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Article(id=1241699616937276337, tenantId=1146029695717560320, journalId=1240670690148397066, issueId=1241699613942543237, articleNumber=null, orderNo=null, doi=10.3963/j.issn.1001-487X.2024.03.015, pmid=null, cstr=null, oa=null, hot=null, price=null, onlineType=0, articleFormat=0, articleType=null, articleTypeStr=null, receivedDate=1713456000000, receivedDateStr=2024-04-19, revisedDate=null, revisedDateStr=null, acceptedDate=null, acceptedDateStr=null, onlineDate=1773973858341, onlineDateStr=2026-03-20, pubDate=1725120000000, pubDateStr=2024-09-01, doiRegisterDate=null, doiRegisterDateStr=null, onlineIssueDate=1773973858341, onlineIssueDateStr=2026-03-20, onlineJustAcceptDate=null, onlineJustAcceptDateStr=null, onlineFirstDate=null, onlineFirstDateStr=null, sourceXml=null, magXml=null, createTime=1773973858341, creator=13701087609, updateTime=1773973858341, updator=13701087609, issue=Issue{id=1241699613942543237, tenantId=1146029695717560320, journalId=1240670690148397066, year='2024', volume='41', issue='3', pageStart='1', pageEnd='260', issueExtLink='null', onlineDate='null', pubDate='null', beforeIssueId=null, nextIssueId=null, price=null, status=1, issueComplete=1, articleOrder=1, issueType=-1, specialIssue=null, createTime=1773973857626, creator=13701087609, updateTime=1773992982583, updator=13701087609, preIssue=null, nextIssue=null, ext={EN=IssueExt(id=1241779829880721843, tenantId=1146029695717560320, journalId=1240670690148397066, issueId=1241699613942543237, language=EN, specialIssueTitle=, coverIllustrator=null, specialIssueEditor=, specialIssueAbout=), CN=IssueExt(id=1241779829880721844, tenantId=1146029695717560320, journalId=1240670690148397066, issueId=1241699613942543237, language=CN, specialIssueTitle=, coverIllustrator=null, specialIssueEditor=, specialIssueAbout=)}, issueFiles=null}, startPage=121, endPage=129, ext={EN=ArticleExt(id=1241699617214100415, articleId=1241699616937276337, tenantId=1146029695717560320, journalId=1240670690148397066, language=EN, title=Analysis on Structural Dynamic Response Characteristics of RC Frame Building under Blasting Failure Condition, columnId=1240702072073548400, journalTitle=Blasting, columnName=EXPLOSIVE DEMOLITION, runingTitle=null, highlight=null, articleAbstract=

To study the continuous collapse resistance and dynamic response characteristics of the structure after the blasting failure of the local columns of the RC frame building, the deformation and stress adjustment process of the beam-column substructures adjacent to the columns were observed in real-time through an on-site blasting test of the central column of a blasting and demolition project of an 8-storey frame building. The PKPM was used to establish the corresponding building model. Meanwhile, the dynamic response characteristics of the remaining structure under the failure of the central column and the resistance to continuous collapse was calculated by the demolition component method and the demolition component method in SAP 2000. The results show that the theoretical value of strain after the blast failure of the center column is about 260 με using the strain-moment theoretical formula. The dynamic strain measured in the field is about 377 με, and the value calculated by the numerical simulation is about 238 με. The results of the dynamic strain by the three methods are relatively close. The computed value of the vertical displacement at the failure point is 3.2 mm, close to the field displacement of 2.67 mm, and the calculated value of the plastic angle is 0.051°, close to the field angle of 0.05°. The remaining structure experienced a significant dynamic impact at the instant of the central column failure, and the acceleration values along the positive and negative directions are roughly the same, with a maximum value of 3.5 m/s. After the failure of the central column, the load redistribution occurs in the remaining structure, and the vertical load originally borne by the central column is shared by the surrounding columns, resulting in a significant catenary effect on the upper beam body.

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为研究RC框架楼房局部立柱爆破失效后结构的抗连续倒塌性能和动力响应特性，依托一栋8层框架结构楼房爆破拆除工程，通过中心立柱现场爆破试验，对立柱邻近梁柱子结构的变形与应力调整过程进行了实时观测。采用PKPM建立实际工程对应的建筑模型，并基于拆除构件法，应用SAP 2000中非线性动力分析方法计算中心立柱失效下剩余结构的动力响应特性以及抗连续倒塌能力。结果表明：中心立柱爆破失效后，采用应变-弯矩理论公式得出的应变理论值约为260 με、动应变现场实测值约为377 με、数值模拟值约为238 με，三种方法得出的动应变结果较为接近；失效点竖向位移计算值3.2 mm，与现场实测位移2.67 mm接近，其导致的塑性转角计算值为0.051°，与现场实测值0.05°接近；中心立柱爆破失效瞬间，剩余结构经历了明显的动力冲击，其正负向加速度值大致相同，最大加速度达到3.5 m/s2；中心立柱失效后，剩余结构产生荷载重分布，原本由中心立柱承担的竖向荷载被周边柱所分担，导致上方梁体出现显著的悬链线效应。本论文的研究成果有助于框架结构楼房在意外爆炸情况下的易损性评估与抗连续倒塌性能分析，同时也可为拆除爆破中的试爆立柱选择提供参考。

