1.3Mcr范围）是在1.38-3.69范围

取平均值2.12

也就是说

平均而言

张力加劲超过50％的一个轻型钢筋板在屈服荷载的瞬间开裂

对于每一个板

在ACI 318的方法低估了瞬间挠度后开裂

特别是对于轻型钢筋板

此外

在这一时刻ACI 318突然不成模型

在起初开裂处

突然改变力矩偏转结果的方向

也没有预测的正确形状矩挠度曲线

在短期挠度的低估使用ACI 318模式是经化验报告在这里在表示实践中相当大的比

不同于Eurocode 2和BS 8110

ACI 318模型不承认或为在开裂的力矩

这将不可避免地减少在实践中出现的由于张力引起的混凝土干燥收缩或热变形

对于许多板

因早期干燥或温度变化在数周内将发生铸件的开裂

以及经常暴露之前

其板全方位服务的负荷

通过限制混凝土拉伸应力水平的拉伸筋只有1.0 MPa

BS 8110的方法对测试板的上下挠度和立即高于开裂力矩的高估

由于约束的早期收缩和热变形

这并非不合理和占损失的刚度发生在实践中

不过

BS 8110提供了一个相对较差模型刚度

并错误地认为

平均拉力混凝土裂缝进行了实际调高M增大和中性轴的上升

因此

BS 8110开裂后力矩偏转斜率图甚至超过了所有板测量斜坡

这种方法使用比Eurocode 2或ACI两种方式更繁琐

在所有情况下

Eurocode 2挠度计算[EPS.(3)-(5)]是在更接近与实测挠度在整个负载范围内协议可以看出在图2

荷载-挠度曲线的形状并使用Eurocode 2是一个比这更好的代表性实际曲线结果

使用EP.(1)

考虑到具体的变异材料性能影响的板

该协议Eurocode 2在运行特征和对开裂的随机性之间的预测和试验结果在如此广泛的受拉钢筋比率是相当显著的

在图2（）0.80和1.39之间的值平均值为1.07

Eurocode 2的方法提供了ACI 318或BS 8110更好地估计短期行为

6.结论

虽然张力加劲只对重钢筋梁挠度的影响相对较小

这是非常重要的对于Iuncr ICR的比例很高的轻型钢筋构件

例如作为最实用的钢筋混凝土楼板

加劲张力的模型纳入ACI(2005)

Eurocode 2(CEN1993)

和BS 8110(1985) 已提交并且轻型钢筋混凝土楼板的适用性已进行评估

计算模型的三个代码瞬时挠度进行了比较与来自11个实验室测试测量挠度在含有不同数量的钢筋板

在Eurocode 2方案EP.(5)已被证明是更准确地模拟了瞬时负载变形的加固构件轻型钢筋构件的波形和ACI 318（EP.(1)比更为可靠的方法

出自：JOURNAL OF STRUCTURAL ENGINEERING (c) ASCE JUNE 2007

参考文献

[1]American Concrete Institute （ACI）.（2005）. "Building code requirements for structural concrete." ACI 318-05

ACI Committee 318

Detroit.

[2]Bischoff

P. H. (2005). "Reevaluation of deflection prediction for concrete beams reinforced with steel and fiber-reinforce polymer bars." J.Struct. Eng. 131(5)

[3]Branson

D. E. (1965). "Instantaneous and time-dependent deflections ofsimple and continuous reinforced concrete beams." HPR Rep. No. 7

1

Alabama Highway Dept.

Bureau of Public Roads

Ala.

[4]British Standards Institution (BS).（1985）. "Structural use of concrete Part 2

code of practice for special circumstances." BS8100: Part2:1985

British Standard

London

England.

[5]European Committee for Standardization（CEN）. （1992）. "Eurocode 2:Design of European Prestandard

Brussels

Belgium.

[6]Gilbert

determination of fcs." Australian J. Structural Engineering

5（1）

61-71.