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姚颖康（1981-），男，博士、教授，主要从事建筑物精细爆破拆除和土木工程防灾减灾等领域的理论与技术研究，（E-mail）。

YAO Ying-kang (1981-), male, Ph. D, professor, mainly engaged in the theoretical and technical research in the fields of fine blasting and demolition of buildings and civil engineering disaster prevention and mitigation, (E-mail) .

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1b.江汉大学　爆破工程湖北省重点实验室，武汉　430056, bio={"content":"

姚颖康（1981-），男，博士、教授，主要从事建筑物精细爆破拆除和土木工程防灾减灾等领域的理论与技术研究，（E-mail）。

YAO Ying-kang (1981-), male, Ph. D, professor, mainly engaged in the theoretical and technical research in the fields of fine blasting and demolition of buildings and civil engineering disaster prevention and mitigation, (E-mail) .

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姚颖康（1981-），男，博士、教授，主要从事建筑物精细爆破拆除和土木工程防灾减灾等领域的理论与技术研究，（E-mail）。

YAO Ying-kang (1981-), male, Ph. D, professor, mainly engaged in the theoretical and technical research in the fields of fine blasting and demolition of buildings and civil engineering disaster prevention and mitigation, (E-mail) .

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articleId=1241699616937276337, language=CN, label=图10, caption=立柱失效后失效节点上方梁弯矩时程曲线, figureFileSmall=eBkYqKW6jYQ9yk7MJDlkHA==, figureFileBig=s2/+ESS+YVGMIvI39/XzGQ==, tableContent=null), ArticleFig(id=1241756531738333722, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 1, caption=

List of beam and column reinforcement

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层数构件配筋/mm
1~8层梁4Φ20/4ϕ 16
梁箍筋 ϕ 8@100/200
1层柱8ϕ 20/4Φ18
柱箍筋 ϕ 8@100
2~4层柱8ϕ 20/4Φ16
柱箍筋 ϕ 8@200
), ArticleFig(id=1241756533269254687, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表1, caption=

梁柱配筋表

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层数构件配筋/mm
1~8层梁4Φ20/4ϕ 16
梁箍筋 ϕ 8@100/200
1层柱8ϕ 20/4Φ18
柱箍筋 ϕ 8@100
2~4层柱8ϕ 20/4Φ16
柱箍筋 ϕ 8@200
), ArticleFig(id=1241756533407666730, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 2, caption=

Evaluation criteria for nonlinear analysis

, figureFileSmall=null, figureFileBig=null, tableContent=
钢筋混凝土框架构件延性转角备注
钢筋混凝土框架梁16° 
钢筋混凝土框架柱  
钢筋混凝土框架2°侧移≤H/25
), ArticleFig(id=1241756533529301548, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表2, caption=

非线性分析评估标准

, figureFileSmall=null, figureFileBig=null, tableContent=
钢筋混凝土框架构件延性转角备注
钢筋混凝土框架梁16° 
钢筋混凝土框架柱  
钢筋混凝土框架2°侧移≤H/25
), ArticleFig(id=1241756533659324979, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 3, caption=

Peak strain at beam bottom

, figureFileSmall=null, figureFileBig=null, tableContent=
工况近端峰值应变/με远端峰值应变/με
A应变片损坏302
B222340
C634333
), ArticleFig(id=1241756533793542716, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表3, caption=

梁底峰值应变

, figureFileSmall=null, figureFileBig=null, tableContent=
工况近端峰值应变/με远端峰值应变/με
A应变片损坏302
B222340
C634333
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Displacement test results

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工况观测前/m观测前平均值/m观测后/m观测后平均值/m观测差/mm塑性转角/°
爆破1.177191.177191.179801.179862.670.05
1.177171.17994
1.177211.17985
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位移测试结果

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工况观测前/m观测前平均值/m观测后/m观测后平均值/m观测差/mm塑性转角/°
爆破1.177191.177191.179801.179862.670.05
1.177171.17994
1.177211.17985
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Stress of the beam above the failure column

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构件梁近端/MPa梁远端/MPa
A梁7.64-7.82
B梁10.08-7.35
C梁6.07-5.26
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失效柱上方梁的应力

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构件梁近端/MPa梁远端/MPa
A梁7.64-7.82
B梁10.08-7.35
C梁6.07-5.26
), ArticleFig(id=1241756536255599194, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 6, caption=

Comparison of strain results

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失效点上方梁实测数据/με理论计算数据/με数值模拟数据/με
梁近端峰值梁远端峰值整体梁近端梁远端
A梁应变片损坏302236246252
B梁222340312325237
C梁634333232196169
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应变结果对比

, figureFileSmall=null, figureFileBig=null, tableContent=
失效点上方梁实测数据/με理论计算数据/με数值模拟数据/με
梁近端峰值梁远端峰值整体梁近端梁远端
A梁应变片损坏302236246252
B梁222340312325237
C梁634333232196169
), ArticleFig(id=1241756537908154980, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 7, caption=

Dynamic displacement and acceleration extremes

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 失效点最大位移/mm最大塑性转角/（°）正向最大值/（m·s-2）负向最大值/（m·s-2）
模拟爆破立柱3.20.0513.5-3.6
实验爆破立柱2.670.05  
), ArticleFig(id=1241756538025595497, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表7, caption=

动态位移和加速度极值

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 失效点最大位移/mm最大塑性转角/（°）正向最大值/（m·s-2）负向最大值/（m·s-2）
模拟爆破立柱3.20.0513.5-3.6
实验爆破立柱2.670.05  
), ArticleFig(id=1241756538126258798, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 8, caption=

Change of axial force of column

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失效工况紧邻柱失效前后轴力变化/kN
A柱 B柱 C柱 D柱
失效前-1889.92-2057.31-2186.28-1653.66
失效后-2425.66-2524.12-2678.92-1838.51
), ArticleFig(id=1241756538239505010, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表8, caption=

立柱轴力变化

, figureFileSmall=null, figureFileBig=null, tableContent=
失效工况紧邻柱失效前后轴力变化/kN
A柱 B柱 C柱 D柱
失效前-1889.92-2057.31-2186.28-1653.66
失效后-2425.66-2524.12-2678.92-1838.51
), ArticleFig(id=1241756538365334137, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=EN, label=Table 9, caption=

Bending moment changes of beams at failed joints

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失效工况失效点梁弯矩变化/（kN·m）
失效前失效后
中心立柱失效21.74-367.71
), ArticleFig(id=1241756538453414527, tenantId=1146029695717560320, journalId=1240670690148397066, articleId=1241699616937276337, language=CN, label=表9, caption=

失效节点梁弯矩变化

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失效工况失效点梁弯矩变化/（kN·m）
失效前失效后
中心立柱失效21.74-367.71
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RC框架楼房中心立柱爆破失效条件下结构动力响应特性分析
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姚颖康 1a, 1b , 章强 1a, 1b , 侯永恒 1a, 1b , 贾永胜 1a, 1b, 2
爆破 | 拆除爆破 2024,41(3): 121-129
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爆破 | 拆除爆破 2024, 41(3): 121-129
RC框架楼房中心立柱爆破失效条件下结构动力响应特性分析
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姚颖康1a, 1b , 章强1a, 1b, 侯永恒1a, 1b, 贾永胜1a, 1b, 2
作者信息
  • 1a.江汉大学　省部共建精细爆破国家重点实验室，武汉　430056
  • 1b.江汉大学　爆破工程湖北省重点实验室，武汉　430056
  • 2.武汉爆破有限公司，武汉　430056

    • 姚颖康（1981-），男，博士、教授，主要从事建筑物精细爆破拆除和土木工程防灾减灾等领域的理论与技术研究，（E-mail）。

    YAO Ying-kang (1981-), male, Ph. D, professor, mainly engaged in the theoretical and technical research in the fields of fine blasting and demolition of buildings and civil engineering disaster prevention and mitigation, (E-mail) .