[7]Standards Australia（AS）. （2001）. "Australian standard for concrete Sydney

Australia.JOURNAL

外文原稿2

The Twelfth East Asia-Pacific Conference on Structural Engineering and Construction

Design of Building Structures to Improve their Resistance

to Progressive Collapse

D A Nethercota

a Department of Civil and Environmental Engineering

Imperial College London

Abstract：It is rare nowadays for a "new topic" to emerge within the relatively mature field of Structural Engineering. Progressive collapse-or

more particularly

understanding the mechanics of the phenomenon and developing suitable ways to accommodate its consideration within our normal frameworks for structural design-can be so regarded. Beginning with illustrations drawn from around

the world over several decades and culminating in the craft based to science based and thus from prescriptive to quantitatively justified approaches This can be traced through concepts such as: permissible stress

ultimate strength

limit states and performance based. As clients

users and the general public of: an assessment of demand

modelling behaviour and identification of suitable failure criteria. The treatment of topics such as fatigue

fire resistance

durability and serviceability can all be seen to built schemes

demolition of several completed buildings

temporary removal of gas in level of recent US Guidance (SEI 2010)

illustrates the principle. The approach is simple to appreciate

requires minimal structural calculation and

in situations where the original provisions are found to be inadequate can be made to work by providing more substantial connections andor additional reinforcement in floor slabs In an interesting recent development that recognizes the link to the generation of catenary action

US Guidance the alternative load path approach presumes the instantaneous loss of a single column and then requires that the ability of the resulting damaged structure to bridge the loss be demonstrated by suitable calculation (Gudmundsson and Izzuddin 2010). The approach may be implemented at varying levels of sophistication in terms of the analysis; for example

recent thinking in the United States (SEI 2010) makes provision for any of: linear static

non-linear static or non-linear dynamic analysis and provides some guidance on the use of each. It may also be used as the basis for more sophisticated numerical studies of particular structures and particular incidents e.g. forensic work; the best of these-which are likely to be computationally very demanding- without separation of key elements.

Figure 2: Simplified multi-level approach for progressive collapse assessment

Additional features

designed to make the approach attractive for use by practicing Engineers level

see Figure 2.

*? A realistic criterion of failure is employed

corresponding to reaching the ductility limits in connections.

*? Quantitative comparisons between alternative structural arrangements may readily be made.

*? The approach may be implemented using only explicit formulae

thereby permitting simple and rapid calculation.

Full details of the method

both in its original form which utilises ADAPTIC to perform the calculations and in its simplified form

may be found in the series of Imperial papers (2-12).

*a) First yielding of the tensile components (top bolt row of the support connection)

*b) Ultimate capacity of the beam flange at one of the connections (support)

*c) Ultimate capacity of the system (failure of the bottom bolt row of the mid-span connection)

*d) The axial load becomes zero (the deflection of the beam where the axial load changes from compressive to tensile)

*e) The deflection of the beam where the axial load becomes equal to the flange capacity of one of the connections (mid-span connection)