Analysis on Structural Dynamic Response Characteristics of RC Frame Building under Blasting Failure Condition
Ying-kang YAO1a, 1b , Qiang ZHANG1a, 1b, Yong-heng HOU1a, 1b, Yong-sheng JIA1a, 1b, 2
Affiliations
  • 1a.State Key Laboratory of Precision Blasting, Jianghan University, Wuhan 430056, China
  • 1b.Hubei Key Laboratory of Blasting Engineering, Jianghan University, Wuhan 430056, China
  • 2.Wuhan Explosions and Blasting Corporation Limited, Wuhan 430056, China
出版时间: 2024-09-01 doi: 10.3963/j.issn.1001-487X.2024.03.015
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摘要
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为研究RC框架楼房局部立柱爆破失效后结构的抗连续倒塌性能和动力响应特性，依托一栋8层框架结构楼房爆破拆除工程，通过中心立柱现场爆破试验，对立柱邻近梁柱子结构的变形与应力调整过程进行了实时观测。采用PKPM建立实际工程对应的建筑模型，并基于拆除构件法，应用SAP 2000中非线性动力分析方法计算中心立柱失效下剩余结构的动力响应特性以及抗连续倒塌能力。结果表明：中心立柱爆破失效后，采用应变-弯矩理论公式得出的应变理论值约为260 με、动应变现场实测值约为377 με、数值模拟值约为238 με，三种方法得出的动应变结果较为接近；失效点竖向位移计算值3.2 mm，与现场实测位移2.67 mm接近，其导致的塑性转角计算值为0.051°，与现场实测值0.05°接近；中心立柱爆破失效瞬间，剩余结构经历了明显的动力冲击，其正负向加速度值大致相同，最大加速度达到3.5 m/s2；中心立柱失效后，剩余结构产生荷载重分布，原本由中心立柱承担的竖向荷载被周边柱所分担，导致上方梁体出现显著的悬链线效应。本论文的研究成果有助于框架结构楼房在意外爆炸情况下的易损性评估与抗连续倒塌性能分析，同时也可为拆除爆破中的试爆立柱选择提供参考。

框架结构  /  抗连续倒塌  /  立柱失效  /  非线性分析  /  动力响应特性

To study the continuous collapse resistance and dynamic response characteristics of the structure after the blasting failure of the local columns of the RC frame building, the deformation and stress adjustment process of the beam-column substructures adjacent to the columns were observed in real-time through an on-site blasting test of the central column of a blasting and demolition project of an 8-storey frame building. The PKPM was used to establish the corresponding building model. Meanwhile, the dynamic response characteristics of the remaining structure under the failure of the central column and the resistance to continuous collapse was calculated by the demolition component method and the demolition component method in SAP 2000. The results show that the theoretical value of strain after the blast failure of the center column is about 260 με using the strain-moment theoretical formula. The dynamic strain measured in the field is about 377 με, and the value calculated by the numerical simulation is about 238 με. The results of the dynamic strain by the three methods are relatively close. The computed value of the vertical displacement at the failure point is 3.2 mm, close to the field displacement of 2.67 mm, and the calculated value of the plastic angle is 0.051°, close to the field angle of 0.05°. The remaining structure experienced a significant dynamic impact at the instant of the central column failure, and the acceleration values along the positive and negative directions are roughly the same, with a maximum value of 3.5 m/s. After the failure of the central column, the load redistribution occurs in the remaining structure, and the vertical load originally borne by the central column is shared by the surrounding columns, resulting in a significant catenary effect on the upper beam body.

frame structure  /  resistance to continuous collapse  /  column failure  /  nonlinear analysis  /  dynamic response characteristic
姚颖康, 章强, 侯永恒, 贾永胜. RC框架楼房中心立柱爆破失效条件下结构动力响应特性分析. 爆破, 2024 , 41 (3) : 121 -129 . DOI: 10.3963/j.issn.1001-487X.2024.03.015
Ying-kang YAO, Qiang ZHANG, Yong-heng HOU, Yong-sheng JIA. Analysis on Structural Dynamic Response Characteristics of RC Frame Building under Blasting Failure Condition[J]. Blasting, 2024 , 41 (3) : 121 -129 . DOI: 10.3963/j.issn.1001-487X.2024.03.015
正文
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自1968年英国Ronan Point公寓因煤气爆炸引发的连续倒塌事件，至2001年纽约世贸大厦在“9·11”事件中的毁坏，一系列严重的连续倒塌案例促使了国内外学者对建筑结构抗连续倒塌设计的关注。与此同时，伴随城市建设的快速进展及城市更新策略的实施，众多老旧建筑物需要更新或拆除，建筑物拆除的主要方法包括人工拆除、机械拆除以及爆破拆除等。在这些方法中，爆破拆除因其高效率、安全性及经济与环保优势，已成为建筑物拆除的首选方式[1]，但是高层楼房爆破拆除设计主要依赖于工程经验，缺乏量化设计理论。鉴于此，建筑物的精细爆破拆除设计理论和高层结构的抗连续倒塌能力分析已成为当前国内外学者研究的重点领域。
近年来，国内外学者在立柱失效方面做了大量的研究，主要研究方法包括模型试验，数值模拟和理论分析等。模型试验方面，金浏对1/2缩尺的2个后张拉无黏结预应力混凝土梁-柱子结构和1个钢筋混凝土梁-柱子结构进行Pushdown加载[2]，探究了中柱失效下不同钢绞线布置形式对UPC结构的倒塌影响。Qian对四个1/2比例的UPPC梁柱子结构在准静态和动态加载下进行了测试[3]。李治对一个2层2跨钢框架梁-柱子结构进行Pushdown加载试验[4]，研究了结构在倒塌过程中的抗力曲线和破坏模式。数值模拟方面，侯永恒建立了一个4×6跨的8层钢筋混凝土框架结构模型[5]，分析了框架结构在不同立柱失效情况下的动力响应特征。陈泽帆利用OpenSees分别建立了典型RC框架梁柱子结构的静力和动力连续倒塌分析有限元模型[6]，并且确定了结构不确定性对RC框架结构抗连续倒塌性能的影响。张雨笛通过外贴FRP布加固混凝土框架[7]，然后采用数值模拟的方式分析了FRP布加固方式对现浇和装配式混凝土框架子结构抗连续倒塌与抗震性能的影响。Yu通过建立钢筋混凝土填充框架有限元模型与已有试验数据进行对比[8]，探究了立柱失效情况下，剩余结构的倒塌性能。理论分析方面，Yu为了快速预测立柱失效的情况下[9]，钢筋混凝土结构的压拱效应的容量，提出了一种工程分析模型。Zhang为了准确预测梁-板结构在大挠度情况下的主要倒塌机制[10]，提出了一种基于能量的方法。Dat P X提出了一种简化的钢筋混凝土结构二次柱损失时的连续倒塌评估方法[11]，并通过12个梁-板分段组合试验验证了理想的静力响应。提供了一个简单的逐步过程和工作示例。甘艺平通过理论推导将现有等跨结构的分析模型推广至不等跨结构[12]，并提出了一个新的修正三折线模型，可以考虑不同工况下结构的抗弯、压拱、悬梁线和拉膜效应机制。
综上所述，国内外学者在围绕建（构）筑物构件失效条件下结构抗连续倒塌性能，以及在试爆立柱选择上都开展了相关的研究，但是大多研究主要以理论分析、数值模拟和缩尺实验等研究手段，其研究对象与楼房真实的整体结构存在一定差异。为此，本文依托一栋八层框架楼房爆破拆除立柱试爆现场实验，利用PKPM软件建立相对应的建筑模型，再通过SAP 2000有限元软件，运用非线性动力分析方法，拆除对应现场试爆实验中的立柱，进而分析其框架结构的变形能力、加速度响应和荷载传递等动力响应特征，并与实际工程中试爆所得出的位移和应变进行对比，论文成果可以为类似工程试爆立柱的选择以及框架结构楼房在单根立柱爆破失效条件下的抗倒塌能力评判提供参考。
1　工程概况
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待爆破拆除楼房为一栋8层框架结构楼房，由4个单元组成。楼房平面总体呈矩形，总长72.9 m，④~⑤轴之间设有沉降缝，缝宽120 mm；楼房总宽12.9~17.6 m，①~④轴宽16.3 m，⑤~⑨轴宽17.6 m，⑩~⑯轴宽12.9 m；楼高31.0 m，总建筑面积约9041 m2。楼房首层平面结构如图1所示，楼房立柱和梁的配筋如表1所示。
2　局部立柱失效分析理论
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2.1　拆除构件法