Figure 3: Non-linear static response for a single beam

中文翻译2

通过建筑结构设计以改善建筑物的抗倒性

D A Nethercota

土木与环境工程学院--伦敦帝国学院

摘要：如今的"新话题"出现在相对成熟的结构工程领域这是一件罕见的事

抗连续倒塌

或者

更特别的是

了解力学的现象和发展适当的方式

以适应我们正常的框架内审议的结构设计

可以这么认为

在过去的几十年

从来自世界各地的插图画开始

到高高的世贸中心倒塌为止

这些功能必不可少的为具有代表性的治疗和早期的设计方法进行了综述

最近的工作是当时的报道

集中精力在过去7年在伦敦大学帝国学院的发展

在一个能使用各种水平和由设计师一直在发展适合直接使用的综合方法

说明性的结果是用来帮助发现一些关键的管理功能

去展示如何定量比较安排现在可能使和说明赫尔墨斯的一些以前的设计概念之间的不同来直接改善抗倒性

关键词：复合结构

渐进式折叠

鲁棒性

钢铁结构

结构设计

1.引言

随着时间的推移各种不同的结构设计原理被提出

他们发展的自然回想：

*? 越来越关注确保足够的性能

*? 改进过的性能的科学知识

*? 加强能力从工艺为基础的移动科学依据从而从规范的定量合理的方法这可以通过追踪的概念

如:容许应力

强度极限

极限状态和基础性能

作为客户：用户和公众已经变得越来越复杂

因此要求更高的期望

因此

它成为必要的设计师代替一个永久的越来越多的结构性问题的范围的主要是通过考虑达到这个条件将或多或少受到不可接受的方法

所以问题不是以前认为（或只允许在一个隐式的

基本上复制过去的令人满意的性能

方式）开始需要显式的形式的关注：需求评估

模型行为和识别合适的失效准则

论题的处理比如疲劳

耐火性

耐久性和适用性都可以被看作是这个模式

举一个具体的例子：设计充分耐火钢框架建筑开始(已经被认可的)和简单的法定规则对混凝土外层脆弱的构件

但是

近年来

发展成为一个复杂的消防工程学科

关心火灾荷载

提供防护系统

如洒水装置

在发生火灾情况下的反应的计算

能够使定量对比结构安排之间选择

不仅导致了在某种意义上不提供防火时明显的经济效益

在它给了只有微不足道的好处的时候；它也导致了消防安全通道更好的调节原则的理解和明智的行事能力在设计适合安排一个合适的评估基础上的需要

在罗南点于1968年在伦敦坍塌之前

鲁棒性原则

抗连续性倒塌

非比例破坏等是不属于工程词汇里的

这栋在18层发生瓦斯爆炸被破坏的22层预制混凝土公寓建筑导致了新的英国建筑法规诞生

取缔了多年来所谓的系统构建方案

拆除了几个完整的建筑物

排除高层建筑物里的临时瓦斯和建立建构安全方面的常务委员会

最终

合理设计的预处理的好处是公认的

安全的方法来安装燃气设计然后开始进入工业

然后

结构设计指导在当时产生--仍然决定了很多现在的条款--是自然本质上的处方式

没有真正的链接到实际的性能

后来

连续倒塌的发生率如同Murragh Building和世贸中心带来增加如何合理地考虑那些结构的设计实际现象和问题的关注

它被认为是合适的

在这样做

当然

至关重要的风险

包括一个触发事件和失败的结果

所以更繁重的结构要求被适当地使用

可以说

一个不成比例的反应在风险后果是很低很轻的地方要求昂贵的附加条款的情况下

也许如同未能解决那些情况在风险后果是高严重的地方一样有害

本文将回顾当前用来设计抵制连续倒塌的方法和对比过去七年在伦敦大学帝国理工学院进行的这些工作

那里的目标是提供一个依据于实际方法适合用在常规设计

方法的基本特征将被提交

它被使用在几个例子的描述和结果来说明它是如何导致更好的连续倒塌的机制的原理和结构工程师的方式能最好的配置结构

以提供增强的抗性

2.设计抵抗连续倒塌

两种最常用设计方法旨在解决连续倒塌这一问题：

*? 提供绑扎能力

*? 检查交替的荷载通道

图1:领带部队在一个框架结构

首先

本质上的规范和包括确保梁

柱

楼梯和楼板（或者屋盖）可以联合起来提供一个规定的低级的水平联系抗力等级；垂直荷载的实际值要求是通常有相关的

图1

这个来自最近的US Guidance

演示了原理

这个方法对于观察是简单的

只需要很少的结构计算和在最初的规定被发现是不充分的的情况下

能通过提供更多的实质性的连接或在一个有趣的近代发展中水泥楼板中施加额外加固

认识到链式反应的连接的生成

US Guidance已经限制可以展示相关的连接可以携带所需的弹性元件同时进行0.2光辉的旋转的情况的结构构件之间的绑扎的使用

这是不可能的

连系材料应该通过楼板和屋盖

无论如何

近代研究（Nethercot et al 2010a; Nethercot et al 2010b）都建议绑扎力相关较弱和实际抗力去抗连续倒塌

此外

它被规范不允许有意义的替代安排的比较--结构设计的一个基本特征在其最频繁使用的形式替代负载路径方法假定一个单柱的瞬时损耗

然后需要这作为结果的被损伤的构件的能力去渡过这个损失已经被合适的计算证明（Gudmundsson and Izzuddin 2010）

该方法可以在分析方面的高度化的不同程度被实现；比如

在美国最近的研究为线性静力分析

非线性静态或非线性动态分析制定规定和为各自的使用提供一些指导

它也可以被使用作为基点为特定结构或特定工作（如法医）的更精致的数字的研究；最好的这些--可能是计算非常苛刻的--已经证明了他们的能力去紧密地复制的真实的可观察的特性

3.抗连续倒塌的基本特性

三个特征已经预先被鉴证出作为任何合理的现实的方法去设计对抗连续倒塌的必要部分：

*? 事件发生在非常短的时间段内和正在的失败是因此动态

*? 它包括总变形

发生大应变

导致非弹性行为和几何效果的改变一样

*? 失败基本上对应于在受损状态下构件的无能通过一个新的没有关键元素的分离的平衡位置

图2:简化的多层次评估方法抗连续倒塌