拆除构件法是指从结构计算模型中拆除选定的受力构件，模拟构件瞬间失效，然后计算剩余结构的内力和变形，根据《建筑结构抗倒塌设计规范CECS392∶2014》规定[13]，拆除构件后，剩余构件可分为线性静力、非线性静力以及非线性动力3种分析方法。线性静力和非线性静力分析方法在分析时，前者并不能反映结构塑性变形过程，并且两种方法都未考虑动力的影响，通常导致分析结果偏向于保守。鉴于结构构件失效所引起的内力变化是一个复杂的动力过程，本文选择了非线性动力分析方法以提供更为准确的结果。
2.2　立柱失效条件下结构荷载模式和破坏准则
自英国Ronan Point公寓连续倒塌事件发生后，各国都改善了结构抗连续倒塌的规范，目前大多以美国的GSA2003和DOD2010为标准。
（1）美国总务局规范GSA2003
美国GSA2003提出在结构倒塌风险较高时[14]，将采用拆除构件法对结构的抗连续倒塌能力进行评估，并且做出了细致的区分，论文所拆除的为首层立柱，根据GSA2003规定，在做静力分析时，在结构上施加2（G+0.25Q）的荷载组合，其中2是动力放大系数；动力分析时施加（G+0.25Q）的荷载组合，其中G是恒载，Q是活载。
采用线弹性静力分析时，GSA2003规定，用以供需Dcr作为结构可能倒塌的破坏准则，并且GSA2003规定，规则结构Dcr>2，不规则结构Dcr<1.5时，结构破坏发生倒塌的概率较高。
采用非线性静力分析时，美国GSA2003规定以塑性转角和位移线性比作为控制条件，非线性分析的评估标准如表2所示。
（2）美国国防部规范DOD2010
与GSA2003相比，DOD2010规定在做静力分析时[15]，DOD规范也采用能力供需比DCR作为评价标准，Dcr的定义与GSA2003规范相同，但DOD2010以Dcr≤1，作为评判标准。
采用非线性静力分析时，与美国GSA2003判别方法相同。
2.3　应变-弯矩理论
楼房立柱试爆时，爆破产生的冲击波和混凝土飞散物会对周边的人和设备造成损害，课题组采用在试爆立柱上方的梁体上布设应变片以获取立柱爆破失效时楼房结构的动力响应特征。同时，受限于SAP2000有限元分析软件计算模块无法直接获取梁的动应变特征，因此论文应用材料力学弯矩-应变转换理论来分析梁体荷载调整特征[16]，并与实测数据和数值模拟数据进行对比。
弯矩-应变理论，假定梁发生弯曲变形后，尾端的两个横截面仍然保持平面，但是绕各自的中性轴旋转一角度dθ，通过该假定可知微端上到中性面距离为y处长度的改变量为
将线段的长度改变量除以原长dx，即为线段的正应变，根据（1）式得到曲率与应变的关系式
式中，ρ为中性面弯曲后的曲率半径，也是梁弯曲后的曲率半径，根据截面法和平衡条件，纯弯曲时，横截面上面只有弯矩一个内力分量，应用积分的方式
式（4）中负号表示坐标y为正值的微面积dA上的力对z轴之距为负值，M为加载平面内的弯矩，由式（3）和式（4）可得出
根据截面惯性矩定义，将（6）式代入（5）式
由此得到了曲率与横截面上的弯矩以及弯曲刚度的关系式
EI为弯曲刚度，因此根据（2）式和（7）式，可以得出应变和弯矩的关系式
3　现场立柱试爆试验
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3.1　试爆立柱上方梁体应变分析
鉴于立柱爆破作用及其导致的立柱承载特性失效具有极短的时间特性，为了准确获取失效立柱上方梁的变形情况及其动力响应特性，本研究选择了失效立柱上方的承重梁作为主要研究对象。在进行实验测量前，为确保应变片能准确地捕捉到数据，对立柱和上方梁体的装修层及腻子层进行了清除并进行了打磨处理，以提高测量精度。在梁体侧面的20 cm位置贴上了纵向和轴向应变片，同时在梁体两端底部的20 cm位置以及梁中间底部位置也布置了轴向应变片。为了系统地捕捉到数据，研究中选取了位于立柱上方的三根梁，编号分别为A、B、C，其中梁A和B的长度为3.6 m，梁C的长度为7.2 m，即每种工况下各有15个应变测点，如图2所示。
从图3、表3的数据可以看出：
（1）从应变时程曲线图可以看出梁近端由于爆炸荷载的作用，梁体在经历短暂的加载和卸载后，由于立柱完全失效，内力进行重新分配，拉应变达到了峰值，并经过一段时间振荡逐渐趋于平稳的状态。梁远端应变也经历了两次加载与卸载的过程，但与梁近端的过程刚好相反，在立柱完全失效后，内力重新分配，压应变达到了峰值后经过振荡逐渐趋于稳定。
（2）动应变测试结果表明，中心立柱爆破失效后，失效点受到重力荷载的作用，梁底应变变化显著，近端应变均值为428 με，远端的应变均值为325 με。
3.2　立柱失效下的动态位移
失效点位移是判断框架结构是否失效的重要指标，GSA2003规定，当塑性转角大于6°时，结构构件失效。根据梁端塑性铰的定义，利用近似的方法计算梁端转角
式中：l0为梁最大转角到最大挠度的近距离；Δ为梁的挠度。
论文通过在试爆立柱相对应的二楼立柱布置沉降观测贴，通过水准仪进行观测沉降变形，再通过算出塑性转角来判断剩余结构的安全性（参照图4）。
从表4的数据可知：中心立柱爆破失效后，失效点下沉距离为2.6 mm，将下沉距离代入式（9）中，得到塑性转角为0.05°，根据前文关于梁塑性转角的讨论可知，立柱不会发生连续倒塌。
4　剩余结构动态响应特征数值模拟
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4.1　结构模型及参数确定
根据楼房结构特征和试爆立柱部位，论文选取第一节中的⑤~⑨轴的局部结构进行建模（参见图5），依据《混凝土结构设计规范》（GB50010—2010）和楼房实际几何尺寸[17]，采用PKPM软件建立模型，一层柱截面尺寸为400 mm×600 mm，二层以上柱子截面尺寸为400 mm×500 mm；梁截面尺寸为400 mm×700 mm；楼板为100 mm；内外墙体荷载按照等效原则施加于框架梁柱上。梁、板、柱混凝土强度等级均为C30；纵向受力钢筋为HRB400，箍筋为HRB335；屋面恒载3.5 kN/m2，活载为0.5 kN/m2，楼面恒载为4.5 kN/m2，活载为2.0 kN/m2，配筋均与实际结构相近，然后将模型导入至有限元软件SAP2000。
本文的结构模型中，框架的梁和柱为杆单元，而楼面则使用薄壳单元来模拟（参见图6）。根据《建筑结构抗倒塌设计规范CECS392∶2014》规定[13]，论文采用非线性直接积分法进行分析，并且考虑结构的几何非线性（即P-Delta效应）和大位移情况。进行时程分析时，积分步长被控制在不大于0.005 s以确保分析的准确性。基于Rayleigh阻尼的剩余结构模型，并假设钢筋混凝土的阻尼比为0.05。分析中设定构件在1.0 s时刻失效，取失效时间为0.05 s。根据GSA2003规定，承重立柱的失效时间被限制在不大于0.1倍的构件失效后剩余结构的自振周期T。
4.2　应变分析
为使数据更符合现场实测数据，采用SAP2000有限元软件得出立柱失效2 s后，失效点上方A梁，B梁，C梁近端20 cm处的最大主应力和远端20 cm处的最小主应力，如表5所示。
通过SAP2000有限元软件得出失效上方A梁，B梁，C梁的弯矩，其中MA=83.63 kN/m，MB=110.09 kN/m，MC=79.07 kN/m；将数据代入式（8）中，转换成应变与数值模拟数据和第三节现场实测应变进行对比，表6为失效点上方应变结果对比。
从表6的数据可以看出：框架结构楼房中心立柱爆破失效后，现场实测梁近端峰值平均值为428 με、梁近端峰值平均值为325 με；理论计算应变值均值为260 με；数值模拟梁近端均值为256 με、梁远端均值为219 με，三者所得出失效点上方梁的应变值基本接近。总体上看，数值模拟结果和理论计算结果基本符合现场实测结果，结构不会发生连续倒塌，这也表明数值模拟的可靠性。
4.3　动态位移和加速度响应
建（构）筑物爆破拆除的失稳倒塌是一个高度复杂的非线性动力学过程。目前，爆破拆除理论研究的最大瓶颈是缺少像结构工程那样有一套完整的力学理论基础。在研究建筑物爆破拆除后剩余结构的动力学特征以及评估其整体抗倒塌性能时，位移和加速度成为一个重要和必需的参数。为了与实际工程所得数据对比，在SAP2000有限元模型中，拆除与实际工程相对应的立柱，图7为失效点位移时程曲线，图8为失效节点加速时程曲线，表7为失效节点的动态位移和加速度极值，其中加速度向上被标记为正值。
分析图7、图8和表7数据分析可知：
（1）由于自重影响失效点位移并不是从0开始，失效点的竖向位移在经历短暂波动后（0.5 s内），趋于一稳定值，模拟数据与实验数据相差19%，两者大致相同。
（2）将数值模拟结果代入公式（9）得到的塑性转角为0.051°，与实验数据0.05°基本接近，根据GSA2003规范可知，结构不会发生连续倒塌。
（3）框架结构楼房立柱爆破失效后，失效节点的重力大于上部拉力，节点加速向下，产生显著的“失重”效应；失效节点的重力小于上部拉力，节点减速向下，产生显著的“超重”效应，其正方向最大加速度为3.5 m/s2，负方向-3.6 m/s2，两者基本相似，且失重和超重的变换速度极快，直至停止。总体上看，立柱拆除情况下，结构荷载对剩余结构的动力冲击作用影响较小，与失效节点位移对应。
4.4　内力调整
在框架结构中，结构变形是一定的，内力分布也是固定的，但是由于立柱失效，其所承受荷载分配不均匀，结构无法服从线弹性关系，进入塑性变形阶段，形成塑性铰，出现塑性内力重分布。研究特定部位承重立柱拆除后框架结构的内力调整情况，是分析框架结构楼房爆破拆除失稳倒塌破坏机理的基础，也可为结构立柱失效后抗倒塌能力评估提供一定的数据支撑。
本文选取失效立柱周围四根相邻立柱作为研究对象，立柱轴力变化计算结果见表8，失效节点梁端弯矩变化见表9。其中图9为立柱失效工况下内外围底层各柱轴力时程曲线，图10为立柱失效后失效节点上方梁弯矩时程曲线。
从上述计算结果表中和图9、图10时程曲线可知：
（1）中心立柱失效后，首层相邻柱的轴力发生急剧变化，继而上下波动衰减，最终趋于一稳定值。
（2）中心立柱失效后，其内力基本由其相邻的立柱分担，其中A柱的轴力增加近28%，B柱和C柱的轴力增加近23%，D柱距离失效立柱较远，其轴力增加近11%。可见，与失效柱相邻的柱是结构内力重分布的关键构件，且受到两柱之间距离的影响。
（3）中心立柱失效前上方梁为受拉状态，弯矩值为21.74 kN/m，失效后上方梁变为受压状态，弯矩值为-367.71 kN/m，这种显著的弯矩变化表明梁经历了悬链线效应，并且悬链线阻力开始抵抗上部载荷。
5　结论
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论文以一栋8层框架结构楼房立柱现场试爆为实验背景，应用SAP2000中建立了与实际工程相同的RC框架模型，研究了中心立柱失效条件下剩余结构的动力响应特征与整体稳定性，现场实测结果与数值计算结果比较接近，主要结论如下：
（1）通过现场实验研究中心立柱失效条件下，失效点的位移和相邻梁结构的应变数据可以得出，失效点的位移下降2.67 mm。梁底应变在377 με左右，并且结构未发生倒塌破坏。
（2）通过数值模拟研究单根中心立柱爆破失效后的动力响应特征，研究发现立柱失效后，失效节点位移为3.2 mm与现场实测相近，由此导致的塑性转角约为0.051°，明显小于GSA2003规范规定的6°，并且梁体应变值与实测应变基本接近。这表明中心立柱爆破失效后，剩余结构仍处于较为安全的状态，同时体现了此数值模拟的可靠性。
（3）立柱爆破失效后，楼房荷载对剩余结构的动力冲击作用影响较小，由此导致的超重与失重加速度绝对值为3.5 m/s2左右。楼房荷载重新分布，临近立柱承担了相应的轴力，距离较近柱立柱的竖向轴力增加25%左右，距离较远立柱的竖向轴力增加11%。上方梁体由受拉状态转变为受压状态，悬链线效应出现。这表明单根中心立柱失效后，通过内力重分布，剩余结构仍有较大承载能力，同时也反映出与失效立柱紧邻的柱和梁是结构进行内力重分布的关键构件。
（4）本文通过实验，理论分析和数值模拟三种方法探究梁柱结构的变形与动态响应。数据表明中心立柱爆破失效后，残余结构承载力较大，因此楼房的连续倒塌风险相对较小。并且此论文的研究成果有助于框架结构楼房在意外爆炸情况下的易损性评估与抗连续倒塌性能分析，同时也可为拆除爆破中的试爆立柱选择提供参考。
基金
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  • 武汉市知识创新专项基础研究项目(2022020801010371)
  • 湖北（武汉）爆炸与爆破技术研究院博士科研启动基金项目(PBSKL-2022-QD-02)
  • 湖北省重点研发计划项目(2020BCA084)
文献
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doi: 10.3963/j.issn.1001-487X.2024.03.015
  • 接收时间：2024-04-19
  • 首发时间：2026-03-20
  • 出版时间：2024-09-01
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  • 收稿日期：2024-04-19
基金
Wuhan Municipal Knowledge Innovation Special Basic Research Project(2022020801010371)
武汉市知识创新专项基础研究项目(2022020801010371)
Project of Doctoral Research Start-up Fund of Hubei (Wuhan) Institute of Explosion and Blasting Technology(PBSKL-2022-QD-02)
湖北（武汉）爆炸与爆破技术研究院博士科研启动基金项目(PBSKL-2022-QD-02)
Key Research and Development Program of Hubei Province(2020BCA084)
湖北省重点研发计划项目(2020BCA084)
作者信息
    1a.江汉大学　省部共建精细爆破国家重点实验室，武汉　430056
    1b.江汉大学　爆破工程湖北省重点实验室，武汉　430056
    2.武汉爆破有限公司，武汉　430056
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https://castjournals.cast.org.cn/joweb/bp/CN/10.3963/j.issn.1001-487X.2024.03.015
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2种不同金属材料的力学参数
科
Family		 属数
Number of
genus		 种数
Number of
species		 占总种数比例
Percentage of
total species (%)		 属
Genus		 种数
Number of
species		 占总种数比例
Percentage of total
species (%)
鹅膏菌科Amanitaceae 2 11 5.26 鹅膏菌属 Amanita 10 4.78
小菇科 Mycenaceae 2 12 5.74 丝盖伞属 Inocybe 5 2.39
多孔菌科 Polyporaceae 8 14 6.70 蜡蘑属 Laccaria 5 2.39
红菇科 Russulaceae 3 23 11.00 小皮伞属 Marasmius 6 2.87
小菇属 Mycena 11 5.26
光柄菇属 Pluteus 5 2.39
红菇属 Russula 17 8.13
栓菌属 Trametes 5 2.39
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